Text

Plant Safety Evaluation for Beaver Valley Power Station Unit 1 Fuel Upgrade & Increased Peaking Factors. Related Info,Including Recommended Mods to Facility Fsar,Encl
	ML20246K454
Person / Time
Site:	Beaver Valley
Issue date:	04/30/1989
From:	Foley J, Novendstern E, Yeh R; WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
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l O j PLANT SAFETY EVALUATION FOR BEAVER VALLEY POWER STATION UNIT 1 FUEL UPGRADE AND INCREASED PEAKING FACTORS APRIL 1989 O Editor: J. V. Foley Contributors:

R.Y.Yeh C. E. Leach M.D.Woodw J. S. Petzold C. A. Bly C. R. Hyatt A.J. Baker Approved:

E. H. Novendstern Thermal-Hydraulic Design and Fuel Licensing WESTINGHOUSE ELECTRIC CORPORATION Commercial Nuclear Fuel Division P. O. Box 3912 Pittsburgh, Pennsylvania 15230 890518~>061 890509 PDR ADOCK 05000334 "

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PLANT SAFETY EVALUATION for BEAVER VALLEY POWER STATION UNIT 1 FUEL UPGRADE AND INCREASED PEAKING FACTORS TABLE OF CONTENTS SECTION TITLE PAGE 1.0 Introduction and Summary 1.1 Upgraded Fuel Features (VANTAGE 5 Hybrid) 1-1 1.2 Increased Peaking Factors 1-2 1.3 Conclusions 1-2 2.0 Design Features 2.1 Introduction 2-1 2.2 VANTAGE SH Fuel Assembly 2-1 2.3 Other Upgraded Fuel Features 2-2 2.4 Fuel Rod Performance 2-3 3.0 Nuclear Design 3.1 Introduction and Summary 3-1 3.2 Increase in FN AH Limit 3-2 3.3 increase in FO(Z) Limit 3-2 3.4 Methodology 3-3 4.0 Thermal and Hydraulic Design 4.1 Introduction and Summary 4-1 4.2 Methodology 4-1 4.3 Hydraulic Compatibility 4-3 4.4 Effects of Fuel Rod Bow on DNBR 4-3 4.5 FuelTemperature Analysis 4-4 4.6 Transition Core Effect 4-4 4.7 Conclusion 4-4 0013AA:6-890420 l

I TABLE OF CONTENTS (Cont.)

SEC110N TITi.E PAGE 5.0 Accident Analysis 5-1 5.1 Non-LOCA Accidents 5-2 5.1.1 Overtemperature and Overpower AT Protection 5-4 5.1.2 Increase in Heat Removal by the 5-5 Secondary System 5.1.3 Decrease in Heat Removal by the 5-6 Secondary System 5.1.4 Decrease in RCS Flow Rate 5-8 5.1.5 Reactivity and Power Distribution Anomalies 5-11 5.1.6 Increase in Reactor Coolant Inventory 5-15 5.1.7 Decrease in Reactor Coolant inventory 5-15 5.1.8 Steamline Break Mass and Energy Release 5-16 for Postulated Ruptures inside Containment and Equipment Environmental Qualification Outside Containment 5.2 LOCA Accidents 5-42 5.3 Accident Analysis Conclusions 5-46 6.0 References 6-1 APPENDIX A Summary of Technical Specification Changes APPENDIX B Recommended Modifications to Beaver Valley Unit 1 FSAR 0013AA:6-890420 II

LIST OF FIGURES FIGURE TITLE PAGE 2.1 Comparison of the 17x17 VANTAGE SH Fuel Assembly and 2-4 the 17x17 STD Fuel Assembly 5.1 -1 Flow Transients for Partial Loss of Flow - Three Loops 5-19 in Operation, One Pump Coasting Down 5.1 -2 Nuclear Power and RCS Pressure Transient for Partial 5 20 Loss of Flow - Three Loops in Operation, One Pump Coasting Down 5.1 -3 Average and Hot Channel Heat Flux Transient for Partial 5-21 Loss of Flow - Three Loops in Operation, One Pump Coasting Down 5.1 -4 DNBR vs Time for Partial Loss of Flow - Three Loops in 5-22 Operation, One Pump Coasting Down 5.1 -5 Core Flow Coastdown vs Time for Complete Loss of Flow - 5-23 Three Loops in Operation, Three Pumps Coasting Down 5.1 -6 Nuclear Power and RCS Pressure Transient for Complete 5-24 Loss of Flow - Three Loops in Operation, Three Pumps Coasting Down 5.1 -7 Average and Hot Channel Heat Flux Transient for Complete 5-25 Loss of Flow - Three Loops in Operation, Three Pumps Coasting Down 5.1 -8 DNBR vs Time for Complete Loss of Flow - Three Loops in 5-26 !

Operation, Three Pumps Coasting Down ootsAA$soono lii i

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LIST OF FIGURES (Continued)

FIGURE TITLE PAGE 5.1-9 Flow Transients for One Locked Rotor - Three Loops in 5-27 Operation 5.1-10 Reactor Coolant System Pressure for One Locked Rotor - 5-28 Three Loops in Operation 5.1-11 Nuclear Power Transient, Average and Hot Channel Heat 5-29 Flux Transients for One Locked Rotor - Three Loops in Operation.

5.1-12 Maximum Clad and Fuel Centerline Temperatures for One 5-30 Locked Rotor - Three Loops in Operation 5.1-13 Neutron Flux vs Time - Uncontrolled Rod Withdrawal 5-31 from a Suberitical Condition 5.1-14 Core Heat Flux vs Time - Uncontrolled Rod Withdrawal 5-32 from a Subcritical Condition 5.1-15 Fuel Temperature vs Time - Uncontrolled Rod Withdrawal 5-33 from a Subcritical Condition 5.1-16 Nuclear Power vs Time - Improper Startup of an inactive 5-34 Reactor Coolant Pump 5.1-17 Core Heat Flux vs Time - Improper Startup of an inactive 5-35 Reactor Coolant Pump 5.1-18 Vessel Average Temperature vs Time - Improper Startup of 5-36 an inactive Reactor Coolant Pump 0013AA$-890420 , ,

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i LIST OF FIGURES (Continued) i FIGURE TITLE PAGE 5.1-19 RCS Pressure vs Time - Improper Startup of an Inactive 5-37 Reactor Coolant Pump 5.1-20 Nuclear Power Transient, BOL-HFP Rod Ejection Accident 5-38 5.1-21 Peak Fuel and Clad Average Temperature, BOL-HFP Rod 5-39 Ejection Accident 5.1-22 Nuclear Power Transient, EOL-HZP Rod Ejection Accident 5-40 5.1-23 Peak Fuel and Clad Average Temperature, EOL-HZP 5 41 Rod Ejection Accident 0

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LIST OF TABLES TABLE TITLE PAGE

! 2-1 Comparison of 17x17 Standard, Optimized, VANTAGE 5 2-5 and VANTAGE 5 Hybrid Fuel Assembly Mechanical Design Parameters 4-1 Beaver Valley Unit 1 Thermal and Hydraulic Design 4-5 Parameters 1

4-2 DNBR Margin Summary 4-7 5.1 -1 Single Reactor Coolant Pump Locked Rotor Accident Results 5-17 5.1-2 Rod Cluster Control Assembly Ejection Accident Results 5-18 '

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ACKNOWLEDGEMENTS The' editor would like to acknowledge the efforts of the contributors to the design Sections and Appendices to this Safety Report. Major contributors included:

Mechanical Design - C. Hyatt -

Nuclear Design - R. Y. Yeh Thermal and Hydraulic Design - M. D. Woods Non-LOCA Analysis - C. E. Leach LOCA Analysis - J. Petzold Technical Specifications - A. J. Baker in addition, acknowledgement is given to B. D. McKenzie for his assistance in providing the necessary interfaces between CNFD, NATD and the Duquesne Light Company.

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1.0 INTRODUCTION

AND

SUMMARY

I The Beaver Valley Power Station Unit 1 plans to refuel and operate with upgraded Westinghouse fuel features and increased peaking factors. This report summarizes the safety evaluations that were performed to confirm the acceptable use of these options for three loop operation. Sections 2.0 through 5.0 of the Plant Safety Evaluation (PSE) provide the results of the Mechanical, Nuclear, Thermal and Hydraulic, and Accident l Evaluations, respectively. Appendix A gives a summary of the Technical Specification l changes required and the corresponding change pages. Recommended changes to the Beaver Valley Unit 1 FSAR(1) are included in Appendix B.

The Beaver Valley Unit 1 Plant Safety Evaluation is to serve as a reference safety evaluation / analysis report for the region-by-region reload transition from the present Beaver Valley Unit 1 core (Cycle 7) to a core containing the upgraded features described below. Thus, the PSE will be used as a basic reference document in support of future Beaver Valley Unit 1 Reload Safety Evaluations (RSEs) for upgraded fuel reloads.

The PSE utilizes the Westinghouse standard reload methodology (2). Consistent with this methodology, parameters are chosen to maximize the applicability of the PSE evaluations for future cycles. The objective of subsequent cycle specific RSEs will be to verify that applicable safety limits are satisfied based on the reference evaluation / analyses established in this safety evaluation.

1.1 Upgraded Fuel Features (VANTAGE 5 Hybrid) l Beaver Valley Unit 1 Cycle 8 and subsequent core loadings will have fuel assemblies that incorporate a low pressure drop Zircaloy grid. This upgraded fuel feature is known as VANTAGE 5 Hybrid (VANTAGE SH) and has been submitted as an Addendum (3) to the " VANTAGE 5 Reference Core Report," WCAP-10444-P-A(4). VANTAGE SH has received generic NRC approval (5),

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in addition to the VANTAGE 5H design feature, Beaver Valley Unit 1 reloads will also contain several VANTAGE 5 design features (4) and other upgraded fuel design features. Beaver Valley Unit 1 Cycle 7 has already incorporated the Integral Fuel Burnable Absorbers (IFBAs) and Axial Blanket design features of VANTAGE 5.

Subsequent cycles will also contain the VANTAGE 5 Reconstitutable Top Nozzle (RTN) design feature, as well as Debris Filter Bottom Nozzles (DFBNs), snag resistant grids and standardized fuel pellets. A brief summary of the upgraded fuel features is presented in Section 2 of this report.

1.2 Increased Peaking Factors The future cycles of operations for .seaver Valley Unit 1 will use increased FNAH and l Fo(Z) peaking factors. The full power F N AH Peaking factor design limit will increase from the current value of 1.55 to 1.62. The maximum Fo(Z) peaking factor limit will increase from the current value of 2.32 to 2.40 and the K(Z) er.velope will be modified.

These increases will permit more flexibility in developing fuel management schemes (i.e.,

longer fuel cycles, improvement of fuel economy and neutron utilization).

1.3 Conclusions The results of evaluation / analysis described herein lead to the following conclusions:

1. The Westinghouse fuel assemblies containing VANTAGE SH and the additional upgraded fuel features for Beaver Valley Unit 1 are mechanically compatible with the current fuel assemblies, control rods, and reactor Intemals interfaces.

The current design bases for Beaver Valley Unit 1 have been changed as described in this report to accommodate the VANTAGE SH design. j

! 2. Changes in the nuclear characteristics due to the transition to upgraded fuel will be within the range normally seen from cycle to cycle due to fuel management.

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3. The reload upgraded fuel assemblies are hydraulically compatible with the fuel ,

assemblies from previous reload cores.

4. The changes in the design full power F N H limit from 1.55 to 1.62 (with appropriate treatment of uncertainties) is supported by design basis safety analyses summarized in this evaluation. The corresponding changes to the Technical Specifications are as defined in Appendix A.

5. The change in the maximum Fo(Z) limit from 2.32 to 2.40 and modification to the K(Z) enve! ape is supported by the design basis safety analyses summarized in this report. The corresponding changes to the Technical Specifications are as defined in Appendix A.

6. The core design and safety analyses results documented in this report show the '

core's capability for operating safely at the rated Beaver Valley Unit 1 design

~ thermal power.

7. This report establishes a reference upon which to base Westinghouse reload safety evaluations for future reloads with the upgraded fuel features and increased peaking factor limits.

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2.0 DESIGN FEATURES i

2.1 Introduction The mechanical design of the upgraded fuel assemblies for Beaver Valley Unit 1 is the I same as previous reload fuel assemblies except that the upgraded fuel assemblies will .)

incorporate several fuel design improvements. These improvements include the l- VANTAGE SH Zircaloy grid, Reconstitutable Top Nozzles, Debris Filter Bottom Nozzles, l l snag resistant grids and standardized fuel pellets. In addition, the upgraded fuel I l assemblies will continue to use the Integral Fuel Burnable Absorber and Axial Blanket l design features.' The design changes are described in more detail in the following sections.

2.2 VANTAGE SH Fuel Assembly The VANTAGE 5H fuel assembly design evolved from the current VANTAGE 5, Optimized Fuel Assembly (OFA) and Standard (STD) fuel assembly designs. It is based on substantial design and operating experience. Design features from each of these previous designs are incorporated into the VANTAGE SH fuel assembly design. The VANTAGE SH design is characterized by the use of Zircaloy grids with 0.374 inch OD standard fuel rods. To accommodate the Zircaloy grids, the VANTAGE 5H thimble tube diameter was modified to be the same as the 17x17 OFA or VANTAGE 5 fuel. A comparison of the STD and VANTAGE SH fuel assembly design parameters is given in Table 2.1. Figure 2.1- demonstrates the similarity of the two designs and shows a comparison of overall dimensions.

Comparative fuel assembly flow testing results indicate that the VANTAGE SH and the STD 17x17 fuel assembly are hydraulically equivalent. Full assembly testing has

)

confirmed that the VANTAGE SH fuel assembly has hydraulic stability and that the fuel i rod contact wear with the spacer grids is within the allowable design limits.

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The major components that determine the structural integrity of the fuel assembly are the grids. Mechanical testing and analysis of the VANTAGE SH Zircaloy grid and fuel assembly have demonstrated that the VANTAGE SH structural integrity under seismic /LOCA loads will provide margins comparable to the STD 17x17 fuel assembly design and will meet all design bases.

The VANTAGE 5H Zircaloy grid is based on the OFA Zircaloy grid design and operating experience. The grid strap thickness, type of strap welding, basic mixing vane design and pattem, method of thimble tube attachment, type of fuel rod support (6 point),

material and envelope are identical to the OFA Zircaloy grid. This evaluation of the VANTAGE SH grid performance is based on the extensive design and irradiation experience with previous grid designs and full grid testing completed with the VANTAGE 5H grid design.

In order to demonstrate early performance of the Zircaloy grid design, fuel assembly '

demonstration programs were conducted inserting OFA fuel assemblies containing Zircaloy grids into 14x14,15x15 and 17x17 cores. Subsequent to the satisfactory performances observed in these programs, the OFA with Zircaloy grids were loaded and have operated successfully since the early 1980's in many Westinghouse cores (6),

2.3 Other Upgraded Fuel Features Beaver Valley Unit 1 Cycle 8 and subsequent reloads will contain fuel assemblies that incorporate Reconstitutable Top Nozzles, Debris Filter Bottom Nozzles, snag resistant grids and staredardized fuel pellets as well as the VANTAGE SH Zircaloy grids described in the previous section. These design changes, described below, are currently part of the licensing basis in other plants and meet all fuel assembly and fuel rod design criteria.

Rebris_EliterEQttom Nozzle (DFBN) - This bottom nozzle is designed to inhibit debris l

from entering the active fuel region of the core and thereby improves fuel performance by l minimizing debris related fuel failures. The DFBN is a low profile bottom nozzle 1

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design made of stainless steel, with reduced end plate thickness and leg height. The DFBN is structurally and hydraulically equivalent to the existing bottom nozzle.

l Beconstitutable Top Nozzle (RTN) - The RTN differs from the current design in two ways: a groove is provided in each thimble thru-hole in the nozzle plate to facilitate attachment and removal; and the nozzle plate thickness was reduced to provide additional space for fuel rod growth. In conjunction with the RTN, a long tapered fuel rod bottom end plug is used to facilitate removal and reinsertion of the fuel rods.

i Standardized Fuel Pellets - The standardized pellet is a refinement to the current pellet design with the objective of improving manufacturability while maintaining or improving performance. This design incorporates a reduced pellet length, modification to the i previous dish size and the addition of a chamfer.

Snag Resistant Grids - The snag-resistant grids contain outer grid straps which are modified to help prevent assembly hangup from grid strap interference during fuel assembly removal. This was accomplished by changing the grid strap corner geometry and the addition of guide tabs on the outer grid strap.

2.4 Fuel Rod Performance The 0.374 inch OD fuel rod used in the VANTAGE SH fuel assembly is the same as that used in the Beaver Valley Unit 1 17x17 STD fuel assemblies. The design bases, methodology, and models are the same as those described previously(4). No changes in fuel rod design criteria, methods, or models are necessary because of the transition to VANTAGE SH fuel or increased peaking factor limits. The STD and VANTAGE SH fuel are designed according to the Westinghouse fuel performance models(7,8). All fuel rod design criteria are satisfied for the planned irradiation life, oo13W890C0 2-3

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3.0 NUCLEAR DESIGN 3 3.1 Introduction and Summary The effects of using upgraded Westinghouse fuel features and increasing the core peaking factor limits on the nuclear design bases and methodologies for Beaver Valley Power Station Unit 1 are evaluated in this section.

The grid material and grid volume of the VANTAGE SH fuel are different than that of 17x17 standard fuel assemblies. The effects of these changes on core physics characteristics are small and are explicitly modeled in the neutronics codes. The specific values of core safety parameters, e.g., power distributions, peaking factors, rod worths, are primarily loading pattern 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. In addition, the present Beaver Valley spent fuel pool criticality analysis is applicable to the upgraded Westinghouse fuel features, including the use of VANTAGE SH fuel.

The increase in peaking factor limits allows more flexibility when developing the fuel management scheme. Specific items concerning the evaluation and necessary Technical Specification changes are given in Section 3.2 and 3.3.

In summary, the change from the current standard fuel core to a core containing the upgraded fuel product will not cause changes to the current Beaver Valley Unit 1' FSAR nuclear design bases. However, the design bases will be modified due to the increases I to the peaking factor limits. Nuclear design methodology is not affected by the use of upgraded fuel features or increased peaking factors.

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= 3.2 increase in FN AH Limit

'The limit on the nuclear enthalpy rise hot channel factor, FNAH, will take the following form in the Technical Specifications:

FN H = 1.62 (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 fuel.

3.3 Increase in Fo(Z) Limit The limit on the heat flux hot channel factor, Fo(Z), will take the following form in the Technical Specifications:

FQ(Z) s (2.40/P) x (K(z)) for P > 0.5, and Fo(Z) s (4.80)x(K(z)) for P s 0.5 where P = THERMAL POWER / RATED THERMAL POWER, and K(Z) - Tne function obtained from Figure 3.2-2 in the Beaver Valley Unit 1 Technical Specifications for a given core height location. A revised K(Z) function appropriate for the increased Fo(Z) limit is provided in Appendix A of this submittal.

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The increased heat flux hot channel factor limit, Fo(Z), will allow additional flexibility in fuel management'and core operation as well as accommodate the increased nuclear

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enthalpy rise hot channel factor limit and use of fuel that incorporates axial blankets.

With the longer cycles and the necessary higher enrichments that will be used in the future, the radial peaking factor (FNAH discussed in Section 3.2) will increase. . This

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increase will result in higher total peaking factors in the Beaver Valley Unit 1 core. Fuel -

using natural uranium axial blankets may also result in slight increases in the axial component of the total peaking factor. However, this results in minimal peaking factor increases, since Integral Fuel Burnable Absorbers (IFBAs) are also incorporated into the fuel design to reduce axial peaking factors.

3.4 Methodology No changes to the nuclear design philosophy or methods are necessary because of the upgraded fuel product or the use of increased peaking factors. The reload design philosophy includes the evaluation of the reload core key. safety parameters which comprise the nuclear design dependent input to the FSAR safety evaluation for each reload cycle (2). These key safety parameters will be evaluated for each Beaver Valley Unit 1 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 RSE for that cycle.

The 0.374 inch diameter fuel rod has had extensive nuclear design and operating experience with the current Beaver Valley Unit 1 17x17 STD fuel assembly design. The Zircaloy grid material has also had extensive nuclear design and operating experience with the current 17x17 VANTAGE 5 and 17x17 OFA fuel assembly designs. These changes have a negligible effect on the use of standard nuclear design analytical models and methods to accurately describe the neutronic behavior of the VANTAGE 5H fuel.

Peaking factors similar.to the new values for Beaver Valley Unit 1 have been used extensively. Analyses of Beaver Valley Unit i reload cores are performed in accordance with standard reload methodology (2) to ensure compliance with the new Technical Specification peaking factor limits. j i

l 0013A 6-890420 3-3

l 4.0 THERMAL AND HYDRAULIC DESIGN {

1 4.1 Introduction and Summary This section ' describes the calculational methods used for the thermal-hydraulic analysis, the DNB performance, and the hydraulic- compatibility during the transition from mixed-fuel cores to an all VANTAGE SH core. Based on minimal hardware design differences and prototype hydraulic' testing of the fuel assemblies, it is concluded (3) that the STD and VANTAGE SH fuel assembly designs are hydraulically compatible. Table 4-1 summarizes the thermal-hydraulic design parameters for Beaver Valley Unit 1 that-were used in this analysis. The thermal-hydraulic design for the upgraded fuel product was analyzed for an increase in the design limit value for the nuclear enthalpy rise hot-channel factor (FNAH) from 1.55 to 1.62. This increase is achieved by removing unnecessary conservatism from the analysis through use of improved methods and DNB correlation as described in the following section. The thermal-hydraulic design criteria and methoas remain the same as those presented in the Beaver Valley Unit 1 FSAR with the exceptions noted in the following sections. All of the current FSAR thermal-hydraulic de' sign criteria are satisfied.

4.2 Methodology The existing thermal-hydraulic analysis of the 17x17 STD fuel used in the Beaver Valley Unit 1 plant is based.on the standard thermal and hydraulic methods and the W-3 (R-Grid) DNB correlation as described in the Beaver Valley Unit 1 FSAR. The DNB analysis of the core containing both 17x17 STD and VANTAGE SH fuel assemblies has i been modified to incorporate the WRB-1 DNB correlation (9) and a conservative application of the Revised Thermal Design Procedure (RTDP)(10) which is called 1 MINI-RTDP(11).

0013t6/890420 4-1

q.

h' The WRB-1 DNB correlation is based entirely on rod bundle data and takes credit for the significant improvement in the accuracy of the critical heat flux predictions over previous DNB correlations. The approval of the NRC that a 95/95 limit DNBR of 1.17 is appropriate for the 17x17 STD fual assemblies has been documented (12),

The WRB-1 DNB correlation is applicable to VANTAGE 5H fuel since, from a DNB perspective, the VANTAGE SH assembly is virtually identical to the 17x17 Inconel R-Grid design. As documented in VANTAGE SH Fuel Assembly Report (3), the use of the WRB-1 DNB correlation with a 95/95 limit DNBR of 1.17 is applicable to the VANTAGE SH fuel assembly.

With MIN 1-RTDP methodology, peaking factor uncertainties are combined statistically

- with the DNB correlation uncertainties to obtain the overall DNBR uncertainty factor which is used to define the design limit DNBR that satisfies the DNB design criterion.

This criterion is that the probability that DNB will not occur on the most limiting fuel rod is at least 95% (at 95% confidence level) for any Condition I or il event.

' The uncertainties included in the combined peaking factor uncertainty are the nuclear enthalpy rise hot channei factor, (FNAH); the enthalpy rise engineering hot channel factor, (FEAH); and uncertainties in the THINC-IV and transient codes. The increase in DNB margin is realized when nominal values of the peaking and hot channel factors are used in the DNB safety analyses.

With MINI-RTDP, uncertainties in the plant primary system parameters (reactor power, flow, temperature and pressure) are excluded from the statistical combination process and, therefore, no additional surveillance of these parameters is required. initial condition assumptions for the DNB safety analyses will use the same conservative values for these plant system parameters that are used in the standard thermal design methods.

l I

I 0013L6/890420 4-2

. l

- _ _ _ _ - _ _ - - _ _ _- _ _ _ _ . )

.~

For this application, the design limit DNBR for typical and thimble cells is 1.21 which applies for both 17x17 STD and VANTAGE SH fuel assemblies. For use in the DNB safety analyses, the limit DNBR is conservatively increased to provide DNB margin to offset the effect of rod bow and any other DNB penalties that may occur, and to provide i flexibility in design and operation of the plant. The safety analysis limit DNBR providing I for 9% margin is calculated as follows:

1.21 Safety Analysis Limit DNBR = Design Limit DNBR , - 1.33 1.0 - Margin 1.0 - 0.09 Table 4-2 summarizes the available DNBR margin for Beaver Valley Power Station Unit 1.

4.3 Hydraulic Compatibility The STD fuel assembly and VANTAGE SH designs have been shown to be hydraulically compatible in the VANTAGE SH Fuel Assembly Report (3),

4.4 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 11 events. Currently, the maximum rod bow penalty is 1.3%

DNBR at an assembly average bumup of 24,000 MWD /MTU. For burnups greater than 24,000 MWD /MTU, credit is taken for the effect of FN AH burndown, due to the decrease in fissionable isotopes and the buildup of fission product inventory. Therefore, no additional rod bow penalty is required at bumups greater than 24,000 MWD /MTU. Based on the similarities between 17x17 STD and VANTAGE SH fuel assemblies, (i.e. fuel rod >

diameter, fuel rod pitch and grid spacing), this penalty is also applicable to VANTAGE SH fuel assemblies.

For this application, the rod bow penalty will be offset with DNB margin retained between the safety analysis and design DNBR limits, I l

l l

ootswsoono 4-3

4.5 FuelTemperature Analysis There is no difference in the fuel temperatures used in the safety analysis calculations between the VANTAGE SH fuel and the STD fuel. The fuel temperatures for the standardized pellets are the same as those for the unchamfered pellets and slightly less than those for the current chamfered pellet design.

4.6 Transition Core Effect The VANTAGE SH hydraulic test program showed identical results for the VANTAGE SH grid and the STD fuel inconel mixing vane grid, therefore, no transition core DNBR penalty is necessary.

4.7 Conclusion The thermal hydraulic evaluation of the fuel upgrade and peaking factor increase for Beaver Valley Unit 1 has shown that 17x17 STD and VANTAGE SH fuel assemblies are hydraulically compatible and that the DNB margin gained through use of the MINI-RTDP methodology and the WRB-1 DNB correlation is sufficient to allow an increase in the design F N AH rom f 1.55 to 1.62. The core limit curves (Technical Specifications Figure 2.1-1 for three loops in operation) remain valid for both STD and VANTAGE SH fuel assemblies at the design FNAH of 1.62. More than sufficient DNBR margin in the safety limit DNBR exists to cover any rod bow penalties. The upgraded fuel features described in Section 2 do not affect the core flow rate, core flow distribution, or any other safety related parameters. All current thermal-hydraulic design criteria are satisfied.

i ooisweso420 4-4

~

k 4

TABLE 4-1 BEAVER VALLEY UNIT 1 THERMAL AND HYDRAULIC DESIGN PARAMETERS Thermal and Hydraulic Design Parameters Design Parameters Reactor Core Heat Output, Mgt 2,652 Reactor Core Heat Output,10 BTU /Hr 9,051 Heat Generated in Fuel, % 97.4 Core Pressure, Nominal, psia 2250 Radial Power Distribution

1.62[1+0.3(1-P)]

Limit DNBR for Design Transients" 1.33 DNB Correlation" WRB-1 HFP Nominal Coolant Conditions Vessel Thermal Design Flow Rate (including Bypass),106 lbm/hr 100.8 GPM 265,500 Core Flow Rate"*

(ewiuding Bypass, based on TDF) 100 lbm/hr 94.25 GPM 248,242 Core Flow Area, ft2 STD 41.5 V5H 41.7 Cofe Inlet My(s Velocity, Based on 10 lbm/hr-ft STDTDF) 2.27 V5H 2.26 Includes 4% measurement uncertainty.

Applies to STD and VANTAGE SH fuel.

" Based on design bypass flow of 6.5% without thimble plugs.

ootsts/890420 4-5

TABLE 4-1 (Continued)

BEAVER VALLEY UNIT 1 THERMAL AND HYDRAULIC DESIGN PARAMETERS Thermal and Hydraulle Design Parameters Design Parameters Nominal Vessel / Core inlet Temperature, *F 542.5 Vessel Average Temperature, F 576.2 Core Average Temperature, F 580.2 VesselOutletTemperature, F 609.9 Average Temperature Rise in Vessel, F 67.4 Average Temperature Rise in Core, F 71.6 Heat Transfer Active Heat Transfer Surface Area, ft 2 STD 48,600 V5H 48,600 Average Heat Flux, BTU /hr-ft2 STD 181,405 V5H 181,405 Average Linear Power, kw/ft 5.20 Peak Linear Power for Normal Operation, kw/ft 12.48 Temperature at Peak Linear Power for Prevention of Centerline Melt, F 4700 Based on maximum Fo of 2.40 l

l ootswsoono 4-6 L

l TABLE 4 DNBR MARGIN

SUMMARY

li i

l 17x17 STD and VANTAGE 5H FUEL DNB Correlation WRB-1 Correlation Umit 1.17 Design Limit 1.21 Safety Limit 1.33 DNBR Margin 9%

Rod Bow DNBR Penalty 1.3%

Available DNBR Margin -7.7%

DNBR margin between the safety limit and the design limit DNBRs.

_ _ - - - _ - ""I'**_--_-_-__--__--__---_-_-__-__--.----4-7 _ _]

5.0 ACCIDENT ANALYSIS The primary effect of the proposed modifications on the LOCA and non-LOCA design basis calculations is due to the introduction of the VANTAGE SH Zircaloy grid and the increases in the FN AH and Fa peaking factor limits. The safety analysis justification for these design modifications is summarized for the non-LOCA and LOCA design basis calculations in Sections 5.1 and 5.2, respectively. The balance of the fuel upgrade features described in this report have been introduced and evaluated for previ = , reload designs or have a negligible effect on non-LOCA and LOCA methods and conclusions.

The effects of these features on the safety analyses are des::ribed as follows.

Two VANTAGE 5 fuel features (4) have been introduced and evaluated for previous Beaver Valley Unit 1 fuel reload designs. These are Axial Blankets and integral Fuel Burnable Absorbers (IFBAs). Axial Blankets reduce power et the ends of the fuel rod which increases axial peaking. Used alone, Axial Blankets reduce associated DNB margin, but the effect may be offset by the presence of IFBAs which flatten the power distribution. The net Mfect on the axial power shape is a function of the number and configuration of IFBAs in the core and time in life. The effects of Axial Blankets and IFBAs on the reload safety analysis parameters are taken into account in the reload design process. These effects are expected to be acceptable with respect to the reload design limitations imposed by the Technical Specifications applicable for any given reload.

An additione.1 VANTAGE 5 feature (4) described in this report is the Reconstitutable Top Nozzle (RTN). The RTN will be introduced in the Beaver Valley Unit 1 reload design in l Cycle 8. Core flow areas and loss coefficients were preserved in the design of the RTN.

As such, no parameters important to non-LOCA and LOCA safety analyses are affected by the RTN design. Therefore, the incorporation of 'Se RTN is justified for the Beaver Valley Unit 1 Cycle 8 reload design and all subsequer ' reload gcle designs.

l 1

ooisunoo42o 5-1

}-

As with the RTN, the Debris Filter Bottom Nozzle (DFBN) design preserves core flow areas and loss coefficients. Similar to the RTN, the DFBN is evaluated to have no impact an non LOCA and LOCA parameters or methods. Therefore, the incorporation of the DFBN is justified for the Beaver Valley Unit 1 Cycle 8 reload design and all subsequent reload cycle designs.

The Snag Resistant Grid is described in Section 2.3. The modifications introduced by this design have no effect on the non-LOCA and LOCA safety analysis assumptions or methods. Therefore, the incorporation of the Snag Resistant Grid is justified for Beaver Valley Unit 1 Cycle 8 and all subsequent teload cycle designs.

The remaining modification described in this report is the introduction of fresh fuel with the standardized fuel pellet design. The existing fuel inventory is composed of an earlier manufacturing version of the chamfered edge pellet fuel. The standardized design, as discussed in Section 2.3, is a further refinement in chamfer design which incorporates a reduced pellet length but restores the original square-edged pellet design dish size. The evolution of the chamfer pellet fuel design has no impact on the fuel rod geometry l parameters explicitly assumed in the safety analyses. The only non-LOCA and LOCA impact is a small modification to the fuel temperature versus linear power density relationship which is assumed in the safety analyses. The data calculated for the standardized pellet is less limiting than that which has been justified for previous reload {

designs. The non-LOCA and LOCA safety analysis assumption related to the standardized pellet are applicable for the Beaver Valley Unit 1 Cycle 8 reload and future cycle reload designs incorporating the standardized pellet. i 5.1 Non-LOCA Accidents .

(

This section summarizes the non-LOCA reanalyses and evaluations performed for the Beaver Valley Unit 1 upgrade to VANTAGE SH fuel and plant operation with increased FNAH and Fo(Z) peaking factors. in addition, this evaluation continues to support steam generator tube plugging, up to a level of 10%

0013A 6/890420 5-2 t

The major effect of changing from STD 17x17 fuel to VANTAGE SH fuel on the non-LOCA transients is the increased design Rod Control Cluster Assembly (RCCA) drop time. The VANTAGE SH fuel assembly has a thimble tube I.D. of 0.442 inches.

STD fuel has a thimble tube I.D. of 0.450 inches. The smaller VANTAGE SH thimble tube will increase the design RCCA drop time from a current maximum of 2.2 seconds to 2.7 seconds. This slower drop time will affect the results of the fast non-LOCA limiting transients such as Loss of Forced Reactor Coolant Flow Locked Rotor, RCCA Bank Withdrawal from Subcritical and Rod Ejection. The balance of the non-LOCA accidents are evaluated for this fuel upgrade.

Also considered is the increase in the design limit value for the nuclear enthalpy rise hot channel factor, FNAH, from 1.55 to 1.62. In general, an increase in FNAH results in a decrease in the DNBR value for a given set of thermal-hydaulic conditions. On this basis it would be expected that all transients for which DNBR is calculated would be affected.

As noted in Section 4, however, the margins obtained through use of the WRB-1 DNB correlation and the MINI-RTDP methodology allow for the increased peaking factor without changing the core thermal limits (Technical Specification Figure 2.1-1 for three loop operation). Therefore, only those transients which explicitly incorporate a value of FN H ni the calculation of the thermal-hydraulic conditions existing at the time of minimum DNBR require reanalysis. These are Partial Loss of Flow, Complete Loss of Flow RCP Underfrequency, RCP Locked Rotor and Startup of an inactive Loop at an incorrect Temperature. The balance of the non-LOCA transients are evaluated for the increased FNAH and explicit reanalysis is not required.

Finally, this section considers a revision in the Technical Specification limit for peak Fo as a function of core height. The revised Technical Specification increases the maximum Fo from 2.32 to 2.4. Two non-LOCA Condition IV accidents incorporate the maximum FO(Z) limit into the safety analysis assumptions to demonstrate that clad integrity and fuel melting acceptance criteria are met. These are RCP Locked Rotor and Rod Ejection. All other non-LOCA transients are evaluated and do not require explicit reanalysis for the revised Fo(Z) limit.

0013M890420 5-3

.~

Non-LOCA events that are not mentioned above did not require reanalysis for one or more of the following reasons:

1) Transient results are insensitive to the rod insertion rate.

2) Reactor trip was not assumed or explicitly modeled in the analysis.

3) Reactor trip has no effect on the minimum or maximum value of the critical parameter of interest.

4) F N H and/or Fo are not explicit analysis assumptions.

A summary of the non LOCA design basis calculations that were performed or evaluated for these modifications follows.

5.1.1 Overtemperature and Overpower AT Protection (FSAR Appendix 14D)

As noted in Section 4.7, the current Beaver Valley Unit 1 Technical Specification core thermal limits (Figure 2.1-1 for 3 loops in operation) are valid for VANTAGE SH and STD fuel assemblies at a design FN H of 1.62. The revision of the Fo(Z) does not affect the core thermal limits. On this basis, the current N-loop Technical Specification Overtemperature and Overpower AT (OTDT/OPDT) setpoint equation constants continue to protect the core safety limits as shown in Figure 14D-1 of the Beaver Valley Unit 1 FSAR. As discussed in the following paragraphs, the system transient responses for the FSAR events that rely on OTDT/OPDT for protection are not affected by the increased rod drop time or increased peaking factors. Therefore, the current Beaver Valley Unit 1 Technical Specification values for the OTDT/OPDT setpoints remain valid for the incorporation of VANTAGE 5H fuel and increased N F AH and FO(Z).

0013A:6/890420 5-4

i 5.1.2 increase in Heat Removal by the Secondary System Excessive Heat Removal duelo Feedwater System Malfunctions (FSAR 14.1.9) ]

1 l

This ANS Condition ll event is analyzed to show that the DNB design basis is met. I Cases are analyzed for both full power and zero power conditions. The zero power case, as discussed in the FSAR, is bounded by the Uncontrolled RCCA Bank Withdrawal from Suberitical event. For the full power case, the transient is effectively terminated by a turbine trip and feedwater isolation on high-high steam generator level. Calculation of the thermal-hydraulic conditions at the time of minimum DNBR is not dependent on the full power F N H design value or Fo. A conservative evaluation of the effects of the 0.5 second increase in control rod insertion time was periormed by extrapolating the transient DNBR results assuming that reactor trip was de. layed by 0.5 second. The extrapolation showed that ample margin to the DNB limit still exists with a 0.5 second delay. Therefore, the FSAR conclusions remain valid.

Excessive Load increase Incident (FSAR 14.1.10)

This ANS Condition ll event is analyzed to show that the DNB design basis is met following a step load increase from rated power. Cases are analyzed at BOL and EOL conditions with and without automatic rod control, in all cases analyzed, the reactor j stabilized without a reactor trip. Therefore, the increased control rod insertion time will have no effect on this event. Calculation of the thermal-hydraulic conditions at the time of minimum DNBR is not dependent on the full power FN AH design value or Fo. l Therefore, the conclusions of the FSAR remain valid.

Accidentaillepressurization of the Main SleamJystem (FSAR 14.1.13) and Steamline Bypipres (FSAR 14.2.11)

The inadvertent opening of a steam generator relief or safety valve is an ANS Condition ll event which is analyzed to show that the DNS design basis 10 met. The steam system piping failure is an ANS Condition IV transient analyzed to show that the core remains intact and in place and that the radiation doses do not exceed the ocuwesono 5-5

l J

guidelines of 10CFR100. This is demonstrated by showing that the DNB design basis is )

met, even though DNB and possible clad perforation are not necessarily unacceptable for a Condition IV event.

l The analyses are performed assuming zero power initial conditions and peaking factors J consistent with the most reactive RCCA stuck out of the core. The transient is started ]

assuming the reactor is tripped and the core is at the minimum design shutdown margin. j Therefore, the 0.5 second increase in rod insertion time will have no effect on the results of this analysis. An increase in the full power FN AH limit may result in an increase of the ]

stuck RCCA peaking factor. The effects of the increased stuck rod F N H value have ]

been evaluated. The safety analysis DNBR limits are met. The conclusions of the FSAR remain valid. !

5.1.3 Decrease in Heat Removal by the Secondary System Loss of External Electrical Load and/or Turbine Trio (FSAR 14.1.7)

This ANS Condition ll event is analyzed to show that the DNB design basis is met and that primary and secondary side system pressures do not exceed 110% of design values. Four cases are a;.alyzed:

Beginning of Cycle (BOC) with pressurizer pressure control BOC without pressurizer pressure control End of Cycle (EOC) with pressurizer pressure control EOC without pressurizer pressure control.

l l

The increased rod insertion time to the dashpot will not result in system pressures exceeding 110% of design values. Pressure transients from the current analysis of l record were evaluated by extrapolation assuming the reactor trip was delayed 0.5 l

seconds. In all cases there was ample margin to account for the slight expected pressure rise due to the increased design rod drop times. The increased rod drop time to the dashpot will not result in DNBR below the design limit. The calculation of the thermal-hydraulic conditions at the time of minimum DNBR is not dependent on FNAH Of Fo.

l 0013AM90C0 5-6

ONBR for the BOC case without pressure control and both EOC cases rises continuously throughout the transients. Therefore, the increased insertion time will have no effect on the minimum DNBR for these cases. DNBR during the BOC with pressure control case initially rises and then decreases to a minimum value well above the safety analysis limit at the time of reactor trip. The margin to the design DNB limit is very large at the time of reactor trip for this case and the increased rod drop time will not result in a DNBR below the design basis. Therefore, the FSAR conclusions remain valid for the introduction of VANTAGE 5H fuel and the increased F N H and Fa peaking factors.

Loss of Normal Feedwater (FSAR 14.1.8)

This ANS Condition 11 event is analyzed to show that adequate heat removal capability exists _ via the Auxiliary Feedwater System to remove core decay heat, stored energy and RCS pump heat following reactor trip. This is demonstrated by ensuring that the RCS heatup is turned around prior to the time when coolant expansion causes the pressurizer to become filled with water. The calculated RCS volumetric expansion is not affected by the VANTAGE SH fuel or peaking factors. The Loss of Feedwater transient is a slow long-term heatup event and is not sensitive to the rate at which control rods are inserted

, ' following a reactor trip. The results of the current analysis of record and conclusions of the FSAR remain valid.

Loss of Offsite Power to the Station Auxiliaries (FSAR 14.1.11)

This ANS Condition ll event is analyzed to show that adequate heat removal capability exists via natural circulation flow as aided by the Auxiliary Feedwater System to remove core decay heat and stored energy following reactor trip. This is demonstrated by ensuring that the RCS heatup is tumed around prior to the time when coolant expansion causes the pressurizer to become filled with water. The calculated RCS volumetric expansion is not affected by the VANTAGE SH fuel or peaking f actors. This transient is a slow long-term heatup event and is not sensitive to the rate at which the rods are inserted during a reactor trip. With respect to the DNB criterion, this event is oots w ooo42o 5-7

---1--_---.-_ - - _ _ - - - . - - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . - - - _ _ _ _ _ _ _ _ - - _ _ _ _ _

bounded by the Complete Loss of Forced Reactor Coolant Flow analysis which was reanalyzed and shown to be acceptable. The results of the current analysis of record and the conclusions of the FSAR remain valid. l Major Rupture of a Main Feedwater Line (FSAR 14.2.5.2)

This ANS Condition IV event is analyzed to show that adequate heat removal capability exists using the Auxiliary Feedwater System to remove core decay heat, stored energy and RCS pump heat following reactor trip. This is demonstrated by ensuring that the RCS heatup is turned around prior to the time at which the hotlegs would become saturated.

I The Feedline Break accident is a long term heatup event and is not sensitive to the rate at which the control rods are inserted following a reactor trip. The heatup transient continues for many minutes following the reactor trip. The 0.5 second increase in control ;

rod insertion time will result in an insignificant increase in the integrated heat produced by the core during the transient. No significant increase in hotleg temperature or system pressures would occur due to the increase in control rod insertion time. As an ANS Condition IV event, the minimum DNBR limit acceptance criterion is not applied. The increased design FNAH and Fo values do not affect the calculated RCS response to the Feedline Break accident. The results of the current analysis of record and the conclusions of the FSAR therefore remain valid.

5.1.4 Decrease in RCS Flow Rate Partial and Comotete Loss of Forced Reactor Coolant Flow (FSAR 14.1.5 & 14.2.9) 9 The Partial Loss of Flow accident is an ANS Condition ll event. The Complete Loss of Flow accident is an ANS Condition lli event. Both of these transients have been reanalyzed in support of N-loop operation. The Fartial Loss of Flow transient assumes the coastdown of one RCP during 3 loop, full power operation while the Complete Loss of Flow transient assumes the coastdown of 3 RCPs. The analyses have incorporated 0013A:6/890420 5-8

the VANTAGE SH design RCCA drop time of 2.7 seconds and the increased design FN AH of 1.62 in the determination of the thermal-hydraulic conditions existing at the time of minimum DNBR. There is no explicit transient assumption for Fo.

The results of these two transients are shown in Figures 5.1-1 through 5.1-4 and 5.1-5 through 5.1-8, respectively. The coastdown transients are shown in Figures 5.1-1 and I 5.1-5. Transient calculations were performed to provide a basis for comparison of the analytical pump coastdown characteristics to the plant startup test data. On the basis of this comparison, the analysis calculations are verified to be conservative with respect to actual plant behavior. For both transients, the FACTRAN code (13) is used to calculate the heat flux transient based upon nuclear power and flow from LOFTRAN(14). The Partial Loss of Flow transient is terminated by a low RCS loop flow reactor trip; the Complete Loss of Flow transient is terminated by reactor trip on reactor coolant pump undervoltage. In both cases, the DNBR safety analysis limit is not violated for the VANTAGE 5H and STD fuel assemblies at the design FN AH of 1.62. Therefore, the safety analysis DNBR limits are met and the conclusions of the FSAR remain valid.

Forced reactor coolant pump frequency decay in all three RCPs was also reanalyzed for the VANTAGE SH fuel and the increased FNAH. The transient assumptions for this case are identical to the Complete Loss of Flow case except for the flow coastdown. The Underfrequency analysis assumed a constant frequency decay rate of 5 Hz/second. No credit is taken for RCP trip on underfrequency. The transient is terminated by reactor trip on RCP underfrequency. The transient results indicate that the safety analysis DNBR limit is not violated for the VANTAGE SH and STD fuel assemblies assuming an FNAH design limit of 1.62. Therefore, the cafety analysis acceptance criteria are met for this loss of flow event. It is determined that the underfrequency event is the limiting loss of flow case for these analyzed conditions.

The recommended FSAR markups for the Partial and Complete Loss of Forced Reactor Coolant Flow accidents are included in Appendix B.

1 0013A 6/890420 5-9 L__ _ _ _ _ _ _

Single _BaacioLCoolant Pump Locked RoloL(FSAR 14.2.7)

Reactor Coolant Pump Locked Rotor is an ANS Condition IV event analyzed for determination of peak RCS pressure and peak fuel clad temperature assuming DNB to occur in the core. The accident is postulated as an instantaneous seizure of one reactor coolant pump rotor at full power conditions with N loops in operation. Flow through the faulted reactor coolant pump is rapidly reduced leading to an initiation of a reactor trip on a low flow signal. If the reactor is not tripped promptly, clad temperature may exceed the limit value of 2700 F and RCS pressure may increase above that which would cause stresses to exceed the faulted condition stress limits. Therefore, the Locked Rotor transient can be sensitive to an increase in RCCA drop time.

The Locked Rotor transient was reanalyzed to incorporate a 2.7 second RCCA drop time, a full power FNAH design limit of 1.62 and a maximum Fo(Z) value of 2.4. The FACTRAN code (13) is used to calculate the core hot spot heat flux transient based upon nuclear power and flow from LOFTRAN(14). The results of the analysis are shown in Figures 5.1-9 through 5.1-12 and Table 5.1-1. Cases with and without offsite power were examined. The case without offsite power assumes coastdown of the unaffected RCPs to be initiated at the beginning of the transient. For both cases analyzed, the peak RCS pressure reached during the transient is less than that which would cause stresses to exceed the faulted condition stress limits. Additionally, the peak clad temperatt e calculated for the hot spot remains less then 2700 F and the amount of zirconium-water reaction is small. Therefore, it is concluded that the integrity of the primary coolant system is not endangered and the core will remain intact with no consequential loss of l

core cooling capability.

l l

The recomended FSAR markups for the Single Reactor Coolant Purnp Locked Rotor accident are included in Appendix B.

l l

ootswooo42o 5-10

5.1.5 Reactivity and Power Distribution Anomalles UnCQDitoHed RCCA Withdrawal from a Subchtical Condition (FSAR 14.1.1)

The RCCA Withdrawal from Subcritical accident is an ANS Condition 11 event performed at zero power conditions. It is characterized by a rapid power increase. The power excursion is retarded by Doptr feedback and the transient is terminated by a reactor trip on the power range high neutron flux low setpoint. Due to the high rate at which the power increases, this transient can be sensitive to RCCA drop time. Calculation of the thermal-hydraulic conditions at the time of minimum DNBR !s not dependent on FN H Of Fo.

The FSAR RCCA Withdrawal from Subcritical accident was reanalyzed with a RCCA drop time of 2.7 seconds to the dashpot. Transient results are shown in Figures 5.1-13 through 5.1-15. Figure 5.1-13 shows the neutron flux transient. The neutron flux overshoots the full power nominal value, but this occurs for only a very short time period.

The energy release and the fuel temperature increases are relatively small. The thermal heat flux response, of interest for DNB considerations, is shown on Figure 5.1-14. The beneficial effect of the inherent thermal lag in the fuel is evidenced by a peak heat flux less than the full power nominal value. There is a large margin to DNB during the transient since the rod surface heat flux remains below the design value, and there is a j high degree of subcooling at all times in the core. Figure 5.1-15 shows the response of j the average fuel and cladding temperature. Tha clad fuel temperature increases to a j value lower than the nominal full power value. The minimum DNBR for the VANTAGE l SH and STD fuel assemblies, as determined by the WRB-1 correlation, at all times remains above the limit value.

The core and the RCS are not adversely affected, since the combination of thermal power and the coolant temperature result in a minimum DNBR well above the limiting l value. Thus no fuel or clad damage will occur. l l

The recommended FSAR markups for the uncontrolled RCCA withdrawal from '

Subcritical Condition event are included in Appendix B.

0013A:6/890420 5-11

\

Uncontrolled RCCA Bank Withdrawal at Power (FSAR 14.1.2)

This ANS Condition ll event is analyzed to show that the DNB design basis is met.

9 Various power levels and reactivity insertion rates for both minimum and maximum reactivity feedback are analyzed. The transients are terminated by an Overtemperature AT or High Neutron Flux reactor trip. As previously noted, the Overtemperature AT setpoints are not changed, so that the time the reactor trip setpoint is reached would remain the same. F N H and Fo are not explicit analysis assumptions for this transient and the core thermal limits remain unchanged; therefore, only the increased rod drop time remains to be evaluated. A conservative evaluation of the effects of the 0.5 second increase in control rod drop time was performed by extrapolating the transient DNBR results assuming that the reactor trip was delayed by 0.5 second. The extrapolations showed that ample margin to the DNB limit still exists with a 0.5 second delay. The conclusions of the FSAR remain valid.

Rod Cluster Control Assembly Misoperation (FSAR 14.1.3.14.2.10)

RCCA Misoperation is categorized into four types of events. Three of these are classified as ANS Condition ll events: dropped RCCA, dropped RCCA bank, and statically misaligned RCCA. The fourth, single RCCA withdrawal, is classified as an ANS Condition ill event. The current calculation of the thermal-hydraulic conditions at the time of minimum DNBR for the Condition 11 events is based on a RCCA insertion time which is greater than the VANTAGE SH design insertion time of 2.7 seconds. The calculation of percent rods in DNB for the Condition lli event conservatively does not credit rod insertion due to reactor trip. The effects of the increased FNAH have been

! evaluated for the appropriate DNB acceptance criterion. The analyses for these events do not explicitly assume a value for Fo. In all cases, the conclusions in the FSAR are verified for each reload. Specifically, the DNB acceptance criteria is met for the Condition ll events and the calculated numter of fuel rods experiencing DNB is confirmed to be within the current safety analysis limit of 5% for the Condition ill event.

l l

0013MB90420 5-12

l I

Startup of an inactive Reactor Coolant Loop at an incorrect Temperature (FSAR 14.1.6)

The Startup of an inactive Loop transient is an ANS Condition ll event analyzed to

]

demonstrate that the DNB design basis is met. The transient has been reanalyzed {

incorporating the VANTAGE SH increased design RCCA drop time of 2.7 seconds and )

an FN AH consistent with a full power design limit of 1.62 and part-power multiplier of !

0.3. The transierit has no explicit dependency on Fo. The FACTRAN code (13) is used to calculate the heat flux transient based upon nuclear power and flow from I LOFTRAN(14). The transient results are shown in Figures 5.1-16 through 5.1-19. The transient is terminated by reactor trip on the High Neutron Flux P-8 trip setpoint. The DNBR safety analysis limit is not violated for the VANTAGE SH and STD fuel assemblies at the increased FN AH.' Therefore, the safety analysis DNBR limit is met and the conclusions of the FSAR remain valid.

Uncontrolled Boron Dilution (FSAR 14.1.4)

This ANS Condition ll event is analyzed to show that adequate time exists for operator action to terminate an inadvertent dilution prior to the loss of shutdown margin. The transient is analyzed for Mode 1 in automatic and manual rod control and in Mode 2.

The Mode 1 case for manual rod control assumes reactor trip on Overtemperature AT.

The impact of a 0.5 second increase in rod drop time on an operator action time of approximately 15 minutes is imperceptible. The Mode 1 automatic rod control case does not assume reactor trip. The Mode 2 calculation of available operator action time starts at the time of reactor trip. Therefore, all the FSAR cases are unaffected by the increase in rod drop time. The FSAR Boron Dilution accident reactivity insertion transients are bounded by those examined for the Rod Withdrawal at Power accident. Therefore, the Boron Dilution transient calculation does not include an explicit evaluation for the Condition ll DNB acceptance criterion. The analysis assumptions have no explicit dependency on FN H or Fo. Therefore, the FSAR conclusions remain valid for these modifications.

0013A:6/890420 5-13

f A

Ruoture of a Control Rod Drive Mechanism Housing (FSAR 14.2.6)

The RCCA ejection accident is an ANS Condition IV event that is characterized by a rapid power burst. Due to the speed at which the power increases this transient can be sensitive to the RCCA drop time. The RCCA ejection transients were reanalyzed with a L RCCA drop time (time to dashpot) of 2.7 seconds and a peak Fo(Z) of 2.4. The transients are not sensitive to FNAH. The limiting criteria for this event are:

1) Average fuel pellet enthalpy at the hot spot below 225 cal /gm for unirradiated fuel and 200 cal /gm for irradiated fuel.

2) Average clad temperature below 2700 F.

3) Fuel melting limited to less than the innermost 10 percent of the pellet at the hot j spot. (Melting is assumed to occur at 4900 F for BOL conditions and 4800 F for EOL conditions)

The FSAR beginning and end-of-life cases at hot full power and hot zero power were analyzed for the above acceptance criteria. Table 5.1-2 provides key analysis assumptions and results for each case. The results of the limiting BOL and EOL cases are also shown in Figures 5.1-20 through 5.1-23. In all cases it is shown that the applicable safety analysis acceptance criteria are met.

The recommended FSAR markups for the RCCA ejection accident are included in i Appendix B.

Inadvertent Loading ang! Ooeration with a Fuel Assembly in an improcer Position (FSAR 14 2.8) 4 This ANS Condition ill event addresses the possibility and consequences of one or more fuel pellets having the wrong enrichment or the loading of a fuel assembly without the prescribed amount of burnable poisons. The FSAR concludes that any significant perturbation from the intended core inventory would be detectable due to the resulting l l

00G M B90420 5-14 i

p

+

effects on power distribution. The described peaking factor increases and VANTAGE SH fuel do not affect the ability of core instrumentation to detect unexpected power shapes.

Therefore, the FSAR conclusions remain valid.

- 5.1'.6 increase in Reactor Coolant inventory .

u Sourious Ooeration of the Safety inlection System at Power (FSAR 14.1.16)

' The Spurious Operation of the Safety injection System is an ANS Condition ll event. The transient produces a negative reactivity transient causing a reduction in core power. The power reduction ;causes a decrease in reactor coolant average ~ temperature and consequent coolant shrinkage. Pressurizer pressure and level decrease until the reactor.

is tripped on the' low pressurizer pressure signal. During the transient the DNB ratio never decreases below the initial value, therefore the 0.5 second increase in control rod insertion time' will have no effect on the minimum DNBR. Calculation of the thermal hydraulic conditions on which the DNBR determination is based is not dependent upon the full power F NAH or Fo design values. Therefore, the conclusions of the FSAR remain valid.-

5.1.7 Decrease in Reactor Coolant inventory

' Accidental Deoressurization of the Reactor Coolant System (FSAR 14.1.15)

This ANS Condition ll event is analyzed to show that the DNB design basis is met. This transient is terminated by a reactor trip on Overtemperature AT. Minimum DNBR occurs immediately following reactor trip. A conservative evaluation of the effect of the 0.5 second increase in control rod drop time was performed by extrapolating the transient DNBR results assuming that the reactor trip was delayed by 0.5 second. The I extrapolations showed that abundant margin to the DNB limit still exists with a 0.5 second increase in rod insertion time. Calculation of the thermal-hydraulic conditions at the time of minimum DNBR is not dependent on the full power FAH design value or Fo.

Therefore, the conclusions of the FSAR remain valid.

oois u eeo42o 5-15

5.1.8 Steamline Break Mass and Energy Releases for Postulated Ruptures inside Containment and Equipment Environmental Qualification Outside Containment The limiting Steamline Break transient for core response is found in FSAR Section 14.2.5. The Steamline Break transients analyzed for containment response and equipment qualification, altematively, are designed to maximize break mass and energy releases. The analysis assumptions for these steamline break calculations are not dependent on FNAH or Fo. Additionally, the calculation results are insensitive to the rate d which control rods are inserted. The 0.5 second increase in RCCA drop time would l increase the integrated energy produced by the core by an insignificant amount.

Therefore, the mass and energy releases used in containment response calculations (21) and the mass and energy releases calculated for equipment qualification outside containment (22) remain valid.

l l

{

l I

t 0013AW890420 5-16

.~

TABLE 5.1-1 SINGLE REACTOR COOLANT PUMP LOCKED ROTOR ACCIDENT RESULTS 3 Loops 3 Loops Operating Operating (with Offsite (without Offsite Power) Power)

Maximum Reactor 2597 2642 Coolant System Pressure (psia)

Maximum Clad Temperature ( F) 1795 1870 Core Hot Spot Amount of Zr-H2O at Core 0.269 0.415 Hot Spot (% by Weight) oo13Am9042o 5-17

1

.~

rg.

)

TABLE 5.1-2 !

Rod Cluster Control Assembly Ejection -

1 Accident Results Time in Life BOL BOL EOL EOL Power level 102 0 102 0 Ejected Rod Worth 0.20 0.70 0.21 1.0

(%A K)

Delayed Neutron Fraction 0.55 0.55 0.47 0.47 Feedback Reactivity 1.3 1.74 1.6 3.55 ,

Weighting Trip Reactivity (%A K) 4.0 2.0 4.0 2.0 Fo before Rod Ejection 2.544 -- 2.544 --

Fo after Rod Ejection 7.10 10.0 7.6 21.75 Number of Operational Pumps 3 2 3 2 Max. Fuel Average 4102 2558 3910 3630 Temperature ( F)

Max. Fuel Center 4966 3045 4862 4149 !

i Temperature ( F)

Max. Clad Average 2337 1818 2231 2671 Temperature ( F) l Max Fuel Stored Energy 180 103 170 155 (cal /gm)

Fuel Melt (%) <10 0 <10 0 4

i 0013AW890420 5-18

Figure 5.1-1 Flow Translants for Partial Loss of Flow i Three Loops in Operation, One Pump Coasting Down i

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Figure 5.1-2 l Nuclear Power and RCS Pressure for Partial Loss of Flow Three Loops of Oper?.tlon, One Pump Coasting Down 1.4

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Figure 5.1-3 Average and Hot Channel Heat Flux Transient for Partial Loss of Flow Three Loops in Operation, One Pump Coasting Down 1.4

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1 Figure 5.1-4 DNBR versus Time for Partial Loss of Flow )

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Figure 5.1-5 Core Flow Coastdown versus Time for Three Loops in Operation,-

Three Pumps Coasting Down, Complete Loss of Flow l

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Figure 5.1-6 Nuclear Power Transient and Pressurizer Pressure Transient l'~ ~

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l Figure 5.1-7 Average and Hot Channel Heat Flux Transients for Three Loops in Operation, Three Pumps Coasting Down, Complete Loss of Flow 1.4

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Figure 5.14 DNBR vs Time for Three Loops in Operation, Three Pumps Coasting Down, Complete Loss of Flow l

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1.6-1.4-1.2-1.0 0 1 2 3 4 5 TIME (SECONDS) 0013A:6/890412 5-26

Figure 5.1-9 Flow Transients for Three Loops in Operation, One Locked Rotor 1

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Figure 5.1-10 Reactor Coolant System Pressure Transient for Three Locos in Operation One Locked Rotor 2700 W/0 0FFSITE POWER W OFFSITE POWER - - - -

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e Figure 5.1-11 l

Nuclear Power Transient, Average and Hot Channel Heat Flux Transients for Three I Loops in Operation, One Locked Rotor 1.2-

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Figure 5.1-12 i I

p Maximum Clad and Fuel Centerline Temperatures at Hot Spot for Three Loops in Operation One Locked Rotor

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Figure 5.1-13 Uncontrolled Rod Withdrawal from a Subcritical Condition Neutron Flux versus Time 102 101 --

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4

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Figure 5.1-14 Uncontrolled Rod Withdrawal from a Subcritical Condition i Core Heat Flux versus Time i 1.0

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Figure 5.1-15 Uncontrolled Rod Withdrawal From a Subcritical Condition, Fuel Temperature versus Time

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1. s Figure 5.1-16 Nuclear Power versus Time improper.Startup on an inactive.

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1 Figure 5.1-17 1

l Core Heat Flux versus Time improper Startup of an inactive Reactor Coolant Pump l

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i Figure 5.1 Vessel Average Temperature Versus Time improper Startup of an inactive Reactor Coolant Pump 1

4 700 680-660-W m 640- !

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Figure 5.1-19 '

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RCS Pressure versus Time - Improper Startup of an inactive Reactor Coolant Pump ;

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) 0013AW890412 5-37

_ _ _ _ _-__-____-____________a

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Figure 5.1-20 Nuclear Power Transient BOL-HFP Ejection Accident 3.0

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a.

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C' Figure 5.1-21 Peak Fuel and Clad Average Temperature versus Time BOL-HFP Rod Ejection Accident 6000.

IMELTING TEMPERATURE 4900 Fl 5000.~_ __ _ __ _

- - h l FUEL CENTER TEMPERATURE sooo. - J I

Y 3000.

?

E IFUEL AVERAGE TEMPERATURE l W

2000. -

7 1ooo. '

l CLAD OUTER TEMPERATURE o.

o. 1. 2. 3. 4. 5. 6. 7. 4. 9. 10.

TIME (SECOEDS) 0013A:6/890412 5-39

Figure 5.1-22 Nuclear Power Transient EOL-HZP Rod Ejection Accident 102 101 --

xE WE gz x es 20 0-- .

-6z dS EG su.

10~2-10~'00

0. .50 1.0 1.5 2.0 2.5 3.0 3.5 4.0 TIME (SECONDS) 0013M890420 5-40

Figure 5.1-23 Peak Fuel and Clad Average Temperature versus Time EOL-HZP Rod Ejection Accident l

6000.

l WELTING TEMPERATURE 4800 Fl 5000. ,

l 4000. , l FUEL CENTER TEMPERATURE l g

E d g

W 3000.

E e

2000.

1000. l FUEL AVERAGE TEMPERATURE l [

<> l CLAD OUTER TEMPERATURE l 0.

0. 1. 2. 3. 4 5. 6. 7. 8. 9. 10.

TIME (SEC0llDS) 1 0013A:6/890412 5-41 l

5.2 LOCA Accidents This section summarizes the evaluations performed to assess the effects of the fuel features described in Sections 2.2 and 2.3 and increased peaking factors on the Beaver Valley Unit 1 LOCA analyses. As noted in the VANTAGE 5H Fuel Assembly Report (3),

the low pressure drop Zircaloy grids have no adverse effect on the LOCA analyses due to the mechanical and hydraulic similarity to 17x17 STD fuel. The Zircaloy grids provide a Peak Clad Temperature (PCT) benefit due to increased rewetting when compared with the 17x17 STD grid. Additional d?ferences introduced by the VANTAGE SH assembly, such as slight flow area changes attributed to thimble tube diameter, have been evaluated as having negligible impact on the LOCA analyses. The other upgraded fuel features described in Section 2.3 have been evaluated to have no impact on the LOCA analyses. The only additional items which can potentially affect the LOCA analyses are the increase in design RCCA rod drop time and peaking factor increases.

LARGE BREAK LOCA (FSAR Section 14.3.2)

The Large Break LOCA analysis (15) for Beaver Valley Unit 1, performed using the Westinghouse 1981 Evaluation Model with BASH (M) for 17x17 standard fuel, resulted in a PCT of 1918 F for the limiting Cd=0.4 maximum safeguards break. This analysis was performed using increased peaking factor limits for a design F N AH of 1.62 and a maximum Fo(Z) of 2.40.

As noted in the VANTAGE SH Fuel Assembly Report (3), LOCA reanalysis is not needed in transitioning from 17x17 STD to 17x17 VANTAGE SH without IFMs if there is available margin to the 10CFR50.46 limits to accommodate any LOCA transition core penalty.

The transition from 17x17 STD fuel to 17x17 VANTAGE SH fuel without IFMs results in no transition core peak cladding temperature penalty (3,5),

The introduction of VANTAGE SH fuel results in the eventual replacement of the 17x17 !

Standard fuel Inconel grid material with Zircaloy-4. The effect of the increase in the mass of zircaloy in the core has been evaluated in terms of the impact on FSAR transient l

l 1

0013A:6/89042o 5-42 l

- - - - - - - - - - _-- _-_a

calculations for metal-water reaction. The limiting non-LOCA and LOCA FSAR ;

calculation for metal-water reaction is associated with the Large Break LOCA transient.

Large Break LOCA analyses using BASH have been performed to explicitly account for the presence of the new grid material. The studies have confirmed that the zircaloy grid temperatures are such that the grid contribution to the total core wide Zr-H2O reaction is insignificant. Therefore, the presence of VANTAGE SH fuel has no significant impact on the expected amount of hydrogen generation due to FSAR non-LOCA and LOCA transient metal-water reaction.

An evaluation has been performed, based upon the 1981 Evaluation Model with BASH, to consider any other effects on the analysis due to the VANTAGE SH fuel. The Large Break LOCA evaluation model does not take credit for the negative reactivity introduced by the control rods. Instead, the reactor is brought to a subcritical condition by the presence of voids in the core caused by the rapid depressurization of the RCS. Since credit is not taken for the negative reactivity introduced by the control rods, the increase in design rod drop time will have no effect on the current Large Break Analysis (15),

Based on the discussion given above, the use of VANTAGE SH zircaloy grids and the increase in peaking factors will not result in an increase in the peak clad temperature for Beaver Valley Unit 1 Large Break LOCA analysis (15). Therefore, these changes are acceptable and the resulting peak clad temperature remains within the regulatory limits.

SMALL BREAK LOCA (FSAR SECTION 14.3.1) 1 The Small Break LOCA analysis for Beaver Valley Unit 1(15) predicted a peak clad i temperature of 1802 F using the NOTRUMP Westinghouse Small Break Evaluation Model(17,18). This analysis was performed using increased peaking factor limits for a l design F N H of 1.62 and a maximum Fo(Z) of 2.40. f The only VANTAGE SH Zircaloy grid feature which affects the Small Break LOCA analysis is the increase in design rod drop time. The Westinghouse Small Break model assumes the reactor core is brought to a suberitical condition by the negative reactivity of the control rods. The increase in the design rod drop time to a maximum value of 2.7 l

l l 0013A:6/890420 5-43

seconds is enveloped by the existing Small Break LOCA analysis (151 which was performed with a conservative rod drop time of 3.2 seconds. Therefore, for small break LOCA, the use of VANTAGE SH Zircaloy grids and the increase in peaking factors is acceptable. i i

STEAM GENERATOR TUBE RUPTURE (FSAR SECTION 14.2A)

The steam generator tube rupture (SGTR) accident is analyzed to ensure that offsite doses remain below the limits defined in 10CFR100. The primary thermal-hydraulic factors affecting this conclusion include the extent of fuel failure that occurs during the event, the primary to secondary flow through the ruptured tube and the mass and energy j released to the atmosphere from the steam generator with the ruptured tube. The amount of fuel failure assumed for the SGTR analysis in the Beaver Valley Unit 1 FSAR is 1% and will not change since this assumption is independent of the fuel grid material and peaking factors. The primary to secondary break flow and the mass released to the atmosphere from the ruptured steam generator are dependent upon the RCS and secondary system operating parameters. The introduction of Zircaloy grids does not affect these parameters. Similarly, the increase of the FNAH and Fo safety limits does not change the RCS and secondary thermal and hydraulic design parameters modeled in the analysis. Finally, since a finite time for rod insertion is not considered for the SGTR analysis (i.e., subcriticality is implicitly assumed at the time of reactor trip), an increase in the design rod drop time will not impact the SGTR results. Thus, these changes will have no effect on the Beaver Valley Unit 1 SGTR analysis.

BLOWRHQWRBEACTOR VESSEL AND LOOP FORCES (FSA3 SECTION 14.3.3 and FSAR APPENDIX B)

The major factors in determining the resulting forces from a postulated LOCA on the vessel and the internals are the initial reactor coolant system primary fluid temperature and pressure. Since VANTAGE SH Zirceloy grids (including the associated increase in design rod drop time) and the increase in peaking factors do not change the primary side design temperatures and pressures which are modeled in the forces analysis 00), there will be no effect on the LOCA hydraulic forces. In addition, since the VANTAGE 5H grid is at least as strong as the STD grid (3), there will be no impact on overall mechanical integrity.

C013 M 890C0 5-44

POST LOCA LONG-TERM CORE COOLING. BORON EVALUATION (telated to FSAR SECTION 14.3.2)

The Westmphouse licensing position for satisfying the requirements of 10CFR Part 50 Section 50.46 Paragraph (b) ltem (5) "Long-Term Cooling" is defined in WCAP-8339(20). The Westinghouse commitment is that the reactor will remain shutdown by borated ECCS water residing in the RCS and sump after a LOCA. Since credit for the control rods is not taken for a large break LOCA, the borated ECCS water provided by the accumulators and the RWST must have a concentration that, when i

mixed with other sources of borated and non-borated water, will result in the reactor core remaining subcritical assuming all control rods out.

Since the use of VANTAGE SH Zircaloy grids (including the associated increase in design rod drop time) and the increase in peaking factors will not affect the sources of borated and non-borated water assumed in the long term cooling calculation, it is concluded that there would be no change to the long term cooling capability of the ECCS system. Further, this licensing commitment is checked by Westinghouse on a cycle by cycle basis ensuring compliance with this requirement independent of this safety evaluation.

tiO_T_LFJ SWITCHOVER TO PREVENT POTENTIAL BORON PRECIPITATION Post-LOCA hot leg recirculation time is determined for inclusion in emergency procedures to ensure no boron precipitation in the reactor vessel following boiling in the core. This recirculation time is dependent on power level, and the RCS, RWST, and accumulator water volumes and Doron concentrations. The VANTAGE SH Zircaloy grids (including the associated increase in design rod drop time) and increased peaking factors will have no effect on the assumptions for the RCS, RWST, and the accumulators in the hot leg switchover calculation. Thus, there is no effect on the post LOCA hot leg switchover time.

0013AM0420 5-45

LOCA Containment Integrity (FSAR 14.3.4)

There is no impact on the short term mass and energy and subcompartment pressure analysis since fuel design changes and upgrades, including the increase in design rod I I

drop time and peaking factor increases, have a negligible affect on the transient. For the short term subcompartment analyses, approximately only the first 3 seconds of the blowdown are considered. Because plant initial conditions are unchanged, the 3 seconds of blowdown are negligibly affectea.

The long term mass and energy and containment peak pressure analysis is not adversely affected by the fuel upgrade, including the associated increase in design rod drop time, or increased peaking factors since the plant Tavg remains the same.

Additionally, the VANTAGE SH fuel rod is the same as that used in the STD 17x17 fuel assembly. Since the fuel rod designs are the same, there is no difference in initial core stored energy, and hence no additional energy would be available for release to containment.

On summary, there is no impact on the FSAR LOCA containment integrity analyses due to the increase in peaking factors and the use upgraded Westinghouse fuel features, including the VANTAGE SH fuel design, for Beaver Valley Unit 1.

5.3 Accident Analysis Conclusion Sections 5.0, 5.1 and 5.2 have summarized the impact on the LOCA and non-LOCA design basis calculations of the introduction of the following fuel features and design upgrades in Cycle 8 of Beaver Valley, Unit 1.

- Reconstitutable Top Nozzle

- Debris Filter Bottom Nozzle

- Standardized Fuel Pellet

- Snag-Resistant Grid

- VANTAGE 5 Hybrid Fuel

- Increased Peak Fo (z) and Modified K(z) Curve

- Increased Full Power Design FNAH oots u seo42o 5-46 I

it is concluded that the primary safety analysis impact is due to the VANTAGE 5 Hybrid fuel and the increased peaking factors. In most cases it was found that all the intended Cycle 8 modifications are supported by the existing licensing basis safety analyses. In these cases it was concluded that specific safety analyses are insensitive to the fuel and design upgrades or have otherwise incorporated bounding analyses assumptions, such as the Small and Large Break LOCA analyses. Explicit reanalyses were required for non-LOCA transients sensitive to the increased design RCCA drop time associated with the VANTAGE 5 Hybrid fuel. The WRB-1 correlation has been introduced to evaluate transient DNBRs for both the standard and VANTAGE 5 Hybrid fuel. The MINI RTDP methodology has been applied to maintain the existing core thermal limits with the increased FN H. Non-LOCA transients sensitive to FN H in the determination of thermal-hydraulic conditions at minimum DNBR were reanalyzed. Non-LOCA transients sensitive to Fo were also reanalyzed.

All transient reanalyses and evaluations demonstrate that all applicable safety analysis acceptance criteria continue to be met for the intended fuel and design upgrades that will be introduced in Cycle 8 of Beaver Valley Unit 1.

0013A:6-890420 5-47

t

6.0 REFERENCES

.1. Updated Final Safety Analysis Report, Beaver Valley Power ~ Station, Unit 1, Revision 6, January,1988.

- 2. Davidson, S. L. ed. et al., " Westinghouse Reload Safety Evaluation Methodology,"

WCAP-9273-NP-A, July 1985.

3. Davidson, S. L. ed. et al., " VANTAGE SH Fuel Assembly," WCAP-10444-P-A, Addendum 2, April 1988 and Letter from W. J. Johnson (Westinghouse) to M. W.

Hodges (NRC), NS-NRC 98-3363, dated July 29,1988; " Supplemental Information for WCAP-10444-P-A Addendum 2, ' VANTAGE 5H Fuel Assembly'".

4. Davidson, S. L. ed. et al., " VANTAGE 5 Fuel Assembly Reference Core Report,"

WCAP-10444-P-A, September 1985.

5. Letters from A. C. Thadani (NRC) to R. A. Wiesemann (Westinghouse):

" Acceptance for Referencing of Topical Report WCAP-10444-P-A, Addendum 2,

' VANTAGE 5H Fuel Assembly'", November 1,1988 and Clarifications on the Safety Evaluation of the Topical Report WCAP-10444-P-A Addendum 2,1/5/89.

6. Foley, J. and Skaritka, J., " Operational Experience with Westinghouse Cores,"

(through December 31,1987), WCAP-8183, Revision 16, August,1988.

7. Miller, J. V. (ed.) " improved Analytical Model used in Westinghouse Fuel Rod Design Computations," WCAP-8785, October 1976.

]

8. Weiner, R. A., (et. al.), " improved Fuel Performance Models for Westinghouse Fuel Rod Design and Safety Evaluations," WCAP-10851-P-A, August 1988.

9. F. E. Motley, K. W. Hill, F. F. Cadek, and J. Shefchek, "New Westinghouse Correlation WRB-1 for Predicting Critical Heat Flux in Rod Bundles with Mixing Vane I Grids," WCAP-8762-P-A, July,1984. I oo1SA:6-890Co 8-1 l

Y_-__-__-_-_________-____ .-_ ___ _. l

, I

, l

10. A. J. Friedland and S. Ray, " Revised Thermal Design Procedure," WCAP-11397, February, 1987. (Also Letter, A. C. Thadani (USNRC) to W. J. Johnson l (Westinghouse), " Acceptance for Referencing of Licensing Topical Report WCAP-11397, Revised Thermal Design Procedure," January 17,1989).

11. S. Ray, " MINI Revised Thermal Design Procedure (MINI RTDP)," WCAP-12178-P, March,1989.

12. Letter, D. F. Ross, Jr. (NRC) to D. B. Vassala (NRC), ' Topical Report Evaluation for l

WCAP-8762", April 10,1978.

13. C. Gunin, "FACTRAN, A FORTRAN IV Code for Thermal Transients in a UO2 Fuel

~

Rod," WCAP-7908, Westinghouse Electric Corporation, June 1972.

14. T. W. T. Bumett, C. J. McIntyre, J. C. Baker, R. P. Rose, "LOFTRAN Code Description," WCAP-7907, Westinghouse Electric Corporation, June 1972.

15. Letter from Duquesne Light to NRC Document Control Desk, " Sixth Refueling Outage Plant Modifications - Additional Information," (Submittal of Westinghouse WCAP-11639), December 7,1987.

16. Besspiata, J. J., et. al., 'The 1981 Version of the Westinghouse ECCS Evaluation Model Using the BASH Code," WCAP-10266-P-A, Revision 2 (Proprietary)

WCAP-10337-A (Non-Proprietary), March 1987.

17. Lee, H. Rupprecht, S., Tauche, W., Schwartz, W. R., " Westinghouse Small Break ECCS Evaluation Model Using the NOTRUMP Code," WCAP-10054-P-A, August 1985.

18. Meyer, P. E., "NOTRUMP, A Nodal Transient Small Break and General Network Code," WCAP-10079-P-A, August 1985.

19. Reactor Pressure Vessel and Internals System Evaluation for the Beaver Valley Unit No.1 Upflow Conversion, WCAP-11556.

0013A:6-890C0 6-2

a 0

l

20. Beidelon, F. M., Massie, H. W., and Zordan, T. A., " Westinghouse ECCS Evaluation l- Model Summary," WC AP-8339, July,1974.

21. Letter from Duquesne Light to NRC, Technical Specification Change Request 1 A-79, Attachment C, item 3, "Results of LOCTIC and REFLO Computer Codes,"

April 4,1983.

l

22. J. C. Butler, D. S. Love, et al, " Steam Line Break Mass / Energy Releases for Equipment Environmental Qualification Outside Containment -

Report to Westinghouse Owner's Group" WCAP-10961-P, Rev.1,-October 1985.

0013A:6-890420 6-3

APPENDIX A Summary of Technical Specification Changes 1 .'-

0013A:6/890414 A-1

l

SUMMARY

OF TECHNICAL SPECIFICATION CHANGES FOR FUEL UPGRADE AND INCREASED PEAKING FACTORS FOR BEAVER VALLEY UNIT 1 Eage Section Description Justification B 2-1 2.1.1 Basis Changed W-3 (R-Grid) This change reflects B 2-4 2.2.1 Basis correlation to WRB-1 the DNB correlathn B 2-6 2.2.1 Basis correlation and used for Standard B 3/4 2-1 3/4.2 Basis added design DNBR limits. and VANTAGE SH fuel.

B 3/4 2-5 3/4.2.2 Basis B 3/4 2-5 3/4.2.3 Basis B 3/4 2-6 3/4.2.5 Basis B 3/4 4-1 3/4.4.1 Basis B 2-2 2.1.1 Basis Changed value for FNAH Design includes 3/4 2-8 3.2.3 increase in FN AH-B 3/4 2-4 3/4.2.2 Basis B 3/4 2-4 3/4.2.3 Basis 3/4 1-22 3.1.3.4 Revised rod drop time This change is a result to less than or equal to of changes in the fuel 2.7 seconds, due to the VANTAGE SH fuel design. The effect of this increase on safety analysis has been considered.

0013M890414 A-2

t'

SUMMARY

OF TECHNICAL SPECIFICATION CHANGES FOR FUEL UPGRADE AND INCREASED PEAKING FACTORS (cont.)

Eage Section Description Justification B 3/4 2-5 3/4.2.3 Basis Revised DNBR margin for This change reflects meeting rod bow penalty. change in DNB correl-ation and methods.

3/4 2-5 3.2.2 Revised Fo(Z) limit Design includes 3/4 2-7 Figure 3.2-2 to 2.40 (P > 0.5) increase in Fo(Z).

B 3/4 2-1 3/4.2.1 Basis and 4.80 (P s 0.5).

New Figure 3.2-2.

B 3/4 9-4 3/4.9.14 Basis Changed description cf Design includes spent fuel criticality changes in the fuel analysis to include all due to VANTAGE SH applicable fuel types. fuel design.

oois w eoo4 4 A-3

i APPENDIX A Summary of Technical Specification Changes l

l l

I i

i l

ootsA:stemt4 A-1

_ _ _ ___ _ _ _ __ _ ___- _ -__ _ _ ___ - __ _--- - --_ - - _ -__ _ ______ _ ___ __ _ _-_ ___ -__ - - _ -_ _ _ _____ _--_ _ _ _ _ ___- a

SUMMARY

OF TECHNICAL SPECIFICATION CHANGES FOR i FUEL UPGRADE AND INCREASED PEAKING FACTORS FOR BEAVER VALLEY UNIT 1 Eage SftcliQD Description Justification B 2-1 2.1.1 Easis Changed W-3 (R-Grid) This change reflects

, B 2-4 2.2.1 Basis correlation to WRB-1 the DNB correlation B 2-6 2.2.1 Basis correlation and used for Standard B 3/4 2-1 3/4.2 Basis added design DNBR limits. and VANTAGE 5H fuel.

B 3/4 2-5 3/4.2.2 Basis B 3/4 2-5 3/4.2.3 Basis B 3/4 2-6 3/4.2.5 Basis

'B 3/4 4-1 3/4.4.1 Basis B 2-2 2.1.1 Basis Changed value for FNAH Design includes 3/4 2-8 3.2.3 increase in FN AH-B 3/4 2-4 3/4.2.2 Basis B 3/4 2-4 3/4.2.3 Basis 3/4 1-22 3.1.3.4 Revised rod drop time This change is a result to less than or equal to of changes in the fuel 2.7 seconds. due to the VANTAGE 5H fuel design. The effect i of this increase on safety analysis has been considered.

ooiswoeom A-2 l

)

l

SUMMARY

OF TECHNICAL SPECIFICATION CHANGES FOR PUEL UPGRADE AND INCREASED PEAKING FACTORS (cont.) l

')

Eage Section Description Justification B 3/4 2-5 3/4.2.3 Basis Revised DNBR margin for This change reflects meeting rod bow penalty. change in DNB correl-ation and methods.

3/4 2-5 3.2.2 Revised Fo(Z) limit Design includes 3/4 2-7 Figure 3.2-2 to 2.40 (P > 0.5) increase in Fo(Z).

B 3/4 2-1 3/4.2.1 Basis and 4.80 (P s 0.5).

New Figure 3.2-2.

B 3/4 9-4 3/4.9.14 Basis Changed description of Design includes spent fuel criticality changes in the fuel analysis to includa all due to VANTAGE SH applicable fuel types. fuel design.

l l

ootsA:e/soo454 A-3

l 1

_2 .1 SATETY LIMITS BASES 2.1.1 ftEACTOR CORE The restrictions of this safety limit prevent overheating of the fuel and possible cladding perforation which would result in the, release of fission producto to the reactor coolant. Overheating of the fuel cladding is prevented by restricting fuel operation to within the nucleate boiling regime where the heat transfer coefficient is large and the t4 adding surface temperature is slightly above the coclant saturation temperature..

Operation above the upper boundary of the nucleate boiling regime could result in excessive cladding temperatures because of the onset i

- of departure. from nucleate boiling (DNB) and the resultant sharp

reduction in heat transfer coefficient. DNB is not a directly measurable parameter during operation and therefore THERMAL POWER and Reactor Coolant Temperature and Pressure have been related to DN3 throuch theM/7m correlation. The AVIMEpid4DNB correlation l bd#S~I has been developed to predict 'the DNB f'.ux and thallocation of DNB .

for axially uniform and non-uniform heat flux dis tributions. The local DNB heat' flux ratio, DNBR, defintui_as_t.he r atio of the heat flux that would cause DNB at a particular core location to the local heat flux, is indicative of the margin to DNB. _ ygg,y T n .um v u o thy'D ri s ia s te p at n, o. .a1 an a 'an a ic at d ra ie es is li.it :

ud5Ed( e- ti a s, a

1. v u e re on o 95 en pr ab i a 5 1 re t

i co .i ne el h D w 1 et ec d c se a a pp op a rg t D fo a' o ra n cc it ns The curves of Figure / 2.1-1, My show the loci of points of THERMAL POWER, Reactor Coolant System pressure and average temperature for which the minimum DNBR is no less than((2}{2 or the average enthalpy at the vessel exit is equal to thehenthalpy cf saturated liquid.

4At & L sin DNBR

[

'I BEAVER VALLEY - UNIT 1 B 2-1 Amendment No. 112

BMES '2.1. I REAcfod CofE INSERT 1 The DNB design basis is as follows there must be at least a 95 percent probability that the minimum DNBR of the limiting fuel rod during Condition I and II events is gre'ater than or equal to the DNBR limit of the DNB correlation being used (the WRB-1 correlation in this application). The correlation DNBR limit is established based on the entire applicable experimental data set such that there is a 95 percent probability with 95 percent confidence that DNB will not occur when the minimum DNBR is at the DNBR limit (1,17 for the WRB-1 correlation).

In meeting this design basis, uncertainties in nuclear and thermal parameters, and fuel fabrication parameters were combined statistically with the DNB correlation uncertainties to determine the plant DNBR uncertainty and establish the design DNBR limit such that there is at least a 95% probability with 95% confidence level that the minimum DNBR for the limiting fuel rod is greater than or equal to the DNBR limit. For this application, the design DNBR limit is 1.21. This DNBR value must be met in plant safety analyses using nominal values of the input parameters that were included in the DNBR uncertainty evaluation. In addition, margir w s been maintained in the design by meeting a safety analysis DNBR limit of ...a in performing safety analyses.

l l

8EM62 \/ALLGY- UAlld L

5AFETY LIMIT 1 .

BASES l.62 The curves are based on an enthalpy hot channel An factor, dg. o and a reference cosine with a peak of'a1.55 for axial pow the expression: g l

(g = @[1 + 0.3 (1-P)]

where P is the fraction of RATED THERMAL POWER These limiting heat flux conditions are higher than those calculated for the range of all control rods fully withdrawn to the maximum allowable control red insertion assuming the axial power imbalance When theis within the limits of the f(AI) function of the Overtemperature trip.

axial power imbalance is not within the tolerance, the a to provide protection consistent with core safety limits.

2.1. 2 REACTOR COOLANT SYSTEM PRES $UR_E The restriction of this Safety Limit protects the integrity of the

. Reactor Coolant System from overpressurization a ,

v the contairunent atmosphere.

The reactor pressure vessel and pressurizer are designed to Section 111 of the ASME Code for Nuclear Power Plant which The permits a maximum transient pressure of 110t (2735 psig) of design pressure.

Reactor Coolant System piping and fittings are d The Safety pressure of 1205 (2g85) psig of component design pre .

, associated code requirements. -

The entire Reactor Coolant System is hydrotested at 3107 psig te demonstrate integrity prior to initial operation.

AME::DMEN NO. SS 8-2-2 SEAVER VALLEY - UNIT 1 e

.bh WM tjLIMITINGStFETYSYS~EMSETTINGS- h,

/

bases rotection to ensore The Power Range Negative Rate trip provide or control rod drop that accidents.

the minimum DN5R is maintained aboveAt high power a single or m cause local flux peaking which, when in conjunction with nuclear power being maintained equivalent to turbine power by action of the automatic red contrel system, could cause an unconservative local DNER to exist.

The Power Range Ne;stive Rate trip will prevent this from occurring by h9e neflnok. 508,yah, ... }. - 6r M es etp e b p h n

..,.. .. ......r.. ...,,

nac for,$o. oaA tripp kngo,ing e*NS on the reactar which r

. fMp,on td yestse , A.,

foshldeb Inter-nediate .&handskom Source LfRance.

% m Nuclear Flux f3,p lfsg /,-,K The Intermediate and Source Range,These Nuclear trips Flux trips redundant provide provide reactor core protection during reactor startup.

protection channels.

to the low setpoint trip of the Power Range, Neutro about active.

10+3 counts per second unless manually blocked when a current level proportional to approximately 25 percent of RATED No THERMAL POWER unless manually blocked when P-10 beccmes active. h tiredit was taken for operation of the trips associatec with either t e Intermediate or Source Range Channels in the accident analyses; however, their functional catability at the specified trip settings is recuired by this specification to enhance the overall reliability of the Reacter Protection System.

Overterersture 6T The Overtemperature AT trip provides core protection to prevent DNB for all combinations of pressure, power, coolant temperature, and axial power distribution, provided that the transient is slow 4 secends), and pressure is within the range between the High and Low Pressure reactor trips. This setpcint includes corrections for changes in der.sity and heat capacity of water with temperature and dynamic compensation for piping delays frcm the core' to the loop temperecure detectors.

With normal axial power distribution, this reactor trip '

limit is always belcw the core safety limit as shown on Figures 2.1-1, If axial peaks are greater than design, as indicated 2.1-2 and 2.1-2.

by the difference between tco and bottem pcwer range nuclea Table 2.2-1.

Amendment No. 21

'B 2 a BEAVER VALLEY - UNIT 1

.y.. ..

LIMITINC. SAFETY SYSTEM SETTINGS BASES through the pressurizer safety valves. No credit was taken for operation of this trip in the accident analyses; however, its functiona.1 capability at the specified trip setting is required by this specification to enhance the overall reliability of the Reactor Protection System.

Loss of Flow The Loss of Flow trips provide core protection to prevent DNB in the event of a loss of one or more reactor coolant pumps.

Above 11 percent of RATED THERMAL POWER, an automatic reactor trip will occur if the flow in any two loops drop below 90% of nominal full loop flow. Above 31% (P-8) of RATED THERMAL POWER, automatic reactor trip will occur if the flow in any single loop drops below 90% of .

fje

nominal full loop flow. This 1 tter trip will prevent the minimum value of the DNBR from acino bel during normal operational transients DNSA and anticipated transients When 2 loops are in operation and the M Overtemperature AT trip set point is adjusted to the value specified for all loops in operation. With the Overtemperature AT trip set point adjusted to the value specified for 2 loop operation, the P-8 trip at 66% RATED THERMAL POWER with loop stop valves open and at 71% RATED

" THERMAL POWER with a loop stop valve closed will prevent the minimum value of the DNBR from going belowM during normal operational transients and anticipated transients iwith 2 loops in operation.

Steam Generator Water Level d h OM84 The Steam Generator Water Level Low-Low trip provides core protec-tion by preventing operation with the steam generator water level below the minimum volume required for adequate heat rpmoval capacity. The specified setpoint provides allowance that there will be sufficient water inventory in the steam generators at the time of trip to allow for starting delays of the auxiliary feedwater system.

Steam /Feedwater Flow Mismatch and Low Steam Generator Water Level The Steam /Feedwater Flow Mismatch in coincidence with a Steam Generator Low Water Level trip is not used in the transient and accident analyses but is included in Table 2.2-1 to ensure the functional capa-bility of the specified trip settings and thereby enhance the overall BEAVER VALLEY - UNIT l B 2-6

REACTIVITY CONTROL SYSTEMS

~

'. ROD DROP TIME-LIMITING-CONDITION FOR OPERATION 2.7 3.1.3.4 The individual full lenoth (s < and contro ) rod drop time from l the fully withdrawn position shall be 4 econds from begirining of decay of stationary gripper coil voltage.to lies et entry with:

~

a. Tavg g 541*F, and

b. All reactor coolant pumps operating, gPLICAEILITY: MODE 3.

3 TION:

'. a. With the drop time of any full leEgth rod determined to exceed the above limit, restore the rod drop time to Within the above limit prior to proceeding to MODE 1 or 2. .

b. With the' rod drop times within limits but determined with 2 reactor coolant pumps operating, operation may proceed provided THERMAL POWER is restricted to:

1. < 61% of RATED THERMAL POWER when the reactor coolant stop valves in

. The nonoperating loop are open, or . /

2. 4 66% of RATED THERMAL POWER when the reactor coolant stop valves In the nenoperating loop are closed.

SURVEILLANCE REQUIREMENTS 4.1.3.4 The rod drop time of full length rods shall b'e demonstrated through l measurem'ent prior, to reactor criticality:

a. Fo,r all rods following each removal of the reactor vessel head.

b. For specifically affected individual rods following any maintenance on or modification to the control rod drive system which could affect the drop time of those spcuific rods, and

c. At least once per 18 months. j I

BEAVER VALLEY - UNIT 1 3/4 l-22 AMENDMENT NO. 51

p q l

. l

~

l POWER DISTRIBUTION LIMITS.

i HEAT FLUX HOT CHANNEL FACTOR-F q (Z)

LIMITING CONDITION FOR OPERATION 3.2.2 F (Z) shall be limited by the following relationships:

q 2.40 Fq (Z) f 93:331 [K(Z)) for P >0.5 P

F9 (Z) i [( ] [K(Z)] for Pd 0.5 where P = THERMAL POWER RATED THERMAL POWER and K(Z) is the function obtained from Figure 3.2-2 for a given core height location.

APPLICABILITY: HODE 1 ACTION:

With F (Z) exceeding its limit:

9

a. Reduce TEERMAL POWER at least 1% for each 1%gF (Z) exceeds the limit within 15 minutes and similiarly reduce the Power Range Neutron Flux-High Trip Satpoints within the next 4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months ; POWER OPERATION may proceed for up to a total of 72 hours8.333333e-4 days 0.02 hours 1.190476e-4 weeks 2.7396e-5 months ; subsequent POWER OPERATION may proceed provided the Overpower

.o T Trip Satpoints have been reduced at least 1% for each 1%

F (Z) exceeds the limit. The Overpower'AT Trip Setpoint rkduction shall be performed with the reactor suberitical.

b. Identify and correct the cause of the out of limit condition prior to increasing THERMAL POWER; THERMAL POWFR may then be increased provided Fq(Z) is demonstrated through incore mapping to be within its limit.

1 Amendment No. 9 BEAVER VALLEY - UNIT 1 3/4 2-5 l

l.2 U

m I.O (0 0.I 0) (s.o,l.o)

O (lo.s o.s4) g O O.8 UJ N

w

]

g O.6 (i2,o,o,s4) _

m o

z 0,4 _

l U

O.2 x

0.0 0 2 4 6 8 10 12 14 CORE HEIGHT (FT) i Figure 3.2-2. K(Z)- Normalized Fo(Z) as a Function of Core Height Af.All&R 1/ ALLEY UAltf i 3l+ 2 ~7

POWER DISTRIBUTION LIMITS N

NUCLEAR ENTHALPY NOT CHANNEL FACTOR - F g LIMITING CONDITION FOR OPERATION 3.2.3 F"g sg1 be limited by the following relationship:

F"g i G1 [1 + 0.3 (1-P)] ,

l where P = THERMAL POWER EXTEU"TIGEE"755 APPLICA8ILITY: PCDE 1 ACTION:

With F" g exceeding its limit:

~

a. ReduceTHElpdLPOWERtolessthan505ofRATEDTHEIMALPOWERwithin 2 hours2.314815e-5 days 5.555556e-4 hours 3.306878e-6 weeks 7.61e-7 months and reduce the Power Range Neutron Flux-Migh Trip setpoints to i 555 of RATED THE M AL POWER within the next 4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months ,

b. Demonstrate thru in-core sapping that F" i 24 hours2.777778e-4 days 0.00667 hours 3.968254e-5 weeks 9.132e-6 months after exceeding the l'ait or bs ewithin itsPOWER THEIMAL limit within to less than 55 of RATED THEIMAL POWER within the next 2 hours2.314815e-5 days 5.555556e-4 hours 3.306878e-6 weeks 7.61e-7 months , and

c. Identify and correct the cause of the out of limit condition prior to increasing THEIMAL POWER, subsequent POWER OPERATION may proceed provided that F"g is demonstrated through in-core mapping to be within its limit at a nominal 505 of RATE TMDMAL POWER prior to exceeding this THEIMAL power, at a nominal 755 of RATED THERMAL POWER prior to exceeding this THEIMAL power and within 24 hours2.777778e-4 days 0.00667 hours 3.968254e-5 weeks 9.132e-6 months after attaining M5 or greater RATED THEIMAL POWER.

SEAVER VALLEY - UNIT 1 3/4 2-8 Amendment No. 97

3/4.2 POWER DISTRIBUTION LIMITS BASES 1

The specifications of this section provide assurance of fuel integ-rity during Condition I (Nomal Operation) and !! (Incidents of Moderate t 2:da. Frecuency) events by: (a) maintaining the minimum DNBR in the core y, DNde M Glimiting during nemal operation and in short term transients, and (b) the fission gas release, fuel pellet temperature & cladding g mechanical properties to within assumed design cH teria. In addition, limiting the peak linear power density during Conditions I events pro-vides assurance that the initial conditions assumed for the LOCA analyses are meet and the ECCS acceptance criteria limit of 2200*F is not exceeced.

The definitions of het channel facters as used in these specifi-cations are as follows:

Fn(Z) Heat Flux Hot Channel Factor, is defined as the maximum local heat flux on the surface of a fuel rod at core elevation I

- divided by the average fuel rod heat flux, allowing for man-ufacturing tolerances on fuel pellets and rods.

F"g Nuclear Enthalpy Rise Het Channel Factor, is defined as the ratio of the integral of linear pcwer alcng the red with the highest integrated power to the average rod power.

3/4.2.1 AXIAL FLUX DIFFERENCE (AFD)

The limits on AXIAL FLUX DIFFERENCE assure that the Fn(2) upper bound envelcoe ef 7221 times the nomali:ed axial peaking Yactor is not gg exceeded during Ether nomal operation or in the event of xenon re-distribution following power changes.

Target flux difference is detemined at equilibrium xenon c:ndi-tiens. The full length r:ds may be positioned within the core in acc rdance with their respective insertion limits and shculd be instated near their nomal position for steady state ope-atien at high power levels. The value of the target flux difference obtained uncer these conditiens divided by the fracticn of RATC THERMAL PCWER is tne tar;a-flux difference at RATED THEMAL pCWER fcr the ass:ciated c:r.e.turm; conditions. Target flux differences fer other THEFNAL FCWER levels tre S 3/4 2-1 fcencment Nc. 3, 7 SEAVER VALLEY - UN T 1 e

L

~

TOL'ER Df STRIBUTION L!t*1TS l

BASES l 3/4.2.2 and 3/4.2.3 HEAT FLUX ANC NUCLEAR ENTHALPY HOT CHANNEL FACTORS Fg (2) and F3g i

The limits on heat flux and nuclear enthalpy hot channel factors ensure that 1) the design limits on peak local power density and minimum DNBR are not exceeded and 2) in the event of a LOCA the peak fuel c{ad temperature will not exceed the ECCS acceptance criteria limit of 2200 F..

Each of these hot channel factors are measurable but will noma 11y only be determined periodically as specified in Specifications 4.2.2 and 4.2.3.

This periodic surveillance is sufficient to insure that the hot channel factor limits are maintained provided:

a. Control rod in a single group move together with no individual rod insertion differing by more than 112 steps from the group demand position,

b. Control rod groups are sequenced with overlapping groups as described in Specification 3.1.3.5.

c. The control rod insertion limits of Specifications 3.1.3.4 and 3.1.3.5 are maintained.

d. The axial power distribution, expressed in tems of AXIAL FLUX DIFFERENCE is maintained within the limits.

N The relaxation in F y as a function of THERMAL POWER allows changes in the radial power shape for all permissible rod insertion limits. F"A H will be maintained within its limits provided conditions a thru d above, '

are maintained.

When an F measurement is taken, both experimental error and manufacturing g

55 is the appropriate experimental error tolerance must be allowed for.

allowance for a full core map taken with the incore detector flux mapping system and 35 is the appropriate allowance for manufacturing tolerance.

The specified limit of F" con Si allowance for uncertainties which means that nomal, full p$wer,tains anthree loop operation will result n

r , s91.0..

Il Beaver Valley-Unit 1 B 3/4 2-4 I Amendment No. 73 l

POWER DISTRIBUTION LIMITS BASES

'T - - i r;d i: Ling _______ . . . ..... .. . . . -

avai to- offset this reduction in the generic generic mar The margins, totaling 9.1% DNBR, and com .

, any rod bow p ties (< 1.3% for the worst c y offsets burnup of 24,000 MWD ich occurs at a y 1 This margin includes the follovi

1. Design Limi R o f 1. 3 0 vs . 1. 2

2. Grid ng (K s ) of 0.046 vs. 0.059 3.

DNBR alMultiplier DiffusionofCoefficient 0.865 vs. 0.88 of 0.038 vs. O.

- 5. " itch r 920ticn -

The radial peaking factor F provide assurance that the net (Z) ls' measured periodically to within its limit. channel factor, Fn The F as provided in the limit.

RaNYal Peaking Factor Limit Report xper) forRatedThermR1 Power (F{y(Z), r specification 6.9.1.14 was determined from expected power control maneuvers over the full range of burnup conditions in the core.

5/4.2.4 OUADRANT POWER TILT RATIO The quadrant power tilt distribution ratio limit assures that the radial power analysis. satisfies Radial the power design values used in the power capability distribution startup testing and periodically during power operation. measurements are made during The and limit of 1.02 at which corrective action is required provides DNB linear heat generation tilts.

rate protection with x-y plane power The two-hour time allowance for operation greater than with a tilt condition

~ 1.02 but less than 1.09 is provided identification and correction of a dropped or misaligned rod.to allev event such action does not correct the tilt, In the uncertainty on Fg is reinstated by reducing the margin for the maximum allowed power by 3 percent for each percent of tilt in excess of 1.0.

{

BEAVER VALLEY - UNIT 1 B 3/4 2-5 Mendment No. ){,139 j i

1 BASES 3/4.2.1 AXIAL FLUX DIFFERENCE Insert 2.

Fuel rod bowing reduces the value of DNB ratio. Margin has been maintained between the DNBR value used in the safety analyses (1.33) and the design limit (1.21) to offset the rod bow penalty and other penalties which may apply.

l 1

1

$6 AVE 2 LIALL6Y~ L/AlIf $

__-_-____-_-__-________-a

\ ""

l POW'ER DISTRIBUTION LIMITS BASES 3/4.2.5 DNB PARAMETERS

The limits on the DNB related parameters assure that each of the parameters are maintained within the norral steady The statelimits envelope are of operation assumed in the transient & accident analyses.

consistent with the initial F5AR assumptions and have b anaiyzed transient. g cLwar p h. 24t, The 12 hour1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months periodic surveillance of these parameters through instrument readout is sufficient to ensure that the parameters are restored within their limits following load changes and other expected transient operation.

flow rate is adequate to detect flow degradation and ensure correlation of the flow indication channels with measured flow such that the ind percent flow will provide sufficient verification of flow rate on a 12 hour1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months basis. ,

AT.encment No. 9 B 3/4 2-6 BEAVER VALLEY - UNIT 1

3/4.4 REACTOR COOLANT SYSTEM 8ASES 3/4.4.1 REACTOR C00LANT LOOPS f UN The plant is designed to operat ith all reactor coolant loops in operation and maintain DN8R above during all nomal operations and anticipated transients. In Modes 1 and , with one reactor coolant loop not in operation. THERMAL POWER is restricted to < 31 percent of RATED THERMAL POWER until the Overtemperature AT ri s reset. Either action ensures that the DN8R will be maintained above A loss of flow in two loops will cause a reactor trip if operating above P-7 (11 percent of RATED THERMAL POWER) while a loss of flow in one loop wil cause a reactor trip if operating above P-8 (31 percent of RATED THERMAL POWER).

g ggg In MODE 3, a single reactor coolant loop provides suff cient heat removal i

capability for removing decay heat; however, due to the initial conditions assumed in the analysis for the control rod bank withdrawal from a suberitical condition, two operating coolant loops are required to meet the DNS design basis for this Condition II event.

. inh 0 DES 4 and 5, a single reactor coolant loop or RHR subsystem provides sufficient heat removal capability for removing decay heat; but single failure considerations require that at least two loops be OPERABLE. Thus, if the reactor coolant loops are not OPERABLE, this specification requires two RHR loops to be OPERA 8LE.

The operation of one Reactor Coolant Pump or one RHR pump provides V

adequate flow to ensure mixing, prevent stratification and produce gradual reactivity changes during bcron concentration reductions in the Reactor Coolant System. The . reactivity change rate associated with boron reduction will, therefore, be within the capability of operator. recognition and control.

The restrictions on starting a Reactor Coolant Pump with one or more RCS cold legs less than or equal to 275'F are provided to prevent RCS pressure ,

transients, caused by energy additions from the secondary system, which could exceed the limits of Appendix G to 10 CFR Part 50. The RCS will be protected against overpressure transients and will not exceed the limits.

}

BEAVER VALLEY - UNIT 1 8 3/4 4-1 Amendment ,No. 92

i j

REFUELING OPERATIONS l W Es [ STD/VMfAGG j CH Asl OM The results of the spe ual pool c ticality analysis (August 1986) for Westinghouse fuel in th e of four storage locations show that there is more than 0.3% ma in to the k limit of 0.95 with all uncertainties included. Base onthesensiINitystudycompleted with this analysis, an increase the maximum allowed enrichment for fuel stored in the spent fu storage racks from 4.00 to 4.05 w/o will increase the maxing r, k by less than 0.002. Therefore, with Westinghouse 17 x 17 w.up uel gfenriched at 4.05 w/o stored in the spent fuel racks in three of four storage locations and with all of the assumptions and conservatism presented in the criticality analysis, the maximum rack k,gg will be less than 0.95.

3/4.9.15 CONTROL ROOM EMERGENCY HABITABILITY SYSTEMS The OPERABILITY of the control room amargency habitability system ensures that the control room will remain habitable for operations

~

personnel during and following all credible accident conditions. The ambient air temperature is controlled to prevent exceeding the allowable equipment qualification temperature for the equipment and instrumentation in the control room. The OPERABILITY of this system in conjunction with control room design provisions is based on limiting the radiation exposure to personnel occupying the control room to 5 rem or less whole body, or its equivalent. This limitation is consistant with the requirements of General Design Criteria 19 of Appendix "A", 10 CFR 50.

BEAVER VALLEY - UNIT 1 B 3/4 9-4 Amendment No. "t99-Letter dated 2/5/39

e:m APPENDIX B Recommcuded Modifica*. ions to Beaver Valley Unit 1 FSAR

BVPS-1-UPDATED FSAR Rev. 6 (1/88) l TABLE OF CONTENTS (CONT'D)

Section Title Page 3.3.2.5.2 Rod Cluster Control Assemblies 3.3-28 I 3.3.2.5.3 Power Shaping With the Part-Length Control Rod Bank 3.3-28 3.3.2.5.4 Burnable reicen*fods ' A 5## ** 3.3-28 1 3.3.2.5.5 Peak Xenon Startup 3.3-29 1 3.3.2.5.6 Load Follow Control and Xenon Control 3.3-29 '

3.3.2.5.7 Burnup 3.3-29 3.3.2.6 Control Rod Patterns and Reactivity Worth 3.3-29 3.3.2.7 Criticality of Fuel Assemblies 3.3-31 3.3.2.8 Stability 3.3-33 l 3.3.2.8.1 Introduction 3.3-33 3.3.2.8.2 Stability Index 3.3-34 3.3.2.8.3 Prediction of the Core Stability 3.3-35 3.3.2.8.4 Stability Measurements 3.3-35 3.3.2.8.5 Comparison of Calculations With Measurements - Cycle 1 3.3-36 3.3.2.8.6 Stability Control and Protection 3.3-37 3.3.2.9 Vessel Irradiation 3.3-38

_,_,. 3.3.3 Analytical Methods 3.3-39 3.3.3.1 Fuel Temperature (Doppler)

Calculations 3.3-39 3.3.3.2 Macroscopic Group Constants 3.3-40 3.3.3.3 Spatial Few-Group Diffusion Calculations 3.3-41 3.4 THERMAL AND HYDRAULIC DESIGN 3.4-1 3.4.1 Design Bases 3.4-1 3.4.1.1 Departure from Nucleate Boiling Design Basis 3.4-1 3.4.1.2 Fuel Temperature Design Basis 3.4-2 3.4.1.3 Core Flow Design Basis 3.4-2 1 3.4.1.4 Hydrodynamic Stability Design Bases 3.4-3 3.4.1.5 other Considerations 3.4-3 3.4.2 Description 3.4-3 3.4.2.1 Summary Comparison 3.4-3 3.4.2.2 Fuel and Cladding Temperatures (Including Densification) 3.4-4 3.4.2.'2.1 UO2 Thermal Conductivity 3.4-6 .

3.4.2.2.2 Radial Power Distribution in UO2 i Fuel Rods 3.4-6 !

3.4.2.2.3 Gap Conductance 3.4-7 j 1 3.4.2.2.4 Surface Heat Transfer Coefficients 3.4-8 3.4.2.2.5 Fuel Clad Temperatures 3.4-8 l 3.4.2.2.6 Treatment of Peaking Factors 3.4-8 l 3.4.2.3 Critical Heat Flux Ratio and ;

Departure from Nucleate Boiling Ratio and Mixing Technology 3.4-9 3-3 l L - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - _ - - - - _ _ ----- _

E.

n-L, BVPS-1-UPDATED FSAR Rav. O (1/82) (

l-LIST OF TABLES I' '

Table Title 3.1 SMAh472AST ble Reactor Design Cz.g;r.;;;

3.1-2 Analytic. Techniques in Core Design-3.1-3 Design-Loading Conditions for Reactor Core Components 3.2-1 Maximum Deflections Allowed For Reactor Internal Support Structures

3. 3 -l' Reactor Core Description-3.3-2 Nuclear Design ~ Parameters 3.3-3 . Reactivity Requirements-for Rod Cluster Control Assemblies

3.3-4 Axia1' Stability Index PWR Core with I

, a 12 ft Height

.3-5 . Typical Neutron Flux Levels (n/cm:-sec) at Full Power (Three Loop)

3.3-6. Comparison of Measured and Calculated Doppler Defects 3.3-7 Benchmark Critical Experiments 3.3-8 Saxton Core II Isotopics Rod MY, Axial Zone.6 3.3-9 Critical Boron Concentrations, HZP, BOL 3.3-10 Comparison of Measured and Calculated Rod Worth 3.3-11 Comparison of Measured and Calculated-Moderator Coefficients at HZP, BOL 3.4-1 ReactorDesign.$0$AAV;WA1.able

..r ar. cn

' 3. 4 Thermal Hydraulic Design Parameters for l One'of Three Coolant uoops Out of Service 3.4-3 Void Fractions at Nominal Reactor Conditions with Design Hot-Channel Factors !

l 3-6 <

__ _ _ _ _ _ _. .m i.-

BVPS-1-UPDATED FSAR Rev. 0 (1/82)

LIST OF FIGURES Figure Title 3.1-1 Power Distribution Limits 3.2-1 Fuel Assembly Cross Section 17 x 17 3.2-2 -Fuel Assembly Outline 47 17 3,R-Aa fuel Aste- bly OuHia Va- fH .

3.2-3 Fuel Rod Schemat4c See tyyf Srp pu / Desig .,

3. 2 -3 e fat / fodClad Typical Scheandm b'c % W-Pellet ge sy fuel beYd, Dimdhsion/

3.2-4 s as 7 a Function of Exposure 3.2-5 Representative Fuel Rod Internal Pressure and Linear Power Density for the Lead Burnup 7od as a Function of Time 3.2-6 Lower Core Support Assembly (Core Barrel Assembly) 3.2-7 Upper Core Support Assembly 3.2-8 Plan View of Upper Core Support Structure 3.2-9 Full Length Rod Cluster Control and Drive Rod Assembly with Interfacing Components 3.2-10 Full Length Rod Cluster Control Assembly Outline 3.2-11 Full Length Absorber Rod 3.2-12 (deleted) 3.2-13 Burnable Poison Assembly (Conceptual) .

l 3.2-14 Burnable Poison Rod - Cross Section 3.2-15A Primary Source Assembly 3.2-15B Secondary Source Assembly l 3.2-16 Thimble Plug Assembly 3.2-17 Full length Control Pod Drive Mechanism 3.2-18 Full Length Control Rod Drive Mechanism Schematic 3.2-19 Part Length Control Rod Drive Mechanism 3-8

BVPS-1-UPDATED FSAR Rev. 0 (1/82)

LIST OF FIGURES (CONT'D)

Figure Title 13.3-14 (deleted) 3.3-15 (deleted) 3.3-16 (deleted) 3.3-17 (deleted) 3.3-18 Plow Chart for Determining Spike Model 3.3-19 Predicted Power Spike Due to Single Nonflattened Gap in the Adjacent Fuel 3.3-20 Power Spike Factor as a Function of Axial Position 3.3-21 Maximum F .

x Power vs Axial Height During Normal O eration 3.3-22 R:di:1 "cching ......, .

" Calcul;':icnc xY, ... .g __..__,_

3.3-23 Peak Power During Control Rod Malfunction Overpower Transients 3.3-23A Peak Power During Boration/ Dilution Over-Power Transients 3.3-24 Comparison Between Calculated and Measured Relative Fuel Assembly Power Distribution 3.3-25 Comparison of Calculated and Measured Axial Shape 3,3-26 Measured Value of F for Full Power Rod Configurations 0 3.3-27 Doppler Temperature Coefficient at BOL and EOL vs T EFF or ycle 1 3.3-28 Doppler Only Power Coefficient vs Power Level at BOL and EOL, of Cycle 1 3.3-29 Doppler Only Power Defect vs Percent Power, BOL and EOL, Cycle 1 3.3-30 Moderator Temperature Coefficient BOL, Cycle 1, No Rods 3-10

1 l

BVPS-1-UPDATED FSAR Rev. 0 (1/82)

LIST OF FIGURES (CONT'D)

Figure Title 3.3-31 Moderator Temperature Coefficient EOL, Cycle 1 \

3.3-32 Moderator Temperature Coefficient as a Function of Boron Concentration BOL, Cycle 1, No Rods 3.3-33 Hot Full Power Temperature Coefficient During Cycle 1 for the Critical Boron Concentration 3.3-34 Power Coefficient vs Percent Full Power for BOL and EOL 3.3-35 Power Defect vs Percent Full Power BOL and EOL 3.3-36 Rod Cluster Control Assembly Pattern 3.3-37

- (dc.lsted)[ff icaldo tioA.<GM/

Ac/cscdeda/ SJu(foeens NiNou'*/

of- 604 //E F -- da"/G 'l4"* M "

Design - Trip / ur, o' Sa e~ flan g_

3.3-38 C 1%eve 3.3-39 Normalized Rod Worth vs Percent Insertion, All Rods But One 3.3-40 Axial Offset vs Time PWR Core with a 12 ft Height and 121 Assemblies 3.3-41 XY Xenon Test Thermocouple Response Quadrant Tilt Difference vs Time 3.3-42 Calculated and Measured Doppler Defect and Coefficients at BOL Two-Loop Plant, t

121 Assemblies, 12 ft Core 3.3-43 Comparison of Calculated and Measured Boron Concentration for Two Loop Plant, 121 Assemblies, 12 ft Core 3.3-44 Comparison of Calculated and Measured C Two-Loop Plant with 121 Assemblies, B

12 ft Core 3.3-45 Comparison of Calculated and Measured C in Three-Loop Plant, 157 Assemblies, l$ ft Core 3.4-1 Peak Fuel Average and Surface Temperatures During Fuel Rod Lifetime vs. Linear Power l

l 3-11

l BVPS-1-UPDATED FSAR Rav. 5 (1/87) l SECTION 3 REACTOR 3.1

SUMMARY

DESCRIPTION l i

This chapter describes:

1.' the mechanical components of the reactor and reactor core -

including the fuel rods and fuel assemblies, reactor internals, and the control rod drive mechanisms )

2. the nuclear design

3. the thermal-hydraulic design. y g17X / 7 5/aMa,w/(STA))a M g(p/fj/

I!* 2 5?.".ff35, if.#f,~__ps

_, comprisedofanarrayofffuelassemblies.7{.ch-o . _v..,

m m...-..

....-.....s~..-.

The core is cooled and moderated by light water at a pressure of 2250 psia in the Reactor Coolant System. The moderator coolant contains boron as a neutron poison. The concentration of boron in the coolant is varied as required to control relatively slow reactivity changes including the effects of fuel burnup. Additional boron, in the form o- burnable pci 0- rods, is employed in the first core to establish e desired initia et vity.

Two hundred and sixty-four fuel rods are mechanically joined in a square array to form a fuel assembly. The fuel rods are supported in intervals along their length by grid assemblies which maintain the lateral spacing between the rods throughout the design life of the assembly. The grid assembly consists of an " egg-crate" arrangement of interlocked straps. The straps contain spring fingers and dimples for fuel rod support as well as coolant mixing vanes. The fuel rods consist of slightly enriched uranium dioxide ceramic cylindrical

! pellets contained in slightly cold worked Zircaloy-4 tubing which is plugged and seal welded at the ends to encapsulate the fuel. All fuel rods are pressurized with helium during fabrication to reduce stresses and strains to increase fatigue life.

Fuel assemblies may also contain non-fueled rods. Non-fueled rods may be used in core locaticns where fuel damage has occurred or may occur. The use of non-fueled rods began when fuel inspections performed during the fifth refueling outage identified leaking fuel rods in a peripheral assembly. It was determined that the fuel rod leakage was attributable to baffle jetting.

Inf / Fael 6 a%fle. Absorbers (IFOA 0 o r 3.1-1 6

_._m._ _ _ _ _ _ _ _ _.

i INSERT A The significant new mechanical design features of the VANTAGE SH fuel assembly design are described in References 1 and 2. These features include the following:

- Integral Fuel Burnable Absorbers (IFBAs)

- Axial Blankets (six inches of natural uranium dioxide at both ends of the fuel stack)

- Replacement of six intermediate Inconel grids with Zircaloy grids

- Slightly longer fuel rods and thinner top and bottom nozzle end plates to accommodate extended burnup

- Reconstitutable Top Nozzles (RTNs)

- Redesigned fuel rod bottom end plug to facilitate reconstitution capability

- Reduction in guide thimble and instrumentation tube diameter

BVPS-1-UPDATED FSAR Rev. 5 (1/87)

The solution to this problem, recommended by Westinghouse and used by other utilities, involves fuel assembly reconstitution as a means to allow the insertion of non-fueled rods into a fuel assembly. In the reconstitution process, the fuel rods in positions subject to problem conditions would be removed and replaced with non-fueled rods. The reconstituted fuel assemblies meet essentially the same design requirement as the original fuel assembly, and the use of reconstituted assemblies will not result in a change to existing safety criteria and design limits.

The center position in the assembly is reserved for the in-core instrumentation, while the remaining 24 positions in the array are equipped with guide thimbles joined to the grids and the top and bottom nozzles. Depending upon the position of the assembly in the core, the guide thimbles are used as core locations for rod cluster control assemblies, neutron source assemblies, and burnable pcicen rods. Othcrwisc, thc guide thimbles arc fittcd with plugging dcvices to limit bypass ficw.

g The bottom nozzle is a box-like structure which serves as a bottom structural element of the fuel assembly and directs the coolant flow distribution to the assembly.

The top nozzle assembly functions as the upper structural element of the fuel assembly in addition to providing a partial protective housing for the rod cluster control assembly or other components.

The rod cluster control assemblies consist of individual absorber rods fastened at the top end to a spider assembly. These assemblies contain full length absorber material to control the reactivity of the core under operating conditions.

The control rod drive mechanisms for the full. length rod cluster control assemblies are of the magnetic latch type. The latches are controlled by three magnetic coils. They are so designed that upon a loss of power to the coils, the rod cluster control assembly is released and falls by gravity to shutdown the reactor.

The components of the reactor internals are divided into three parts consisting of the lower core support structure (including the entire core barrel and neutron shield pad assembly), the upper core support structure and the in-core instrumentation support structure. The reactor internals support the core, maintain fuel alignment, limit fuel assembly movement, maintain alignment between fuel assemblies and control rod drive mechanisms, direct coolant flow past the fuel elements and to the pressure vessel head, provide gamma and neutron shielding, and provide guides for the in-core instrumentation, abGothers The nuclear design analyses and locations for evaluations [ establish physical control rods and burnable p;; sons and physical parameters such as fuel enrichments and boron concentration in the coolant such that the reactor core has inherent characteristics which, together with corrective actions of the reactor ' control, 3.1-2

1 BVPS-1-UPDATED FSAR Rev. 5 (1/87) protective and emergency cooling systems, provide adequate reactivity control even if the highest reactivity worth rod cluster control I assembly is stuck in the fully withdrawn position. '

The design also provides for inherent stability against diametral and azimuthal power oscillations.

The thermal-hydraulic design analyses and evaluation establish coolant flow parameters which assure that adequate heat transfer is assured between the fuel cladding and the reactor coolant. The thermal design takes into account local variations in dimensions, power generation, flow distribution and mixing. The mixing vanes incorporated in the fuel assembly spacer grid design induce additional flow mixing between the various flow channels within a fuel assembly as well as between adjacent assemblies.

Instrumentation is provided in and out of the core to monitor the nuclear, thermal-hydraulic, and mechanical performance of the reactor and to provide inputs to automatic control functions.

The reactor core design, together with corrective actions of the reactor control, protection and emergency cooling systems can meet the reactor performance and safety criteria specified in Section 3.2.

E Tc illuctretc the effect arsemblier, Table 3.1-1 presents thermal-hydraulic and mechanical design parameters bet.:cen the Beaver cftheuceofthe17x17 fuel [lw

-c compariccn of- the princiq 1 nuclear, Valley 17 x 17 fuel assemblies, including the effects of fuel densification, -the cacc if Ocaver Valicy . cre to usc 15 x 15 fuc-1

_ _-_ - - . _u.. , s . - , .. .. s. u.

z...,

.s.. > - > ... . > . ,

The effects of fuel densification were evaluated with the methods described in Reference 4Er i t 3 dbse/ vet The analysis techniques employed in he core design are tabulated in Table 3.1-2. The loading conditic ns considered in general for the core internals and components are tabulated in Table 3.1-3. Specific or limiting loads considered for design purposes of the various components are listed as follows:j fuel assemblies in Section 3.2.1.1.2; reactor internals in %ction 3.2.2.3 and Table 4.1-10; neutron absorber rods, burnable pcfcen rods, neutron source rods and thimble plug assemblies in Section 3.2.3.1.3; full-length control rod drive mechanisms in Section 3.2.3.1.4. The dynamic analyses, input forcing functions, and response loadings are presented in Section B.3.

3.1-3 l

l

__- __________m_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ .

l BVPS-1-UPDATED FSAR Rev. 0 (1/82)

REFERENCES FOR SECTION 3.1 J. M. Hellman, et. al., " Fuel Densification, Experimental 3 Results and Model for Reactor Application", WCAP-8218, Westinghouse Engineering Corporation (October 1973).

1. Davidson, S. L. (Ed.), et al, " VANTAGE SH Fuel Assembly,"

WCAP-10444-P-A, Addendum 2-A, February 1989.

2. Davidson, S. L. (Ed.), et al, " VANTAGE 5 Fuel Assembly Reference Core Report," WCAP-10444-P-A, September 1985.

4 3.1-4 :

________________________.______.__.__._____________._____.___m_______________._______________.____________._____m__.__.________.________.__.____________________._a

BVPS-1-UPDATED FSAR R;v. 5 (1/87)

5. The reactor internals, in conjunction with the fuel assemblies, direct reactor coolant through the core to achieve acceptable flow distribution and to restrict bypass flow so that the heat transfer performance requirements can be met for all modes of operation. In addition, the internals provide core support and distribute coolant flow to the pressure vessel head so that the temperature difference between the vessel flange and head do not result in leakage from the flange during the Condition I and II modes of operation. Required in-service inspection can be carried out as the internals are removable. and provide access to.the inside of the pressure vessel.

3.2.1 Fuel 3.2.1.1 Desi n Bases for bof( e 17 x 17 STD ad VANTA(rE CH Suel assemb/Sy b

-The-A f uel rod and fuel assembly design bases are established to entisfy the general performance and safety criteria presented in S ction 3.2 and specific criteria noted below. Fuel rods may be rcplaced by non-fueled rods. For a description of non-fueled rods, cce Section 3.1.

3.2.1.1.1 Fuel Rods The integrity of the fuel rods is ensured by designing to prevent txcessive fuel temperatures, excessive internal rod gas pressures due co fission gas releases, and excessive cladding stresses and otrains. This is achieved by designing the fuel rods so that the following conservative design bases are satisfied during condition I cnd condition II events over the fuel lifetime:

1. Fuel Pellet Temperatures -

The center temperature of the hottest pellet is to be below the melting temperature of the Uo2 (melting point of 5,080*F(1) unirradiated and reducing by 58'F per 10,000 MWD /MTU). While a limited amount of center melting can be tolerated, the design conservatively precludes center melting. A calculated centerline fuel temperature of 4,700*F has been selected as an overpower limit to assure no fuel melting. This provides sufficient margin for uncertainties, as described in Sections 3.4.1.2 and 3.4.2.10.1.

2. Internal Gas Pressure -

The internal gas pressure is less than the nominal coolant design pressure. This conservative limit precludes primary tensile stresses in the clad.

3. Clad Stress - The effective clad stresses are less than that which would cause general yield of the clad. While the clad has some capability for accommodating plastic strain, the yield stress has been accepted ,as a conservative design basis.

3.2-2

BVPS-1-UPDATED FSAR Rev. 4 (1/86) J 3.2.1.1.2 Fuel Assembly Structure

-Structural integrity of the ' fuel -assemblies is' assured by setting limits on stresses and deformations due to- various loads an'. by determining of other components.

that the assemblies do not interfere with the functioning Three types of loads ~are considered.

1. Non-operational loads such as those due to shipping and handling,

2. Normal and abnormal loads which are defined for Conditions I and II,

3. Abnormal loads which are defined for conditions III and IV.

1 i

These criteria are applied to the design and evaluation of the top "

and bottom nozzles, the guide thimbles, the grids and the thimble joints. i l

The design bases for evaluating the structural integrity of the fuel 0** *******

tl avial a d Gg la}etal

1. Non-Operational - loading with dimensional stability,

2. Normal Operation (Condition I) and Incidents of Moderate Frequency (Condition II),

For the normal operating (Condition I) and upset conditions (condition II), the fuel assembly. component structural design criteria are classified into two material categories, i

namely, austenitic steels and zirealoy. The stress categories and strength theory presented in the ASME Boiler and Pressure Vessel Code,Section III, are used as a general guide. The maximum shear-theory (Tresca criterion) for combined stresses is used to determine the stress intensities for the austenitic steel components. The stress intensity is defined as the numerically largest difference l

between the various principal stresses in a three dimensional field. The allowable stress intensity value for austenitic steels, such as nickel-chromium-iron alloys, is given by the lowest of the following:

a. 1/3 of the specified minimum tensile strength or 2/3 of the specified minimum yield strength at room temperature; i 3.2-4

-____ _____ - __ _____ _ __ _ _ 2

BVPS-1-UPDATED FSAR Rev. 4 (1/86) '

4

. Zircaloy components the ' stress limits are set at two-thirds of the material yield strength, Su, at reactor operating temperature. This results in zircaloy stress intensity limits being the smaller of 1.6 sy (0.2% offset yield strength) or .70 Su for primary membrane and 2.4 Sy or 1.5 su for primary membrane plus bending.- For conservative purposes, the Zircaloy unirradiated properties are used to define the stress limits. The grid component strength criteria are based on experiment,al tests.____,Th:___mlimit i,: ::::blich t .9 m___ _ _m_ _, __, ____ ,__m co o+

Yned Yeue men s+teY~ 5N15 Is ~$'s550[oNe55&5celcn' a

13bikbe:, ofpi crus N $$7en9k Wobi of P,0 eArbe, e level 4:'

fuel assembly stress and Ioad limits usdd for the ' w rat-c feu d condition for a combined seismic and blowdown accident e as follows.

1. Top No 0,000 psi (elastic analysi

2. Bottom Nozzle 00 psi (elast analysis)

3. Thimble = 1.5 Sm

4. Ft$elRod=1.5Sm

5. Grid = elas stability (no buckling) considering -analysis methods described in WC 5000, it-is conclu -thtt the BVPS-1 fuel assembly stresses duri seismic and b own conditions will be less than the stress a oad 1 . The BVPS-1 has DBE forces substantially less than the O.

analyzed in WCAP-7950.

3.2.1.2 Design Description The :tc.dr'. A fuel 17% 17 STD and ass MMG f//

y and fuel rod design data are given in Table 3.3-1.

]

Two hundred and sixty-four fuel nds, twenty-four guide thimble tubes and one instrumentation thimble etbe are arranged within a supporting structure to form a fuel assembly. The instrumentation thimble is located, in the center position and provides a channel for insertion of- an incore neutron detector if the fuel assembly is located in an instrumented core position. The guide thimbles provide channels for

'incertion of either a rod cluster control assembly, a neutron source I accembly, a burnable pei;;. assembly or a plugging device, depending on the position of the particular fuel assembly in the core. Figure as.,

3.2-6

i BVPS-1-UPDATED FSAR Rev. 4(jfgg) i 3.2-1 shows a cross-section of the fuel assembly array, and Figure 3.2-2 shows a fuel assembly full length view. The fuel rods are loaded into the fuel assembly structure so that there is clearance ,

between the fuel rods ends and the top and bo.ttom nozzles. All ,

strrdsrd fuel :: --5110: in the :::: ::: idcatical- in sechanical T u.mnn b. . ..

D J S W A -^>

Each fuel assembly is installed vertically in the reactor vessel and stands upright on the lower core plate, which is fitted with alignment pins to locate and orient the assembly. After all fuel assemblies- are set in place, the upper support structure is installed. Alignment pins, built into the upper core plate, engage and locate the upper ends of the fuel assemblies. The upper core plate then bears downward against the fuel assembly top nozzle via the holddown springs to hold the fuel assemblies in place.

3.2.1.2.1 Fuel gods /

ad JPh I?M 17 SYD and VhifME SW - - l The A fuel rods consist of uranium ioxide ceramic pellets contalh9 d in slightly cold worked Zirealoy- tubing which is plugged and ' seal welde t the ends to encapsul te the fuel. ,XJbchematicrof the)ffuel 1 rods shown in Figures 3.2-3. The fuel pellets are right circular cylinders consisting of : lightly enriched uranium-dioxide powder which has been compacted by cold pressing and then sintered to the required density. The ends of each pellet are dished slightly to allow greater axial expansion at the center of the pellets.

IAIS&it7~ C ->

To avoid overstressing of the cladding or seal welds, void volume and clearances are provided within the rods to accommodate fission gases released from the fuel, differential thermal expansion between the cladding and the fuel, and fuel density changes during burnup.

shifting of the fuel within the cladding during handling or shipping prior to core loading is prevented by a stainless steel helical :

spring which bears on top of the fuel. At assembly the pellets are '

stacked in the cladding to the required fuel height, the spring is then inserted into the top end of the fuel tube and the end plugs pressed into the ends of the tube and welded. All fuel rods are internally pressurized with helium during the welding process in order to minimize compressive clad stresses and creep due to coolant operating pressures. The helium pre-pressurization may be different for each fuel region. Fuel rod pressurization is dependent on the plahued fuel burnup as well as other fuel design parameters and fuel characteristics (particularly densification potential).

The cold helium design pressure for current Westinghouse PWR fuel rods is several hundred psi. The precise design pressure values for-BVPS-1 fuel regions depend on detailed performance evaluations. Such detailed design information is proprietary to Westinghouse and will not be included in the Updated Final Safety Analysis Report.

3.2-7

INSERT 8 The VANTAGE SH fuel assembly design is shown in Figure 3.2-2a.

The design changes from the 17x17 STD to the VANTAGE SH design include reduced guide thimble and instrumentation tube diameters and replacement of the six intermediate (mixing vane) onel - Do**k grids with Zircaloy grids. The debris filter bottom DFBN) design has been incorporated into the VANTAGE 5H fuel assemblies. The DFBN sign is similar to the standard bottom nozzle design except that 's it thinner and has a new pattern of smaller flow holes.

The DFBN ielps to minimize passage of debris particles that could cause fretting damage to fuel rod cladding.

The VANTAGE SH assembly has the same cross-sectional envelope as the 17x17 STD fuel assembly. However, the VANTAGE SH assembly overall length has been increased to accommodate extended burnup.

INSERT C The VANTAGE SH fuel rod is of the same design as the 17x17 STD fuel rod except that the VANTAGE SH fuel rod is longer to provide a longer plenum and bottom end plug. The bottom end plug has an internal grip feature to facilitate rod loading on both designs.

The bottom end plug is also longer to provide an improved lead-in for the removable top nozzle reconstitujon feature.

The Beaver Valley Unit 117x17 STD and VANTAGE SH fuel has axial olankets and Integral Fuel Burnable Absorbers (IFBAs) and will use a standardized fuel pellet design.

The standardized fuel pellet design is a refinement to the .

chamfered pellet design. The standard design helps to improve manufacturability while maintaining or improving perfomance (e.g.,im handling) proved pellet chip resistance during manufacturing and The IFBA coated pellets are identical to the enriched uranium dioxide pellets except for the addition of a thin boride coating on the pellet cylindrical surface. Coated pellets occupy the central portion of the fuel column. The number and pattern of IFBA rods within an assembly may vary depending on specific application. The ends of the enriched coated pellets and enriched uncoated pellets are dished to allow for greater axial expansion at the pellet centerline and void volume for fission gas release.

An evaluation and test program for the IFBA design features are given in Section 2.5 of Reference 20.

The axial blankets are a nominal six inches of natural uranium dioxide pellets at each end of the fuel rod pellet stack. Axial blankets reduce neutron leakage and improve fuel utilization. The axial blanket pellets are of the same design as the enriched and i IFBA pellet designs exce)t for an increase in length. The length difference in the axial 31anket pellets will help prevent accidental mixing with the enriched and IFBA pellets.

.-----_----_---J

BVPS-1-UPDATED FSAR Rev. 2 (1/84) 'l The fuel rods are designed such that: +

1. The internal gas pressure will not exceed the nominal system coolant pressure even during anticipated transients (Condition II) 2.- clad' flattening will not occur during the fuel core life.

3.2.1.2.2 Fuel Assembly Structure The fuel- assembly structure consists of a bottom nozzle, top nozzle, guide' thimbles and grids, as shown in Figure 3.2-2.

Bottom Nozzle The bottom nozzle is a box-like structure which serves as a bottom

, atructural element of'the fuel assembly and directs the coolant flow distribution to .the assembly. The square nozzle is fabricated from type 304 stainless steel and consists-of a perforated plate and four angle legs with bearing plates as shown in Figure 3.2-2. The legs form a plenum for the inlet coolant flow to the fuel assembly. The plate itself acts to prevent a downward ejection of the fuel rods from their fuel assembly. The bottom nozzle is fastened to the fuel assembly guide. tubes by locking cup or weld-locked screws which penetrate..through the nozzle and mate with an inside fitting in each guide tube.

1~AlSERV b -9

- oolant flow through the fuel assembly is directed from the plenum in the bottom nozzle upward through the penetrations in the plate to the channels. between the- fuel rods. The penetrations in the plate are L positioned between the rows of the fuel rods.

Axial loads (holddown) imposed on the fuel assembly and the weight of the fuel assembly are transmitted through the bottom nozzle.to the lower core plate. Indexing and positioning of the fuel assembly is controlled by alignment holes in two diagonally opposite bearing plates which mate with locating pins in the lower core plate. Any lateral loads on the fuel assembly are transmitted to the lower core plate through the locating pins.

Top Nozzle The top nozzle assembly functions as the upper structural element of the fuel assembly in addition to providing a partial protective housing for the rod cluster control assembly or other components. It

-consists' of an adapter plate, enclosure, top plate, and pads. The

~ integral velded assembly has holddown springs mounted on the assembly as shown in Figure 3.2-2. The sp-ings and bolts are mado of ~7eonel 718 and Inconel 600 respectively, whereas other components are made of type 304 stainless steel.

ZAJSEA7 5 4 3.2-8 i

INSERT D The Debris Filter Bottom nozzle (DFBN) design has been introduced into the Beaver Valley Unit 1 Region 10 fuel assemblies to help reduce the possibility of fuel rod damage due to debris-induced fretting. The 304 stainless steel DFBN is similar to the conventional bottom nozzle design used previously for Beaver Valley. However, the DFBN design incor hole size and pattern (described below)porates and a decreaseda modified nozzle flow height and thinner top plate to accommodate the high burnup fuel rods. The DFBN retains the design reconstitution feature that facilitates easy removal of the nozzle from the fuel assembly.

The relatively large flow holes in a conventional bottom nozzle are replaced with a new pattern of smaller flow holes in the DFBN.

The holes are sized to minimize passage of debris particles large enough to cause damage. The hole sizing was also designed to provide sufficient flow area, comparable pressure drop, and continued structural integrity of the nozz?:. Tests to measure pressure drop and demonstrate structural integrity have been performed to verify that the DFBN is totally compatible with the current design.

INSERT E The Reconstitutable Top Nozzle (RTN) design for the VANTAGE SH fuel assembly differs from the conventional design in two ways:

a groove is provided in each thimble thru-hole in the nozzle plate to facilitate attachment and removal; and the nozzle plate thickness is reduced to provide additional axial space for fuel rod growth. Additional details of this design feature, the design bases and evaluation of the reconstitutable top nozzle are given in Section 2.3.2 in Reference 20.

BVPS-1-UPDATED FSAR Rev. 2 (1/84)

The square adapter plate is provided with penetrations to permit the flow of coolant upward through the top nozzle. Other round holes are provided to accept sleeves which are welded to tubes.

The ligaments in the plate cover the tops of the fuel rods and prevent their upward ejection from the fuel assembly. The enclosure is a sheet metal shroud which sets the distance between the adapter plate and the top plate. The top plate has a large square hole in the center to permit access for the control rods and the control rod spiders. Holddown springs are mounted on the top plate and are fastened in place by bolts and clamps located at two diagonally opposite corners. On the other two corners integral pads are positioned which contain .;

alignment holes for locating the upper end of the fuel assembly.

Guide and Instrument Thimbles dhSo/bte The guide thimbles are structural members which also provide channels for the neutron absorber rods, burnable poi cn rods or neutron source assemblies. Each one is fabricated from Zircaloy-4 tuoing having two different diameters. The larger diameter at the top provides a relatively large annular area to permit rapid insertion of the control rods during a reactor trip '

as well as to accommodate the flow of coolant during normal operation. Four holes are provided on the thimble tube above the dashpot to reduce the rod drop time. The lower portion of the guide thimbles has a reduced diameter to produce a dashpot action and to accommodate the outflow of water from the dashpot during a reactor trip. The dashpot is closed at the bottom by means of an end plug which is provided with a small flow port to avoid fluid stagnation in the dashpot volume during normal operation. The top end of the guide thimble is fastened to a tubular sleeve by three expansion swages. The sleeve fits into and is welded to the top nozzle adapter plate. The lower end of the guide thimble is fitted with an end plug which is then fastened into the bottom nozzle by a locking cup or veld-locked screw.

The central instrumentation thimble of each fuel assembly is not attached to either the top or bottom nozzles, but the thimble is constrained by its seating in counterbores of each nozzle. The thimbles internal diameter dces not vary, and in-core neutron detectors pass through the bottom nozzle's large counterbore into the center thimble.

TMM7' f - '>

Grid Assemblies The fuel rods, as shown in Figure 3.2-2, are supported laterally at intervals along their length by grid assemblies which maintain the lateral spacing between the rods throughout the design life of the assembly. Each fuel rod is afforded lateral support at six contact points within each grid by the combination of support dimples and springs. The grid assembly consists of individual slotted straps interlocked and brazed in an " egg-crate" arrangement to join the straps permanently at their points of 3.2-9

INSERT F With the exception of a reduction in the guide thimble diameter and increased length above the dashpot, the VANTAGE SH guide thimbles are identical to those in the 17x17 STD design. A reduction in the guide thimble outside and inside diameters is required due to the thicker Zircaloy grid straps. The VANTAGE SH guide thimble tube ID provides adequate clearance for the control rods. The reduced VANTAGE SH thimble tube also provides sufficient diametral clearance for burnable absorber rods, source rods, and dually compatible thimble plugs. The thimble plugs used before Region 10 are not the dually compatible type and cannot be inserted into the VANTAGE 5H guide thimbles.

The VANTAGE SH instrumentation tube diameter has also been decreased relative to the 17x17 STD assembly instrumentation tube.

This decrease still allows sufficient diametral clearance for the flux thimble to traverse the tube without binding.

The top Inconel grid sleeve, insert, and thimble of the VANTAGE SH design are joined together using three bulge joint mechanical l attachments similar to that used in the 17x17 STD design. This bulge joint connection was mechanically tested and was found to meet all applicable design criteria.

The intermediate Zircaloy grids of the VANTAGE 5H design employ a single bulge connection to the sleeve and thimble as compared to a double bulge connection used in the Inconel grids. Mechanical testing of this bulge joint connection showed that it meets all applicable design criteria.

INSERT G y kk The VANTAGE SH fuel assemblies utilize Zircaloy intermediate grids. The top and bottom (non-mixing vane) g ids are the conventional Inconel grid design. The VANTAG" 5H Zircaloy grid (Reference 21) is based on the OFA Zircaloy gted design and operating experience. The grip strap thickness, type of strap welding, basic mixing vane design and pattern, method of thimble tube attachment, type of fuel rod support (6 point), material and envelope are identical to the 0FA Zircaloy grid.

BVPS-1-UPDATED FSAR Rev. 0 (1/82) intersection. The straps contain spring fingers, support dimples and mixing vanes.

17 % I7 STD Fuel ASS em blv TheAgrid material is Inconel 718, chosen because of its corrosion resistance and high strength properties. The magnitude of the grid restraining force on the fuel rod is set high enough to minimize possible fretting, without overstressing the cladding at the points of contact between the grids and fuel rods. The grid assemblies also allow axial thermal expansion of the fuel rods without imposing restraint sufficient to develop buckling or distortion of the fuel rods.

Two types of grid assemblies are used in each fuel assembly. One type, with mixing vanes projecting from the edges of the straps into the coolant stream, is used in the high heat flux region of the fuel assemblics to promote mixing of the coolant. The other type, located at the ends of the assembly, does not contain mixing vanes on the internal straps. The outside straps on all grids contain mixing vanes which, in addition to their mixing function, aid in guiding the grids and fuel assemblies past projecting surfaces during handling or during loading and unloading of the core.

TNSE1E7 G -=>

3.2.1.3 Design Evaluation 3.2.1.3.1 Fuel Rods The fuel rods are designed to assure the design bases are satisfied for Condition I and II events. This assures that the fuel performance and safety criteria (Section 3. 2) are satisfied.

~~

Materials - Fuel Cladding The desired fuel rod cladding is a material which has a superior combination of neutron economy (low absorption cross section), I high strength (to resist deformation due to differential pressures and mechanical interaction between fuel and clad), high corrosion resistance (to coolant, fuel and fission products), and high reliability. Zircaloy-4 has this desired combination of cladding properties. As shown in Reference 8, there is considerable PWR operating experience on the capability of Zircaloy as a cladding material. Clad hydriding has not been a significant cause of clad perforation since current controls on fuel contained moisture levels were instituted.(81 I/JS BC T % -+ From Page 3,5L - U Materials - Fuel Pellets Sintered, high density uranium dioxide fuel reacts only slightly with the cladding, at core operating temperatures and pressures.

In the event of cladding defects, the high resistance of uranium dioxide to attack by water protects against fuel deterioration although limited fuel erosion can occur. As has been shown by 3.2-10 i

BVPS-1-UPDATED FSAR Rev. 0 (1/82) operating experience and extensive experimental work, the thermal design parameters conservatively account for changes in the thermal performance of the fuel elements due to pellet fracture which may occur during power operation. The consequences of defects in the cladding are greatly reduced by the ability of uranium dioxide to retain fission products including those which are gaseous or highly volatile.

Observations from several operating Westinghouse PWR's has shown that fuel pellets can densify under irradiation to a density higher than the manufactured values.l81 Fuel densification and subsequent incomplete settling of the fuel pellets results in local and distributed gaps in the fuel rods.

An extensive analytical and experimental effort f'I was underway by Westinghouse in 1973 to characterize the fuel densification phenomenon and identify improvements in pellet manufacturing to eliminate or minimize this anomaly (see Section 1. 5) .

The effects of fuel densification have been taken into account in the nuclear and thermal-hydraulic design of the reactor described herein in Sections 3.3 and 3.4, respectively.

Metallographic examination of irradiated commercial fuel rod 7 have shown occurrences of fuel / clad chemical interaction.

Reaction layers of < 1 mil in thickness have been observed between fuel and clad at limited points around the circumference.

Westinghouse metallographic data indicates that this interface IN5&F layer remains very thin even at high burnup. Thus, there is no X indication of propagation of the layer and eventual clad g penetration.

fa$8 Stress corrosion cracking ir another phenomenon related to p_g fuel / clad chemical interaction. Out-of-reactor tests as well as operational experience have shown that in the presence of high clad tensile stresses, large concentrations of iodine will chemically attack the Zircaloy tubing and can lead to eventual clad cracking.

Materials - Strength Considerations one of the most important limiting factors in fuel element duty is the mechanical interaction of fuel and cladding. This fuel-cladding interaction produces cyclic stresses and strains in the cladding, and these in turn consume cladding fatigue life. The reduction of fuel-cladding interaction is therefore a principal goal of design. In order to achieve this goal and to enhance the cyclic operational capability of the fuel rod, the technology for using pre-pressurized fuel rods in Westinghouse PWR's has been developed.

3.2-11

BVPS-1-UPDATED FSAR Rev. 0 (1/82) 3.2.3 Reactivity Control System 3.2.3.1 Design Bases Bases for temperature, stress on structural members, and material compatibility are imposed on the design of the reactivity control components.

3.2.3.1.1 Design Stresses The reactivity. control system is designed to withstand stresses originating from various operating conditions such as those summarized in Table 4.1-10, as well as from conditions which directly involve the components of the-reactivity control system.

Allowable Stresses For normal operating conditionsSection III of the ASME Boiler and Pressure Code is used. All components are analyzed as Class I components under Article NB-3000.

Dynamic Analysis The cyclic stresses due to dynamic loads and deflections are combined with the stresses imposed by loads from component weights, hydraulic forces and thermal gradients for the determination of the total stresses of the reactivity control system.

3.2.3.1.2 Material Compatibility Materials are selected for compatibility in a pressurized water reactor environment, for adequate mechanical properties at room and operating temperature, for resistance to adverse property changes in a radioactive environment, and for compatibility with interfacing components.

3.2.3.1.3 Reactivity Control Components The reactivity control components are subdivided into two categories:

1. Permanent devices used to control or monitor the core

2. Temporary devices used to control or monitor the core.

The permanent ;;S cocm nents are the full length rod cluster control assemblies, control rod drive assemblies, neutron source assemblies, and thimble plug assemblies. Although the thimble plug assembly does not directly contribute to the reactivity ' control of the reactor, it is presented as a reactivity control system component in this document because it is n :d:d to restrict bypass flow through those thimbles not occupied by absorber, source or burnable poison rods.

may be usel 3.2-40

BVPS-1-UPDATED FSAR Rev. 3 (1/85) abset N The temporary component is the burnable ;;inc sembly which is normally reload as to used in thefuel optimize initial core andBurnable loading. could pe used for any r cicen assemblies described in Section 3.2.3 utilize borosilicate glass absorber material. Wet Annular Burnable Absorber (WABA) rods may be used instead of the standard borosilicate glass absorber rods. The WABA red design consists of annular pellets of aluminum oxide-boron carbide (A17 03-B 4C) burnable absorber material encapsulated within two concentric Zircaloy tubings. The reactor coolant flows inside the inner tubing and outside the outer tubing of the annular rod. Details of ' the WABA design are described in Reference 19. The design bases for the rest of the mentioned components are in the following paragraphs.

Absorber Rods The following are considered design conditions under Article NB-3000 of the ASME Boiler and Pressure Vessel Code Section III. The control rod which is cold rolled 304 stainless is the only non code material used in'the control

rod assembly.

The stress intensity limit Sm for this material is defined at 2/3 of the 0.2% offset yield stress.

1. The external pressure equal to the reactor coolant system operating pressure

2. The wear allowance equivalent to 1.,000. reactor trips

3. Bending of the rod due to a misalignment in the guide tube

4. Forces imposed on the rods during rod drop

5. Loads caused by accelerations imposed by the control rod drive mechanism

6. Radiation exposure for maximum core life.

The absorber material temperature shall not exceed its melting point, which is (1,470 F for Ag-In-Cd.(21 (30y ss for As fcroSll1cale lQ SS des ,

/)bSofher Me4'f n) yn Burnable !'ci:c.- Rods aw/ ahica/,y / for de M e

The absetkt' 2 burnable pel;;n rod clad /lis designed as a Class 1 component under Article NB-3000 of the 1971 ASME Boiler and Pressure Vessel Cc~de ,Section III, for Conditions I and II. For abnormal loads during Conditions III and IV, code stresses are not considered ;

limiting. Failures of the burnable p;ican rods during these conditions must not interfere with reactor shu own or emergency i

cooling of the fuel rods.

gggy The burnable p c i r c.- absorber material, borosilicate glass, is non-structural. The structural elements of the burnable 4ee-rod are designed to maintain the absorber geometry even if 1

3.2-41

BVPS-1-UPDATED TSAR Rev. 3 (1/85) botos)llca}s the abso material' is fractured. The- rods are designed so that the absorber material is below its softening temperature,

_1,492 F for reference 12.5 weight percent (w/o) boron rods. In addition, the structural elements are designed to prevent excessive slumping. Borosilicate glass is accepted for use in burnable pc ' cr. rods if the softening temperature is 1,510 1-18 F. The softening temperature is defined in ASTM C-338-73, Test for the softenJng point of glass.

abSotbet Neutron Source Rods The neutron source rods are designed to withstand the following:

1. 'The external pressure equal to the reactor coolant system operating pressure

2. An internal pressure equal to the pressure generated by I released gases over the source rod life.

Thimble Plug Assembly

The thimble plug assemblies satisfy the following:

1. Accommodate the differential therm'al expansion between the fuel assembly and the core internals

2. Maintain positive contact with the fuel assembly and the core internals

3. Be inserted into or withdrawn from the fuel assembly by a force not exceeding 25 pounds.

3.2.3.1.4 Control Rod Drive Mechanisms The mechanisms are Class I components designed to meet the stress requirements for normal operating conditions of Section III of the ASME Boiler and Pressure Vessel Code. Both static and alternating stress intensities are considered. The stresses originating from the required design transients are included in the analysis.

A dynamic seismic analysis is required on the full length control rod drive mechanism when a seismic disturbance has been postulated - to confirm the ability of the mechanigm to meet ASME Core,Section III, allowable stresses and to confirm its ability to trip when subjected to the seismic disturbances.

The control rod drive mechanism (CRDM) design used for the 17x17 fuel assembly control rod is identical to the 15 x 15 control rod drive mechanism. The seismic analysis and response of the 17x17 control rod drive mechanism is identical to those of the 15x15 mechanism plants.

The part length control rod drive mechanism meets only the stress limits defined in the ASME Code, Section III, in order to 3.2-42 4

BVPS-1-UPDATED FSAR Rev. 3 (1/85) l maintain structural integrity when l since it is a non-tripping mechanism. subjected to seismic loads Full Length Control Rod Drive Mechanism operational Requirements The basic i operational ' requirements for the full length control rod drive mechanisms are:

1. 5/8-inch step

2. 150-inch travel i

3. 360-pound maximum load 4.

Step in or out at 45 inches / min (72 steps / min) 5.

Power assembly interruption shall initiate release of drive rod i

6.

Trip delay of less than 150 ms - Free fall of drive rod assembly shall !

begin less than 150 ms power interruption no after action is being executed matter what holding or stepping temperatures of 100 F to 550 F. with any load and coolant

7. 40-year design life with normal refurbishment

8. 28,000 complete travel excursions which is 13 x 106 steps with normal refurbishment.

3.2.3.2 Design Description Reactivity control is soluble chemical neutron absorberprovided by neutron abscrbing rods and a concentration is (boric acid). The boric acid such as: varied to control long-term reactivity changes 1.

Fuel depletion and fission product buildup 2.

Cold to hot, zero power reactivity change 3.

Reactivity change produced by intermediate-term fission products such as xenon and samarium absptbet

4. Burnable pciccn-depletion.

Chemical and volume control is covered in Section 9.1.

The rod cluster control for: assemblies provide reactivity centrol

1. Shutdown

2. Reactivity changes due to coolant temperature changes 3.2-43

BVPS-1 UPDATED FSAR Rev. 3 (1/85) l The absorber rods are fastened securely to the spider to assure i

trouble free service. The rods are first threaded into the l spider fingers and then pinned to maintain joint tightness, after which the pins are welded in place. The end plug below the pin position is designed with a reduced section to permit flexing of the rods to correct for small operating or assembly misalignments.

The overall length is such that when the assembly is withdrawn through its full travel the tips of the absorber rods remain engaged in the guide thimbles so that alignment between rods and thimbles is always maintained. Since the rods are long and slender, they are relatively free to conform to any small mis-alignments with the guide thimble.

Burnable Msob ciccr Assembly gc fagg ,g , f Q M T M ch burnable poison assembly consists of burnable poison ro at ched to a hold down assembly. Burnable poison assemblies re shown 'n Figure 3.2-13.

The poiso rods consist of borosilicate glass tubes contained within Type 04 stainless steel tubular cladding whi is plugged and seal weld at the ends to encapsulate the gl s. The glass is also support along the length of its insi diameter by a thin wall tubular 'nner liner of Type 304 st 'nless steel. The top end of the line is open to permit th diffused helium to pass into the void vol e and the liner ov hangs the glass. The liner has an outward fl e at the bo om end to maintain the position of the liner with e glass. A typical burnable poison ;

rod is shown in longitudinal and ransverse cross-sections in Figure 3.2-14.

i The rods are statically su ended nd positioned in selected

~

guide thimbles within sp ified fue assemblies, location of these assemblies are sh in Figure 3. 5. The poison rods in each fuel assembly ar grouped and attache together at the top end of the rods to hold down assembly by flat, perforated retaining plate w ch fits within the fuel as mbly top nozzle and rests on t '

adaptor plate. The retaining late (and the poison rods) held down and restrained against ve ical motion through a ring pack which is attached to the pl e and is compresse by the upper core plate when the react upper interna assembly is lowered into the reactor. This arra. ement assur s that the poison rods cannot be ejected from the cor by fl forces. Each rod is permanently attached to the base pla a nut which is lock weldep into place.

botoGiliort abso/ber The clad in the $ d assemblies is 10 percent cold worked Type 304 stainless steel. All other structural materials are 304 or 308 stainless steel xcept for the springs which are Inconel 718.

The borosilicate glass tube provides sufficient boron content to meet the criter discossed in Section 3.3.1.2 and 3.3.1.5.

rke tMe owl chal is e;,stoy V 3.2-46

1 Insert "H" (to page 3.2-46)

Each burnable absorber assembly consists of borosilicate or WABA burnable absorber rods attached to a hold down assembly. Burnable absorber assemblies containing borosilicate poison are shown in Figure 3.2-13. WABA rods may be used in place of the borosilicate absorber rods.

The borosilicate absorber rods consist of borosilicate glass tubes contained within Type 304 stainless steel tubular cladding which is plugged and seal welded at the ends to enscapulate the glas.s. The glass is also supported along the length of its inside diameter by a thin wall tubular inner liner of Type 304 stainless steel. The top end of the liner is open to permit the diffused helium to pass into the void volume and the liner overhangs the glass. The liner has an outward flange at the bottom nd to maintain the position of the liner with the glass. A typical borosilicate burnable absorber rod is shown in longitudinal and transverse cross-sections in Figure 3.2-14.

A WABA rod consists of annular pellets of alumina-boron carbide (Alp-03-84C) burnable absorber material contained within two concentric Zircaloy tubes. These Zircaloy tubes, which form the inner and outer clad for the WABA rod, are plugged and welded at each end to encapsulate the annular stack of absorber material. The assembled rod is then internally pressurized to about 650 psig and seal welded. The l absorber stack lengths are positioned axially within the WABA rods by the '

ute of Zircaloy bottom-end spacers. An annular plenum is provided within the rod to accommodate the helium gas released from absorber material depletion during irradiation. The reactor coolant flows inside the inner tube and outside the outer tube of the annular rod. Further design details are given in Section 3.0 of Reference 22.

The burnable absorber rods are statically suspended and positioned in selected guide thimbles within specified fuel assemblies; location of these assemblies is shown in Figure 3.3-5. The absorber rods in each assembly are attached together at the top end of the rods to a hold down assembly by a flat, perforated retaining plate which fits within the fuel assembly top nozzle and rests on the adapter plate. The absorber rod assembly it held down and restrained against vertical motion through a spring pack which is attached to the plate and is compressed by the upper core plate when the reactor upper internals assembly is lowered into the reactor. This arrangement assures that the absorber rods cannot be ejected from the core by flow forces. Each rod is permanently attached to the base plate by a nut which is locked into place.

BVPS-1-UPDATED FSAR Rev. 3 (1/85) pellets stacked to a height of approximately 88 inches. )

The rods )

in each assembly are permanently fastened at the top end to a l holddown assembly, which is identical to that of the burnable peicer assemblies.

absetbst i The other structural members are constructed of Type 304 stain-less steel except for the springs. The springs exposed to the reactor coolant are wound from an age hardened nickel base alloy for corrosion resistance and high strength. The springs, when contained within the rods where corrosion resistance is not necessary, are oil tempered carbon steel. -

Thimble Plug Assembly In order to limit bypass flow through the rod cluster control guide thimbles in fuel assemblies which do not contain either control rods, source rods, or burnable cisen rods, thimble plug assemblies are utilized.. dbehu The thimble plug assemblies as shown in Figure 3.2-16 consist of a flat base plate with short rods suspended from the bottom sur-face and a spring pack assembly. The twenty-four short rods, called thimble plugs, project into the upper ends of the guide thimbles to reduce the bypass flow area. Similar short rods are also used on the source assemblies and burnable poison assemblies to plug the ends of all vacant fuel assembly guide thimbles. At installation in the core, the thimble plug assemblies interface with both the upper core plate and with the fuel assembly top nozzles by resting on the adaptor plate. The spring pack is com-pressed by the upper core plate when-the upper internals assembly is lowered into place. Each thimble plug is permanently attached to the base plate by a nut which is locked to the threaded end of the plug by a small lock-bar welded to the nut.

I All components in the thimble plug assembly, except for the springs, are constructed from Type 304 stainless steel. The springs are wound from an aged hardened nickel base alloy for corrosion resistance and high strer.gth.

3.2.3.2.2 Control Rod Drive Mechanism All parts exposed to reactor coolant are made of metals which resist the corrosive action of the water. Three types of metals are used exclusively: stainless steels, Inconel-X and cobalt based alloys. Wherever magnetic flux is carried by parts exposed to the main coolant, 400 series stainless steel is used. Cobalt based alloys are used for the pins and latch tips. Inconel-X is used for the springs of both latch assemblies and 304 stainless steel is used for all pressure containing parts. Hard chrome plating provides wear surfaces on the. sliding parts and prevents galling between mating parts.

A position indicator assembly slides over the full length control rod drive mechanism rod travel housing. It detects the drive rod 3.2-48

BVPS-1-UPDATED FSAR Rev. 0 (1/82) .

I

a. Due to temperature differences

b. Due to expansion of different materials

7. Interference between components

8. Vibration (mechanically or hydraulically induced)

9. All operational transients listed in Table 5.2-2.

10. Pump Overspeed

11. Seismic Loads (operation basis earthquake and design basis earthquake).

The main objective of the analysis is to satisfy allowable stress limits, to assure an adequate design margin, and to establish deformation limits which are concerned primarily with the func-tioning of the components. The stress limits are established not only to assure that peak stresses will not reach unacceptable values, but also limit the amplitude of the oscillatory stress component in consideration of fatigue characteristics of the materials. Standard methods of strength of materials are used to establish the stresses and deflections of these components. The dynamic behavior of the reactivity control components has been studied using experimental test data (D-Loop, Section 1.5) and experience from operating reac ors.

ke lace di ZNShir 2 design of reaftiity component rods provides a suffic4 nf col 'd volume within the burnable poison and source s to limit th 'nternal pressures to a value which s '

fies the criteria in on 3.2.3.1. The void volume f the helium in the burnable poison as is obtained throug e use of glass in tubular form which pro s a central d along the length of the rods. Helium gas is not eas y the neutron absorber rod material, thus the absorber r sustains an external pres-sure during operating co ions. internal pressure of source rods continues increase from

t until end of life at which time the ernal pressure never excee that allowed by the criteri in Section 3.2.3.1. The stress lysis of j reactivi component rods assumes 100% gas release to e rod void uma, considers the initial pressure within the rod, d l

umes the pressure external to the compone.n nt rod is zeao. bee Based on available data for properties of the borosilicat glass and on nuclear and thermal calculations for burnable r eizen rods, gross swelling or cracking of the glass tubing is not expected during operation. Some minor creep of the glass at the hot spot on the inner surface of the tube could occur but would continue only until the glass came in contact with the inner liner. The wall thickness of the inner liner is sized to provide adequate support in the event of slumping and to collapse locally before rupture of the exterior cladding if unexpected large volume changes due to swelling or cracking should occur. The top 3.2-55

Insert "Z" (to page 3.2-55)

- The design of reactivity component rods provides a sufficient cold void volume within the burnable absorber and source rods to limit the internal pressures to a value which satisfies the criteria in Section 3.2.3.1. The void volume for the helium in the borosilicate glass burnable absorber rods is obtained through'the use of glass in tubular form which provides a central void along the length of the rods. For the WABA rods, an annular void volume is provided between the two tubes at the top and along the

- length of each WABA rod. Helium gas is not released by the neutron ,

absorber rod material, thus the absorber rod only sustains an external i pressure during operating conditions. The internal pressure of source rods continues to increase from ambient until end of life at which time the internal pressure never exceeds that allowed by the criteria in Section 3.2.3.1. Except for the WABA rods, the stress analysis of reactivity component rods assumes 100% gas release to the rod void volume, considers the initial pressure within the rod, and assumes the pressure external to the component rod is zero. The stress analysis for the WABA rods assumed a maximum 30% gas release.

9 O

1 BVPS-1-UPDATED FSAR Rev. 0 (1/82) corresponding decreased ductility (as measured by tensile tests) but energy absorption (as measured by impact tests) remain quite high. Corrosion of the materials exposed to the coolant is quite low and proper control of C1 and 0-2 in the coolant will prevent the occurrence of stress corrosion. All of the austenitic stain-less steel base materials used are processed and fabricated to preclude sensitization. Although the control rod spiders are fabricated by furnace brazing, the procedure used requires that the pieces be rapidly cooled so that the time-at-temperature is minimized. The time that is spent by the control rod spiders in the sensitization range, 800-1,500 F, is not more than 0.2 hours2.314815e-5 days 5.555556e-4 hours 3.306878e-6 weeks 7.61e-7 months , as a maximum, during fabrication to preclude sensitization. The 17-4 PH parts are all aged at the highest standard aging tempera-ture of 1,100 F to avoid stress corrosion problems exhibited by aging at lower temperatures.

Analysis of the full length rod cluster control assemblies show that if the drive mechanism housing ruptures the rod cluster control assembly will be ejected from the core by pressure differential of the operating pressure and ambient pressure across the drive rod assembly. The ejection is also predicted on the failure of the drive mechanism to retain the drive rod / rod cluster control assembly position. It should be pointed out that a drive mechanism housing rupture will cause the ejection of only one rod cluster control assembly with the other assemblies remaining in the core. Analysis also showed that a pressure drop in excess of 4,000 psi must occur across a two- fingered vane to break the vane / spider body joint causing ejection of two neutron absorber rods from the core. Since the greatest pressure of the primary system coolant is only 2,250 psi, a pressure drop in excess of 4,000 psi could not be expected to occur. Thus, the ejection of the neutron absorber rods is not possible.

absoder Ejection of a burnable poison thimbi plug assembi is con-ceivable based on the postulation that he hold down ba fails and that the base plate and burnable poison rods are se erely deformed. In the unlikely event that failure of the. hol down bar occurs, the upward displacement of the burnable poison assembly only permits the base plate to contact the upper core plate. Since this displacement is small, the major portion of the bcrcailicete glesa tubing remains positioned within the core.

In the case of the thimble plug assembly, the thimble plugs will partially remain in the fuel assembly guide thimbles thus main-taining a majority of the desired flow impedence. Further displacement or complete ejection would necessitate the square base plate and burnable poison rods be forced, thus plastica 11y deformed, to fit up through"a smaller diameter hole. It is expected that this conditi a requires a substantially higher force or pressure drop than at of the hold down bar failure.

Experience with control rods,absorber burnable3psicon rods, and source rods are discussed in Reference 8.

3.2-57

i BVPS-1-UPDATED FSAR Rev. 0 (1/82)

Burnablek.bsorbet cinen Assemblies absetber The burnable p :: n assemblies are static temporary reactivity control elements. The axial position is assured by the hold down assembly which bears against the upper core plate. Their lateral position is maintained by the guide thimbles of the fuel assemblies.

The individual rods are shouldered against the underside of the retainer plate and securely fastened at the top by a threaded nut which is then locked in place by a welded pin. The square ,

dimension of the retainer plate is larger than the diameter of '

the flow holes through the core plate. Failure of the hold down bar or spring pack therefore does not result in ejection of the burnable poisen d from the core.

G other The only incident at could potentially result in ejection of the burnable poiso/n rods is a multiple fracture of the retainer plate. This is not considered credible because of the light loads borne by this component. During normal operation the loads borne by the plate are approximately 5 lb/ rod or a total of 100 lb. distributed at the points of attachment. Even a multiple fracture of the retainer plate would result in jamming of the plate segments against the upper core plate, again preventing ejection. Excessive reactivity increase due to burnable poison ejection is therefore prevented.

.T NS&dT ~3 -+

same type of stainless steel clad used on rod cluster co trol also used on the burnable poison rods. In this i-cation t is even less susceptibility to mechanica damage since these a static assemblies. The guide thi es of the fuel assembly af the same protection from da e due to fuel rod failures as that ribed for the rod clu r control rods.

The consequences of clad brea are a similarly small. The poison material is borosilicate g s which is maintained in position by a central hollow e. the event of a hole developing in the clad for postulated son the expected consequence is only the oss of the helium duced by the absorption process i the primary coolant. Th lass is chemically inert an emains remote from high coolant velo ies, therefore sign ' cant loss of poison material resulting reactivity i ease is not expected.

Rods this design have performed very well in actual service wi no failures observed through full life of one fuel cycle.

Drive Rod Assemblies All postulated failures of the drive rod assemblies either by ,

fracture or uncoupling lead to the fail safe condition. If the I drive rod assembly fractures at any elevation, that portion remaining coupled falls with, and is guided by the rod cluster 3.2-60

l 1.

INSERT J l

l Burnable absorber rods are clad with either stainless steel or l Zircaloy-4. The burnable absorber'is either a borosilicate glass tube whose position is maintained by a central hollow stainless steel tube or Al203-B4C annular pellets contained within two concentric Zircaloy tubes. . Burnable absorber rods are placed in static assemblies and are not subjected to motion that might' damage the rods. Further, the guide thimble tubes of the fuel assembly afford additional protection from damage.

During the accumulated thousands of years of burnable absorber rodlet operating experience, only one instance of penetration of the stainless steel burnable absorber cladding has been observed.

The consequences of clad breach are small. It is anticipated that upon clad breach, the B4C or berosilicate glass would be leached by the coolant water and that localized power peaking of a few percent would occur. However, no design criteria would be violated.. Additional information on the consequences of oostulated WABA rod failures is presented in Reference 22.

I

BVPS-1-UPDATED FSAR Rev. 3 (1/85)

References for Section 3.2 (Cont'd)-

11. Deleted, Revision 0.

12. W. J. O'Donnell, B. F. Langer, " Fatigue Design Basis for Zircaloy Components", Nuclear Science and Engineering, 20, 1-12, 1964.

13. E. E. DeMario, S. Nakazato, " Hydraulic Flow Test of the 17 x 17 Fuel Assembly", WCAP-8279, Westinghouse Electric Corporation (February, 1974).

14. L. T. Gesinki, " Fuel Assembly Safety Analysis for Combined Seismic and Loss-of-Coolant Accident", WCAP-7950, Westinghouse Electric Corporation (July 1972).

15. " Westinghouse PWR Internals Vibration Summary Three-Loop Internals Assurance", WCAP 7765-AR, Westinghouse Electric Corporation' (November 1973) .

16. Trojan Final Safety Analysis Report, Appendix A.12, Docket No. 50-344, Portland General Electric Company.

17. M. B. Aycock, " Internals Vibration Testing on Westinghouse 3

&, 4 Loop Plants", AEC Meeting Summary for RESAR-3,

-(July 3, ).974).

18. '.WRods",

F. Knowles, " Beaver Valley SAM-1078, Westinghouse Removal of Part Length Electric Corporation (Proprietary), (February 23, 1978).

19. " Westinghouse Wet Annular Burnable Absorber Evaluation Report", WCAP-10021-P-A Revision 1, Westinghouse Electric Corporation (October 1983)

20. Davidson, S. L. (Ed.), et al, " VANTAGE 5 Fuel Assembly Reference Core Report," WCAP-10444-P-A, September 1985.

21. Davidson, S. L. (Ed.), et al, " VANTAGE SH Fuel Assembly,"

WCAP-10444-P-A, Addendum 2-A, February 1989.

22. Skaritka, J. (Ed.), " Westinghouse Wet Annular Burnable Absorber Evaluation Report," WCAP-10021-P-A, Revision 1, October 1983.

3.2-70

_ _ _ _ _ _ _ _ _ . - _ _ _ - _ _ _ - _ _ _ _ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - _ _ - _ - _ - _ _ _ _a

BVPS-1-UPDATED FSAR Rev.4 (1/86) operating conditions, thereby providing negative reactivity feedback characteristics. The design basis meets GDC 11.

Discussion When compensation for a rapid increase in reactivity is considered, there are two major effects. These are the resonance absorption effects (Doppler) associated with changing fuel temperature and the

spectrum effect resulting from changing moderator density. These l basic physics characteristics are often identified by reactivity i coefficients. The use of slightly enriched uranium insures that the Doppler coefficient of reactivity is negative. This coefficient provides the most rapid reactivity compensation. The core is also designed to have an overall negative moderator temperature coefficient of reactivity so that average coolant temperature or void contents provides another, slower compensatory effect. Nominal power operation is permitted only in a range of overall negative moderator temperature coefficient. The negative moderator temperature coefficient can be achieved through use of fixed burnabl p; ice..

and/or control rods by limiting the reactivity held down by cluble boron.

abset et 3.3.1.3 Control of Power Distribution Basis The nuclear design basis are that, with at least a 95% confidence level:

1. The fuel will not be operated at greater than 11.9 kw/ft under normal operating conditions including an allowance of 2% for calorimetric error and not including a power spike factor due to densification.

2. Under abnormal conditions including the maximum overpower condition, the fuel peak power will not cause melting as defined in Section 3.4.1.2.

t

3. The fuel will not operate with a power distribution that violates the departure from nucleate boiling (DNB) design basis (i.e., the DNBR shall not be less than 4vseg as discussed in Section 3.4.1) under Condition I and I events including the maximum overpower condition. gg

4. Fuel management will be such as to produce rod powers and burnups consistent with the reload safety evaluation for each fuel cycle.

The above bases meet GDC 10.

3.3-3

BVPS-1-UPDATED FSAR Rev. 5 (1/87) of a hypothetical failure to trip following anticipated transients

.will be performed to show that no significant core damage would result and system peak pressures would be limit'ed to acceptable

' values and no failure of the Reactor Coolant System would. result.

These analyses -were in accordance with the AEC policy outlined in WASH-1270 " Technical Report on Anticipated Transients Without Scram for Water-Cooled Power Reactors," September, 1973.

3.3.2 Description 3.3.2.1 Nuclear Design Description The reactor core consists of a specified number of fuel rods which are held in bundles by spacer grids and top and bottom fittings. The fuel rods are constructed of Zircaloy cylindrical tubes containing i UO2 fuel pellets. .The bundles, known as fuel assemblies, are arranged in a pattern which approximates a right circular cylinder.

Some fuel assemblies may contain non-fueled rods. For a description of non-fueled rods see Section 3.1.

Each fuel assembly contains a 17 x 17 rod array composed of 264 fuel rods, 24 rod cluster control (RCC) thimbles and an in-core instrumentation thimble. Figure 3.2-1 shows-a cross sectional view of a 17 x 17 fuel assembly and the related RCC locations. Further details of the fuel assembly are given in Section 3.2.1.

The fuel w a given assembly have the same uranium enrichment in h:th [rodktklackthiniMS th: redi:1 and ;;ici pi;;;;. ZAISMT*

three different enrichments are used.in the initial core loading to Fuel assemblies ff of establish a favorable radial power distribution. Figure 3.3-1 shows the fuel loading pattern used in the first core. Two regions consisting of the two lower enrichments are interspersed so as to j form a checkerboard pattern in the central portion of the core. The i

third region is arranged around the periphery of the core and contains the highest enrichment. The enrichments for the first core are shown in Table 3.3-1. The cycle 4 core reload used a low leakage loading pattern t; t;h; si;;nt;;; cf an increased cycle length and a reduced fluence [$ edtf $5 ockleveIate at the p riphery of the core.

The core average enrichment is determined by the amount of fissionable material required to provide the desired core lifetime and energy requirements. The physics of the burnout process is such that operation of the reactor depletes the amount of fuel available due to the absorption of neutrons by the U235 atoms and their subsequent fission. The rate of U235 depletion is directly proportional to the power level at which the reactor is operated. In addition, the fission process results in the formation of fission products, some of which readily absorb neutrons. These effects, depletion and the buildup of fission products, are partially offset by the buildup of plutonium, shown in Figure 3.3-2 for the 17 x 17 fuel assembly, which occurs due to the non-fission absorption of neutrons in U-238 Therefore, at the beginning of any cycle a 3.3-7 l

A Insert A. p. 3.3-7 4 in the rarlial plane. However, the uranium enrichment may change with fuel height (e.g. the. fresh fuel asserelies loaded in cycle 7 used unenriched uranium fuel in the top and botta six inches of the fuel rods. 'Jhe i

middle 120 indes of each feed assembly used the enriched urania fuel) .

e f ,I 9

,g,.

BVPS-1-UPDATED FSAR Rev. 4 (1/86) reactivity reserve equal to the depletion of the fissionable fuel and the buildup of fission product poisons over the specified cycle life must be " built" into the reactor. This excess reactivity is controlled by removable neutron absorbing material in the form of boron dissolved in the primary coolant and burnabl3gpci;cn rods.

ncen ra [on of boric acid in the primary coolant is varied to provide control and to compensate for long-term reactivity requirements. The concentration of the soluble neutron absorber is varied to compensate for reactivity changes due to fuel burnup, fission product poisoning including xenon and samarium, burnable p;iscn depletion, and the cold-to-operating moderator temperature change. Using its normal makeup path, the Chemical and volume Control System (CVCS) is capable of inserting negative reactivity at a rate of approximately 30 pcm/ min when the reactor coolant boron concentration is 1000 ppm and approximately 35 pcm/ min when the reactor coolant boron concentration is 100 ppm.

If the emergency boration path is used, the CVCS is capable of inserting negative reactivity at a rate of approximately 65 pcm/ min when the reactor coolant concentration is 1000 ppm and approximately 75 pcm/ min when the reactor coolant baron concentration is 100 ppm.

The peak burnout rate of xenon is 25 pcm/ min (Section 9.1 discusses the capability of the CVCS to counteract xenon decay). Rapid transient reactivity requirements and safety shutdown requirements are met with control rods.

ho v'ert As the boron concentration is increased, the moderato , temperature coefficient becomes less negative. The use of ,e solub r oi;cn alone would result in a positive moderator coefficient at L for the first cycle. Therefore, burnable p;icen rods were us in the first core to reduce the soluble boron ncentration suffi ently to insure that the moderator temperature co fficient is neg ve for power operating conditions. During opera on, the psi;ca content in these rods is depleted thus adding posi ive reactivity to offset some of the negative reactivity from f1 el depletion and fission product buildup.

The depletion rate of the burnable p-i an rods is not critical since chemical shim is always availabl nd flexible enough to cover any possible deviations in the expect d burnab p:i:On depletion rate.

Figure 3.3-3 is a graph of a typi al cor epletion with and without burnable pei :: ods. Note that even end-of-life conditions some residua p;i :: ins i th bur b p i cn rods resulting in a net de ease in t firs cy le time. Upon completion of the first cycle, most o the ur a --icen rods are normally removed becau e the . moderator suff ciently negative.

m ra M efficient g ,g s in ~ reload cores is boten Tables 3.3-1 through 3.3-3 contain a summary of the reactor core design parameters for the fir:t fuci 07:10, including reactivity coefficients, delayed neutron fraction and neutron lifetimes.

Sufficient information is included to permit an independent 3.3-8

I BVPS-1-UPDATED FSAR Rev. 2 (1/84) then F (N, Z)

Factor.

is the Core Average Axial Peaking l

To include the allowances made for densification effects, which are height dependent, the following quantities are defined.

S (Z) = height the allowance made for densification effects at Z in the core. See Section 3.3.2.2.5.

P(Z) = ratio of power per unit core height in the horizontal plane at height Z to the average value of power per unit core height.

Then F (Q) = Total peaking factor

= Maximum kw/ft Average kw/ft Including densification allowance F (Q) = max (F (N, XY) (Z) x P(Z) x S (Z) ) x F(N,U) x F (E,Q)

(3.3-3) 3.3.2.2.2 Radial Power Distributions The power shape in horizontal sections of the core at full. power is a function of the fuel and burnable poisod leading patterns, and the presence or absence of a single bank of full length control rods. Thus, at any time in the cycle any horizontal section of the core can be characteri::ed as (a) unrodded or (b) with group D control rods. These two situations combined with burnup effects determine the radial power shapes which can exist in the core at full power. The effect on radial power shapes of power level, xenon, samarium and moderator density effects are considered also but these are quite small. The effect of non-uniform flow distribution is negligible. While radial power distributions in various planes of the core are often illustrated, the core radial enthalpy rise distribution as determined by the integral of power up each channel is of greater interest. Figures 3.3-6 through 3.3-11 show representative radial power distributions for one eighth of the core for representative operating conditions. These conditions are (1)

Hot Full Power (HFP) at Beginning of Life (BOL) -

unrodded -no xenon, (2) HFP at BOL - unrodded - equilibrium xenon, (3) HFP at BOL - Bank D in - equilibrium xenon, (4) HFP at Middle of Life -

unrodded -

quilibrium xenon, and (5) HFP at End of Life -

unrodded - quilibrium xenon.

fMYally 3.3-11 l

l

BVPS-1-UPDATED FSAR Rev. 2 (1/84)

Since the position of the hot channel varies from time to time, a single reference radial design power distribution is selected for DNB calculations. This reference power distribution is chosen conservatively to concentrate power in one area of the core, minimizing the benefits of flow redistribution. Assembly powers are normalized to core average power.

3.3.2.2.3 Assembly Power Distributions For the purpose of illustration, assembly power distributions from the BOL and EOL conditions corresponding to Figures 3.3-7 and 3.3-10, respectively, are given for the same assembly in Figures 3.3-12 and 3.3-13, respectively. l Since the detailed power distribution surrounding the hot channel varies from time to time, a conservatively flat assembly power distribution is assumed in the DNB analysis, described in Section 3.4, with the rod of maximum integrated- power artificially raised to the design value of F (N, aH) . Care is taken in the nuclear design of all fuel cycles and.all operating conditions to ensure that a flatter assembly power distribution does not occur with limiting values of F(N,aH),

3.3.2.2.4 Axial Power Distributions The shape of the power profile in the axial or vertical direction is largely under the control of the operator through the automatic motion of full length rods responding to manual operation of the Chemical and Volume Control System. Nuclear >

effects which cause varietions in the axial power shape include moderator density, Doppler effect on resonance absorption, spatial xenon and burnup. Automatically controlled variations in total power output and full length rod motion are also important in determining the axial power shape at any time. Signals are available to the operator from the ex-core ion chambers which are l long ion chambers outside the reactor vessel running parallel to !

the axis of the core. Separate signals are taken from the top and bottom halves of the chambers. The difference between top i and bottom signals from each of four pairs of detectors is displayed on the control panel and called the Flux Difference, AI. Calculations of core average puaking factor for many plants and measurements from operating plants under many operating situations are associated with either AI or axial offset in such a way that an upper bound can be placed on the peaking factor.

For these correlations axial offset is defined as:

axial offset = ( *

~

t b Where e and e are the top and bottom detector readings, respectively. b The radial power involving distribution the partial [ionshown insert in Figurerods of control / 3.3-8 and 2.2 a represents 3.3-12

_ _ _ - _ _ _ _ _ _ _ _ _ - _ _ _ _A

BVPS-1-UPDATED FSAR Rev. 2 (1/84) 3.3.2.2.6 Limiting Power Distributions According to the ANSI classification of plant conditions, Condition I occurrences are those which are expected frequently or regularly in the course of power operation, maintenance, or maneuvering of the plant. As such, Condition I occurrences are accommodated with margin between any plant parameter and the value of that parameter which would require either automatic or i manual protective action. Inasmuch as Condition I occurrences occur frequently or regularly, they must be considered from the point of view of affecting the consequences of fault conditions (Conditions II, III and IV). In this regard, analysis of each fault condition described is generally based on a conservative set of initial conditions corresponding to the most adverse set of conditions which can occur during Condition I operation.

The list of steady state and shutdown conditions, permissible deviations (cuch cc enc ccclent lecp cut of cercice) and opera-tional transients is given in Section 14.D. Implicit in the definition of normal operation is proper and timely action by the reactor operator. That is, the operator follows recommended operating procedures for maintaining appropriate power distri-butions and takes any necessary remedial actions when alerted to do so by the plant instrumentation. Thus, as stated above, the worst or limiting power distribution which can occur during normal operation is to be considered at the starting point for analysis of ANSI Conditions II, III and IV events.

Improper procedural actions or errors by the operator are assumed in the design as occurrences of moderate frequency (ANSI Condition II). Some of the consequences which might result are listed in Section 14.1. Therefore, the limiting power shapes which result from such Condition II events, are those power shapes which deviate from the normal operating condition at the recommended axial offset band, e.g. due to lack of proper action by the operator during a xenon transient following a change in power level brought about by control rod motion. Power shapes which fall in this category are used for determination of the reactor protection system setpoints so as to maintain margin to overpower or DNB limits.

The means for maintaining power distributions within the required hot channel factor limits are described in the Technical Specifi-cations. A complete discussion of power distribution control in Westinghouse PWRs is included in Reference 4. Detailed infor-mation on the design constraints on local power density in a Westinghouse PWR, on the defined operating procedures and on the measures taken to preclude exceeding design limits is presented in the Westinghouse Topical Reports and distribution

_. control and load following procedures. (2 9gowerThe following paragraphs summarize these reports and describe the calculations used to establish the upper bound on peaking factors.

3.3-14

BVPS-1-UPDATED FSAR Rev. 2 (1/84)

3. The control full length bank insertion limits are not violated

4. Axial power distribution procedures, which are given in terms of flux difference control and control bank position, are observed.

The axial power distribution procedures referred to above are part of the required operating procedures which are followed in normal operation. Briefly they require control of the axial offset (flux difference divided by fractional power) at all power levels within a permissible operating band of a target value corresponding to the equilibrium full power value.

Calculations, as described in Reference 30, are performed for normal operation of the reactor including load following maneuvers. Beginning, middle and end of cycle conditions are included in the calculations. Different histories of operation are assumed prior to calculating the effect of load follow transients on the axial power distribution. These different histories assume base loaded operation and extensive load following. These calculated points have been synthesized from axial calculations combined with radial factors appropriate for rodded and unrodded plares in the first cycle. The calculated values have been increased by a factor of 1.05 for measurement i error, a factor of 1.03 for the manufacturing tolerances.

The envelope drawn over the calculated F (Q) points in Figure -

3.3-21 represents an upper bound envelope on local power density versus elevation in the core. It should be emphasized that this envelope is a conservative representation of the bounding values of local power density. Expected values are considerably smaller and, in fact, less conservative bounding values may be justified with additional analysis or surveillance requirements. For example, Figure 3.3-21 bounds both BOL and EOL conditions but without consideration of radial power distribution flattening with burnup, i.e., both BOL and EOL points presume the same radial peaking factor. Inclusion of the burnup flattening effect would reduce the local power densities corresponding to EOL conditions which may be limiting at the higher core elevations.

P c analyci: indic te the plant rey be operated up to 05.5 p '

cent 11 rated power for the first 2,270 Effective Ful ower Hours an to 100 percent full rated power ther . er with reference to -core monitoring only.

For the analysis give. Figure 3.3-21 . radial power distri-butions are assumed to be varian Ith burnup but vary with core elevation. In general, r power distributions for the unrodded and rodded porti o of core will be elevation, burnup, and power level pendent.

The power depe '

nce can be bounded by an express of the form 1.0 xy Fxy [1 + 0.2 (1-P)] (3. 5 3.3-16

l l

BVPS-1-UPDATED FSAR Rev. 2 (1/84) I ere: Pisgercentreactorpowerleveland 1-F given in Figure 3.3-22 for unrodded planes and r pY s covered by control Bank D.

For the first cycle, ounding values of F 1.0 ere assumed to be 1.58 for elevations pe . ated by con rods. This becomes 1.71 when the factor for co'nt vat' - and the engineering factor are included. The adjusted f to or unrodded planes are:

1. F $ 1.6 up to 2.4 ft elevati

2. I 1.75 from 2.4 ft elevation to 7.8 evation xy

- 3. F xy ; 1.53 abcvc ~.S ft clevaticr. s <

Finally, as previously discussed, this upper bound envelope is based on procedures of load follow which require the operator to operate within an allowed deviation from a target equilibrium value of axial flux difference. These procedures are detailed in the Technical Specifications and predicated only upon ex-core surveillance supplemented by the normal monthly full core map requirements and by computer based alarms on deviation and time of deviation from the allowed flux difference band.

Allowing for fuel densification effects the average kw/ft at 2,652 MWt is 5.20 kw/ft. From Figure 3.3-21, the conservative upper bound value of normalized local power density, including uncertainty allowances i s --h-3 corresponding to a peak local power density of 11.

kw/ftat(102percentpower.

To determine reactor protection system set points, with respect to power distributions, three categories of events are considered, namely rod control equipment malfunctions, operator errors of commission and operator errors of omission.

The first category comprises uncontrolled rod withdrawal (with rods moving in the normal bank sequence) for full length banks.

Also included are motions of the full length banks below their insertion limits, which could be caused, for example, by uncon-trolled dilution or primary coolant cooldown. Power distribu-tions were calculated throughout these occurrences assuming short term corrective action, that is no transient xenon effects were considered to result from the malfunction. The event was assumed to occur from typical normal operating situations which did include normal xenon transients. It was further assumed in determining the power distributions that total power level would be limited by reactor trip to below 118%. Since the study is to determine protection limits with respect to power and axial off-set, no credit was taken for trip set point reduction due to flux difference. The results are given in_ Figure 3.3-23 in units of kw/ft. The peak power density which can occur in such events, assuming reactor trip at or below 1185, is thus limited to 18.0 kw/ft for the three loop plant including uncertainties and densification effects.

3.3-17

BVPS-1-UPDATED FSAR Rev. 2 (1/84)

The second category also appearing in Figure 3.3-23, assumes that the operator mis-positions the full length rod bank in violation of the insertion limits and creates short term conditions not included in normal operating conditions.

The third category assumes that the operator fails to take action to correct a flux difference violation. The results shown on Figure 3.3-23A are F (Q) multiplied by 102% power including an allowance for calorimetric error. The figure shows that, provided the assumed error in operation does not continue for a period which is long compared to the xenon time constant, the maximum local power does not exceed 20.0 kw/ft including the above factors. However, the Technical Specifications restrict operation such that this peak linear power density is less than 18.0 kw/ft. These events are considered Condition II events.

It should be noted that a reactor overpower accident is not assumed to-occur coincident with an independent operator error.

Analyses of possible operating power shapes for the reactor des-cribed herein show that tM appropriate hot channti factors F(Q) and F(N,aH) for peak local power density and for DNB analysis at full power are the reluer ci - in Table 3. 3-2 =d addrcs;cd in theTechnicalSpecificatioy I descdbed s F (Q) can be increased with decreasing power as shown in the Technical Specifications. Increasing F(N,6H) with decreasing power is permitted by the DNB protection set points and allows radial power shape changes with rod insertion to the insertion limits as described in Section 3.4.3.2. It has been determined !

that provided the above conditions 1 through 4 are observed, the Technical Specification limits are met.

When a situation is possible in normal operation which could result in local power densities in excess of those assumed as the pre-condit. ion for a subsequent hypothetical accident, but which would not itself cause fuel failure, administrative controls and alarms are provided for returning the core to a safe condition.

These alarms are described in detail in Sections 7.

3.3.2.2.7 Experimental Verification of Power Distribution Analysis This subject is discussed in depth in Reference 2. A summary of this report is given here.

In a measurement of peak local power density, F(Q), with the movable detector system described in Section 7.7.1 and 3.4.5, the i following uncertainties have to be considered:

i 1. reproducibility of the measured signal j 2. errors in the calculated relationship between detector l

current and local flux Anal erked hr Cyd K & i-caase th RQ ad F(MANYes ka,e beeakQ em 'NiN hc o AOM* W k a.e dw gu$nh4 3_1g

BVPS-1-UPDATED FSAR Rev. 2 (1/84)

. \

3.3.2.3 Reactivity Coefficients The kinetic characteristics of the reactor core determine the response of-the core to changing plant conditions or to operator adjustments made during normal operation, as well as the core response during abnormal or accidental transients. These kinetic characteristics are quantified in reactivity coefficients. The reactivity coefficients reflect the changes in the neutron multi-plication due to varying plant conditions such as power, moderator or fuel temperatures, or less significantly due to a change in pressure or void conditions. Since reactivity coefficients change during the life of the core, ranges of coefficients are employed in transient analysis to determine the response of the plant throughout life. The results of such simulations and the reactivity coefficients used are presented in Chapter 14. The analytical methods and calculational models used in calculating the reactivity coefficients are given in Section 3.3.3. These models have been confirmed through extensive testing of abnormal or accidental transients. These kinetic characteristics are quantified in reactivity co-efficients. These models have been-confirmed through extensive testing of more than thirty cores similar to the plant described herein; results of these tests are discussed in Section 3.3.3.

Quantitative information for calculated reactivity coefficients, i including fuel-Doppler coefficient, moderator coefficients l (density, temperature, pressure, void) and power coefficient is given in the following sections.

3.3.2.3.1 Fuel Temperature (Doppler) Coefficient The fuel temperature (Doppler) coefficient is defined as the change in reactivity per degree change in effective fuel temperature and is primarily a measure of the Doppler broadening of U- 238 and Pu-240 resonance absorption peaks. Dopplerj adening of other isotopes such as U-236, Np-237, etc., are also considered but their contributions to the Doppler effect are small. An increase in fuel temperature increases the ef fective resonance absorption cross . Sections of the fuel and produces a corresponding reduction in reactivity.

The fuel temperature coefficient 4s calculated by performing two-group X-Y calculations using!$an updated version of the TURTLE llO codeg Moderator temperature is held ccnstant and the ower level is varied. Spatial variation of fuel temperature is taken into account by calculating the effective fuel temperature as a function of power density as discussed in Section 3.3.3.1.

The Doppler temperature coefficient is shown is Figure 3.3-27 as a function of the effective fuel temperature (at beginning-of-life and end-of-life conditions). The effective fuel temperature is lower than the volume averaged fuel temperature since the neutron flux distribution is non-uniform through the pellet and gives preferential we ght to the s rface temperature. The ce h fke Ac/va ,ced Mo 4/ Cd @ADN 3.3-21

BVPS-1-UPDATED FSAR Rev. 2 (1/84)

Doppler-only -contribution to the power coefficient, defined later, is shown in Figure 3.3-28 as a function of relative core power. The integral of the differential curve on Figure 3.3-28 is the Doppler contribution to the power defect and is shown in Figure 3.3-29 as a function of relative power. The Doppler coefficient becomes more negative as a function of life as the Pu-240 content increases, thus increasing the Pu-240 resonance absorption but less negative as the fuel temperature changes with burnup as described in Section 3.3.3.1. The upper and lower limits of Doppler coefficient used in accident analyses are given in Section 14. -

3.3.2.3.2 Moderator coefficients The moderator coefficient is a measure of the change in reactivity due to a change in specific coolant parameters such as density, temperature, pressure or void. The coefficients so obtained are moderator density, temperature, pressure and void coefficients.

Moderator Density and Temperature Coefficients The moderator temperature (density) coefficient is defined as the change in reactivity per degree change in the moderator tempera-ture. Generally, the effect of the changes in moderator density

.as well as the temperature are considered together. A decrease in moderator density means less moderation which results in a negative moderator coefficient. An increase in coolant temperature, -keeping the density constant, leads to a hardened neutron spectrum and results in an increase in resonance absorption in U-238, Pu-240 and other isotopes. The hardened spectrum also causes a decrease in the fission to capture ratio in U-235 and Pu-239. Both of these effects make the moderator coefficient more negative. Since water density changes more rapidly with temperature as temperature increases, the moderator temperature (density) coefficient becomes more negative with increasing temperature.

The soluble boron used in the reactor as a means of reactivity control also has an effect on moderator density coefficient since the soluble boron poison density as well as the water density is ,

decreased when the coolant temperature rises. A decrease in the '

l soluble poison concentration introduces a positive component in I the moderator coefficient.

Ch50tbet khus if the concentration of soluble poison is large enough, the i hnetvalueofthecoefficientsmaybepositive.

pcircr rods present, however, the initial hot boron concentration With the burnable i is sufficiently low that the moderator temperature coefficient is negative at operating temperatures. The effect of control rods is to make the moderator coefficient more negative by reducing the required soluble boron concentration and by increasing the

" leakage" of the core.

3.3-22

. .- . _ _ _ . - _ . . _ _ . _ . . _ _ _ _ _ -_m.___-__.___________.-_____.__m___________

BVPS-1-UPDATED FSAR Rev. 2 (1/84) i 3.3.2.3.3 Power Coefficient The combined effect of moderator temperature and fuel temperature change as the core power level changes is called the total power coefficient and is expressed in terms of reactivity change per percent power change. The power coefficient at BOL and EOL conditions is given in Figure 3.3-34.

It becomes more negative with burnup reflecting the combined effect of moderator and fuel temperature coefficients with burnup. The power defect (integral reactivity effect) at BOL and EOL is given in Figure 3.3-35.

3.3.2.3.4 Comparison of Calculated and Experimental Reactivity Coefficients Section 3.3.3 describes the comparison of calculated and experi-mental reactivity coefficients in detail. Based on the data pre-sented there, the accuracy of the current analytical model is:

1. .2%6P for Doppler and power defect

2. 2 2 pcm/'F for the moderator coefficient Experimental evaluation of the calculated coefficients was done during the physics startup tests described in Section 13.

3.3.2.3.5 Reactivity Coefficients Used in Transient Analysis Table 3.3-2 gives the representative ranges for the reactivity coefficients used in transient analysis. The exact values of the coefficient used in the analysis depend on whether the transient of interest is examined at the beginning or end of life, whether most negative or the most positive (least negative) coefficients are appropriate, and whether spatial nonuniformity must be con-sidered in the analysis. Conservative values of coefficients, considering various aspects of analysis, are used in the transient analysis. This is completely described in the Technical Specifications.

The values listed in Table 3.3-2 and illustrated in Figures 3.3-27 through 3.3-35 apply to thc corc dcccribcd in C c y/e /,

Tabic 3.3 1. The coefficients appropriate for use in subsequent cycles depend on the core operating history, the number and !

enrichment of fresh fuel assemblies, the loading pattern of (

burned and fresh fuel, the number and location of any burnable i poison rods, etc. The need for a reevaluation of any accident in a subsequent cycle is contingent upon whether or not the coefficients for that cycle fall within the identified range used in the analysis presented in Section 14. Control ro requirements are given in Table 3.3-3 for the ccrc dcccribc and for a hypothetical equilibrium cycle since these are markedly different. These latter nurters are provided for information 3.3-24

BVPS-1-UPDATED FSAR Rev. 2 (1/84) coefficient becomes more negative as the fuel depletes because ~

the boron concentration is reduced. This effect is the major .

contributor to the increased requirement at end of life. l 3.3.2.4.3 Redistribution During full power operation the coolant density decreases with core height. This, together with partial insertion of control rods, results in less fuel depletion near the top of the core.

Under steady state conditions, the relative power distribution will be slightly asymmetric towards the bottom of the core. On the other hand, at hot zero power conditions, the coolant density is uniform up the core, and there is no flattening due to Doppler. The result will be a flux distribution which at zero power can be skewed toward the top of the core. The reactivity insertion due to the skewed distribution is calculated with an allowance for the most adverse effects of xenon distribution.

3.3.2.4.4 Void Content A small void content in the core is due to nucleate boiling at full power. The void collapse coincident with power reduction makes a small reactivity contribution.

3.3.2.4.5 Rod Insertion Allowance At full power, the control bank is operated within a prescribed band of travel to compensate for small periodic changes in boron concentration, changes in temperature and very small changes in tw xenon concentration not compensated for by a change in boron concentration. When the control bank reaches either limit of this band, a change in boron concentration is required to compensate for additional reactivity changes. Since the insertion limit is set by a rod travel limit, a conservatively high calculation of the inserted worth is made which exceeds the normally inserted reactivity.

3.3.2.4.6 Burnup Excess reactivity of 10%AP to 25%AP (hot) is installed at the beginning of each cycle to provide sufficient reactivity to compensate for fuel depletion and fission products throughout the cycle. This reactivity is controlled by the addition of soluble boron to the coolant and by burnable poison. The soluble boron concentration for several core configurations, the unit boron worth, and burnable poic worth are given in Tables 3.3-1 and 3.3-2. Since the excess'qnreactivity for burnup is controlled by soluable boron and/or b rnable poison, it is not included in control rod requirements. dyb 3.3.2.4.7 Xenon and Samarium Poisoning Changes in xenon and samarium concentrations in the core occur at a sufficiently slow rate, even following rapid power level 3.3-26

BVPS-1-UPDATED FSAR Rev. 2 (1/64) changes, that the resulting reactivity change is controlled by changing the soluble boron concentration.

3.3.2.4.8 pH Effects Changes in reactivity due to a change in coolant pH, if any, are sufficiently small in magnitude and occur slowly enough to be controlled by the boron system. Further details are available in Reference 6.

3.3.2.4.9 Experimental Confirmation Following a normal shutdown, the total core reactivity change during cooldown with a stuck rod has been measured on a 121 assembly, 10 ft high core and a 121 assembly, 12 ft high core.

In each case, the core was allowed to cool down until it reaches criticality simulating the steamline break accident. For the ten foot core, the total reactivity change associated with the cooldown is overpredicted by about 0.3% with respect to the measured result. This represents an error of about 5% in the total reactivity change and is about half the uncertainty allow-ance for this quantity. For the 12 ft core, the difference between the measured and predicted reactivity change was an even smaller 0.2%AP. These measurements and others demonstrate the ability of the methods described in Section 3.3.3 to accurately predict the total shutdown reactivity of the core.

3.3.2.5 Control Core reactivity is controlled by means of a chemical poison dissolved in the coolant, Rod Cluster Control Assemblies, and burnabled seebes

{"circr rods as described below, 3.3.2.5.1 Chemical Poison Boron in solution as boric acid is used to control relatively slow reactivity changes associated with:

1. The moderator temperature defect in going from cold shutdcwn at ambient temperature to the hot operating temperature at zero power

2. The transient xenon and samarium poisoning, such as that following power changes or changes in Rod Cluster Control position

3. The excess reactivity required to compensate for the effects of fissile inventory depletion and buildup of long-life fission products

4. Q The~hsoket urna. c poison depletion.

The boron concentrations for various core conditions are presented in Table 3.3-2.

3.3-27

BVPS-1-UPDATED FSAR Rev. 2 (1/84) 3.3.2.5.2 Rod Cluster Control Assemblies The Rod Cluster Control Assemblies employed are full-length assemblies. The number of full-length assemblies are shown in Table 3.3-1. The full-length Rod Cluster Control Assemblies are used for shutdown and control purposes to offset fast reactivity changes associated with:

1. The required shutdown margin in the hot zero power, stuck rods condition

2. The reactivity compensation as a result of an increase in power above hot zero power (power defect including Doppler, and moderator reactivity changes)

3. Unprogrammed fluctuations in boron concentration, coolant temperature, or xenon concentration (with rods not exceeding the allowable rod insertion limits)

4. Reactivity ramp rates resulting from load changes.

The allowed full-length control bank reactivity insertion is limited at full power to maintain shutdown capability. As the power level is reduced, control rod reactivity requirements are also reduced and more rod insertion is allowed. The control bank position is monitored and the operator is notified by an alarm if the limit is approached. The determination of the insertion limit uses conservative xenon distributions. In addition, the Rod Cluster Control Assembly withdrawal pattern determined from these analyses is used in determining power distribution factors and_in determining the maximum worth of an inserted Rod Cluster Control Assembly ejection accident. For further discussion, refer to the Technical Specifications on Rod Insertion Limits.

Power distribution, rod ejection and rod misalignment analyses are based on the arrangement of the shutdown and control groups of the Rod Cluster Control Assemblies shown in Figure 3.3-36.

All shutdown Rod Cluster Control Assemblies are withdrawn before withdrawal of the control banks is initiated. In going from zero to 100 percent power, control banks, A, B, C and D are withdrawn sequentially. The limits of rod positions and further discussion on the basis for rod insertion limits are provided in the Technical Specifications.

3.3.2.5.3 Power Shaping With The Part-Length Control Rod Bank Part-length rods are no longer permitted in the core and have been removed from the reactor vessel.

3.3.2.5.4 Burnable Abserhet conon Rods alsother The burnable poison rods provide partial control of the excess reactivity available during the first fuel cycle and may be used for any reload to optimize fuel loading. In doing so, these rods 3.3-28

I BVPS-1-UPDATED FSAR Rev. 2 (1/84) .

I abscthef prevent the modera or temperature coefficient from being positive l l

at normal operati.g conditions. They perform this function by j reducing the requ'rement for soluble poison in the moderator at l the beginning of the first fuel cycle as described previously.

The burnable p o i _ycn rod pattern rc=ining in the initial j core "

together with the number of rods per assembly is shown in Figure 3.3-5, while the arrangements within an assembly are displayed in Figure 3.3-4. The initial reactivity worth of these rods is.

shcun in Table 3.3-1. The boron in the rods is depleted with burnup but at a sufficiently slow rate so that the resulting critical concentration of soluble boron concentration is such that the moderator temperature coefficient remains negative at all times for power operating conditions.

3.3.2.5.5 Peak Xenon Startup Compensation for the peak xenon buildup is accomplished using the l boron control system. Startup from the peak xenon condition is l accomplished with a combination of rod motion and boron dilution.

The boron dilution may be made at any time, including during the shutdown period, provided the shutdown margin is maintained.

3.3.2.5.6 Load Follow Control and Xenon Control During load follow maneuvers, power changes are accomplished using control rod motion and dilution or boration by the boren i system as required. Control rod motion is limited by the' control '

red insertion limits on the full-length rods as provided in the Technical Specifications and discussed in Sections 3.3.2.5.2.

Reactivity changes due to the changing xenon concentration can be centrolled by rod moti*on and/or changes in the soluble boron concentration.

3.3.2.5.7 Burnup Control of the excess reactivity for burnup is accomplished using soluble boron. The boron concentration must be limited during operating conditions to insure the moderator temperature coef-ficient is negative. Sufficient burnable poison was installed at the beginning of Cycle 1 to give the desired Cycle 1 lifetime without exceeding the boron concentration limit. The practical minimum boron concentration is 10 ppm.

3.3.2.6 Control Rod Patterns and Reactivity Worth The full-length Rod Cluster Control Assemblies are designated by function as the control groups and the shutdown groups. The terms " group" and " bank" are used synonymously throughout this report to describe a particular grouping of control assemblies.

The Rod Cluster Assembly pattern is displayed in Figure 3.3-36 which is not expected to change during the life of the plant.

The control banks are labeled A, B, C cnd D and the shutdown banks are labeled SA, SB, etc., as applicable. Each bank, although cperated and controlled as a unit, is comprised of two 3.3-29

BVPS~1-UPDATED FSAR Rev. 6 (1/88) 1 The effective U-238 temperature for resonance absorption is obtained from the radial temperature distribution by applying a radially ;

dependent wei function. The weighting function was determined from REPAD("ghting > Monte Carlo calculations of resonance escape probabilities in several steady state and transient temperature distributions. In each case a flat pellet temperature was determined which produced the same resonance escape probability as the actual distribution. The weighting function was empirically determined from these results. i i

(

'The effective Pu-240 temperature for resonance absorption is I determined by a convolution of the radial distribution of pu-240 number densities from LASER burnup calculations and the radial f '

weighting function. The resulting temperature is burnup dependent, but the difference between U-238 and pu-240 temperatures, in terms of reactivity effects, is small.

The effective pellet temperature for pellet dimensional change is that value which produces the same outer pellet radius in a virgin pellet as that obtained from the temperature model. j The effective '

clad temperature for dimensional change is its average value.

The temperature calculational model has been validated by plant Doppler defect data as shown in Table 3.3-6 and Doppler coefficient data as shown in Figure 3.3-42. Stability index measurements also provide a sensitive measure of the Doppler coefficient near full power (See Section 3.3.2.8). It can be seen that Doppler defect data is typically within 0.2% of prediction.

3.3.3.2 Macroscopic Group Constants g/cgyh TNWAT C I -+ kruse f 5 Macroscopic few-group constants and analogous microscopic cross sections (needed for feedback and microscopic d epletion calculations) generated for fuel cells by a rcccnt version of the LEOPARD (") and CINDER (") Codes, which are linked "SY internally and provide burnup dependent cross sections. Normally a be simplified approximation of the main fuel chains is used; however, where needed, a complete solution for all the significant isotopes in the fuel chains from Th-232 to Cm-244 is available("). Fast and thermal cross section library tapes contain microscopic cross l sections taken for the most part from the ENDF/F(") library, with a few exceptions where other data provide better agreement with critical experiments, isotopic measurements, and plant critical boron values. The effect on the unit fuel cell of non-lattice components in the fuel assembly is obtained by supplying an appropriate volume _

fraction of these materials in an extra region which is homogenized t with the unit cell in calculations.

the fast (MUFT) and thermal (SOFOCATE) flux In the thermalcalculation, the fuel rod, clad, and moderator are homogenized by energy-dependent disadvantage factors derived from an analytical fit to integral transport theory results.

3.3-40 d

Jnsert "C1" (to page 3.3-40)

There are two lattice codes used for the generation of macroscopic group constants for use in the spatial few group diffusion codes. The first code is a linked version of LEOPARD (Reference J5) and CINDER (Reference 16) and-the second code is PHOENIX-P (Reference K ). A description of each code follows. 90-Insert "C2" (to page 3.3-41)

The PHOENIX-P computer code is a two-dimensional, multigroup, transport based lattice code and capable of providing all necessary data for PWR analysis. Being a dimensional lattice code, PHOENIX-P does not rely on pre-determined spatial / spectral interaction assumptions for a heterogeneous fuel lattice, hence, will provide a mere accurate multi-group flux solution

' than versions of LEOPAR0/ CINDER. The PHOENIX-P computer code is approved by the USNRC as the lattice code for generating macrosc pic and microscopic few group cross sections for PWR analys ence The solution for the detailed spatial f x an 6

nergy distribution is divided into two major steps-in PHOENIX-P ). In the first step, a two-dimensional fine energy group nodal solution is obtained which couples individual subcell regions (pellet, clad and moderator) as well as surrounding pins. PHOENIX-P uses a method based on the Car 1vik's collision probability approach and heterogeneous response fluxes which preserves the heterogeneity of the pin cells and their surroundings. The nodal solution provides accurate and detailed local' flux distribution which is then used to spatially homogenize the pin cells to fewer groups.

The second step in the solution process solves for the angular flux distribution using a standard S4 discrete ordinates calculation. This step is based on the group-collapsed and homogenized cross sections obtained from the first step of the solution. The S4 fluxes are then used to normalize the detailed spatial and energy nodal fluxes. The normalized nodal fluxes are used to compute reaction rates, power distribution and to deplete the fuel and burnable absorbers. A standard B1 calculation is employed to evaluate the fundamental mode critical spectrum and to provide an improved fast diffusion coefficient for the core spatial codes.

The PHOENIX-P code employs a 42 energy group library which has been derived mainly from ENDF/B-V files. The PH0ENIX-P cross sections library was designed to properly capture integral properties of the multi-group data during group collapse, and enabling proper modeling of important resonance 1 parameters. The library contains all neutronic data necessary for modeling i

fuel, fission products, cladding and structural, coolant, and control / bur,nable absorber materials present in Light Water Reactor cores.

Group constants for burnable absorber cells, guide thimbles, instrument thimbles, control rod cells and other non-fuel cells can be obtained directly from PHOENIX-P without any adjustments such as those required in the cell or ID lattice codes.

- - - - _ _ _ 1

BVPS-1-UPDATED FSAR Rev. 6 (1/88)

Group constants for burnable poison cells, guide thimbles, instrument thimbles, and interassembly gaps are generated in a manner analogous to the fuel cell calculation. Reflector group constants are taken from infinite medium LEOPARD calculations. Baffle group constants are calculated from an average of core and radial reflector microscopic group constants for stainless steel.

Group constants for control rods are calculated in a linked version of the HAMMER (20 and AIM (22) codes to provide an improved treatment of selfshielding in the broad resonance structure of these isotopes at epithermal energies than is available in LEOPARD. The Doppler broadened cross sections of the control rod materials are represented as smooth cross sections in the 54-group LEOPARD fast group structure and in 30 thermal groups. The four-group constants in the rod cell and appropriate extra region are generated in the coupled space-energy transport HAMMER calculation.

A corresponding AIM calculation of the homogenized rod cell with extra region is used to adjust the absorption cross sections of the rod cell to match the reaction rates in HAMMER. These transport-equivalent group constants are reduced to two-group constants for use in space-dependent diffusion calculations. In discrete X-Y calculations only one mesh interval per cell is used, and the rod group constants are further adjusted for use in this standard mesh by reaction rate matching the standard mesh unit assembly to a fine-mesh unit assembly calculation.

Validation of the cross section method is based on analysis of critical experiments as shown in Table 3.3-7, isotopic data as shown in Table 4.3-8, plant critical boron (C) values at HZP, BOL, as shown in Table 3.3-9 and at HFP as a function of burnup as shown in Figures 3.3-43 through 3.3-45. Control rod worth measurements are shown in Table 3.3-10.

Confirmatory critical experiments on burnable poisons are described in Reference 23.

3:45647' C2 -+ !

3.3.3.3 Sp tial Few-Gr,9up Diffusion Calculations place ein C3 tial ( eW-group diffusion calculations consist primarily <dI' two- diffusion X-Y calculations using an updated version-of,the TURTLE c two-group nodal calculations using PALADON erence 31 and 32) and -group axial calculations using an u ed version of the PANDA code. .

Discrete X-Y calculations mesh _ cell) are carried out to determine critical boron conce ons and power distributions in the X-Y plane. An axial age the X-Y plane is obtained by synthesis from unrodded a odded planes. ial effects in unrodded depletion calculation e accounted for by the al buckling, which varies with burnu nd is determined by radial deple calculations which are m ed in reactivity to the analogous depletion calculati . The moderator coefficient is evaluated by var ' the inle emperature in the same X-Y calculations used for po ribution and reactivity predictions.

3.3-41

Insert "C3" (to page 3.3-41)

Spatial few-group diffusion calculations have primarily consisted of two-group X-Y calculations using an updated version of the TURTLE code, and two-group axial calculations using an updated version of the PANDA code.

Howeter, with the advent of VANTAGE 5 fuel and, hence, axial feetures such as axial blankets and part length burnable absorbers, there will be a greater reliance on three dimensional nodal codes such as 3D ANC (Advanced Nodal Code) (Reference ). The three dimensional nature of the nodal codes provides both the radial and axial power distributions.

3'l Nodal three dimensional calculations are carried out to determine the critical boron concentrations and power distributions. The moderator coefficient is evaluated by varying the inlet temperature in the same calculations used for power distribution and reactivity predictions.

BVPS-1-UPDATED FSAR Rev. 6 (1/88)

Validation of TURTLE reactivity calculations is associated with the alidation of the group constants themselves, as discussed in Section

.3.3.2. Validation of the Doppler calculation is associated with the fuel temperature validation discussed in Section 3.3.3.1.

Validation of the moderator coefficient calculations is obtained by comparison with plant measurements at hot zero power conditions as shown in Table 3.3-11.

Axial calculations are used to determine differential control rod worth curves (reactivity versus rod insertion) and axial power shapes during steady state and transient xenon conditions (flyspeck curve).

Group constants and the radial buckling used in the axial calculation are obtained from three dimensional calculations (3-D TURTLE or

-PAfrABON from which group constants are homogenized by flux weighti g7'\ggg validation of the spatial codes for calculating power distributions Irvolves the use of in-core and ex-core detectors and is discussed in Faction 3.3.2.2.7.

Based on comparison with measured data it is estimated that the accuracy of current analytical methods is:

1. i 0.2% for Doppler defect

2. i 2 x 10-5/*F for moderator coefficient

3. i 50 ppm for critical boron concentration with depletion

4. i 3% for power distributions

5. t 0.2% for rod bank worth

! A K is asecI s hio -Lessi~ a t a u/ Ae - d,-a s ui I ca lcu lahu s . Rnc ca , he useal h>< safely a alyses ca/oc taf, '

Ctihibalboten CNCe b " NH3 C"YNl 3 W W Ibh i na cbwh coe Efo 'e n d d,1, c $c . -

1 I

l l

3.3-42 l

l l

l

BVPS-1-UPDATED FSAR Rev. 6 (1/88)

References for Section 3.3 (Cont'd)

24. Ocicted by Revicion 0.

25.-Ocicted by Revicien 0. '

[e ace MN6

26. Delcted by Rcvision 0. k*bE#t"CES A Y, # # # A

27. J. M. Hellman, (Ed.), " Fuel Densification Experimental Results and Model for Reactor Application," WCAP-8219, Westinghouse Electric Corporation (October, 1973).

j

28. J. M. Hellman and J. W. Yang, " Effects of Fuel Densification Power Spikes on Clad Thermal Transients," WCAP-8359, Westinghouse Electric Corporation (July, 1974).

29. T. Morita, et al., " Topical Report, Power Distribution Control and Load Following Procedures," WCAP-8385, Westinghouse Electric Corporation (September, 1974).

30. C. Eicheldinger, " Westinghouse Letter to D. B. Vassato," l NS-CE-687, Westinghouse Electric Corporation (July, 1975).

31. Ccaden, T. M., ct. cl., "PALADON Ucstinghcusc Nodci Computcr l

-Codc", WCAr--04SSA (Proprictcry) cnd WCAr-04SSA (Nen-rreprictcry-)r !

--December, 1070.

32. Ankney, R. O., "rA..ADON Ucstinghcusc Ncdci Ocaputcr Codc, GuyylcmuuL l' , WCAF- 94 85 enyylcineuL, ScyLembus, 1981.

33. Nuclear Regulatory Commission, Letter to All Power Reactor Licensees, from B. K. Grimes, April 14, 1978, "OT Position for Review and Acceptance of Spent Fuel Storage and Handling Applications." l

34. W. E. Ford III, et al., "CSRL-V: Processed ENDF/B-V 227-Neutron-Group and Pointwise Cross-Section Libraries for criticality Safety, Reactor and Shielding Studies," !

ORNL/CSD/TM-160 (June 1982).

35. N. M. Greene, et al., "AMPX: A Modular Code System for !

Generating Coupled Multigroup Neutron-Gamma Libraries from ENDF/B," ORNL/TM-3706 (March 1976). i

36. L. M. Petrie and N. F. Cross, " KENO IV--An Improved Monte Carlo Criticality Program," ORNL-4938 (November 1975). ;

37. M. N. Baldwin, et al., " Critical Experiments Supporting Close :

Proximity Water Storage of Power Reactor Fuel," BAW-1484-7, (July '

1979).

1

38. J. T. Thomas, " Critical Three-Dimensional Arrays of U (93.2) --

Metal Cylinders," Nuclear Science and Engineering, Volume 52, .

pages 350-359 (1973). I 3.3-45 Nd 1~UJ2A7~ E /Ge&en c e5 J 9', V o a V/

i

. Insert E References for Section 3.3 p 3.3-44

24. Strawbridge, L.E., and Barry, R. F., " Criticality Calculation for Uniform Water-Moderated Lattices," Nuclear Science and Eng. 23, 58 (1965).

25. Nodvik, R,J., "Saxton Core II Fuel Performance Evaluation,"

WCAP-3385-56 Part II, " Evaluation of Mass Spectrometric and Radiochemical Analyses of Irradiated Saxton Plutonium Fuel," July,

'1970.

26. Leamer, R D., et.al., "PU02-U02 Fueled Critical Experiments,"

WCAP-3726-1, July, 1967.

37 . Liu, Y. S., et.al., "ANC: A Westinghouse Advanced Nodal Computer Code," WCAP-10965-P-A, September, 1986, f/O . Nguyen, T.Q. , et.al . , " Qualification of the PH0ENIX-P/ANC Nuclear Design System for Pressurized Water Reactor Cores," WCAP-11596-P-A, June, 1988.

yf . Mildrum, C.M. , Mayhue, L.T. , et.al. , " Qualification of the PH0ENIX/POLCA Nuclear Design and Analysis Program for Boiling Water Reactors," WCAP-10841 (Proprietary) and WCAP-10842 (Non-proprietary),

June, 1985.

9

BVPS-1-UPDATED FSAR Rev. 0 (1/82) 3.4 THERMAL AND HYDRAULIC DESIGN 3.4.1 Design Bases The overall objective of the thermal and hydraulic design of the reactor core is to provide adequate heat transfer which is compatible with the heat generation distribution in the core such that heat removal by the Reactor Coolant System or the Emergency Core Cooling System (when applicable) assures that the following performance and safety criteria requirements are met:

1. Fuel damage, the penetration of the fission product barrier (i.e., the fuel rod clad), is not expected during normal operation and operational transients (Condition I) or any transient conditions arising from faults of moderate frequency (Condition II). It is not possible, however, to preclude a very small number These will be within the capability of rod failures. '

of the plant cleanup system and are consistent with the plant design bases.

2. The repctor can be brought to a safe state following a Condition III event with only a small fraction of fuel rods occur damaged although sufficient fuel damage might to preclude resumption of operation without considerable outage time.

3. The reactor can be brought to a safe state and the core can be kept suberitical with acceptable heat transfer geometry following transients arising from Condition IV events.

In order to satisfy the above criteria the following design bases have been established for the thermal and hydraulic design of the reactor core.

3.4.1.1 Departare from Nucleate Boiling Design Basis Basis Departure from nucleate boiling will not occur on at least 95% of the limiting fuel rods during normal operation and operational transients and any transient conditions arising from faults of moderate frequency (Condition I and II events) at a 95%

confidence level. -Meteorically this he; b;;n c;n:crvatively met by limiting the minimen deperLere frem nucleete 1~iling retic -

-- (DNBR) Le 1.00 end fes Lhie epplic.Livu a minimum DiGR vf 1. 3 0 -

mill centinue to be usedr- g pgg g Discussion By preventing departure from nucleate boiling, adequate heat transfer is assured between the fuel clad and the reactor coolant, thereby preventing clad damage as a result of inadequate 3.4-1

INSERT A:

For this application, this criterion is met by limiting the minimum departure from nucleate boiling ratio (DNBR) to a design value of 1.21. Plant specific margin to accomodate rod bow and other DNB penalties and allow for flexibility in the design, operation and analysis of the plant is provided by performing the safety analyses to a DNBR limit value of 1.33.

l INSERT B:

fiistorically, the DNBR limit has been 1.30 for Westin this application, the WRB-1 correlation ln significant improvement in the accuracy o(Reference is used. With the 37)ghouse using limit ofthis 1.17correlation applies. instead of the previous correlation, a DNBR correlation th The design method used to meet the DNB design basis is the MINI-Revised Thermal Design Procedure (Reference 38) which is a conservative application of the Revised Thennal Design Procedure (ReferenceB4 In the MIN 1-RTDP method, uncertainties in .he nuclear peaking factors and fuel fabrication parameters are combined statistically with the DNB correlation uncertainties to define the DNBR design limit such that there is at least a 95 percent probability (with 95 percent confidence) that DNB will not occur when the calculated minimum DNBR is equal to or greater than the design limit.

The uncertainties included in the MINI-RTDP method are for the nuclear enthalpy hot-channel factor, F(N,AH); the enthalpy rise engineering hot-channel factor, F(E,4H); and the THINC-IV and transient codes. Since the uncertainties in these parameters are considered in detennining the design DNBR value, the plant safety analyses are perfanned using input values without uncertainties for these parameters. For this application, the DNBR design limit value is 1.21.

In addition to the considerations above, a specific plant allowance has been considered in the present analysis. In particular, a DNBR limit value of 1.33 has been used in the safety analyses for the plant. The difference between the DNBR value used in the safety analyses and the design DNBR value (1.33 vs. 1.21) penalties plant provides specific that may DNB margin to offset the rod bow penalty and other DNB occur.

the design, operation or analysa,.s of the plant.This 0NB margin may also be used INSERT C:

iar conditions outside the range of parameters for the WRB-1 correlation (refer

!. o Section 3.4.2.3.1), the W-3 correlation is used with a DNBR correlation h.r.it of 1.30 for pressure equal to or greater than 1000 psia. For low pressure applications (500-1000 psia), the W-3 DNBR correlation limit is 1.45 (Reference 8l!P) .

B'/PS-1-UPDATED FSAR Rev. O (1/82) cooling. Maximum fuel rod surface temperature is not a design basis as it will be within a few degrees of-coolant temperature during operation in the nucleate boiling region. Limits provided by'the nuclear control and protection systems are such that this design basis will be met for transients associated with Condition II events including overpower transients. There is an additional large DNBR margin at ' rated power operation and during normal operatin transients.

$h5 U bW0 3.4.1.2 Fuel Temperature Design Basis Basis During modes of operation associated with condition I and condi-tion II events, the maximum fuel temperature shall be less than '

the melting temperature of UO 2. The UO2 melting temperature for at least 95% of the peak power fuel rods will.not be exceeded at the 95% confidence level. The melting temperature of UO2 taken as 5080 F IO unirradiated and is considered to decline 58 is- F per 10,000 MWD /MTU. By precluding UO2 melting, the fuel geometry is preserved and possible adverse effects of molten 002 on the cladding are eliminated. To - preclude center melting and as a basis for overpower protection system setpoints, a calculated centerline fuel temperature of 4700 F has been selected as the overpower limit. This provides sufficient margin for uncertain-ties in the thermal evaluations as described in Section 3.4.2.10.1.

Discussion Fuel rod thermal evaluations are performed at rated power, maximum -overpower and during transients at various burnups.

These analyses assure that this design basis as well as the fuel integrity design bases given in Section 3.2 are met. They also provide input for the evaluation of Condition III and IV faults given in Chapter 14.

3.4.1.3 Core Flow Design Basis

~ )

i Basis i

I A minimum of 95.5% of the thermal flow rate will pass through the fuel rod region of the cc,re and be effective for fuel rod cool-ing. Coolant flow through the thimble tubes as well as the )

leakage from the core barrier-baffle region into the core are not 3 considered effective for heat removal.

Discussion l Core cooling evaluations are based on the thermal flow rate (minimum flow) entering the reactor vessel. A maximum of 4.5% of this value is allotted as bypass flow. This includes RCC guide thimble cooling flow, head cooling flow, baffle leakage, and leakage to the vessel outlet nozzle.

3.4-2

BVPS-1-UPDATED FSAR Rev. 0 (1/82) 3.4.1.4 Hydrodynamic Stability Design Bases Modes of operation associated with Condition I and II events shall not lead to hydrodynamic instability.

3.4.1.5 Other Considerations The above design bases together with the fuel clad and fuel assembly design bases given in Section 3.2.1.1 are sufficiently comprehensive so additional limits are not required.

Fuel rod diametral gap characteristics, moderator-coolant flow velocity and distribution, and moderator void are not inherently limiting. Each of these parameters is incorporated into the thermal and hydraulic models used to ensure the above mentioned design criteria are met. For instance, the fuel rod diametral gap characteristics change with time (see Section 3. 2.1. 3.1) and the fuel rod integrity is evaluated on that basis. The effect of the moderator flow velocity and distribution (see Section 3.4.2.3) and moderator void distribution (see Section 3.4.2.5) are included in the core thermal (THINC) evaluation and thus affect the design bases.

Meeting the fuel clad integrity criteria covers possible effects of clad temperature limitations. As noted in Section 3. 2.1. 3.1, the fuel rod conditions change with time. A single clad tempera-ture limit for Condition I or Condition II events is not appro-priate since of necessity it would be overly conservative. A clad temperature limit is applied to the loss-of-coolant accident (Section 14.3.1) control rod ejection accident,(2) and locked rotor accident.(j'I 3.4.2 Description 3.4.2.1 Summary Comp:ricca-

'T e design of the BVPS-1 reactor with the 17 x 17 fuel rod arr ;

pe ssembly has the following identical thermal and hyd lic parame a as the 15 x 15 fuel rod array reactor design:

1. Co ower

2. Vessel lo flow rate

3. System pressure

4. Coolant inlet temp e

5. Core and ve average and - 't coolant temperatures

6. Open ttice fuel rod array.

Values each parameter are presented in Table 4-1 for all coo loops in service and in Table 3.4-2 for a but one lant loop in service. It is also noted that in thz ower I capability evaluation, there has not been any change in 3.4-3

I BVPS-1-UPDATED FSAR Rev. 0 . (1/82)

Meef /ka DA6 desof USS of Seeh4 3.yU design criterin. The reactor is still designed toAa~Linimum

-n=n >

operation, 1.30 as operational well as no fuel centerline melting during normal transients and faults of moderate frequency.

e basis for the improvements of this. design is an increase i th number of fuel rods in the reactor core and the result red tion in average and maximum linear heat generation g

ate (LHG kW/ft) and heat flux. This has a direct impact on re cing the m imum expected cladding temperature during a 1 ss-of-coolant the maxim ccident fuel temperatures.

(see Sections 14.2 and 14.3) and also reduces The peak to average va e of the cosine axi heat flux distribution used for design NB calcu-lations is ibility as di ussed in Section 3.4.3.2.

creased to 1.55 to provide greater op ating flex-A compari on of the two designs is summ rized in Table 3.4-1.

Approximately 15% 'n DNBR has been retained as argin in all DNB analyses performed 'n this application, (i. ., a multiplier of 0.865 as discussed 1 Section 3.4.2.3.2). he above margin is being retained for the following reasons:

1. To incorporate final results of the DNB and mixing tests describe in Section .5 which are expected to confirm the app ' cation the DNB correlation and mixing model descr bed i Section 3.4.2.3.3 to the 17 x 17 fuel geometry a d id type

2. To incorporate final' sults of experimental D-Loop hydraulic tests o the 17 x 17 fuel assembly to confirm the pres ure dr characteristics used to establish the pr' ary loop ow rate

3. To allow for ny fabrication tolerances larger than presently us d in design analys s that could occur in the manufa uring of-a 17 x 17 fu 1 assembly.

The results of the 7 x 17 geometry DNB tests en are discussed in Section 3.4.2. .1, and indicate a DNBR multip er of 0.88 (for 26" grid spacin is required for the "R" grid NB correlation (Sections 3.4. .3.1 and 3.4.2.3.2). Preliminary sults of the D-Loop hydra ic testsI'O (i.e., pressure drops) a. lower than predicted, nd hence no specific DNB margin is exp ted to be required. There is, therefore, remaining margin whi can be used as scussed in item 3 above if required.

'The e ects of fuel densification on DNB have been eva ated l- util' ing the methods and models described in detail in Refer ce 80 nd summarized in the following sections. The net effect f f 1 densification is a reduction of 0.2% in the DNBR. This i scribed in Section 3.4.2.2.1.

3.4.2.2 Fuel and Cladding Temperatures (Including Densification)

Consistent with the thermal-hydraulic design bases described in Section 3.4.1, the following discussion pertains mainly to fuel 3.4-4

\VPS-1-UPDATED FSAR Rev. 0 (1/82) and Zircaloy has been measured and found to be dependent on the contact pressure, composition of the gas at the interface and the surface roughness of the fuel and cladding.f2'i(381 This infor-mation together with the surface roughness found in the fuel and clad ' installed in Westinghouse reactors leads to the following correlation:

K h = 0.6p + gas (3.4-4) 14.4 x 10~5 where h = contact conductance, Btu /hr-ft2 _op p = fuel clad contact pressure, psi 3.4.2.2.4 Surface Heat Transfer Coefficients The fuel rod surface heat transfer coefficients during subcooled forced convection and nucleate boiling are presented in Section 3.4.2.8.1.

3.4.2.2.5 Fuel Clad Temperatures The outer surface of the fuel rod clad at the hot spot operates at a temperature of approximately 660F for steady state operation at rated power throughout core life due to the onset of nucleate boiling. Initially (beginning-of-life) this temperature is that of the clad metal outer surface.

During operation over the life of the core, the buildup of oxides and crud on the fuel rod surface causes the clad surface tempera-ture to increase.

evaluation for Allowance is made in the fuel center melt this temperature rise. Since the thermal-hydraulic design limits DNB, adequate heat transfer is provided between the fuel clad and the reactor coolant so that the core thermal output is not limited by considerations of the clad temperature. Figure 3.4-4 shows the axial variation of average clad temperature for the average power rod both at beginning and end-of-life.

3.4.2.2.6 Treatment of Peaking Factors The total heat flux hot channel factor, F (Q) , is defined by the

ratio of the maximum to core average heat flux. ?.c precented in

-Tabic 3.3 2 and discucccd in Ocction 3.2.2.2.C, t.he design velue

-for F (^) for noraci cpcrction 1: 2.22 including fuel densifica tien cffcctc. Thic result; in c pcck iccal power of 11.0 kW/ft-f ct full pcucr ccaditienc. As described in Section 3.3.2.2.6 the peak local power at the maximum overpower trip point is 18.0 kW/ft. The centerline temperature at this power level must be below the UO2 melt tem over thp lifetime of the rod, (1~he pe.al( lcy/,0ewert hl/0$f&et perature cdNes is y O t r A 73b/e JM-/

a><d h basec( n fka des l*s va/a e of W$ & nonal ofe&As.

3.4-8

1 BVPS-1-UPDATED FSAR Rev. 0 (1/82) subchannels are also divided into axial steps such that each may be treated as a control volume. By solving simultaneously the

. mass, energy, and momentum equations, the local fluid conditions in each control volume are calculated. The W-3 correlation, developed from single channel data, can be applied to rod bundles by using the subchannel local fluid conditions calculated by the THINC code.

It was shown by Tong l8 M that.the above approach yielded conser-vative predictions particularly in rod bundles with mixing vane grid spacers. Hence : correctic.; facter v : d celcped tc adapt' he W-3 correlation, which was developed based on single channe

.d ta, to rod bundles with spacer grids. This correction fact ,

te .ed the " Modified Spacer Factor", was developed as a multi i-er the W-3 correlation.

The " dified Spacer Factor" was developed from rod b die DNB test r ults conducted in the Westinghouse high pres re water loop at Columbia University. These tests were e ducted on non-unifo axial heat flux test sections to dete ine the DNB performanc of a low parasitic, top-split mix g vane grid design, her fter referred to as the "R" grid. description of this test pr ram and a summary of the results re given below.

The grid to b used in the 17 x 17 fuel asse y will be similar in design to th "R" grid.

"R" grid rod bun e DNB tests I8 U I8 O wer conducted over a wide range of simulated eactor conditions. ese conditions were:

1. Axial grid pacing 20 nches, 26 inches, and 32 i ches

2. Local DNB qual y -15 percent to +15 percent

3. Local mass veloci y 1.6 x 10' to 3.7 x 10' 2

lbs/hr-ft

4. Local inlet tem rat e 440F to 620F

5. Pressure 1490 to 2440 psia

6. Local hea flux .3 x 10 5 to 1.1 x 10 6 2

B u/hr-ft

7. Axial eat flux di ribution Non- iform (Cos u and u Sin u)

8. Heated length 8 feet a d 14 feet 9 Heater rod O.D. 0.422 inch Replae d roseer p 3.4-10

INSERT 0:

The WRB-1 DNB correlation (Reference 77) was developed based exclusively on the large bank of mixing vane grid rod bundle CHF data (over 1100 points) that Westinghouse has collected. The WRB-1 correlation, based on local fluid conditions, represents the rod bundle data with better accuracy over a wide range of variables than the previous correlation used in design. This correlation accounts directly for both typical and thimble cold wall cell ,

effects, uniform and non-uniform heat flux profiles and variations in rod !

heated length and grid spacing. The WRB-1 correlation is applicable to the 17x17 STD and VANTAGE SH fuel.

The applicable range of variables is:

Pressure : 1440 $ P 5 2490 psia 6 2 Local Mass Velocity : 0.9 5 Gloc/10 53.7lb/ft-hr Local Quality :

-0.2 5 X1oc 5 0.3 Heated length, Inlet to : Lh 5 14 feet ;

CHF Location Grid Spacing :

13 5 gsp g 32 inches Equivalent Hydraulic Diameter : 0.37 5 d, 5 0.60 inches _

Equivalent Heated Hydraulic : 0.46 5 dh 1 o.s8 inches Diameter Figure 3.4-5 shows the measured critical heat flux plotted against predicted critical heat flux using the WRB-1 correlation.

In order to meet the design criterion that DNB will not occur at a 95 percent probability with a 95 percent confidence level, a limiting value of DNBR is determined by the method of Owen (Reference &f). Owen has prepared tables which give values of K such that "at least a proportion P of the population is greater than 5 - K s pwith confidence 7," where E[P and s are the sample mean

~

and standard deviation. When this method was carried out using the data on Figure 3.4-5, the results indicated that a reactor core with these fuel geometries may operate with a minimum DNBR of 1.17 and satisfy the design criterion.

I J

I BVPS-1-UPDATED FSAR Rev. 0 (1/82) e experimental program consisted of a DNB test series for bot a all and/or partial channel surface heated condition in a 6 ro bundle arranged in a 4 x 4 array. A radial power profile as simu ated by operating the central 4 rods of the bundle a 15 perce t higher power than the other rods. Two test series were conduc ed on 26 inch axial grid spacing: 1) all channel rface heated ondition (typical cell), 2) partial channel surface heated c dition (thimble " cold wall" cell).

For the thi le cold wall tests series, one of the c tral four heater rods was replaced by an unheated rod. T simulated unheated thi e was made up of a thin steel rod ov r which were placed cerami cylinders with an outer diameter equal to the thimble outer di meter. These thimbles are atta ed to the grid in the same mann as in the reactor core using sleeve which is brazed into the g 'd and then bulged out above nd below the grid to connect to the t imble.

These rod bundle DNB data have been analy ed and a " modified" spacer factor (37)(381 s been developed t conservatively incor-porate the "R" mixing va e grid benefit f r both typical and cold wall cells. This "modifi d" spacer fact r is:

F'(S) = ( P )'*5 ( .445 - O. 71 L) [e (x + 0.2) 2 - 0.73) 225.896

+K(S) G (TDC)0+83 6

10 0.019 (3.4-6) where:

P = the primary syst m pressu e, psia L = the total heate core leng , feet x = the local qua ty expressed n fractional f orm G = the local ma velocity, lbs r-ft 2 TDC = the thermal iffusion coeffici nt K(S) = the axial grid spacing coeff'cient which has the following values:

Grid Spacing, inch K(S) l 32 0.027 26 0.046 20 0.066 1

Figure 3.4- shows all the "R" grid typical cell da a. Figure I 3.4-6 show all the "R" grid thimble cell data. The redicted heat flux in Figures 3.4-5 and 3.4-6 incorporates the modified

) spacer f ctor per Equation 3.4-6 for typical cells and quation 3.4-7 f thimble cold wall cells:

PRED " B, N x F'(S) Q.b W-3 3.4-11

DVPS-1-UPDATED FSAR Rev. 0 (1/82) ere 9" W-3 DNB,N is the predicted non-uniform DNB heat flux usi g The W-3 correlation as described in Reference 39.

PRED " W-3 B, N, CW x F'(S) (3.4-8) where 9"g (DNB, N, CW) is the predicted non-unifo ' DNB heat flux with flow cel having a cold (unheated) wall eva ated with the W-3 cold wall c Reference 40.

relation described in the Secti n 3.4.2.3.2 and F ' (S) as defined in quation (3.4-6) is th same in both Equa-tions (3.4-7) and (3. -8) for both typica and thimble cold wall cells.

Effect of 17 x 17 Geo etry on DNB A test program similar to th one scribed above was conducted at the Westinghouse high pres re ater loop at Columbia Univer-sity. In this test program, D data was obtained for 17 x 17 fuel assembly geometry re I in a 5 rod bundle array.

Test results were obtained r typi 1 cells (all walls heated) in 8 foot and 14 foot bund s, and for thimble cold wall cells in an 9 foot bundle. All b dies had ' uni rm axial heat flux, and mixing vane grids on 26 'nch spacing. '

The data obtained we analyzed with the e isting "R" grid DNB correlation describ above to determine the fect of DNB of the 17 x 17 fuel asse ly gecmetry. Plots of rat'o of measured to predicted DNB he.. flux versus flow parameters ow that the "R" grid correlati properly accounts for local fluid parameters.

However, the ' " grid correlation consistently ove predicts the DNB heat fl x. Hence, a multiplier of 0.88 on e modified spacer fac r F'(S) is required to correctly predict he magni-tude of t e DNB heat flux for 17 x17 geometry.

Figure 3. 4-6A , shows the 17 x 17 data obtained in thi test prog m .- The predicted heat flux includes the 0.88 multipli on the modified spacer factor, Equation (3.4-5), as noted abo e.

T DNB probability curve, based on'the~ statistics of the dat shown in Figure 3.4-7.

Effects of Fuel Densification on DNB Effect of Heat Flux Spikes An discussed in Section 3.3.2.2 and Reference 80 a gap or combi-nation of gaps results in a heat flux spike on the individual or adjacent fuel' rods. Recent Westinghouse high pressure DNB Water TestsI'll on a 14 foot axially non-uniformly heated 4 x 4 rod bundle were carried out to measure the effect of heat flux 3.4-12

i BVPS-1-UPDATED FSAR Rev. 0 (1/82) spikes.

inch spacing.

The rod bundle incorporated mixing vane grids on a 26 A 20% heat flux spike was placed on three adjacent rods at the axial location where DNB is most likely to occur.

This test series was run at the same conditions as those of two earlier test series which had unspiked rods so that a comparison of spiked and unspiked data could be made and the spike effects isolated.

three spiked rods.

Figure 3.4-23 shows the relative positions of the the spiked rods.

Figure 3.4 24 shows the heat flux profile of The test facilities consisted of a high pressure loop capable of supplying water at pressures up to 2400 psia with flow rates up to 400 GPM and inlet temperatures in excess of 600F. The power supply was capable of delivering up to 4.5 MW.

Using these test facilities, a 14 foot, 16 rod test section can be operated over a wide range of test parameters. For the present tests, these ranges were:

1. Pressure 1500-2400 psia

2. Inlet temperature 401-569F

3. Mass velocity 1.5-3.5 x 10' lb/hr-ft.2

-Figurc 3.d.-25 chouc the mercured vc. the predicted DNE ". cat flun

-27.

of the cpiked data plotted 'cith the uncpiked data cf Pcference Thic uncpiked mining vanc grid data ucc taken 'eith varicut-tect ccctienc for "hich va ricd. -

-nu-ber cf important parameters were-The results of the spike test series indicated that the spike effect on DNB is -cmaller than had bcca predicted .cith the non uni vum um . mo .m m . r, m.m - in the '? 3 ONO corrciation.

Further crc, the =cacured cpike cffect is so small that it lies

) within the repeatability of DNB mea surements ,-which arc included-margin leading to a accign minimum DNOR cf 1.30. The

-in th spike geometry modeled in the above rod bundle experiment was also more severe than that presently ascribed to fuel densifica-I tion effects, and hence, the absence of a spike effect indicated that a special spike factor in DNB need not be incorporated into the BVPS-1 reactor design.

(

Effects of Pellet Eccentricity and Clad Ovality on DNB Individual fuel pellets can be eccentrically located in the clad at BOL. The clad may also assume an oval shape at some later time in life. Both of these cases will produce azimuthal varia-tion of the pellet clad gap. However, these local heat flux peaks will have limited axial lengths at any azimuthal angle.

For the eccentric case the local heat flux peak at a given azi-muthal angle will have a maximum length equal to several pellet lengths. This is due to the randomness of the angle of contact of 3.4-13

BVPS-1-UPDATED FSAR Rev. 0 (1/82) the pellets in the rod at BOL.

The randomness has been verified by observation of radiographs of Beznau 1 fuel rods and is due in part to the variation in pellet diameter.

For the clad ovality case, the local heat flux peak also has a maximum angle.

length equal to several pellets at a given azimuthal of the cracked This is pellet due tofragments the randomness in the axial of thedirection.

azimuthal location The recent spike DFB tests ten described previously indicate that

'for.360' circumferential heat flux spikes at 20% magnitude and 6" long,-a special spike factor on DNB need not be incorporated into Westinghouse reactor designs which incorporate the Westinghouse type mixing vane grids. Since the 6 inch length is equivalent to

~10 pellet lengths, no' reduction in DNBR due to pellet eccentric-ity or clad ovality is ' applied in DNB evaluations. Similarly, the heat flux engineering hot channel factor, F(E,0,) of 1.03 which allowed for variations in manufacturing tolerances and was used at a point, to determine the local maximum linear heat generation rate the " hot spot", is no longer considered in DNB evaluations. This subfactor was determined by statistically combining the tolerances for the fuel pellet diameter, density, enrichment, eccentricity and the fuel rod diameter. F (E, Q) continues to be applied in determining the peak power and in fuel pellet temperature evaluations.

The effect of manufacturing tolerances which affect the inte-grated . values along a channel, i.e., enthalpy rise engineering hot channel subfactor corresponding to pellet diameter, density and - enrichment, and fuel rod diameter, pitch and bowing, are still considered in all DNB evaluations as described in Section 3.4.2.3.4.

I 3.4.2.3.2 Definition of DNB Heat Flux Ratio (DNBR)

The DNB heat flux ratio as applied to this design when all flow l cell walls are heated is:

DNBR = q"(DNB), Nh/h/ (3.4-9) q " (loc ) /IU where fNS&& T E i q" (DNB) , N e q"(DNB), EU F

(3.4-10) and q"DNB,EU is the uniform DNB heat flux as predicted by the W-3 DNB correlation (s o when all flow cell walls are heated.

F is the flux factor to account for non-uniform axial heat flux distributions (an with the "C" term modified as in Reference 40.

F' 'C) is thc modified cpecer fcctor dcfined by Equction 2.M in Sectic.- 2.4.2.2.1 2nd using an axial grid crecing coefficient, 3.4-14

INSERT E:

9"DNB, N is the heat flux predicted by the applicable DNB correlation

'For the W-3 correlation,

' INSERT F:

For the WRB-1 correlation,

9"WRB-1

' 4 DNB,.N p

where F is the same flux shape factor that is used with the W-3 correlation

f l

BVPS-1-UPDATED FSAR Rev. 0 (1/82)

K(S), b2ced en the 26 inch grid cpacing dat; juct dcccribcd.

Cince the actual grid cpacing ic apprcximatcly 20 inchec, thic value ic conservatively 10.- cince the DNS performance .::c found to improve Oc 2.Mial grid cpacing ic decresced f 3 M .-

q" loc is the actual local heat flux.

The DNB heat flux ratio as applied to this design when a cold wall is present is:

DNBR = q"(DNB), N, CW[O//$/ (3.4-11) q" (loc) U/// // /

where q" = q"DNB, EU, Dh x CWF ( 's . 4 -12 )

DNB, N, CW F where />7IX q"DNB,EU,Dh is the uniform DNB heat flux as predicted by the W-3 cold wall DNB correlation"'I when not all flow cell walls are heated (thimble cold wall cell).

CWF "' = 1.0-Ru [13.76-1.372 1.7e 4.732 (G) - 0.0535 (3.4-13)

W i

-0,0 T

-0.0619 ( P )*14 -8.509Dh 0.107]

1000 g

-g and Ru = 1 - De/Dh 7Als EtT F +

' (S) defined by Equation 3.4-5 in Section 3.4.2.3.1 is the sa as ed for typical cell.

Values o inimum DNB provided in Section 3.4.3.3 ar he limit-ing value ob 'ned by applying the above two defi ions of DNBR to the appropri cell (typical cell) with a walls heated, or a thimble cold wal 11 (with a partial h ed wall condition).

Approximately 15% i DNBR margin s been retained in all DNB analyses performe in this pplication. Specifically, all DNBR's computed by Equ 'o 3.4-9 and 3.4-11 have been multiplied by 0.865. Hence, he value 1.30 is quoted, the actual calculated number .ing eit Equation 3.4-9 or 3.4-11) is 1.50. The basis fo etaining this argin, and its use, are discussed in detail Section 3.4.2.1.

1 Histograms of oth the "R" grid (15 x 15 geomet data and the 17 x 17 y etry data satisfy the criterion of be' obtained from a 6rmal distribution just as does the data used t develop the riginal W-3 DNB correlation; however, the means and st ard viations are slightly different from those of the W-3.

l 3.4-15

l BVPS-1-UPDATED FSAR Rev. 0 (1/82) obability distribution curves for the "R" grid data and t 17 x data (including the 0.88 multiplier) are compared at of the -

orrelation in Figure 3.4-7. From this fi it can be seen that e approach which was valid for t original W-3 DNB correlation is servatively applicable the "R" grid and 17 x 17 data as repre ed by Equation .4-7 and 3.4-8 (with the 0.88 multiplier for t 7 x 1 ata). Westinghouse will continue to design to a minimum R of 1.30 for the peak rod or rods in the core. Based on W-3 s 'stics, the proportion of such peak rods that wil at reach DNB is 5 or greater at a 95 percent confidence evel. Since the stati s of the W-3 correlation ar conservatively retained to define t robability of DNB, lines (95% probability-at 95% confidence 1 for the W- are shown on the plots of the "R" grid data, Fi s 3 and 3.4-6 and the 17 x 17 data, Figure 3.4-6A.

3.4.2.3.3 Mixing Technology The rate of heat exchange by mixing between flow channels is pro-portional to the difference in the local mean fluid enthalpy of the respective channels, the local fluid density and flow veloc-ity. The proportionality is expressed by the dimensionless thermal diffusion coefficient, TDC, which is defined as:

TDC = W' (3.4-14)

Pva where:

W' = flow exchange rate per unit Jength, 1bs/ft-sec P = fluid density, lbm/ft 3 V = fluid velocity, ft/sec a = lateral flow area between channels per unit length, fta/ft The application of the TDC in the THINC analysis for determining the overall mixing effect or heat exchange rate is presented in Reference 36.

As a part of an ongoing R&D program Westinghouse has s and directed mixing tests at Columbia University D M.ponsored These series of tests, using the "R" mixing vane grid design on 13, 26 and 32 inch grid spacing, were conducted in pressurized water loops at Reynolds numbers similar to that of a PWR core under the following single and two phase (subcooled boiling) flow condi-tions:

1. Pressure 1500 to 2400 psia

2. Inlet temperature 332 to 642F

3. Mass velocity 1.0 to 3.5 x 10 8 lbm/hr ft 2 5

4. Reynolds number 1.34 to 7.45 x 10 3.4-16

in BVPS-1-UPDATED FSAR Rev. 4 (1/86) resistance in the channel due to the local or bulk L bciling. The effect of the non-uniform power distribution is- inherently considered in the THINC analysis for every operating condition which is evaluated.

4. Flow Mixing:

The subchannel mixing model incorporated in the THINC code- and used in reactor design is based on experimental data ("2) discussed in Section 3.4.3.4.1. The mixing vanes incorporated in the spacer grid ' design induce additional flow mixing between the various flow channels in a fuel assembly as well as between adjacent 1 assemblies. This mixing reduces the enthalpy rise in the hot channel resulting from local power peaking or unfavorable mechanical tolerances.

3.4.2.3.5 Effects of Rod Bow on DNBR The phenomenon of fuel rod bowing, as described in Reference 86 must be accounted for in the DNBR safety analysis of Condition I and Condition II . events for each plant application. Applicable generic credits for margin resulting from retained conservatism in the evaluation of DNBR and/or margin obtained from measured plant operating parameters (such as F(N,AH) or core flow), which are less 1Luiting than those required by the. plant safety analysis, can be used to offset the effect of rod b "*

as dscussed A Sech M/,/

Thefsafety y analysis ,

foz _

BVPS-1 ,

maintained sufficient margin -f 0.{

I'.

v .' *, * " 5._ .' "" ". *. n. s',,. , * **"r.E.

su m."*.J "m N 1. . ._ wuC " _ ._ .%. .m_ M.

. ' .* u' "m ' w . ' '".v . "!.'"*. .'". M. . @, 'mm multiplic; cf 0.005 vs. 0.00, ;nd pit;h ;;ductien, to acc mmodate full and ics flow DNBR penalties identified in Reference 87 (< p rcent for the worst case which occurs at a burnup of ,000 MWD /MTU). 3 The maximum rod bow pen es accounted for in the design safety analysis are based on an asembly average burnup of ,000 MWD /MTU. At

.burnups greater than 000 MWD /MTU, credit is taken for the effect of F(N,aH) burndown due to the decrease in. fissionable isotopes and'the buildup oof fission product inventory, and no additional rod bow penalty is. required.

3.4.2.4 Flux Tilt Considerations Significant quadrant power tilts are not anticipated during normal operation since this phenomenon is caused by some asymmetric perturbation. A dropped or misaligned RCCA could cause changes in hot-channel factors; however, these events.are analyzed separately in Chapter 14. This discussion will be confined to-flux tilts caused by X-Y xenon transients, inlet temperature mismatches, enrichment variations within tolerances and so forth.

The design value of the enthalpy rise hot-channel factor F(N,aH) j which includes-an 8% uncertainty (as discussed in Section 3.3.2.2.7), 1 3.4-19 i

BVPS-1-UPDATED FSAR Rev. 4 (1/86) 8 8

rates used in the design, which were determined in part from the pressure losses calculated by the method described here, are conservative.

3.4.2.8.3 Void Fraction Correlation There are three separate void regions considered in flow boiling in a PWR as illustrated in Figure 3.4-12. They are the wall void region (no bubble detachment) , the subcooled boiling region (bubble detachment) and the bulk boiling region.

In the wall void region, the point where local boiling begins is determined when the clad temperature reaches the amount of super-heat predicted by Thom's("O correlation (discussed in Section 3.4.2.8.1). The void fraction in this region is calculated using Maurer's(50 relationship. The bubble detachment point, where the superheated bubbles break away from the wall, is determined by using Griffith's(s0 re3ationship.

The void fraction in the subcooled boiling region (that is, after the detachment point) is calculated from the Bowring(50 correlation. This correlation predicts the void fraction from detachment point to the bulk boiling region.

The void fraction in the bulk boiling region is predicted by using homogeneous flow theory and assuming no slip. The void fraction in this region is therefore a function pnly of the thermodynamic quality. fu Mdff na/M nan or ega/ // #e Sdd anal is D M R I,M F 3.4.2.9 Ther al E fects of Operational Transients ,

-f Safefy a.talysLs IN'$ WAA DNB core sa ty limits are g erated as a function of coolant temperature, essure, core power and axial power imbalance.

Steady-state ope ation within ese safety limits insures that the minimum DNBR is not less than 1.30. Figure 14D-1 shows the DNDR =

1.31 limit lines and the resulting overtemperature Delta-T trip lines (which become part of the Technical Specifications) plotted as aT

) vs T-average for various pressures. This system provides adequate protection against anticipated operational transients that are slow with respect to fluid transport delays in the primary system. In addition, for fast transients, e.g., uncontrolled rod bank withdrawal at power incident (Section 14.1.2) specific protection functions are provided as described in Section 7.2 and the use of these protection functions are described in Chapter 14 (see Table 14D-2). The thermal response of the fuel rod is discussed in Section 3.4.3.7.

3.4.2.10 Uncertainties in Estimates 3.4.2.10.1 Uncertainties in Fuel and Clad Temperatures As discussed in 3.4.2.2, the fuel temperature is a function of crud, axide, clad, gap, and pellet conductances. Uncertainties in the fuel temperature calculation are essentially of two types: fabrication 3.4-24 L__ _ ___ __ _

BVPS-1-UPDATED FSAR Rev. 4 (1/86) 4 level is also a function of variables such as control rod worth and position, and fuel depletion throughout lifetime. Radial power ,

distributions in various planes of the core are often illustrated for

. general interest, however, the core radial enthalpy rise distribution es determined by the integral of power up each channel is of greater importance for DNB analyses. These radial power distributions, characterized by F(N,aH) (defined in Section 3.3.2.2.2) as well as axial heat flux profiles are discussed in the following two sections.

, 3;4.3.2.1- Nuclear Enthalpy Rise Hot-Channel Factor, F(N,AH)

The Nuclear Enthalpy Rise Hot-Channel Factor, F(N,AH) is given by the equation:

H F(N,aH) = ' hot rod power = Max / q'(xo,yo,z)de (3.4-19)

H average rod power _,

l I /oq (x,y,z)dz N

Where: q'(x,y) is the local power density at a point x,y,z N is the number of fuel rods H is the total core height.

The way in which F(N,aH) is used in the DNB calculation is important. The location of minimum DNBR depends on the axial profile and the value of DNBR depends on the enthalpy rise to that point.

Basically, the maximum value of the rod integral is used to identify

he most- likely rod for minimum DNBR. An axial power profile is obtained which when normalized to the design value of F(N,aH) recreates the axial heat flux along the limiting rod. The surrounding rods- are assumed to have the same axial profile with rod average powers- which are typical of distributions found in hot assemblies. In this manner worst case axial profiles can be combined with worst case radial distributions for reference DNB calculations.

It would be u7ted again that F(N,aH) is an integral and is used as such in the m 'B calculations. Local heat fluxes ara obtained by.

using hot channel and adjacent channel explicit power shapes which take into account variations in horizontal power shapes throughout the core. 'The sensitivity of the THINC-IV analysis to radial power

. shapes is discussed in Reference 43.

For operation at a fraction of full power, P, the design F(N,aH) used is given by:

/d R F(N,aH) = tifdr [ 1 + 0. 3 ( 1-P ) ] (3.4-20)

The permitted relaxation of F(N,aH) is included in the DNB protection setpoints and insertion to the allowslimits insertion radial ("

p)owerthus shapeallowing changes greater with rod flexibility in the nuclear design.

3.4-30 l

l BVPS-1-UPDATED FSAR Rev. 4 (1/86) i As discussed in Section 7.2, factors included in establishing the .

Overtemperature AT and Overpower AT trip setpoints include the reactor coolant temperature in each loop and the axial distribution of core power through the use of the two section excore neutron detectors.

3.4.5.3 Instrumentation to Limit Maximum Power Output The output of the three ranges (sourec, intermediate, and power) of detectors, with the electronics of the nuclear instruments, are used to limit the maximum power output of the reactor within their respective ranges.

There are six radial locations containing a total of eight neutron flux detectors installed around the reactor in the primary shield, two proportional counters for the source range installed on opposite

" flat" portions of the core containing the primary startup sources at an elevation approximately one quarter of the core height. Two compensated ionization chambers for the intermediate range, located in the same instrument wells and detector assemblies as the source range detectors are positioned at an elevation corresponding to one half of the core height; four dual section uncompensated ionization chamber assemblies for the power range installed vertically at the

-four corners of the core and located equidistant from the reactor vessel at all points and, to minimize neutron flux pattern distortions, within one foot of the reactor vessel. Each power range detector provides two signals corresponding to the neutron flux in the upper and in the lower sections of a core quadrant. The three ranges of detectors are used as inputs to monitor neutron flux from a completely shutdown condition to 120 percent of full power with the capability of recording overpower excursions up to 200 percent of full power.

The difference in neutron flux between the upper and lower sections of the power range detectors are used to limit the overtemperature aT and Overpower aT trip setpoints and to provide the operator with an indication of the core power axial offset. In addition, the output of the power range channels are used for:

1. The rod speed control function

2. To alert the operator to an excessive power unbalance between the quadrants

3. Protect the core against rod ejection accidents

4. Protect the core against adverse power distributions resulting from dropped rods.

Details of the neutron detectors and nuclear instrumentation design and the control and trip logic are given in Chapter 7. The limit '

an neutron operation and trip setpoints are given in the Technical Specificat s.

fgy Nh he><s 3.4-44

f BVPS-1-UPDATED FSAR Rev. 4 (1/86)

References for Section 3.4 (Cont'd)

35. L.- S. Tong,- " Critical Heat . Fluxes .on Rod Bundles", in "Two-Phase Flow and Heat Transfer in Rod Bundles", 31-41, p American Society of Mechanical Engineers, New York.(1969).

D .

36. H. Chelemer, J. Weisman and L. S. Tong, "Subchannel Thermal Analysis of Rod Bundle Cores", WCAP-7015, Revision 1, Westinghouse Electric Corporation (January 1969).

37. T. $ ef_ face uhyf ZASEL7'

" ticy,

. :nd F. T.I fe4kvence.

Cadek, "0NE37 T;;t: R :ults-for New Mixing Vene Cride (R) , WCAP-7005-L, We;tinghou;e Electric 0;;perati;; (Es;prict;ry), (July, 1072) and TCAT 7050, S;;tingh;;;; 21;;tri; 0;rporation (O;tcher, 1072).

38. 7.gef./

aceM;ti;y udA(:nd %#T.SEATT. C d. / lk,efence "0NS Tc:t31(2;;;1t: for "R" Crid

}

-Thimble Ccid ';.';11 011&", WCAP-7005-L, Addendem I, )

Westinghouse Electric Corporation (Freprietary), (October, j

~ 1972), and WCAF-7050 Addendum I, 'c.';;tingh u e Electric :

I

- Corporatica (Oct:b r, 1972).

39. L. S. Tong, " Prediction of Departure from Nucleate Boiling for an Axially Fon-Uniform Heat Flux Distribution", Journal of Nuclear Energy, 21, 241-248 (1967).

40. L. S. . Tong, " Boiling Crisis and Critical Heat Flux", AEC. i critical Review Series, TID-25887, U.S. Atomic Energy ~

Commission (1972).

41. F. F. Cadek, F. E. Motley and D. P. Dominicis, "Effect of Axial Spacing on Interchannel Thermal Mixing with the "R" Mixing. Vane Grid", WCAP-7941-L, (June, 1972), Westinghouse Electric Corporation (Proprietary), and WCAP-7959, Westinghouse Electric Corporation (October, 1972).

4:2. F. F. Cadek, "Interchannel Thermal Mixing with Mixing Vane-Grids", WCAP-7667-L,- (May, 1971), Westinghouse Electric Corporation -(Proprietary), and WCAP-7755, Westinghouse Electric Corporation (September, 1971).

43. L. E. Hochreiter, " Application of the THINC-IV Program to PWR- Design", WCAP-8054, (October, 1973), Westinghouse Electric Corporation (Proprietary), and WCAP-8195, Westinghouse Electric Corporation (October, 1973).

44. F. W. Dittus and L. M. K. Boelter, " Heat Transfer in Automobile Radiators of the Tubular Type", California {

Univiversity Publication in Enc., 2, No. 13, 443-461 T'1930).

l

45. . 3. Weisman, " Heat Transfer to Water Flowing Parcilel to Tube 1' Bundle s "., , Nuclear Science and Engineering, Vol.6, 78-79 (1959).

3.4-49 .

l

N BVPS-1-UPDATED FSAR R3v. 4 (1/86)

, References for Section 3.4 (Con't)

78. D. S. Rowe, C. W. Angle, " Crossflow Mixing Between Parallel Flow Channels During Boiling, Part III Effect of Spacers on Mixing Between Two . Channels," BNWL-371, part 3, Battell )

Pacific Northwest Laboratory (January, 1969).

(

79. J. M '. Gonzalez-Santalo and P. Griffith, "Two-Phase Flow Mixing . in - Rod Bundle Subchannels", American Society. of. i Mechanical Engineers Paper 72-WA/NE-19.

80. J. . M. 'Hellman (Ed.), " Fuel Densification Experimental Results and Model For. Reactor Application", WCAP-8219, Westinghouse Electric Corporation (October, 1973).

81. K. W. Hill, F. E. Motley and F. F. Cadek, "Effect of Local' Heat Flux Spikes on DNB in Non-Uniform Heated. Rod Bundles",.

WCAP-8174, (August, 1973), Westinghouse Electric Corporation st, 1973).

r(0 Proprietary)k,andWCAP-8202,h(Au

82. . ././a Cl C 4)Ll, $MT /

Y . 'f . Mill , F . E . M tley , .}. . H . '.1: 2::1, CE $

"Effect

- tf 17 ;; 17 ru;l A;;r.hly Cec .;tri su OC" , UCA;-0207, "Westingheu;; Electric C rper; tic. W..:;h, 1974).

83. S. Nakazato, E. E. DeMario, " Hydraulic Flow Test of the 17 x 17 Fuel Assembly", WCAP-8297, (February, .1074 ) .

84. F. E. Motley, A. H. Wenzel, F. F. Cadek, "The Effect of 17 x 17-. Fuel Assembly Geometry on Interchannel' Thermal Mixing",

WCAP-8299, Westinghouse Electric Corporation (March, ' 374) .

85. T. M. Burke, C. E. Meyer, and J. Shefcheck, " Analysis of Data from .the Zj WCAP-8453, (December, m%,4 07 (Proprietary) Unit 1) and THINC Verification Westinghouse WCAP-8454, Test",

Electric C tion (December, 1974).

86. Skaritka, J.~ (Ed.) 1979. Fuel Rod Bow Evaluation. WCAP-8691, Revision 1 (Proprietary) and WCAP-8692, Revision 1 (Non-proprietary).

87. Westinghouse 1981 letter. E.P. Rahe, Jr. (Westinghouse) to J.

R. Miller (USNRC), NS-EPR-2515, dated October 9, 1981, entitled: Partial Response to Request Number 1 for Additional Information on WCAP-8691, Revision 1, and letter, E.P. . Rahe, Jr. (Westinghouse) to J.R. Miller (USNRC), NS-EPR-2572, dated March 16, 182, entitled: Remaining Response to Request Number l' for Additional Information on WCAP-8691, Revision 1.

N 'AtY .D $bt f } fe k sce W n /W2Mri 6%ce e7 3.4-53 4

..l l

l INSERT 1

37. Motley, F., E., Hill, K. W., Cadek, F. F. and Shefchek, J.,

.I "New Westinghouse Correlation WRB-1 for Predicting Critical Heat Flux in Rod Bundles with Mixing Vane Grids,"

WCAP-8762-P-A, July 1984.

38. S. Ray, " MINI Revised Thermal Design Procedure (MINI RTDP)," I WCAP-12178-P, March 1989. l l

82. Friedland, A. J. and Ray, S., " Revised Thermal Design Procedure,"

WCAP-11397 (Proprietary), February 1987 and Letter, A. C. Thadani l (USNRC) to W. J. Johnson (Westinghouse), " Acceptance for l Referencing of Licensing Topical Report WCAP-11397, Revised Thermal Design Procedure," January 1989.

88. Letter from A. C. Thadani (USNRC) to W. J. Johnson (Westinghouse), -

" Acceptance for Referencing of Licensing Topical Report WCAP-9226-P/9227-NP, Reactor Core Response to Excessive Secondary Steam Releaser," January 31, 1989.

89. Owen, D. B., " Factors for One-Sided Tolerance Limits and for Variable Sampling Plans," SCR-607, March 1963.

5 l

vyp4

, EVPS-1-UPDATED FSAR Rev. 0 (1/82) l TABLE 3.1-1 grggs REACTOR DESIGN f_ _y?"!OO!!

._ T?_"L" DesofNaIue

, , _ _ , , _ , _ _ _ _ _ _ .. m ,. ,- i

' ". ?. .'._.~.~ 0_". a'." ~,. ' . 7 7. 5.._,~.m _...

1 Thermal and Hydraulic 0;;;ificetica O . ifi;;ti;..

Design Parameters Off;;t; Eff ;t:

Reactor Core Heat Output, 2,652 0,552 MWt g Reactor Core Heat Output, 9,051 x 10' 9,051 ; 10' Btu /hr g Heat Gencrated in Fuel, % 97.4 07.4 g System Pressure, Nominal 2,250 2,250 _

psia

/ System Pressure, Min.

Steady State, psia 2,220 2,0:0 3

g Minimum DNBR for Design W* I 1.00-

Transients

/(,' ., ~ , Coolant Flow . :

1)es,g % !

g. Total ThermalAFlow Rate, 100.9 x 10 8 100.0 x 10' i 44.

% Effective Flow Rate for Heat Transfer, lb/hr M x 10' 0 0. 3 x 101-i

/ Effective Flow Area for Heat Transfer, ft 8 41.5 (51D)

41. y (v 85H) 41.0 ~

[ Average Velocity Along _

Fuel Rods ft/sec -Mr 315 32'1 gg,3 (y,gg (STP)) 14.4 g Average Mass Velocity, :PMRL x 10' STD) 2. 31 = 10 5

lb/hr-ft* g.g x go 6 ((yg.g)

Coolant Temperature, 'F M Nominal Inlet 542.5 540.5 M Average Rise in Vessel 67.4 57.4 K Average Rise in Core 70.2 i

s

% Average in Core M 0.2 377, g Average in vessel 576.2 070.2 1 of 5

~rapy\

s BVPS-1-UPDATED FSAR Rev. 0 (1/82)

.a TABLE 3.1-1 (CONT'D) f REACTOR EESIGN _f_}KgW!00 :1EfGLS

_ _ . . . T0;.::

Deb <n Va lk e-0

'. 17 . :f;nr.2 := 1 "1;;t Sith 10:10 Witheet.

Thermal and Hydraulic 0;;;ificati;; ::.;ifi;;ti;n Design Parameters Off;;;; ::ffect; Heat Transfer

% Active Heat Transfer, Surface Area, ft" 48,600 40,i!0 l M Average Heat Flux, 181,400 007,000 Btu /hr-ft 8 455,4co

% Maximum Heat Flux for Normal Operation, M l1I 000,000_

Btu /hr-ft*

Average Thermal output, 5.2 0.7 kW/ft -

l'A.5 Maximum Thermal Output 1b4 f11 10.0

('-

~ g.

' ""/'

Anu' ' ' " ' " " " '

Peak linear power ~for 18.O I21 01.1 determination of protection set points, kW/ft *

^

9.4a

g. Heat Flux Hot Channel < 3=42. 0.70 Mi Factor F(Q)

Fuel Central Temperature, 'F 3,580 g Peak at 100t Power ):p0&tL 4,000

% Peak at Mavinum Thermal Output for Nazimum 4,15.0 4,000 overpower Trip Point Core Mechanical Design Parameters Fuel Assemblies

% Design RCC Canless 000 0;;1;;;

1

[ Number of Fuel Assemblies 157 -159-

[ UO Rods per Assembly 264 004

g. Rod Pitch, inches

. 0.496 0. 00 ';

2 of 5

i BVPS-1-UPDATED FSAR Rev. 0 (1/82)

TABLE 3.1-1 (CONT'D) i

&G ""E REACTOR DESIGN k f1&T6AS

_.Q.?" ISO" T?_ ELE //

I

,,..,, n c _ ___ no n e. _ ,

{ott dec h cg[ blbnt Eth$ 15x[bbi5hout

...crm:1 and Hydraulic- Densificatier Dcncificction Design Parameters Effects Effcct:

Fuel Assemblies K Overall Dimensions, 8.426 x 8.426 0.425 x S.425 inches E Fuel Weight (as UO 181,205 175,200 pounds (C c,/e /) 2) ,

7 J'I. Z(ircaloy lbs. 38,230 25,200 Cy ele dWeight, g_ g n y . ,y y

.Pf. Number of Grids per 64y- Type R/YV 7 Assembly Ac,. eel nf -57D

&loy - VSH X. Loading Technique 3 Region 2 ncgicn Non-Uniform Ncn "nifer.

Fuel Rods i Jfr. Number 41,448 22,02S J6'. Outside Diameter, inches 0.374 0.422 ff.

Diametral Gap,[inche,guncmifr/

Regicnc 1, 2 pelh/5) 0.0065 0.0055 Regicn 3 0.0005 0.00S5

)6'. Clad Thickness, inches 0.0225 0.024 1 M. Clad Material Zircaloy-4 - 'irccicy '

Fuel Pellets JM Material UO2 -907 Sintered Cintcred

g. Density (% of 95 Theoretical) f[. Diameter, ncgicnc 1,inches, 2 64*cM/et/g//eS) 0.3225 0.3550-negion 0.222: 0.;04 inches 0.530 0.000

[. Length,D

<'C ch y

3 of 5

BVPS-1-UPDATED FSAR Rev. 0 (1/82)

TABLE 3.1-1 (CONT'D)

V8 !"E-REACTOR DESIGN f@f"i &T&A$ -' "' " 9 4517 V

'.".^.'

17xl? .9eference E"PS-1 Core /1eckaa,lCa/ Plent '3ith 15:15 Mithcut Thermcl and Hydraulic Ocncificcticn Ocncificctica Design Parameters Effect: Effect:

Rod Cluster Control Assemblies

[ Neutron Absorber Ag-In-Cd p._ Cladding Material Type 304SS Tipc 204SS Cold Worked Cold !!crkcd

[ Clad Thickness, inches 0.0185 0.010 g Number of Clusters, Full 48 Length fff. Number of Absorber Rods per Cluster 24 -

,4r9 . Core Barrel, I.D./O.D., 133.85/137.875 133.05/137.875 inches ,

1 M Thermal Shield, I.D./O.D.,

inches 143.63/148.000 143.53/14S.000 Nae l eat 1ksku h amelds Structure Characteristics Core Diameter, inches 119.7 119.7 (Equivalent)

/NO

[ Core Average Active Fuel 143.' 144

Height, inches l

Reflector Thickness and Composition Top - Water plus Steel, N10 ^10 -

inches

[ Bottom - Water plus Steel, inches N10 ^10

[ Side - Water plus Steel, N15 %15 inches

[ H20/U, Cold Molecular Ratio (Lattice) 3.43 Volume 2.41 Molecular 3.55 Vol'me-4 of 5

_ _ _ _ _ _ _ _ _ _ _ _ ________ _____ _ m

BVPS-1-UPDATED FSAR Rev. 0 (1/82)

TABLE 3.1-1 (CONT'D) I

(}( Dest 9

l REACTOR DESIGN }^ Q'ft[TGt26

" " ^ " " ' ' " ' "

U 17 17 ".-ference- B"DS-1 ,

Th;= h. c.h_/"yd::ulic Plet wie -15r.15 ith ut j hn-ification D:::ific: Men- l Design Parameters Effect: -Effect; { ;

Feed Enrichment, w/o (Cycl 4 I)

[ Region 1 2.10 2.00 JHf. Region 2 2.60 2.70 JH!I. Region 3 3.10 2.25 fII 2 10 (2TThis limit is associated with the value of F(Q)=M.

S ee Section 3.3.2.2.6.

I The 15x15 de;ign velue Of PMF=2.72 ::: spicyed for-nuclear-

.-d thermal-hyd: uli: desig.. He":ver, to meet the Final

..: : : p t,_ r, .:,, : Criteri for : :rgency core eccling, the aller:d -

__ - , m, e gM e se aeme.

(2) dasec/ o, 4.r % cm lypass F/oe 5 of 5

d 2 ne 3 oc in 1.123 te cr eo 3 3.323 1.333...

2 3 33 2

3 2 3 3

3 Sfe n

33 4B 2.333... 3 33 3 3 3 1 3 3 33 3 3 3

. l ae rr ue y

,tct h r e Euo d a >

d Cr e r

r E{

o C

Rtd b i

P Y C r

Osn F

,t a'

l e

u l

B Tp, M e ne f / Rg t eed l

) u dmo F E .

2 p oec ee D D d.

8 m cl sd uo N E

N b c

/% M

/ I A o es om C I s 1 C n i h d A A

(

0 wes oty dil gn ng ii i f

e D R .

Af

_ R E

E r

wna ts i Pw M L

T 0 v oin se do EOa M R e lFa B

ed W M La A U 3 R H T N

G I

S E

D E

R R O A C S 2 h -

F - N ltlts f

D" 1 I l n D e aieaa mwueg ee e o E 3 8 u r fh tr s i T E q c ed n no r s A E U i i hof ,o t ac oh u 2 D L Q n m trosl ft f P B I h a ea a sd i f f U A N c n llngs.

t a oz ne sw t

i D

o

- T !

i e y aeonic 1

S P

V C

E T

C T D dng an cuiaft ifth ac,e r

ifr poeyee r,

i d

c ce p po ci ig n

ns ad to srg o n n r p

u 1

N 4

B I i m dtfs o ome oli C T cl eliisa c con cod s sho y

Y ie - essne rl 2r i L o o i pte o A td ao idnnal r rrg uni ,e p mooere 2%e N tM emcdtr c coe oohrcs i afr Dh _

A S S MM T 3 G 2 0

0 se s

cr t_

l a

n r

s i u ts ss i e rr p

u o

,lax ec y

- s tg n

h ,%% 42 y t

f u

h#

e ep r ni , '

t t ote,e n cl G I d aa d tronc C s

e a r

o e n rn ar u

Qtnce).

n a

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BVPS-1-UPDATED FSAR Rev. 0 (1/82)

.( TABLE 3.3-1 REACTOR CORE DESCRIPTION-

, Active Core Equivalent Diameter, inches 119.7 Core Avg. Active Fuel'Hgt., O?d 7 14.o Fir:t 2::, inch Height-to-Diameter Ratio 1.20 Total Cross-Section Area, ft* 78.14 Reflector Thickness and Composition Top - Water plus Steel, inches 10 Bottom -' Water plus Steel, inches 10 Side - Water plus Steel, inches 15 Fuel Assemblies Number 157 Rod Array 17 x 17 Rods per Assembly 264 Rod Pitch, inches 0.496 overall Transverse Dimensions, 8.426 x 8.426 inches Fuel Weight (as UOz), lbs [Gyc/e h 181,205 zircaloy weight, lbs (Cyt/c I) 38,230 Number of Grids per Assembly FR type- % - trvn mais k v b c-Composition of Grids 3NGMS- 7

/*C0"el 7/8 - Sfa./,mf Weight of Grids (E factive in 1,885 7 Y - fo

4je s-#

Core) lbs (Cyc/t I-Number of Guide Thimbles per 24 Assembly Composition of Guide Thimbles Zircaloy 4

(

Diameter of Guide Thimbles (upper 0.450 I.D. x part), inches 0.482 0.D. ( 57D)

O , '/ 4 k I.D,y 1 of 3 0, y ?V 0,b, (4J,ge pg) i;

f i i

l BVPS-1-UPDATED FSAR Rev. 0 (1/8'2) l TABLE 3.3-1 (CONT'D)

REACTOR CORE, DESCRIPTION Fuel Assemblies (Cont'd)

Diameter of Guide Thimbles (lower 0.397 I.D. x part), inches 0.429 0.D. 1 Diameter of Instrument Guide 0.450 I.D. x Thimbles, inches

0. 482 0.D. (STD)

Fuel Rods 08#) I'D' X 0A 74 0.D. Ola~fa,)a [b Number 41,448 Outside Diameter, inches 0.374 Diameter Gap, inches (b cod /'d //e ) 0.0065 Clad Thickness, inches 0.0225 Clad Material Zircaloy-4 Fuel Pellets Material U02 Sintered Density (percent of Theoretical) 95 FuelEnrichmentsw/o[Cyc/el) 4 Region 1 2.10 Region 2 2.60 Pagion 3 3.10 Diameter, inches Regions 1,2,3 0.3225 Length, inches (C y c,lt /) 0.530 Mass of UO r Foot of Fuel Rod, 0.364 lt/ft (Cyc O Rod cluster Control Assemblies Neutron Absorber Ag-In-Cd Composition 80%, 15%, 56 2 of 3

BVPS-1-UPDATED FSAR Rev. 0 (1/82)

TABLE 3.3-1 (CONT'D)

REACTOR CORE DESCRIPTION Rod Cluster Control Assemblies (Cont'd)

Diameter, inches 0.341 Density, lbs/ inches: 0.367 Cladding Material Type 304, Cold-Worked.

Stainless Steel Clad Thickness, inches 0.0185 Number of Full Length Clusters. 48-Number of Absorber Rods per 24 Cluster Full Length Assembly Weight 147 (dry), ib.

C BurnableA.h:4the..e;;. Rods (. .et s

.yc/eCere)

I)

Number - T-iti:1 C^ 1,072 Material Borosilicate Glass Outside Diameter, inches 0.381 Inner Tube, O.D., inches 0.1815 Clad Material Stainless Steel Inner Tube Material Stainless Steel Boron Loading (w/o B Os in glass rod) 12.5 Weight,of B 8 ' per foot of .000419 rod, ib/ft Initial Reactivity Worth, top 7.0 (hot), 5.5 (cold)

Excess Reactivity (C ye /f /)

Naw4=nm Fuel Assembly k <1.6 (Cold, clean, UnborateI Water) t Maximum Core Reactivity (Cold, Eero 1.25 4.

Power, Beginning of Cycle 1) 3 of 3

BVPS-1-UPDATED FSAR Rav. 0 (1/82)

TABLE 3,3-2 NUCLEAR DESIGN PARAMETERS Core Average Linear Power, kW/ft, Including Densification Effects 5.20

< oL W Total Heat Flux Hot-Channel Factor <2.32 Nuclear Enthal?y Rise Hot-Channel / (,7 Factor, F (N, A d) - 1. 5 5 --

Reactivity Coefficients Doppler Coefficient See Figures 3.3-27 and 3.3-28 Moderator Temperature Coefficient at See Figures Operating Conditions, pcm/*F(21 Boron Coefficient in Primary Coolant, 3. 3-305E57-31,-32 m.-33, pcm/ppmf21 -16 to -8 Rodded Moderator Density Coefficient <+0.43 x los at Operating Conditions, pcm/gm/cc ~

Delayed Neutron Fraction and Lifetime ( c/t M S BOL, (EOL) 0.0075 (0.0044) cffBOL, I (EOL), y see 19.4 (18.1)

Control Rod Worths Rod Requirements See Table 3.3-3 Maximum Bank Worth, pcm <2,300 Maximum Ejected Rod Worth See Technical Specifications 1

(11N ote: 1 pcm = (percent mille) 10 sop where op is calculated from two statepoint values of k,ggy Ln (km /k t ) .

)

l i

1 of 2 E__ _ _ _ _ _ _ - - - _ _ - - - - - - - _ . -_

BVPS-1-UPDATED FSAR Rev. 0 (1/82)

TABLE 3.3-2 (CONT'D)

NUCLEAR DESIGN PARAMETERS C

- Boron concentrations (Three Loop Operation, ii.Cle lQ:t ?; 10)

Refueling k 0.95, Cold, Rod Cluster 2000 C8$$ro<lAssembliesIn 1196 Zero Power k 0.99, Cold, Rod Cluster C8dkr=olAssembliesOut 1429 Zero Power !

k 0.99, Hot, Rod Cluster C8$kr=lAssembliesOut o 1419 Full Power, No Xenon, Beginning of Cycle k 1.0, Hot Rod ClusterCo8bko=lAssembliesOut 1195 Full Power, Equilibrium Xenon, k 1.0, Hot, Rod Cluster C8dkr=lAssembliesout o 906 ReductionWithFuelBurnug)

First Cycle, pps/GWD/MTU See Figure 4.3-3 Reload Cycle, ppm /GWD/MTU N100 f

(2) Gigawatt Day (GWD) = 1,000 Megawatt Day (1,000 MWD). During the first cycle, fixed burnable poison rods were present which which significantly reduce the boron depletion rate compared to reload cycles. The values quoted are representative of averages only.

I-i 9

O 2 of 2

t.

BVPS-1-UPDATED FSAR Rev. 0 (1/82) s T I TABLE 3.4-1

_ REACTOR DESIGNVf0Wfkb C.."?a!SOM T?.EI,L Thermal and 17 x 17 With 15 x 15 Without Hydraulic Design Parameters Densification DC. ifi:: tion Reactor Core Heat Output,, MWt 2,652 2,052 Reactor core Heat Output, Btu /hr 9,051 x 10' 0,001 x 10' Heat Generated in Fuel, % 97.4 97.4 System Pressure, Nominal psia 2,250 2,250 i System Pressure, Minimum Steady State, psia 2,220 2,220 Minimum DNBR at Nominal Conditions Typical Flow Channel 2.02 Thimble (Cold Wall) Flow g,$p Channel 1.70 Minimum DNBR for Design 1,33 Transients >

                                                                                                                             >1.20 DNB Correlation                                            '

w ' " 1, " "c' 2 with

                                                                                           \AJRS-1g".1 ~ e'           o     m:difi d :p::cr f        )           f::ter)

Coo', ant Flow , PcGign Total Thermal / Flow Rate, lb/hr 100.9 x 10' 100.0 x 10 -- 5 Effective Flow Rate for Heat Transfer, lb/hr N3) 6 i M x 10' 95.2 x 10' l Effective Flow2 Area for Heat Transfer, ft 41.5 (STP) 41.0 A na (v-ss) I l 1 of 3 ,

BVPS-1-UPDATED FSAR Rev. 0 - (1/82) TABLE 3.4-1 (CONT'D) REACTOR DESIGN fM A11&TfdS w0eJARISOi; TAO;,0 Thermal and 17 x 17 With 15 ; 15 "i2.;;;. Hydraulic Design Parameters Densification O...;ifi;;tien Average Velocity Along Fuel 13."5 (3TD) Rods, ft/sec M j3,3(v g}}) 10.4 Average Mass Velocity, 4 17 2 lb/hr-ft

                                                                        -h9t x 12 a.s.c.       wos (v-sa)

(frD) 2.01 x 10' Coolant Temperature Nominal Inlet, F 542.5 04 2. "; Average Rise in Vessel, F 67.4 07.4 Average Rise in Core, F g* 70.0 Average in Core, F M " 57^.0 Average in Vessel, F 576.2 570.2 Heat Transfer Active Heat Transfer, Surface Area, ft 2 NMO 4;,3;; Average Heat Flux, Btu /hr-ft 181,400 207,000 Maximum Heat Flux, for normal operation 4MAdo l Btu /hr-ft* M (2) .;;;,;;; Average Thermal output, kw/ft 5.2 -ib4-Maximum Thermal Output, for normal operation, kw/ft (21 10.0 Maximum Thermal output at Maximum overpower Trip Point, kw/ft 18.0("I 21.0

                                                                              .                                    l i                                                             .                                             1

(

                                          .                        2 of 3 1
          - _ - - _ - _ _ _ _ _ _ _                                                                              .A

BVPS-1-UPDATED FSAR Rsv. 0 (1/82) TABLE 3.4-1 (CONT ' D) REACTOR DESIGN _f"t4..METERSf ." .". ::00N TALL : Thermal and 17 x 17 With 15 x 15 "id. ut-- Hydraulic Design Parameters Densification 0-. ific: tion Fuel Central Temperature (DOL) Peak at 100% Power, F t,250 Peak at Maximum Thermal Output for Maximum Over-power Trip Point, F 4,150 -t,500 Pressure Drop Across core, psi f3i 10.1 1 3.S /2i Ja.7 .t 4.I Across Vessel, including nozzle psi W ~ 4c,1 1 6.0 35.4 1 5.5 (1i Based on 65% com.

2. .. : bypqs4/ca

w. f1:w ::.t; :: di:cu:::d in-0:: tion 3.4.2.7.1.

rat a.4c This limit is associated with the value of F(Q) =M (s) Based on best estimate reactor flow rate as discussed in Section 4.1. ("I sae Section 3.3.2.2.6. 4 e 1 3 of 3

LVr3-1-UPDATED FSAR Rev. 0 (1/82) TABLE 3.4-3 VOID FRACTIONS AT NOMINAL REACTOR CONDITIONS WITH DESIGN HOT-CHANNEL FACTORS Average Maximum Core Queet 0,33*[ -- not S,,3&m = 4.ox - t +.4 */. 1 I L l l . I I ( l of 1

i.

                                    .                                                                                                                                                                                                              i C                                                               159.975

M 3.47 2.383 -> +

c - 152.200 = 5 8.426 b O ! k i 5k lfl t-t g g : - ll a .- ll

.~

lhl .- ll z - i: r=1 ac d W ----< ---< _ _; '= J 7 : : : : T = 1.522 + +~ 153 60 133 10 ft2 55 92 00 71 45 50 90 30 35 5 B35 17XI7 VANTAGE 5-H FUEL ASSEMBLY FIGURE 3.2-2a BEAVER VALLEY UNIT I 17XI7 VANTAGE 5-H FUEL ASSEMBLY

DIM 17X17 V5-H A 152.200 B 7.350 C 144.00 DIA D .329 DIA E .374 DIMENSIONS ARE IN INCHES BOTTOM END PLUG SHOWS INTERNAL GRIP TYPE

      -                                                            FOR V5-H FUEL RODS.

TOP DIA D - BOTTOM U

YY -. bYYY - - ," C' Myl V. - n 01RWR v

                   =   DIM B                = =                  DIM C              =

PLENUM FUEL STACK LENGTH 1 DIM A =

FUEL ROD LENGTH 17XI7 VANTAGE 5-H FUEL ROD

FIGURE 3.2-3a BEAVER VALLE.Y UNIT I 17Xt7 VANTAGE S-H FUEL ROD ASSEMBLY [ l

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tl! a 5 M 4 3 1 1.6 - l l 1.4 - 1 1.2 - 1.0 2 4 6 8 10 12 l TTOM yop CORE HEIGHT (FT.) FIGURE 3.3-21 MAXIMUM FO X POWER VS AXIAL HEIGHT DURING NORMAL OPERATION SEAVER VALLEY POWER STATION UNIT NO.1

UPDATED FINAL SAFETY ANALYS18 REPORT l -

WESTN3 MOUSE PROPRIETARY CLASS 2 2.8 l (0, 2.40) l l (6, 2.40) 2.4

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l (10.8. 2.g i 2.0 \ E W

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POWER VS. AXtAL HEIGHT DURING NORMAL OPERATION O

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REV. 0 (1/82) - l 26PL&C6 h)11H Nfn) 1

                                                   \                          ,

C l 4U2E 3.3-3'7 NrTEseo) I (DELETED) , FIGURE 3 3-37 ACCIDENTAL SIMULTANEOUS WITHDRAWAL OF TWO CONTROL BAEKS EOL, NZP BANKS A AND 8 MOVING IN THE SAME PLANE, PL AT 140 STEPS

BE AVER VALLEY POWER STATION UNIT NO.1 UPDATED FINAL $AFETY ANALYS!5 REPORT l.

70.0 60.0 - E

                                                                   $  S0.0  -

e a S I 40.0 - 8 m 8 Q: 300 - Y p b

                                                                   $  20.0  -

t' s 10.0 - 0.0 ' ' ' ' O SO 100 150 200 250 STEPS WITHOR AWN FIGURE 3.3-37 TYPICAL ACCIDENTAL SIMULTANEOUS WITHDRAWAL OF TWO CONTROL BANKS AT BOL, HZP, BANKS "A" AND "B" MOVING IN THE SAME PLANE BEAVER VALLEY POWER STATION-UNIT i FINAL SAFETY ANALYSIS REPORT

REV. 0 (1/82) 4500

                            .                                        4000         -

3500 - C w E 3000 k W 5 g 2500 - 5 v J

                               $                                                                                                                           i i

2000 - i 1500 - 1000 - Sooo i I I I I I I I I I 0 2 4 6 8 10 12 14 16 18 20 22 LINEARPOWER(KW/FT) FIGURE 7 4-2 PEAK FUEL CENTERLINE TEMPERATURE SURING FUEL 200 LIFETIME VS LINEAR POWER,w r ~ -..-. ? " - DE AVER V ALLE T POWEN $1Allph UNIT NO. I UPDATED FINAL $AFETY ANALY$l$ REPCRT l l

1 REV. 0 (1/82) { k \ lY A6ues s.44 (ern m)

          .4                                                -

LENGTN SPACING FLUX f 8 20" u SIN u h l.2 N' 20" 26" 32' cos e u Stu u O u SIN u O 14' 26" u SIN u 8 32" u SIN u V Mass VEL l'.Y 1.57-3.72 X 106 (gfun.ry2 O INLET TEM RATURE 440-6200F l.0 PRES 5URE 14 -2440 P$la o a CxF QUALITY- . 5 - + 15%

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D DATA FOR TYPICAL BE LE Y P OW E R T10N UNIf NO. !

                                          ,                             D FINAL 3AFETY ANAL            REPORT

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                                                              .o LESEN0; Q 8 FT. 0.3M" DI A. - TYPICAL
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RE 3 81-6A COM ALL 17 B DATA BEAVER VALLE STA!!0h UMIT NO. 1 l U y O NAL $AFE . ' S REPORT t

REV. 0 (1/82) D QGLtT&}

                                                     .99 99.90                                                                                                                ,

99.00 d

                                               $ 99.00     -           7 s 17 pas DATA l 98.00     -

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                                                   .0.00   -

30.00 - 20.00

10. ;

l.2 1.1 1.0 0.9 0.8 0.7 0.6 0.5 0. 1 MEASURED / PREDICTED NEAT FLUX RATIC (1/Mlli. D1184) I DELGt t l 34-7 " PROSA CURVES 3 AND R GRID DNS C0 Oss BEAVER E V P OW E N UNIT NO. 1 ED FINAL $AFETY ANALf31$ t

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g BVPS-1-UPDATED FSAR Rev. 0 (1/82) 14.1 CORE AND COOLANT BOUNDARY PROTECTION ANALYSIS 14.1.1 Uncontrolled Rod Cluster Control Assembly Bank Withdrawal from a Subcritical Condition 14.1.1.1 Identification of Causes and Accident Description i 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 reactor control or control rod drive systems'. This could occur with the reactor either suberitical, hot zero power or at power. The "at power" case is discussed in Section 14.1.2. Although the reactor is normally brought to power from a subcritical condition by means of rod cluster control assembly withdrawal, initial startup procedures with a clean core call for boron dilution. The maximum rate of reactivity increase in the case of boron dilution is less than that assumed in this analysis (Section 14.1.4', Uncontrolled Boron Dilution.) The rod cluster control assembly drive mechanisms are wired into preselected bank configurations which are not altered during reactor life. These circuits prevent the assemblies from being withdrawn in other than their respective banks. Power supplied to the banks is controlled such that no more than two banks can be withdrawn at-the same. time. The rod cluster control assembly drive mechanisms are of the magnetic latch type and coil actuation is sequenced to provide-variable speed travel. The maximum reactivity insertion rate analyzed in the detailed plant analysis is that occurring with the simultaneous withdrawal of the combination of the two control banks having the maximum combined worth at maximum speed. GDThe neutron flux response to. a continuous reactivity insertion is d characterized ~ by a very fast rise terminnted by the reactivity kY feedback effect of the negative Doppler coefficient. This self limitation of the power burst is of primary importance since it limits the power to a tolerable level during the delay time for protection. action. Should a continuous rod cluster control assembly withdrawal accident occur, the transient will be terminated by the i

following automatic features of the Reactor Protection System: '

1. Source Range High Neutron Flux Reacte: Trip - actuated when either of two independent source range channels indicates a neutron flux level above a preselected manually adjustable setpoint. This trip function may be manually bypassed when either intermediate range flux ~ channel indicates a flux level above a specified level. It is automatically ,

reinstated when both intermediate range channels indicate a flux level below a specified level. 14.1-1

y- { i 5 , BVPS-1-UPDATED FSAR Rav. 6.(1/88)

2. Intermediate Rance Hich Neutron Flux Reactor Trip. - ,

actuated when either of two independent intermediate range ! channels. indicates a flux level'aboveLa preselected manual'.y adjustable setpoint. This trip function.may be manuaJiy - bypassed when two of- the four power range channels are. reading above. approximately 10 percent of full power and'is j automatically reinstated when three of the:four channels J indicate a power below this value. E 3. Power Rance High Neutron Flux Reactor Trip (Low settino). - actuated when two out of.the four power range . channels indicate a power level above approximately 25 'l percent 'of full power. This trip function may be manually bypassed 'when two of the four power' range channels indicate a power level above approximately 10 percent of full power and is automatically- reinstated when three of -the four channels indicate a power level below this value.

4. Power Range High Neutron Flux Reactor Trip (Mich Settina) -

actuated when two out of the four power range channels indicate' a power level above a preset setpoint. This trip function is always active. tev @ - mar In- addition, control rod stops on high l'.termediate range flux level (one of two) and high power range flux level (one out of four) serve to discontinue rod withdrawal and prevent the need to actuate the intermediate range flux level trip and the power range flux level trip, respectively. m qg 14.1.1.2: Analysis of Effects and Consequences, Method of-Analysis D - transient is analyzed by two digital computer codes. F e UR three operation c a s e's- the WIT-6(O code i ed to . h calculate t eactivity transient and hence, t uclear r:x,. transient. This e includes the simulation x delayed neutron k groups and the core al and hydrau eedback equatisns. For

     -j   the case with one -loop                     ted       ore nuclear calculations are l

g performed using the spatial on kinetics code TWINKLED O o to determine the power ration time including the various f 4 total core feedback ects i.e., Doppler tivity and moderator g reactivity. ACTRAN C O , code is then use calculate the W thermal h lux transient based on the nuclear transient MI calcu by the WIT-6 and TWINKLE codes. FACTRAN also ca ates uel, clad and coolant temperatures. , GN In order to give conservative results for a startup accident, the following assumptions are made concerning the initial reactor conditions:

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 14.1-2 l

BVPS-1-UPDATED FSAR R2v. 6 (1/88) coefficient conservative values (low absolute magnitude) as a function of power are used, 'CO: 00:ti:n 140.0f fer the uu. w. lwwg m...)..wwu...Loui ..e 4 L.wu .e med fes iho L e-

                                                                                                                   -h      eye.etien case in the nucleos wwd ) .DmuEzTI l

2. Contr56ut' ion of the moderator reactivity coefficient is f 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 moder to reactivity i coefficient. 7:: th: th::: 1 ;p :::: conservative i value, gi"en ir FSAF Table 110-2 is used i the analysis to ;

yield the maximum peak heat flux. Fe. the two icep-ca-se-e-highly conservative value ia obtained by adjueting the

                                                                                                                   ^ initial b;;;n ::n;;ntr:ti:n in th: nuc1::: ::f:.                               ,

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3. The reactor is assumed to be at hot zero power. This assumption is more conservative than that of a lower initial system temperature. The higher initial system temperature yields a larger fuel-water heat transfer coefficient, larger specific heats, and a less negative (smaller absolute !

magnitude) Doppler coefficient, all of which tend to reduce i the Doppler feedback effect, thereby increasing the neutron flux peak. The initial effective multiplication factor is assumed to be 1.0 since this results in maximum neutron flux peaking. Studies made with various initial values for effective multiplication factors have shown that a larger neutron flux peak occurs for larger initial values. Since i k= 1.0 is the upper limit to the suberitical region, it is I the value used in the analysis.

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, as well as delays for trip signal actuation and rod cluster control assembly release, is taken into account. A 10 percent increase is assumed for i power range flux trip setpoint raising it from the nominal value of 25 percent to 35 percent. Previous results, however, show that rise in the neutron flux is so rapid that the effect of errors in the trip setpoint on the actual time at which the rods are released is negligible. In addition, the reactor trip insertion characteristic is based on the j assumption that the highest worth rod cluster control !

assembly is stuck in its fully withdrawn position. See Section 14D.5 for rod cluster control assemmly insertion characteristics.

5. The maximum positive reactivity insertion rate assumed is greater than that for the simultaneous withdrawal of the combination of the two control banks having the greatest combined worth at maximum speed (45 inches / minute). Control rod drive mechanism design is discussed in Section 3.2.3.

[NSEO:T 4] *N. magnNde dCES hek. C. OPP 4144C dlP4ckig o t AguP4 ND*E he.C4USe. the TWNKLE code.y on wMch the. neutronic.s anm% sis is base.d, is a thermal dhian code. ra%er t% a y:Ank ktnatics sprecnnmation. 14.1-3 - _ - - _ _ _ _ _ _ _ _ _ _ - . - - _ _ _ - _ _ _ _ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ i

BVPS-1-UPDATED FSAR Rev. 6 (1/88) tr a t e $ W v. ..e .w 1 3

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(.igh;;y,r:p;tivity inye.Livn rei. ud lewest g - KE9 yw.m uw.m . ... Results Figure 14.1-1 through 14.1-3 show the transient behavior for the (d)indi;;t;d r;;;tivity ins;rti;n rate with the accident terminated by reactor trip at 35 percent nominal power. ;;;asing thsee lee es ep;r; tin;- This insertion rate is greater than that for the two E" highest worth control banks, both assumed to be in their highest incremental worth region. l Figure 14.1-1 shows the neutron flux transient. The neutron flux overshoots the full power nominal value but this occurs for only a j very ort time period. Hence, the energy release and the fuel I tempev ure increases are relatively small. The thermal flux REV response, of interest for DNB considerations, is shown on Figure 14.1-2. The beneficial effect on the inherent thermal lag in the fuel is evidenced by a peak heat flux less than the full power nominal value. There is a large margin to DNB during the transient q since the rod surface heat flux remains below the design value, and y there , is a high degree of subcooling at all times in the core. 71v .. 3 4,.. , ouw-. wuw ...yvu.= vu uus avutava tuwa, v4.uutuw, anu vvvAsus t;;;;;;tur;. r-Th; ;;;r:g; fuel t;;p;;;tur; incr;;;;; ; e value icwei

Li..u who uvmuual Iwil yv.wt value. E REv p Fiw us.

wuw 14.1-1A LL vuyh 14.1-;A she, the sea; tron;i;nt.per;;;t;;; for y m..w ith twe ups..;ing r;;;ter ;;;1;nt 1;;;;, aith th; ;;;id;nt eg.in te.minated .h> reacter trip ;; 25 p;;;;nt n;;in:1 ;;w;;. n Th; ! inserti;n ret; e;;used is g;;;;;; then that for the tu; high;;t worth

         ;;nizel                                  bank;,             b;th e;;;;;d t; i; in th;ir high;;t in;;;;;nt:1 w;;th                                                        ($)
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REPLACE W 4 TH (MsERT C R2V Conclusions In the event of a rod cluster control assembly withdrawal accident from the suberitical condition, the core and the Reactor Coolant System are not adversely affected, since the combination of thermal power and the coolant temperature result in a Departure from Nucleate Boiling Ratio (DNBR) ;;11 :h;;; th; limiting 7:1;; cf 1.00. Thi; cen;16.ieu ruv,_.rding

        .muw-
                          . . _ .                        _ _ _    _, , _ , , __ the OL*CO i; Le;ed en who feci ihei ihe enely5e5 wu w uvmau.4 au6                                        yw-v. va    uw.        . v .- . wuw ywam wvv4-uw wumys.aww.w,       _ _ _ _ . _ _                  ,

1 PPV therLe1 p;w;r ;;.d h;;t fl^,.; ;;; not ;;;;;d:d. ItEPLAcre WiTH IMSERT@ 14.1.2'*-Uncontrolled Roo Cluster Control Assembly Bank Withdrawal at Powe

                                                                                          -[HLS SECT LOM MOT AAobtst E'D 14.1.2.1 Identification of Cau                                                                      Accident Description Uncontrolled rod cluster control assembly                                                                                        hdrawal at power results in an increase in the core heat                                                                            flux.                        the heat 14.1-4 i

BVPS-1 UPDATED FSAR 14.1.1 Uncontrolled RCCA Withdrawal from Subcritical Inserts for Zirc Grids and Increased Peaking Factors Insert A: This event is classified as an ANS Condition II incident. Insert B: I

5. Hiah Neutron Flux Rate Trio - Actuated when the positive rate of I I

change of neutron flux on two out of four nuclear power range channels indicate a rate above the preset setpoint. This trip function is always active. 1 Insert C: i 14.1.1.2 Analysis of Effects and Consequences Method of Analysis The analysis of the uncontrolled RCCA bank withdrawal from subcritical accident is performed in three stages: 1) an average core nuclear power transient calculation; 2) an average core heat transfer calculation; and 3) the departure from nucleate boiling ratio (DNBR) calculation. The average core nuclear calculation is performed using a spatial neutron kinetics code, TWINKLE (Reference 10), to determine the average power generation with time including the various total core feedback effects, i.e., Doppler reactivity and moderator reactivity. The average heat flux and temperature , transients are determined by performing a fuel rod transient heat j transfer calculation in FACTRAN (Reference 2). The average heat flux is next used in THINC (described in FSAR Section 3) for the transient DNBR calculation. Insert D: .

6. The most limiting axial and radial power shapes, associated with having the two highest combined worth sequential control banks in their highest worth position, is assumed in the departure from nucleate boiling (DNB) analysis.

7. The initial power level was assumed to be below the power level expected for any shutdown condition (10E-09 of nominal power).

The combination of highest reactivity insertion rate and lowest initial power produces the highest peak heat flux.

8. Two RCPs are -assumed to be in operation. This lowest initial flow minimizes the resulting DNBR.

Insert E: uncontrolled RCCA bank withdrawal Insert F: Figure 14.1-3 shows the response of the hot spot fuel and cladding temperature. The hot spot fuel average temperature increases to a value lower than th.e nominal full power hot spot value. The minimum DNBR at all time remains above the safety analysis limit value. __L_-___-__. _ - - _ _ _ - _ . _ - - _ _

BVPS-1 UPDATED FSAR 14.1.1 Uncontrolled RCCA Withdrawal from Subcritical Inserts for Zirc Grids and Increased Peaking Factors Insert G: Since only two RCPs were assumed to be in operation, these results are conservative compared to assuming RCPs in operation for N loops. The calculated sequence of events for this accident is shown in Table 14.1-2. With the reactor tripped, the plant returns to a stable condition. The plant may subsequently be cooled down further by following normal plant shutdown procedures. Insert H: which is always greater than the limit value. Thus, the DNB design basis as described in FSAR Section 3 is met. j

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LVPS-1-UPDATED FSAR Rev. 6 (1/88) extraction from the steam 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 would eventually result in DNB. Therefore, in order to avert damage to the cladding the Reactor ProtectionSystemisdesigyedtoterinateav>:-Ak 1 30 . A e nSurE R D4474 Ouch Safely transient bcfcrc[the q u /ps/3 accep a. req.ONDE CN/ce"4' ells 44 bcivwWe/h The automatic features of the Reactor Protection System which prevent core damage following the postulated accident include the following:

1. Power range neutron flux instrumentation actuates a reactor trip if two out of four channels exceed an overpower setpoint.

2. Reactor trip is actuated if any two out of three 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 three AT channels exceed an overpower AT setpoint. This setpoint is automatically varied with axial power imbalance to ensure that the allowable linear heat generation rate (kW/ft) is not exceeded.

4. A high pressurizer pressure reactor trip actuated from any two out of three pressure channels which is set at a fixed point. This set pressure is less than the set pressure for l the pressurizer safety valves.

5. A high pressurizer water level reactor trip actuated from any two out of three level channels which is set at a j fixed point.

In addition to the above listed reactor trips, there are the following rod cluster control assembly withdrawal blocks:

1. High neutron flux (one out of four)

2. Overpower AT (two out of three) 3.- Overtemperature AT (two out of three) -

The manner in which the combination of overpower and overtempera-l ture AT trips provide protection over the full range of Reactor Coolant System conditions is described in Section 7. This includes a plot (Figure 14D-1) presenting allowable reactor coolant loop average i temperature and AT for the design power distribution and flow as a function of primary coolant pressure. The boundaries of operation defined by the overpower. AT trip and the overtemperature aT trip ! are represented as " protection lines" on this diagram. The 14.1-5 l 1 l

BVPS-1-UPDATED FSAR Rev. 6 (1/88) protection retpoint errors lines are so drawn to include all adverse instrumentation and that under nominal conditions trip would occur well within the area bounded by these lines. The utility of this diagram is in the fact that the limit imposed by any given DNBR can be represented as a line. The DNB lines represent the locus of conditions for which the DNBR equals 1,00. All points below and to the left of a DNB line for a given presShre have a DNBR greater than

                  .00.                                                  The diagram shows that DNB is prevented for all cases if the 2rea enclosed with the maximum protec the applicable N R line at any, point.                                                                   on lines is not traversed by pe. sa    y analysa h m sf The                              area                      of permissible operation (power, pressure, and temperature) is                  bounded by                                    the        combination         of reactor trips: high neutron flux (fixed                                                setpoint); high pressure (fixed setpoint); low pressure (fixed cetpoint); overpower and temperature AT (variable setpoints).

14.1.2.2 Analysis of Effects and Consequences, Method of Analysis This transient is analyzed by the LOFTRAN(s) code. cimulates the neutron kinetics, Reactor Coolant System, pressurizer, This code pressurizer relief and safety valves, pressurizer spray, steam generator and steam generator safety valves. The code computes plant variables including temperatures, pressures, and power level. core limits as illustrated in Figure 14D-1 are used as input The to LOFTRAN to determine the minimum DNBR during the transient. The cote limits are calculated "" '--'-- ' "- W-3 ra;E cen eletion. o h" us,'"'

                                                                                                /es  Opp M  7;*in thc fec"k, grid 3,9,spa--

J. J, /, In order to obtain conservative values of DNBR the following assumptions are made:

1. Initial conditions of maximum core power and reactor coolant average temperatures and minimum reactor coolant pressure resulting in the minimum initial margin to DNB.

2. Reactivity Coefficients - Two cases are analyzed:

a. Minimum Reactivity Feedback -

A zero moderator coeffic'3at of reactivity is assumed corresponding to the be w .nning of core life. 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 - A conservatively large '

positive moderator density coefficient and large (in absolute magnitude) negative Doppler power coefficient l are assumed.

3. The reactor trip on high neutron flux is assumed to be actuated at a conservative value of 118 percent of nominal full power. The AT trips include all adverse instrumentation and setpoint errors, while the delays ;

14.1-6

BVPS-1-UPDATED FSAR Rev. 0 (1/82) for the trip signal actuation are assumed at their maximum values. 4 .- The rod control cluster assembly trip insertion characteristic is based on the assumption that the highest worth assembly is stuck in its fully withdrawn position.

5. The maximum reactivity insertion rate assumed is greater than that for the simultaneous withdrawal of the combination of the two control banks having the maximum combined worth at maximum speed.

The effect of rod cluster control assembly movement on the axial core power distribution is accounted for by causing a decrease in overtemperature and overpower AT trip setpoints proportional to a decrease in margin to DNB. - Results Figu;es 14.1-4 and 14.1-5 show the response of neutron flux, pressure, average coolant temperature, and DNBR to a rapid rod withdrawal incident starting from full power. Reactor trip on high neutron flux occurs after start of the accident. Since this is rapid with respect to the thermal time con'stants of the plant, small changes in T and pressure result and large margin to DNB is maintained. av9 The response of neutron flux, pressure, average coolant temperature, and DNBR or a slow rod control assembly withdrawal from full power is shown in Figure 14.1-6 and 14.1-7. Reactor trip on overtemperature AT occurs after a longer period and the rise in temperature and pressure is consequently larger than for the rapid rod cluster control assembly withdrawal. Figure 14.1-8 shows the minimum DNBR as a function of reactivity insertion rate from initia'. i full power operation for the minimum and maximum reactivity feedoack. It can be seen that two reactor trip channels provide protection over the whole range of reactivity rates. These are the high neutron flux and overtempera e AT trip channels. "hc r.inir m. ONO1 i ncycr lecc -thart-1. ;C . hte 'DAdd safe}y a tal y &S aCCefErce Cfikfid aM me.h Figures 14.1-9 and 14.1-10 show the minimum DNBR as a function of reactivity insertion rate for rod cluster control assembly withdrawal incidents starting at 60 and 10 percent power respectively. The results are similar to the 100 percent power case, except as the initial power is decreased, the range over which the overtemperature AT trip is effecti is increased. -ffr-ncith:2 cacc do s th; ON 0 e DA/dA .safedy a.talys is accefA nce uil0'1 entfall arebclow

                                         ~ef1.N. be                                caes.

14.1-7

[ BVPS-1-UPDATED FSAR Rev. 0 (1/83) 1>46R safe aulysis accep}ance-Conclusions M' " '

                                                                                                  .. g The high neutron flux and                   (

overtemperature AT trip channels provide adequate protection over the4 entire range of possible reactivity insertion rates, i.e., the minimum valuc of ON0n ic alwayc larger than 1.30 The DNBR predictions shown on Figure 14.1-10 to 14.1-12 apply to the hottest channel in the core and are the minimum values for any point in the core. 14.1.3 Rod Cluster Control Assembly Misalignment 14.1.3.1 Identification of Causes and #ccident Description Rod cluster control assembly misalignment accidents include:

1. A dropped ful:.-length assembly

2. A dropped ful!.-length assembly bank -

3. Statistically misaligned full length assembly (Table 14.1-1).

Each rod cluster control assembly has a position indicator channel which displays position of the assembly. The displays of arsembly positions are grouped for the operator's convenience. Fully inserted assemblies are further indicated by a rod bottom light. Group demand position is also indicated. The full length assemblics are always moved in preselected banks and the banks are always moved in the same preselected sequence. A dropped assembly or assembly banks are detected by:

1. Sudden drop in the core power level is seen by the Nuclear Instrument System

2. Asymmetric power distribution as seen on out of core neutron detectors or core exit thermocouple

3. Rod bottom light (s)

4. Rod deviation alarm

5. Rod position indication.

Misaligned assemblies are detected by:

1. Asymmetric power distribution as seen on out of core neutron detectors or core exit thermocouple

2. Rod deviation alarm

3. Rod position indicators

4. Power range negative rate trip.

14.1-8

l BVPS-1-UPDATED FSAR Rev. 0 (1/82) The resolution of the rod position indicator channel is 5 percent of span (17.2 inches). Deviation of any assembly from its group by twice this distance (10 percent of span, or 14.4 { inches) will not cause power distributions worse than the design ! limits. The deviation alarm alerts the operator to rod deviation ' with respect to group demand position in excess of 5 percent of span. If the rod deviation alarm is not operable, the operator is required to log the rod cluster control assembly positions in a prescribed time sequence to confirm alignment. If one or more rod position indicator channels should be out of service, detailed operating instructions shall be followed to assure the alignment of the non-indicated assemblies. These operating instructions require selected pairs of core exit thermocouple to be monitored in a prescribed time sequence and following significant motion of the non-indicated assemblies. The operating instructions also call for the use of moveable in-core neutron detectors to confirm core exit thermo' couple indication of assembly misalignment. 14.1.3.2 Analysis of Effects and Consequences, Method of Analysis Steady state power distributions are analyzed for this event using the TURTLE IO code. The peaking factors calculated by TURTLE are then used by the THINC code to calculate the DNBR. For the transient response to a dropped RCCA or RCCA bank the LOFTRAN N code is used. The code simulates the neutron kinetics, Reactor Coolant System, pressurizer, pressurizer relief and safety valves, pressurizer spray, steam generator, and steam generator safety valves. The code computes pertinent plant variables including temperatures, pressures, and power level. Results A dropped rod cluster control assembly typically results in a reactivity insertion of -0.15 percent ok/k. Analyses have shown , that with the core power distribution which exists following the drop of a single rod cluster control assembly, the reactor may be returned to full power with the full power Reactor Coolant System temperature without the DNSn g^'ng "I This is verified by the results in Table 14.1-1.Y abcw 1.00. ff

e. DN64 /

GCCejUdnce. C.fllSa e yweh PJ 5 Q u't. aga y.s1.3 Extensive analyses were performed to show that the minimum DNBR occurs near the end of the transient when the system has essentially returned to a new steady state equilibrium condition. Without automatic rod control, the system will return to a new equilibrium condition at a reduced primary temperature as a result of the moderator reactivity feedback. As typical of PWR uncontrolled response, the return of power is monotonic and therefore, power overshoot is not a condition for this case. A power overshoot after a dropped rod cluster assembly incident can only result from the action of the automatic rod controller. 14.1-9

BVPS-1-UPDATED FSAR Rev. 0 (1/82) to its insertion limit with any one assembly fully withdrawn H thc"t the N?SP f" lca 1.20 a shown i Table 14.1-1. k k $lt a bp6R Sab. ling

                           }y anbelySIS act        ace Chhik a/t >*1ef The insertion limits in the Tech ical Specifications may vary from time to time depending on a number of limiting criteria.                                                            It is preferable, therefore, to analyze the misaligned' assembly case at full power for a position of the control group as deeply inserted as the criteria on minimum DNBR and power peaking factor will allow. The full power insertion limits on control group D must then be 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.

DNB calculations have not been performed specifically for assemblies missing from other banks, however, power shape calculations have been done as required for the ro?. cluster control assembly ejection analysis. Inspection of the power shapes shows that the DNB and peak kW/ft situation is less severe than the group D case discussed above assuming insertion limits on the other groups. 14.1.3.3 Conclusions a e U' F M g,My ana /ys

   .It is shown that in al        cases of dropped single assemblies, the DNBR rcmaine     greatcr     than    1.20 at powcr and, consequently, dropped single assemblies do not cause core damage.

For all cases of dropped groups the reactor is tripped by the power range negative neutron flux rate trip and consequently dropped banks do not cause core d ' Safe age.lySIS 0" # 'C'/' " " ' *" For all cases of any group [ insert d to its rod insertion limit

with any single rod clusteri/ control assembly in that group fully withdrawn, the DNBR rcm;;n; gractcr than 1.20. Thus, rod misalignments do not result in core damage. 14.1.4 Uncontrolled Boron Dilution 14.1.4.1 Identification of Causes and Accident Description Reactivity can be added to the core by feeding primary grade water into the Reactor Coolant System via the reactor makeup portion of the Chemical and Volume Control System. Boron dilution is a manual operation under strict administrative controls with procedures calling for a limit on the rate and duration of dilution. A boric acid blend system is provided to permit the operator to match the boron concentration of reactor coolant makeup water during normal charging to that in the l Reactor Coolant System. The Chemical and Volume Control System is designed to limit, even under various postulated failure modes, the potential rate of dilution to a value which, after indication through alarms and instrumentation, provides the operator sufficient time to correct the situation in a safe and orderly manner. 14.1-11

BVPS-1-UPDATED FSAR Rev. 6 (1/88) TABLE 14.1-2 TIME SEQUENCE OF ly,ENT,5 FOR CONDITION II EVENTS Accident Uncontrolled RCA Withdrawal from a subcritical condition 3 Lvvy Cy .L en -2 Lwwy Ces r a t. en Event Time (see) Timm 'sec;- Initiation of uncontroll d 0 u rod withdrawal 7.5 x ' thl ;;. tic.; :::: z;acL from eam 10- M osaz f nominal power Power range high neutron 0.0 LOA 13.6 flux low setpoint reached Peak nuclear power occurs -7. O 10.6 13.o Rods begin to fall into -7 rs- lO M 14.1

                                             ' core F
                                         -i Minimum bune ecc.em                                                                   12 9 Peak heat flux occurs                                                         M 62.4                  15.5 l

1

                                       <     _     Peak average                    .5. ladtempera-
                                                                                     --                                          er2- l 2,,2 *j          13.7 uture occurs Peak average e                          tempera-                              M i2M                   15.0 ture occurs rc;h recr;;; ;;;1-n; memyera                                                  t:                   N/A m es e ec c .. e I

CEV l 1 of 11

t REV. 0 (1/82) 10.0 , 9.0 8.0 7.0 - 6.0 - l

n0TE: NEUTROM FLUX STARTS l- 4.0 - AT 10- 0 AT m t ZERo l 3.0 - 1 l

                                                                                                                                                           +%(
                           " 2.0 7

0@ 8 5 51.0 - a 0.9 -

6 0.8 - { g 0.7 - ' d 0.6 - 0.5 - S

                           = 0.4          -

0.3 - 0.2 - 0.1 5.0 10.0 5.0 TIME (SECONDS) FI6URE 14 1-1 UNCONTROLLED R0D WITHDRAWAL FROM A SU8 CRITICAL CONDITION NEUTRON FLUX VS. TIME BE AVER VALLEY POWER ST ATION UNIT NO. 1 UPDATED FINAL $AFETY ANALV515 REPORT

J' Figure 184.t-1 Uncontrolled Rod Withdrawal from a Suberitical Condition Neutron Flux versus Time 10' 101 -- 5 a c. g 100 - b 8 z 10~2

14 16 18 20 22 24 26 28 30 0 2 4 6 6 10 12 TIWE (SECONDS)

                                                                                              .- g

[* REV. 6 (l- 88) [. 10... ; ; ; l l .. g .. ... z -. .. E NOTE: NEUTRON FLUX STARTS o -- 33 z AT 10 AT TIME ZERO o 1.0:: :: z : - E :: :: 3 . .- .. g .. .. b .. X o

                                $     B'I ' ?            '

i z g .. r .. g .. .. z- .. __ 10 k5 b0 25 TIME (SECON S) FIG.14.1-1 A (TWO LOOP OPERATI ) UNCONTROLLED ROD THDRAWAL FROM A SUBCRITICAL ONDITION NEUTRON FLUX VS. TIME BEAVER VALLEY POWER STATION UNIT I UPDATED FINAL SAFETY ANALYSIS REPORT - - _ _ _ _ _ _ - = _ - - - _ _-_____-________--______---_-_-_--_-_-___-________-___--_--___Q

r_ I i REV. 0 (1/82) 4 I 1 i l.0 l REACTivlTY INSERTION RATE = 75. x 10-5 AK/SEC 0.8 -

o.s 5

k(

                       ;-                                                                                                                                        7 E                                                                                                                                          9 e
                       -                                                                                                                                            D E o.4    -

d a m 0.2 - 1 i i I i 0 0 4 8 12 16 20 TIME (SECOMDS) FIGURE 14 1-2 UNCONTROLLED R0D WITHDRAWAL FROM A SUBCRITICAL CONDITION THERMAL FLUX VERSUS TIME BE AVER V ALLEY POWER ST ATION UNIT NO. 1 UPDATED FINAL $AFETY ANALYSIS REPORT

Fic,ues 14.1-2. UMO.OWrrROLLEb Roo WimseAwAt. PROM A SVEbC.ItlTICAL CNbmOM MEAT FLUY VsTIME 1.0 i

B w

                                                          .50-
                                                         .40-
                                                         .30-
                                                         .20-
                                                         .10-O.00                           ~  ~     ~  ~   ~
                                                                                                                 ~
                                                                                                                                     )

0 2 4 6 $ 10 12 14 16 18 20 22 24 26 26 ,10 TlWE (SECONDS)

 --------.--__-.m  _ . . _ _ _ _ , . . . , _ _ _ _ _ .

L l i ItEV. 6 C l SS)' - < .j L 1.0000 : ; ; ;

         'S g

Z- .80000-- --

       .$  z
       .o                                             h z          .60000--                                    -                             --

9 , h 1 E~

                      .40000-   -                                                               --

s i d a .20000- - -- I E 0.0 , ; l l l 0 5 10 15 20 25 TIME (SECON ) FIG.14.1-2 A ! (TWO LOOP OPERATION) UNCONTROLLED ROD WITHORAWAL FROM A SUBCRITICAL CONDITION THERMAL FLUX VERSUS TIME , BEAVER VALLEY POWER STATION UNIT I ! UPDATED FINAL SAFETY ANALYSIS REPORT

1 1 i REV. 0 (1/82) l

1000

 ,                                                                REACTIVITY IN RTION RATE : 75 I 10-5 6K/SEC 900  -

b M

                                         !=       -
                                         %                                                                                                                                                                            FUEL (

i

%+

y 700 - 5 5 CLAD 600 - ORE WATER

                                             ,                                                                I                                                                                          I                    I O                                                         4                                                                                           8                     12              1             20 TIME (SECONOS) l FIGURE 14 1-3 UNCONTROLLED R0D WITHDRAWAL FROM A SUBCRITICAL CONDITION, TEMPERATURE VS TIME BE AVER VALLEY POWER ST ATION UNIT NO. 1 UPDATED FINAL SAFETY ANALYS!$ REPORT

Figure 14. l-3 Uncontrolled Rod Withdrawal From a Subcritical Condition, Fuel Temperature versus Time 70 W .. iii G=

                                    -g..

En E .. t

                                            ....                                                                                         ]

Tiuc (sccoNos)

                                            .000 t..O=

9. .

m l i i ils...

                                      .     ..                                                                                                                                                                                                                                       I
                                                                                                                                              )                                                                                                                    __

a .. .. .. .. == .. .. a. o . . . . .. .. Ysue (seconds) 1

REV.6 (l-83) l l 1000.00 l l l l 900.00 - - -- C O 800.00 -- -- 3 e b h i N 700.00 -- . FUEL u e W G 600.00 -- CLAD -- COR WATER 500.00 l l l l 0 5 10 15 20' 25 TIME (SECON ) i FIG.14.1-3A t (TWO LOOP OPERATION) UNCONTROLLED ROD WITHORAWAL FROM A SUBCRITICAL CONDITION TEMPERATURE VERSUS TIME BEAVER VALLEY POWER STATION UNIT I UPDATED FINAL SAFETY ANALYSIS REPORT

BVPS-1-UPDATED FSAR Rev. 0 (1/82) Therefore, adequate time is available following the alarms for the operator to determine the cause, isolate the primary grade water source, and initiate reboration. With the reactor in manual control and if no operator action is taken, the power and temperature rise will cause the reactor to reach the overtemperature AT trip setpoint. The boron dilution accident in this case is essentially identical to a rod cluster control assembly withdrawal accident. The maximum reactivity insertion rate for boron dilution is shown in Figure 14.1-12 and is seen to be within the range of insertion rates analyzed for a I rod cluster control assembly withdrawal accident. Prior to the overtemperature AT trip, an overtemperature AT alarm and turbine runback would be actuated. There is ample time available (approximately 14 minutes) after a reactor trip for the operc or to determine the cause of dilution, isolate the primary grade water sources and initiate reboration before the reactor can return to criticality assuming a 1 percent shutdown margin at the beginning of dilution. ggyg 14.1.5 PartialLossofForcedReactorCohlantFlow f 14.1.5.1 Identification of causes and Acc dent Description A partial loss-of-coolant flow accident an result from a mechanical or electrical failure in a react coolant pump, or from a fault in the power supply to the pump. 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 is not tripped promptly.

c. css EM 6

                                             >The necessary protection against a partial loss-of-coolant flow accident is provided by the low primary coolant flow reactor trip which is actuated by two out of three low flow signals in any reactor coolant loop.          Above approximately 30 percent power (Permissive 8) , low flow in any loop will actuate a reactor trip.

Between approximately 10 percent power (Permissive P-7) and the power level corresponding to Permissive P-8 low flow in any two loops will actuate a reactor trip. 6

                                                                                      'TN S G R 7~ C-Normal power for the pumps is supplied through buses connected to the main generator.        Each pump is on separate bus.                When a generator trip occurs, the pumps are automatically transferred to a bus supplied from external power lines, and the pumps will continue to supply coolant flow to the core.                    Following any turbine trip where there are no electrical faults which require tripping the generator from the network, the generator remains connected to the network for approximately 30 seconds.                      The reactor coolant pumps remain connected to the generator thus ensuring full flow for approximately 30 seconds after the reactor trip before any transfer is made.

14.1-15

l BVPS-1-UPDATED FSAR Rsv. 0 (1/82) 1

14. 5.2- Analysis of Effects and Consequences, Method of l Analysis j l'

The follow three cases have been analyzed: Loop i tially Pumps coasting Loop stop operat down valv 3 1 n ! 2 1 open

       @                                                               2                                                                                 1                 closed                       )

g.

        /2   This transient is analyzed by fo the PHOENIX 'I code is used to                       f                                                                                    alculat i al computer codes. First the loop and core flow f-during the transient.

calculate the time of' reactor trip, based on 3 by PHOENIX, and the n,pdIsar power transient The FTRAN(s) c e is then used to flows calculated lowing reactor u trip. The FACTRAN

I code is then used to cale ate the heat flux transient ha on the nuclear power from LOFT and flow

{I from PHOENIX. ally the THINCf 'I code is used to cal ate the h' minimum DNBR ring the transient based 'on the heat fl from M FACTRAN flow from PHOENIX. The R grid spacer facto is applied the W-3 correlation. The DNBR transients present repre t the minimum of the typical or thimble cell. REV Initial-Conditions Initial operating conditions assumed are the most adverse with respect to the margin to DNB, i.e. , maximum steady state power level, minimum steady state pressure, and maximum steady state-coolant average temperature. (See Section 14D.2 for explanation f initial conditions.,) 1 r. eg.7.;.iu{,

                                                                                                                                            "i { g e.; ef t;g. _,
                                                                                                                                                                           .. ... m -~.

oo. - - r .. ..... s. ..... , ......, ...

           ,ef ;;;                                        ti;e i; ;;;1 '.                                                                    (.bEWTE7 N-1 MCT Cousmelt EN Reactivity Coefficients A conservatively large absolute value of the Doppler-only power coefficient is used (See Table 14D-2).                                                                                                                   The total integrated Doppler reactivity from 0 to 100 percent is assumed to be 0.016Ak.

The lowest absolute magnitude of the moderator temperature coefficient (0.0Ak/F) is assumed since this results in the maximum hot-spot heat flux during the initial part of the transient when the minimum DNBR is reached. Flow Coastdown The flow coastdown analysis is based on a momentum balance around ! I each reactor coolant loop and across the reactor core. This momentum balance is combined with the continuity equation, a pump 14.1-16 E _ _ _ _ _ _ _ = _ _ ___ _ _ . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _

BVPS-1-UPDATED FSAR Rsv. 0 (1/22) momentum balance and the pump characteristics and is based on high estimates of system pressure losses. g Results The d sequence of events is shown on Table u h the three cases d. Figure 14.1-13 +ht, .1-21 show r {$ the loop coastdowns, the - M the nuclear power coastdowns, and the av i ot c t flux coastdowns bf

     "#      for each of                       a cases. The minimum DNBR                                                    f the es is not less than 1.3.

tav , Conclusions . h acs _ . _ _ ___, _ _ m .._ m_ ... Wrg a ssw maamay w a s . ..w w - m...- w - w... am w 3.

                                                                       =*** aaw w wwwsumes wwAww baas
                                                                                                                      . .     .. 4 i

mssu limiting velee ef 1.00 et eny ti.T.e L;ing the ;.. .ne ien t . Thus, I

there will be no cladding damage and no release of fission products to the Reactor Coolant System.

14.1.6 Startup of an Inactive Reactor Coolant Loop 14.1.6.1 Identification of Causes and Accident Description 14.1.6.1.1 With Loop Stop Valves open If the plant is operating with one pump out of service, there is reverse flow through the 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 reactor core inlet temperature). If the reactor is operated at power, 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 unit be brought to a load of less than 25 percent of full power prior to starting a pump in an inactive loop in order to bring inactive loop hot leg temperature closer to the core inlet temperature. Starting 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 causes a rapid reactivity insertion and subsequent power increase. 14.1.6.1.2 With Loop Stop valves clo::ed i In the case of the plant operated at reduced pcwor with a reactor coolant loop out of se.rvice and with the loop stop valves of one of its loops closed there is no flow from the reactor vessel and active loops to the inactive loop and the plant operates much as if it were a unit lacking one loop. With the stop valves in one loop closed, the isolated section of the loop would be cooler than the temperature of the active 14.1-17 _-___-_____-_ _ __ ___--__-__ _ _ _ _-_ - _ A

BVPS-1-UPDATED FSAR Rev. 0 (1/82) loops. Administrative procedures require that the plant be brought to zero load, the temperature of the isolated loop brought to within 20 F of the active loops, and the boron concentration of the isolated loop verified prior to opening the loop stop valves and returning the loop to service. Interlocks are provided to ensure that an accidental startup of an isolated loop which has a lower temperature or lower boron concentration than the core and active loops will be a relatively slow event. The interlocks insure that flow from the isolated loop to the remainder of the Reactor Coolant System takes place through the relief line bypassing the cold leg stop valve for a

                               . period of approximately one hour before the cold leg stop valve can be opened. The flow through the relief line is made low (no g     'more                                                                                                   tnan' 000     gpm)   so that the temperature and boron g                                   concentration in the isolated loop are brought to equilibrium with the remainder of the system at a relatively slow rate should the administrative procedures be violated and an attempt made to open stop valves when the isolated loop temperatures or boron concentration is lower than that in the core and active loops.

Interlocks are provided to:

1. Prevent opening of a hot le loop stop valve unless the cold leg stop valve in the same loop is fully closed.

2. Prevent starting a reactor coolant pump unless:

a. The cold leg loop stop valve in the same loop is fully closed, or j b. Both the hot leg loop stop valve and cold leg loop stop valve are fully open.

3. Prevent opening of a cold leg stop valve unless

a. The hot leg loop stop valve in the same loop has been fully opened for 1 hour.

b. The bypass valve in the loop has been opened for 1 Mu.

c. Flow has existed through the relief line for 1 hour.

d. The cold leg temperature is within 20 F of the highest cold leg temperature in the other loops and the hot leg temperature is within 20 F of the highest hot leg temperature in the other loops.

The interlocks are a part of the Reactor Protection System and include the following redundancy: 14.1-18

.o.

I BVPS-1-tiPDATED FSAR Rev. 0 (1/82) ). 1. Two independent limit switches to indicate that a valvs is fully open.

2. Two independent limit switches to indicates that _a valve is fully closed.

3. Two differential pressure switches in each line which bypasses a cold leg loop stop valve to determine that flow exists in the line. Flow through the line indicates:-

a. The valves in the line are open.

i. b. The pump in the isolated loop is running.

The interlocks meet the IEEE 279-1971 criteria and, therefore, cannot be negated by a single failure. The interlock on hot leg temperatures is a backup for the interlock on cold leg. temperatures. Thus, the single failure criterion applies to the combination and not to each separately. With the above protection system interlocks, the following procedure is necessary in order to reopen loop stop valves once either stop valve in a loop has left the fully open position.

1. The cold leg loop stop valve must be fully closed-before the tot leg stop valve can be returned to its-fully open position.

1 2. Flow must have existed fyom the isolated portion of the system to the remainder of the system (maximum rate is l g ,' approximate gym) for at least 1 hour through the R2J line bypass ing the cold leg stop _ valve and the isolated loop and active loop temperatures must agree within 20 F before the cold leg loop stop valve can be opened. 14.1.6.2 Analysis of Effects and Consequences 14.1.6.2.1 Method of Analysis A detailed digital simulation of the plant, including heat transfer to the steam generators of the active and inactive loops and Reactor Coolant System flow transit times, is used to study the transient following the startup of an idle pump. Icop Stop valves open Assumptions are:

1. Initial conditions of maximum core power and reactor '

coolant average temperatures and minimum reactor coolant pressure resulting in minimum initial margin to DNB. These values are to be consistent with maximum steady state power level allowed with all but 14.1-19

BVPS-1-UPDATED FSAR Rev. 0 (1/82) one loop in operation including appropriate allowances for calibration and instrument errors. The high initial power gives the greatest temperature difference between the core inlet temperature and the inactive loop hot leg temperature. 2.. Following the start of the idle pump, the inactive loop flow reverses and accelerates to its nominal full flow value.

3. A conservatively large (absolute value) negative moderator coefficient associated with the end of life.

4. A conservatively low (absolute value) negative Doppler power coefficient is used. -

5. The initial reactor coolant loop flows are at the appropriate values for one pump out of service, g

e analysis reactor trip is conservatively assumed to

            -              actuate             the      high neutron flux reactor tri                    . trip 3               setpoint was ass                   be 118 percent of no              ull power.      In ul6            practice, however, react                -

w expected to occur on F power range neutron flux Permissive P-8 setpoint yE (set at approx percent of n 1 power). The Permissi setpoint will remain active unti in the h$ e loop reaches 90 percent of its nominal value. D Loop Stop valves closed The start-up of an inactive reactor coolant loop with the loop stop valves initially closed has been analyzed assuming the m ( inactiveportion active loop to be ' ofat- athe boron concentration system of 0 ppm is at 1,500' ppm,while a theconservatively'ggoo high value for beginning of life. The flow through the relief line is assumed at its maximum value of approximatley -&G4- gym. gg 250 14.1.6.2.2 Results With Loop Stop va.1ves open The results following the startup of the idle pump, with the loop stop valves initially open are shown in Figure 14.1-22. The minimum DNBR during the transient is never less than- 1.00 The sequence of events for the accident is sh on l calculated Table 14.1-2. i %,% ar*\ fis With Loop Stop Valves Closed LiW E. Even with the assumption that administrative procedures are violated to the extent that an attempt is made to open the loop stop valves with 0 ppm in the inactive loop _while the remaining,1800 portion of the system is at -1,500rppm, the dilution of the boron ga in the core is slow. The initial reactivity insertion rate is 14.1-20

BVPS-1-UPDATED FSAR Rev. 0 (1/82) U calculated to be 2.1 x 10 ~5[Ak/second to 1.7 x 10 ~5 Lk/sc cnd, considerably less than the rqpctivity insertion rates considered in Section 14.1.2. It takes , to 10 minutes after the beginning of the dilution before the total shutdown margin is lost assuming one percent shutdown margin at the beginning of the accident. This is ample time for the operator to recognize a high count rate signal and terminate the dilution by turning off the pump in the inactive loop or by borating to counteract the dilution. n l chculd bc further pointed cut that the cxpccted shutdcr. margin eveileble at the beginning of life is of the crder of 2 percent r;ther than the 1 perc nt 00 umcd above. The reactivity addition at end of life due to an attempt to open stop valves when the inactive loop temperature is less than the core temperature is smaller than the reactivity addition considered in the above beginning of life case. , 14.1.6.2.3 Conclusions . With Loop Stop Valves open The transient results show that the core is not adversely affected, i.e., there is considerable margin to limiting =Er of 1.00. fke Safe fy a nelySis /i A U. With Loop Stop Valves Closed The redundant interlocks provided in the Reactor Protection System insure that the temperature and boron concentration in an isolated loop are brought to equilibrium with the remainder of the system at a slow rate. Should administrative procedures be violated and an attempt made to open stop valves when the isolated loop temperature or boron concentration is lower than that in the core, the reactivity addition rate is slow enough to allow the operator to take corrective action before shutdown margin is lost. 14.1.7 Loss of External Electrical Load and/or Turbine Trip 14.1.7.1 Identification of Causes and Accident Description Major load loss on the platt can result from loss of external electrical load or from a turbine trip. For either case offsite ; power is available for the continued operation of the plant components such as the reactor coolant pumps. The case of loss of all AC power (station blackout) is analyzed in Section 14.1.11. For a turbine trip, the reactor would be tripped directly (unless below approximately 10 percent power) from a signal derived from the turbine autostop oil pressure (Westinghouse Turbine) and turbinn stop valves. The automatic steam dump system would accommodate the excess steam generation. Reactor coolant temperatures and pressure do not significantly increase if the 14.1-21

BVPS-1-UPDATED FSAR Rev. 0 (1/82) TABLE 14.1-2 (CONT'D) TIME SEQUENCE OF EVENTS FOR CONDITION II EVENTS f Acci t Event Time (seconds) Partial Los of Forced Reactor Cool Flow

1. All loops o rating, Coastdown begins 0 one pump coas ng down Iow flow reactor trip 1.70 Rods begin to drop 2.70 imum DNBR occurs 3.25

2. Two out of three Coas own begins 0 loops operating, one pump coasting Iow f1 eactor trip 2.16 down, loop stop valves open Rods begin drop 3.16 Minimum DNBR oc urs 3.6

3. Two out of three Coastdown begins 0 loops operating, one pump coasting Low flow reactor trip 2.17 down, loop stop valves closed Rods begin to move 3.17 Minimum DNBR occurs 3.65 REPt.Acg WtTH kUSERT h e

4 of 11

BVPS-1 UPDATED FSAR 14.1.5 Partial loss of Forced Reactor Coolant Flow a Inserts for Zirc Grids and Increased Peaking Factors , Insert A: Each RCP is supplied by a separate bus. Insert B: This event is classified as an ANS Condition 11 incident. Insert C: Above P-7, two or more RCP circuit breakers opening will actuate a reactor trip which serves as a backup to the low flow trip. Insert D: . 14.1.5.2 Analysis Effects and Consequences Method of Analysis The loss of one reactor coolant pump with three loops in operation has been analyzed. This transient is analyzed by three digital computer codes: 1) the LOFTRAN Code (Reference 3) is used to calculate the loop and core flow during the transient, the time of time of reactor trip based onthe calculated flows, the nuclear power transient, and the primary system pressure and temperature transients; 2) the FACTRAN Code (Reference 2) is then used to calculate the heat flux transient based on the nuclear power and flow from LOFTRAN; and 3) the THINC Code (Reference 9) is used to calculate the departure from nucleate boiling ratio (DNBR) during the transient based on the heat flux determined by FACTRAN and the flow determined by LOFTRAN. The DNBR transient presented represents the minimum of the trypical or thimble cell. Insert E: Results Figures 14.1-13 through 14.1-16 show the transient response for the loss of a reactor coolant pump with three loops in operation. Figure 14.1-16 shows the DNBR to be always greater than t'ae safety analysis limit. 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 tempreatures d7 not increase significantly above their respective initial values. The calculated sequence of events for the case analyzed is shown in Table 14.1-2. The affected reactor coolant pump will continue to coast down and the core flow will reach a new equilibrium value associated with the two remaining operating pumps. With the reactor tripped, a stable plant condition will eventually be attained. Normal plant shutdown may then proceed.

! BVPS-1 UPDATED FSAR 14.1.5 Partial Loss of Forced Reactor Coolant Flow Inserts for Zirc Grids and Increased Peaking Factors Insert F: The analysis shows that the DNBR will not decrease below the limit value at any time during the transient. Thus, the DNB design basis as described in Section 3 is met. Insert G: TABLE 14.1-2 (CONT'D) TIME SE00ENCE OF EVENTS FOR CONDITION II EVENTS Accident Event Time (sec) Partial Loss of Forced Reactor Coolant Flow (3 Loops Operating, 1 RCP Coasting Down) Coastdown Begins 0.0 Low Flow Reactor Trip 1.4 Rods Begin to Drop 2.4 Minimum DNBR Occurs 3.6

BVPS-1 UPDATED FSAR 14.1.6 Startup of an Inactive Reactor Coolant Loop Inserts for Zirc Grids and Increased Peaking Factors Insert 1: The reactor trip is assumed to occur on low coolant loop flow when the power range neutron flux exceeds the P-8 setpoint. The P-8 setpoint is conservatively assumed to be 79 percent of rated power, which corresponds to the nominal setpoint plus 9 percent for nuclear instrumentation errors. Insert 2: TABLE 14.1-2 (CONT'D) TIME SE0VENCE OF EVENTS FOR CONDITION II EVENTS Accident fvent Time (sec) Startup of an Inactive Reactor Coolant Loop (Loop Stop Valves Open) Initiation of RCP 0.0 Power Reached P-8 Trip Setpoint 2.7 Rods Begin to Drop 3.2 Minimum DNBR Occurs 4.4

l REY. 0 (1/82) l.2 . i 1.0 CORE FLOW

                       $ 0.8  -

E 8 0.6 -

                                                   'C0' '

5 l G ! W g 0.4 - 0.2 - I I I I 0 0 2.0 4.0 6.0 8.0 10.0 . i TIME (SECONDS) I l ggPLA I 1 5160AE 14 1-13 ALL 1.00PS OPERATING, ONE PUMP COASTING DOWN FLOW COASTDOWNS VS. TIME BE AVER VALLEY POWER STATION UNIT NO. 1 UPDATED FINAL 5AFETY ANAltSIS REPORT I _ -_ - _ -- - - - - - - -- 3

REV. 0 (1/82) l l 1.2 1.0 - AVERACE CHANNEL HEAT FLUX

                                                           -                                                                            AND g                                          NOT CHANNEL HEAT FLUX
                                                           -0 .8   -

I g mEpinom FLUX E 0.6 - p 4 2

                                                           - 0.4   -

5 d 0.2 - l I I I o 0 2.0 4.0 6.0 8.0 10.0 TIME (SECONDS) 6 g.cl FIGURE 14 1-14 ALL LOOPS OPERATING, ONE PUMP C0ASTING DOWN FLUX TRANSIENTS VS. TIME BE AVER VALLEY POWER STATION UNIT No.1 UPDATED FINAL 5AFETY ANALYSIS REPORT mammmme e

REV. 0 (1/82) 2.1 2.0 - 1.9 - l l 1.s - 1 l 1.7 1.6 - I I I I I i.s ! O l.0 2.0 3.0 4.0 5.0 1 TIME (SECONOS) i i J i 9.4 FIGURE 14 1-15 ALL LOOPS OPERATING, ONE PUMP COASTING DOWN DNBR VS. TIME BE AVER VALLEY POWER STATION UNIT N0. 1 UPDATED FINAL SAFETY ANALYSIS REPORT

REY. 0 (1/82) a 1.2 l 1.0 MC CORE FLOW E ! o.s - El 8g - LOOP FLOW

                                                                           ! ! o.s
                                                                           *                                                       'N
                                                                           $ *U 3 { o.4   -

88 UU 0 .2 - W' 8! I I I I O 0 2.0 4.0 6.0 8.0 10.0 TIME (SECONOS) FIGURE 14 1-16 TWO OUT OF THREE LOOPS OPERATING, ONE PUMP COASTING DOWN, FLOW COASTDOWNS VS. TIME (NO !.00PS ]$0 LATED) BE AVER WALLEY POWER STATION UNIT NO. 1 UPDATED FINAL SAFEff ANALf515 REPORT

REV. 0 (1/82) i i 1.2 1.0 - U I

                                   ! .s g                                                      AVERAGE CNANNEL NEAT FLUX AND g 0.6   -

NOT CNANNEL NEAT FLUX U E w { 0.4 a w NEUTRON FLUX

                                                    !                           !           !               l 0

0 2.0 4.0 6.0 3.0 lo,o TIME (SECONDS) 1 6 4 FIGURE 14 1-17 TWO OUT OF THREE LOOPS OPERATING, ONE PUMP C0ASTING DOWN, FLUX TRANSIENT YS. TIME (NO LOOPS ISOLATED) BE AVit VALLEY POWER ST ATION UNIT NO. I UPDATED FINAL 5AFEff ANAlf515 REPORT

e 4

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BVPS-1-UPDATED FSAR Rev.'O (1/82) l I

6. Feedwater Flow -

main feedwater flow to the steam i generators is assumed to be maintained. The effect of l loss of main feedwater coincident with loss of load is analyzed in Section 14.1.8. l I 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. J Results The transient responses for a total loss of load from 102 percent full power operation are shown for four cases; two cases . for the beginning of core life and two cases for the end of core life, in Figure 14.1-23 through 14.1-31. Figure 14.1-23 and 14.1-24 show the transient. responses for - the total loss of' steam load at beginning of life with zero moderator temperature coefficient assuming full credit for the pressurizer spray and pressurizer power operated relief valves. No credit is taken for the steam dump. The reactor is tripped by the overtemperature.AT trip c annel. The minimum DNBR is r- " " - th; 1.3C vnise,- ulG% h DM6d Sah/y oealcs {feMy y GQf et Figure 14.1-25 and 14.1-26 show the responses for the total loss of load at end of life assuming a large (absolute value) negativa moderator temperature coefficient. All other plant parameters are the same as the case above. The reactor is tripped by the overtemperature AT trip channel. The pressure decreases rapidly after about 20 seconds as a result of the large reduction in neutron flux. The DNBR increases throughout the transient. and never drops below its initial value. The pressurizer safety valves are not actuated in the transients shown-in Figures 14.1-23 through 14.1-26. The total loss of load accident was also studied assuming the plant to be initially operating at 102 percent of full power with no credit taken for the pressurizer spray, pressurizer power operated relief valves, or steam dump. The reactor is tripped on the high pressurizer pressure signal. Figures 14.1-27 and 14.1-28 are the beginning of life transients with zero moderator coefficient. The neutron flux remains constant at 102 percent of full power until the reactor is tripped. The DNBR generally increases throughout the transient. In this case the pressurizer safety valves are actuated. The peak pressurizer pressure is 2,566 psia and the maximum surge rate is 20.1 ft s/second. This is compared to a pressurizer safety valve capacity of approximately 35.5 f t 3 /second. The steam generator safety valve setpoint is reached at 10 seconds after the start of the transient and the valves are required to relieve a peak flow of 72 percent of the flow at rated full power at 11.5 seconds. Figures 14.1-29 and 14.1-30 are the transients at end of life with the other assumptions being the same as in Figures , 14.1-27 and 14.1-28. 14.1-24

DVPS-1-UPDATED FSAR Rev. 0 (1/82) Again the. DNBR increases throughout the transient 'and the pressurizer safety valves are actuated. Section 14.1.8 presents additional results of analysis for a complete loss of heat sink including loss of main feedwater. This report shows the overpressure protection that is afforded by the pressurizer and-steam generator safety valves. 14.1.7.3 Conclusions Results of the analyses, including those in Section 14.1.8, show that the plant design is such that a total loss of external electrical load without a direct or immediate reactor trip presents no hazard to the integrity of the Reactor Coolant System 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 integrity of the core is maintained by operation of the Reactor Protection System, i.e., the DNBR will be maintained cvc the 1.30 ccluc. Thus there will be no cladding damage and no,r leas of fission produpts to the Reactor Coolant System. WI J e D)l/6f acc efra .tce_ uik h b . 14.1.8 Loss of Normal Feedwater 14.1.8.1 Identification of Causes and Accident Description A loss of normal feedwater (from pump failures, valve malfunctions, or loss of offsite AC power) results in a reduction in capability of the secondary system to remove the heat generated in the reactor core. If the reactor were not tripped , during this accident, core damage would possibly occur from a l sudden loss of heat sink. If an alternate supply of , feedwater were not supplied to the plant, residual heat following reactor trip would heat the primary system water to the point where water relief from the pressurizer occurs. Significant loss of water from the Reactor Coolant System could conceivably lead to core damage. Since the plant is tripped well before the steam generator heat transfer capability is reduced, the primary system variables never approach a DNB condition. The following provide the necessary protection against a loss of normal feedwater:

1. Reactor trip on low-low water level in any steam generator

2. Reactor trip on low feedwater flow signal in any steam generator. (This signal is actually a steam flow-feedwater flow mismatch in coincidence with low water level) 14.1-25

BVPS-1-UPDATED FSAR Rev. 0 (1/82)

                               'The capacity of the auxiliary feedwater pump is such that the water level in the steam generators being fed does not recede below the lowest level at which sufficient heat transfer area is available to dissipate core residual heat without water relief from the Reactor Coolant System relief or safety valves.

From Figure 14.1-31 it can be seen that at no time is the tube sheet uncovered in the steam generators receiving auxiliary feedwater flow and that at no time is there water relief from the pressurizer. If the auxiliary feed delivered is greater than that of one motor driven pump, the initial reactor power is less than 102 percent of the engineered safeguards design rating or the steam generator water level in one or more steam generators is above the low-low level trip point at the time of trip then the result will be a steam generator minimum water level higher than shown and an increased margin to the point at which reactor coolant water relief would occur. - 14.1.8.3 Conclusions Results of the analysis show that a loss of normal feedwater does not_ adversely affect the core, the Reactor Coolant System, or the steam system since the auxiliary feedwater capacity is such that the reactor coolant water is not relieved from the pressurizer relief or safety valves, and the water level in the steam generators receiving feedwater is maintained above the tube sheets. 14.1.9 Excessive Heat Removal Due to Feedwater System j Malfunctions 14.1.9.1 Identification of Causes and Accident Description Reductions in feedwater temperature or additions of excessive feedwater are means of increasing core power above full power. Such transients are attenuated by the thermal capacity of the secondary plant and of the Reactor Coolant System. The overpower

                              - overtemperature protection (neutron power, overtemperature and overpower 4T trips) te a ONOR 16 7 than 1.30.prevents/tSu        any/E /" M L/10 a/?dnpower increa e which couJd lcad saFefy adysis accephece vilem'a.                                                                                              of 7ht DA/df One example of excess heat removal from the primary system is the transient associated with the accidental opening of the feedwater bypass valve which diverts flow around the low pressure feedvs.er heaters. In the event of an accidental opening of the 1ypass valve, there is a sudden reduction in feedwater inlet temperature to the steam generators.                   The increased sub-cooling will create a greater load demand on the Reactor Coolant System.                                                                                                                                                 j Another 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 Reactor Coolant System due to    increased    subcooling               in the steam                       generator.                                                                            With the 14.1-28

_ _______ _____ -- _ _ ___ _ ____ - -- - O

BVPS-1-UPDATED FSAR Rev. 0 (1/82) on the reactor. Steam generator level rises until the feedwater is terminated as a result of the high-high steam generator level trip. Since the generator fill rate is slow, before the trip occurs the plant reaches equilibrium power conditions commensurate with the increased gepctor );hermal loac) . The DNB ratic doc 5 not drop bslow 1.0. S & Iy a a lysc accefraace c,^jfeni Conclusions a tt men Results show that the consequences of excess load increases due to opening the low pressure heater bypass valve are more moderate than those considered for the Excessive Load Increase Accident. Additionally, it has been shown that the reactivity insertion rate which occurs at no load following excessive feedwater addition is less than the maximum value considered in the analysis of the rod withdrawal from a subcritical condition. Also, the DNB ratios encountered for excessive feedwater addi3 ion at power ar ucli abcvc th; l'miting v 'uc of 1.00. d./4 W 274 DN6tt sa au ly,5 is DCC Ljohc e cpikh , 14.1.10 Excessive Load Increase Incident 14.1.10.1 Identification of Causes and Accident Description An excessive load increase incident is defined as a rapid increase in the steam flow that causes a power mismatch between the reactor core power and the steam generator load demand. The RCS is designed to accommodate a 10 percent step load increase or a 5 percent per minute ramp load increase in the range of 15 to , 100 percent full power. Any loading rate in excess of these I values may cause a reactor trip actuated by the reactor protection system. 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. During power operation, steam dump to the condenser is controlled by reactor coolant conditions signals; i.e., high reactor coolant temperature indicates a need for steam dump. A single controller malfunction does not cause steam dump; an interlock is provided which blocks the opening of the valves unless a large turbine load decrease or a turbine trip has occurred. Protection against an excessive load increase accident is provided by the following reactor protection system signals:

1. Overpower AT l 2. Overtemperature AT

! 3. Power range high neutron flux. l l l 14.1-31 m

BVPS-1-UPDATED FSAR Rev. 0 (1/32) 14.1.10.2 Analysis of Effects and Consequences I Method of Analysis This accident is analyzed using the LOFTRAN code. The code simulates the neutron kinetics, Reactor Coolant System, pressurizer, pressurizer relief and safety valves, pressurizer spray, steam generator, and steam generator safety valves. The code computes pertinent plant variables including temperatures, pressure, and power level. 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. Manually controlled reactor at beginning-of-life

2. Manually controlled reactor at end-of-life '

3. Reactvr in automatic control at beginning-of-life

4. Reactor in automatic control at end-of-life.

At beginning of life the core has the least negative moderator temperature coefficient of reactivity and therefore the least inherent transient capability. At end of life the moderator temperature coefficient of reactivity has its highest absolute value. This results in the largest amount of reactivity feedback due to changes in coolant temperature. A conservative limit on the turbine valve opening is assumed and all cases are studied without credit being taken for pressurizer heaters. Initial operating conditions are assumed at extreme values consistent with the steady state full power operation allowing for calibration and instrument errors. This results in minimum margin to core DNB at the start of the accident. Results Figures 14.1-33 through 14.1-36 illustrate the transient with the reactor in the manual control mode. As expected, for the beginning of life case there is a slight power increase, and the average temperature shows a large decrease. This results in a DNBR which increases slightly above its initial value. For the end of life manually controlled case, there is a much larger increase in reactor power due to the moderator feedback. A eduction eve /,fte in t DNBR is experienced, y g gafej, ana/vsjs but ON'"1ce accef5 rem;in&above Cfi It aos e 1.30.b we Figures 14.1-37 through 14.1-4d illustrate the transient assuming the reactor is in automatic control mode. Both the beginning of li'fe and end of life cases show that core power increases, thereby reducing the rate of decrease in the coolant average temperature and pressurizer pressure. For both the beginning of life and end of life cases, the minin - ON0n remaina abcVe 1.30. e Q^ S}S a cc efa.rc e. 14.1-32

BVPS-1-UPDATED FSAR Rev. 0 (1/82) 14.1.10.3 conclusions ' It has been _4_4_.._

mme demonstrated 2..4._ a- ... that for an excessive _4... ..4,, __m <- <_,___ load increase the mt_ ,2_2m _, 55'bi8[ s a k S5 Ybbef b C5lt55"d5 A 5 14.1.11 Loss of Offsite Power to the Station Auxiliaries (Station Blackout)

                                               ,14.1.11.1                       Identification of Causes and Accident Description In the event of a complete loss of offsite power and a turbine trip           there will be a loss of power to the plant auxiliaries, i.e.,            the reactor coolant pumps, condensate pumps, etc.

The events following a loss of a-c power with turbine and reactor trip are described in the sequence listed below: .

1. Plant vital instruments are supplied by emergency power sources.

2. As the steam system pressure rises following the trip, the steam system power operated relief valves are automatically opened to the atmosphere. Steam dump to the condenser is assumed not to be available. If the I steam flow rate through the power relief valves is not available, the steam generator self-actuated safety valves may lift to dissipate the sensible heat of the fuel and coolant plus the residual heat produced in the reactor.

3. As the no load temperature is approached, the steam system power relief valves (or the self-actuated safety valves, if the power relief valves are not available) are used to dissipate the residual heat and to maintain the plant at the hot shutdown condition.

4. The emergency diesel generators started on Icss of voltage on the plant emergency buses begin to supply plant vital loads.

The auxiliary feedwater system is started automatically as discussed in the loss of normal feedwater analysis. The steam driven auxiliary feedwater pump utilizes steam from the secondary system and exhausts to the atmosphere. The motor driven auxiliary feedwater pumps are supplied by power from the diesel generators. The pumps take suction directly from the primary

                                                . plant demineralized water storage tank for delivery to the steam generators.                  The auxiliary feedwater system ensures a feedwater supply of at least 340 gpm upon loss of power to the station auxiliaries, since the steam driven auxiliary feedwater pump' has a capacity of 700 gpm and the motor driven auxiliary feedwater pumps have a normal capacity of 350 gpm each.

14.1-33

BVPS-1-UPDATED FSAR Rev. 0 (1/82) Upon the loss of power to the reactor coolant pumps, coolant flow necessary for core cooling and the removal of residual heat is maintained by natural circulation in the reactor coolant loops. 14.1.11.2 Analysis of Effects and Consequences Method of Analysis l A detailed analysis using the BLKOUT A code is done to obtain the natural circulation flow following a station blackout. The simulation describes the plant thermal kinetics, Reactor Coolant System including the natural circulation, pressurizer, steam generators and feedwater system. The digital program computes pertinent variables including the steam generator level, pressurizer water level, and reactor coolant average temperature. The first few seconds of the transient will closely resemble a simulation of the complete loss of flow incident (Section 14.2.9), i.e., core damage due to rapidly increasing core temperatures is prevented by promptly tripping the reactor. After the reactor trip, stored and residual heat must be removed to prevent damage to either the RCS or the core. The natural circulation flow as a function of residual reactor power is presented in Table 14.1-3. Conclusions Results of the " Complete Loss of Forced Reactor Coolant Flow" analysis (Section 14.2.9) and the " Loss of Normal Feedwater" analysis (Section 14.1.8) show that for a loss of all a-c power no adverse conditions occur in the reactor core. The DN B R -+e-- Ir.aintained ;bevc 1.2^. The RCS is not overpressurized and no water relief will occur through the pressurizer relief or safety valves. Thus there will be no claddin dam =ge and no release fission products to the RCS. , f 14.1.12 Turbine-Generator Accidents Ja4eh a"d&sts d,/[fe// d are M6f; aCCef ence-The consequences of a turbine generator failure in which missiles are generated have been reviewed. Westinghouse turbine-generator units have never experienced a major structural failure of a I rotating part that resulted in missile-like pieces leaving the turbine casing (See Appendix 14A) . The design, construction, and factory test procedures ensure sound rotating disks that equal or exceed specified design requirements. The standard Westinghouse control system includes three separate speed sensors mounted on the turbine stub shaft located in the turbine front pedestal as follows:

1. Mechanical overspeed trip weight (spring loaded bolt) i 2. Electro-magnetic pickup for main speed governing i channel I

14.1-34

j BVPS-1-UPDATED FSAR Rev. 0 (1/82) I I pressurizer, pressurizer relief and safety valves, pressurizer l spray, steam generator, and steam generator safety valves. The code computes pertinent plant variables including temperatures, pressures, and power level. In calculating the DNBR the following conservative assumptions are made:

1. Initial conditions of maximum core power and reactor coolant temperatures and minimum reactor coolant pressure resulting in the minimum initial margin to DNB (See Appendix 14D).

2. A zero' moderator coefficient of reactivity conservative for beginning of life operation in order to provide a conservatively low amount of negative reactivity feedback due to changes in moderator temperature. The spatial effect of void due to local or subcooled boiling is not considered in the analysis with respect to reactivity feedback or core power shape.

3. 'A high (absolute value) Doppler coefficient of reactivity such that the resultant amount of positive feedback is conservatively high in order to retard any power decrease due to moderator reactivity feedback.

It should also be noted that in the analysis power peaking factors are kept constant at the design values while, in fact, the core feedback effects would result in cons.iderable flattening of the power distribution. This would significantly increase the calculated DNBR; however, no credit is taken for this effect. Results Figure 14.1-44 illustrates the flux transient following the accident. Reactor trip on overtemperature AT occurs as shown in Figure 14.1-44. The pressure decay transient following the accident is given in Figure 14.1-45. The rc;ultir-go;; belov 1. 2 0 := ch=r ir Figure 11.1 16.lWOR Sah ^!!On n/c/crdeta ySU 14.1.15.3 Conclusions ac C efita e C IN**/!7 a/C Me$ The pressurizer low pressure and the overtemperature AT reactor protection system signals provide adequate protection against j this accident, and the minicur O!!On reasin; in cxecs: cf 1.20. Thus there will be no ladding d mage pnd no relepse of fipsion products to the RCS. DAldk SaQ analy ti.s accefranc e c,* ire /id aM me.n 14.1.16 Spurious Operation of the Safety Injection System at Power 14.1.16.1 Identification of Causes Spurious Safety Injection System (SIS) operation at power could be caused by operator error or a false electrical actuating 14.1-41 ___-_______A

BVPS-1-UPDATED FSAR Rav. 0 (1/82) The sequence of events is shown in Table 14.2-2. J M- Ti- te Critic 1 " Ort Fl'"' f D"" ---1 7 e15 "se prfer cd for th: th::: ::::: :::t criti el te 0"". It =: f:=d that 211 :_::: h:d : =i..ir = 0"!? ;recter the 1.00. Radiological Consequences The source' of radiation discharged to the environment from a steam line rupture presumes a primary to secondary system leak. The absence of such a leak precludes the discharge of any radioactive material to the environment from the secondary system. For purposes of analyzing the environmental consequences of a hypothesized steam line rupture, it is assumed that a 10 gpm primary to secondary system leak rate exists simultaneously with equilibrium reactor coolant activity due to 1 percent failed fuel. This model is very conservative since continued base load operation with a 10 gpm primary to secondary leak rate with

1 percent failed fuel is unlikely; however, the activity values a used reflect continued power operation. The halogen activity, -

which is assumed to be released during the steam line break, includes a reduction factor of 10 resulting from expected plateout. i Two mechanisme of secondary coolant release to the environment are investigated. The first case considers a main steam line rupture upstream of the main stream trip and nonreturn valves (see Figure 10.3-1). The second case presumes a main steam line break downstream of the main steam trip valve and external to the leak collection structure. For this case, the activity release is for 5 seconds duration which is the time before the main steam line trip valve may be closed. In the calculations, it has been conservatively assumed that the secondary side pressure drops almost instantaneously to atmospheric pressure and the primary system pressure decreases linearly to ::*350 psia from 2,250 psia in 4 hours. This corresponds tp.a cooldown rate on the order of 50 F per hour. At this time, the residual heat removal system is placed in operation to continue cooling the reactor coolant system. The reactor coolant system pressure during this period is assumed to be maintained constant at 350 psia. The residual heat removal system is capable of cooling the reactor coolant system to less ) than 200 F in 4 hours. Eight hours following the accident the reactor coolant pumps can be shut off and the reactor coolant ) l , 14.2-24 e

i BVPS-1-UPDATED FSAR Rev. 0 (1/82) low-low level insertion monitors with visual and audio signals. Operating instructions require boration at low level alarm and emergency boration at the low-low alarm. Reactor Protection The reactor protection in the event of a rod ejection accident has been described in Reference 3. The protection for this accident is provided by the power range high neutron flux trip (high and low setting) and high rate of neutron flux increase trip. These protection functions are described in detail in Section 7.2. Effects on Adjacent Housings Disregarding the remote possibility of the occurrence of a rod cluster control assembly mechanism housing failure, investigations have shown that failure of a housing due to either longitudinal or circumferential cracking is not expected to cause damage to adjacent housings leading to increased severity of the initial accident. A control rod drive mechanism assembly is shown in Figure 3.2-17. The operating coil stack assembly of this mechanism has a 10.719-inch by 10.718-inch cross section and 39.875 inch length. The position indicator coil-stack assembly (not shown) is located above the operating coil stack assembly. It surrounds the rod travel housing over nearly its entire 163.25 inch length. The rod travel housing outside diameter is 3.75 inches and the position indicator coil stack assembly inside and outside diameters are 3.75 inches and 7.0 inches respectively. This assembly consists of a steel tube surrounded by a continuous stack of copper wire coils. The assembly is held together by two end plates, an outer sleeve and four axial tie rods. 1%etT @ g ,-+I f a longitudinal failure of the rod travel housing should occur, the region of the position indicator assembly opposite the break would be stressed by the reactor coolant pressure of 2,250 psia. The most probable leakage path would be provided by the radial deformation of the position indicator coil assembly, resulting in the growth of axial flow passages between the rod travel housing and the steel tube. If failure of the position indicator coil assembly should occur, the resulting free radial jet from the failed housing could cause it to bend and contact adjacent rod housings. If the adjacent housings were on the periphery, the might bend outward from l their bases. The housing material /yis quite ductiler plastic trvhinging without cracking would be expected. Housings adjacent to a f ailed housing , in locations other than the periphery, would not be bent because of the rigidity of multiple adjacent housings. 14.2-32

BVPS-1-UPDATED FSAR Rev. 0 (1/82)

           , ilJSE't.T @

E If circumferential failure of a rod travel housing should occur, the broken-off section of the housing would be ejected vertically because the driving force is vertical and the position indicator coil stack assembly and the drive shaft would tend to guide the broken-off piece upwards during its travel. Travel is limited by the missile shield, thereby limiting the projectile acceleration. When the projectile reached the missile shield it would partially penetrate the shield and dissipate its kinetic energy. The water jet from the break would continue to push the broken-off piece l against the missile shield. l If the broken-off piece of the rod travel housing were short enough to clear the break when fully ejected, it would rebound after impact with the missile shield. The top and plates of the position indicacion coil stack assemblies would prevent the l broken piece from directly hitting the rod travel housing of a i second drive mechanism. Even if a direct hit by the rebounding piece were to occur, the low kinetic , energy of the rebounding l -+ projectile would not be expected to cause significant damage. g From l hisER.T the a @ bove discussion, the probability of damage to an l adjacent housing must be considered remote. However, even if damage is postulated, it cannot lead to a more severe transient since RCCA's are inserted in the core in symmetric patterns, and control rods immediately adjacent to worst ejected rods are not in the core when the reactor is critical. Damage to an adjacent housing could, at worst, cause the RCCA not to fall on receiving a trip signal; however, this is already taken into account in the analysis by assuming a stuck rod adjacent to the ejected rod. The considerations given, above lead to the conclusion that failure of a control rod housing, due either to longitudinal or circumferential cracking, would not cause damage to adjacent housings that would increase the severity of the initial accident. , i 14.2.6.1.2 Limiting Criteria wwtT @ [Due to the extremely low probability of a rod cluster control gg assembly ejection accident, some fuel damage could be considered an acceptable consequence. Comprehensive studies of the threshold of fuel failure and of the l threshold of significant conversion of the fuel thermal energy to mechanical energy, have been carried out as part of the SPERT project by the Idaho Nuclear Corporation.IO Extensive tests of UO zirconium clad fuel rods representative of those in pressurized water reactor type cores have demonstrated failure thresholds in the range of 240 to 257 cal /gm. However, other rods of a slightly different design have exhibited failures as low as 225 cal /gm. These results differ significantly from the TREAT R results, which indicated a failure threshold of 14.2-33

BVPS-1-UPDATED FSAR Rev. 4 (1/86) 3.4.3.2.2 Axial Heat Flux Distributions As discussed in Section 3.3.2.2.4, the axial heat flux distribution can vary as a result of rod motion, power change, or due to spatial xenon transients which may occur in the axial direction. Consequently, it is necessary to measure the axial power imbalance by means of the ex-core nuclear detectors (as discussed in Section 3.3.2.2.7) and protect the core from excessive axial power imbalance. The Reactor Trip System provides automatic reduction of the trip setpoint in the overtemperature aT channels on excessive axial power imbalance; that is, when an extremely large axial offset corresponds to an axial shape which could lead to a DNBR which is less than that calculated for the reference DNB design axial shape. The reference DNB design axial shape used in this amendment is a chopped cosine shape with a peak average value of 1.55. The use of a 1.55 cosine instead of the 1.48 cosine 4eced in the origincl EvrO-1 cubmitt:1 (15 x 15 design) results in increased operating flexibility. There are fewer axial power shapes which give DNBRs less than the DNBR for a 1.55 cosine than there are shapes which give DNBRs less than the DNBR for a 1.48 cosine. Thus, greater axial power imbalances can be allowed when the reference DNB design axial shape is a 1.55 cosine. 3.4.3.3 Core Thermal Response A general summary of the steady-state thermal-hydraulic design parameters including thermal output, flow rates, etc., is provided in Table 3.4-1 for all loops in operation, cnd in Tabic 3.4 2 for cperation trith-cr.c coolant iccp cut of cervice. As stated in Section 3.4.1, the design bases of the application are to prevent departure from nucleate boiling and to prevent fuel melting for condition I and II events. The protective systems described in Section 7 (Instrumentation and Controls) are designed to meet these bases. The response of the core to Condition II transients is given in Section 14. 3.4.3.4 Analytical Techniques 3.4.3.4.1 Core Analysis The objective of reactor core thermal design is to determine the maximum heat removal capability in all flow subchannels and show that the core safety limits, as presented in the Techniedl Specifications, are not exceeded while compounding engineering and nuclear effects. The thermal design takes into account local variations in dimensions, power generation, flow redistribution, and mi.x! ng. THINC-IV is a realistic three-dimensional matrix model which l's been developed to account for hydraulic and nuclear effects on the enthalpy rise in the 3.4-31 l l l j

i BVPS-1-UPDATED FSAR Rev. 4 (1/86) thermal aspects such items as clad creep, fuel swelling, fission gas release, release of absorbed gases, cladding corrosion and elastic deflection, and helium solubility. A detailed description of the thermal model can be found in Reference 76 with the modi'ications for time dependent densification given in Reference 80. 3.4.3.4.3 Hydrodynamic Instability The analytical methods used to access hydraulic instability are discussed in Section 3.4.3.5. 3.4.3.5 Hydrodynamic and Flow Power Coupled Instability i In steady state, two-phase, heated flow, a potential for flow instability in parallel closed channels exists. Although such a 4 potential may not exist in the Westinghouse open lattice array core, it has been evaluated on a conservative basis. The instability may , be a flow excursion from one state to another, or it may be a self-sustained oscillation about one state ('") . Either type is undesirable in a nuclear reactor. First, sustained flow oscillations may cause undesirable forced mechanical vibration of components. Second, flow oscillation may cause system control problems by varying the moderator coefficient. Third, it has been f ound( 5 5 )( 55 ) that during flow oscillations the critical heat flux necessary for DNB may be considerably lower than in steady flow. When the flow channels of a reactor core having common inlet and outlet plenums operate hydraulically in parallel, all have the same pressure drop. An instability in one or a few of the channels where boiling occurs does not significantly change the overall pressure drop because the flow is redistributed to a large number of other stable channels, and the condition of constant imposed pressure drop is satisfied. This type of oscillation is thermohydrodynamic and can best be understood by realizing that a boiling channel constitutes a time-varying, spatial distributed parameter system. In a two-phase flow, the hydrodynamic coupling with a potential for positive feedback may lead to sustained oscillations having sizable amplitudes. Thus, a temporary reduction in the inlet flow rate to a boiling channel will increase the rate of evaporation, thereby raising the average void fraction. This disturbance affects the elevation, acceleration, and frictional pressure drop as well as the heat transfer behavior. For certain conditions of channel geometry, thermal properties of the heated wall, flow rate, inlet enthalpy, heat flux, etc., resonance may occur and sustained oscillations then result. The flow instability of a group of parallel flow channel 5having i l common plenums at the inlet and exit has been investigated analytically. Westinghouse has developed the HYDNA(571 digital computer program for predicting the hydrodynamic stability of parallel closed channels. To verify the capability of HyDNA to 3.4-37

BVPS-1-UPDATED FSAR Rev. 4 (1/86) I I rod internal pressures do not exceed the nominal coolant pressure even at the overpower condition (Section 3.2.1.1.1). The potential effects of operation with waterlogged fuel are discussed in Section 3.4.3.6 which concluded that waterlogging is not a concern during operational transients. Clad flattening, as noted in Section 3.2.1.3.1, has been observed in come operating power reactors. Thermal expansion (axial) of the fuel rod stack against a flattened section of clad could cause failure of the clad. This is no longer a concern because clad flattening is precluded during the fuel residence in the core (see section 3.2.1.3.1). There can be a differential thermal expansion between the fuel rods cnd the guide thimbles during a transient. Excessive bowing of the fuel rods could occur if the grid assemblies did not allow axial movement of the fuel rods relative to the grids. Thermal expansion of the fuel rods is considered in the grid design so that axial loads imposed on the fuel rods during a thermal transient will not result in excessively bowed fuel rods (see Section 3.2.1.2.2). 3.4.3.8 ' Energy Release During Fuel Element Burnout Ts discussed in Section 3.4.3.3 the core is protected from going chrough DNB over the full range of possible operating conditions.-At-full poucr acminal operation, the minimum DNDR wee fennd to be 1.00. In rigura 3.4-7, it can be ;cca that for theec condition;, th; probability of a red going through DNE is ic;; then 0.1 pcrcent at 05 percent confidence level ba;;d on the statistics of the 'J-3 ccrrelatien. In the extremely unlikely event that DNB should occur, the clad temperature will rise due to the steam blanketing at the rod surface and the consequent degradation in heat transfer. During this time there is a potential for a chemical reaction between the cladding and the coolant. However, because of the relatively good film boiling heat transfer following DNB, the energy release resulting from this reaction is insignificant compared to the power produced by the fuel. DNB With Physical Burnout Westinghouse (58) has conducted DNB tests in a 25-rod bundle where physical burnout occurred with one rod. After this occurrence, the 25 rod test section was used for several days to obtain more DNB data from the other rods in the bundle. The burnout and deformation of the rod did not affect the performance of neighboring rods in the test section during the burnout or the validity of the subsequent DNB data points as predicted by the W-3 correlation. No occurrences of flow instability or other abnormal operation were observed. 3.4-40 l _ __ __ ___ _ _____.. ___ _____ _ _ _ _ _ _

BVPS-1-UPDATED FSAR Rev. 4 (1/86) DNB With Return to Nucleate Boiling Additional DNB tests have been conducted by Westinghouse (78) in 19 and 21 rod bundles. In these tests, DNB without physical burnout was experienced more than once on single rods in the bundles for short periods of time. Each time, a reduction in power of approximately 10 percent was sufficient to reestablish nucleate boiling on the surface of the rod. During these and subsequent tests, no adverse effects were observed on this rod or any other rod in the bundle as a consequence of operating in DNB. 3.4.3.9 Energy Release or Rupture of Waterlogged Fuel Elements A full discussion of waterlogging including energy release is contained in Section 3.4.3.6. It is noted that the resulting energy release is not expected to affect neighboring fuel rods. 3.4.3.10 Fuel Rod Behavior Effects from Coolant Flow Blockage Coolant flow blockages can occur within the coolant channels of a fuel assembly or external to the reactor core. The effects of fuel assembly blockage within the assembly on fuel rod behavior is more pronounced than external blockages of the sane magnitude. In both cases the flow blockages cause local reductions in coolant flow. The amount of local flow reduction, where it occurs in the reactor, and how far along the flow stream the reduction persists are considerations which will influence the fuel rod behavior. The effects of coolant flow blockages in terms of maintaining rated core performance are determined both by analytical and experimental methods. The experimental data are usually used to augment analytical tools such as computer programs similar to the THINC-IV program. Inspection of the DNB correlation (Section 3.4.2.3 and Reference 39) shows that the predicted DNBR is dependent upon the local values of quality and mass velocity. The THINC-IV code is capable of predicting the effects of local flow blockages on DNBR within the fuel assembly on a subchannel basis, regardless of where the flow blockage occurs. In Reference 54, it is shown that for a fuel assembly similar to the Westinghouse design, THINC-IV accurately predicts the flow distribution within the fuel assembly when the inlet nozzle is completely blocked. Full recovery of the flow was found to occur about 30 inches downstream of the blockage. With the reference reactor operating at the nominal full ' power conditions specified in Table 3.4-1, the effects of an increase in enthalpy and decrease in mass velocity in the lower portion of the fuel asse@ly woyld not result in the reactor reaching a minimum DNBR cf 1.30. balod the safefy Ikif Va/ae. From a review of the open literature it is concluded that flow blockage in "open lattice cores" similar to the Westinghouse cores cause flow perturbations which are local to the blockage. For instance, A. Oktsubo, et al . ( 7 0 show that the mean bundle velocity is approached asymptotically about 4 inches downstream from 3.4-41

BVPS-1-UPDATED FSAR Rev. 4 (1/86) i l I l 4 a flow blockage in a single flow cell. Similar rcaults were also found for 2 and 3 cells completely blocked. Basmer, et al(75), tested an open lattice fuel assembly in which 41 percent of the subchannels were completely blocked in the center of the test bundle between spacer grids. Their results show the stagnant zone behind the flow blockage essentially disappears after 1.65 L/De or about 5 inches for their test bundle. They also found that leakage flow through the blockage tended to shorten the stagnant zone or in essence the complete recovery length. Thus, local flow blockages within a fuel assembly have little effect on subchannel enthalpy rise. The reduction in local mass velocity is then the main parameter which affects the DNBR. If the BVPS 1 plant were operating at full power and nominal steady state conditions as specified in Table 3 4-1, a reduction in local mass velocity greater than 70 percent would be required to reduce the DNBR frcr '.35 to 1.30. The above mass velocity effect on the DNB correlati was based on the essumption of fully developed flow along the c annel length. In reality, a local flow blockage is expected to pro te turbulence and thus would likely not affect DNBR at all. - beSaft ltmb, coolant flow blockages induce local crossflows as well as promote turbulence. Fuel rod behavior is changed under the influence of a sufficiently high crossflow component. Fuel rod vibration could occur, caused by this crossflow component, through vortex shedding or turbulent mechanisms. If the crossflow velocity exceeds the limit 1stablished for fluidelastic stability, large amplitude whirling results. The limits for a controlled vib ation mechanism are established from studies of vortex shedding and turbulent pressure fluctuations. Crossflow velocity above the established limits can lead to mechanical wear of the fuel rods at the grid suppert locations. Fuel rod wear due to flow induced vibration is considered in the fuel rod fretting evaluation (Section 3.2). 3.4.4 Testing and Verification 3.4.4.1 Tests Prior to Initial Criticality A reactor coolant flow test, as noted in Item 8 of Table 13.1-2, is performed following fuel loading but prior to initial criticality. Coolant Loop pressure drop data is obtained in this test. This data 1 in conjunction with coolant pump performance information allows I determination of the coolant flow rates at reactor operating conditions. This test verifies that proper coolant flow rates have been used in the core thermal and hydraulic analysis. 3.4.4.2 Initial Power and Plant Operation core power distribution measurements are made at several core power levels (see Section 3.3.2.2.7). These tests are used to insure that l conservative peaking factors are used in the core thermal and I hydraulic analysis. 3.4-42

BVPS-1-UPDATED PSAR Rev. 0 (1/82) Average Core Analysis The spatial kinetics computer code TWINKLE IM- (described in Section 14D.10.7) is used for the average core transient ; analysis. This code solves the two group neutron diffusion i theory kinetic equations in one, two or three spatial dimensions (rectangular coordinates) for six delayed neutron groups and up to 2,000 spatial points. The computer code includes a detailed multiregion, transient fuel-clad-coolant heat transfer model, for I calculationgdpointwise Doppler and moderator feedback effects. i esVIn this analysis, the code is used primarily as a one dimensional axial kinetics code since it allows a more realistic representation of the spatial effects of axial moderator feedback and rod cluster control assembly movement and the elimination of axial feedback weighting factors. However, since the radial dimension is missing, it is still necessary to employ very 1 conservative methods (described below) of calculating the ejected , rod worth and hot channel factor. . I Hot Spot Analysis The average core energy addition, calculated as described above, is. multiplied by the appropriate hot channel factors, and the hot spot analysis is performed using a detailed fuel and clad transient heat transfer computer code, FACTRAN IO. This computer code calculates the transient temperature distribution in a cross section of a metal clad 00 fuel rod, and the heat flux at the surface of the rod, using as input the nuclear power versus time and the local coolant conditions. The local coolant conditions before the start of the transient are input to the code. This consists of the system pressure, flow rate and bulk coolant temperature and corresponding density appropriate to the power level. Due to the excellent heat transfer, the fuel and clad temperature is not particularly sensitive to these parameters. The zirconium-water reaction is explicitly represented, and all material properties are represented as functions of temperature. A parabolic radial power distribution is used within the fuel rod. . FACTRAN uses the Dittus-Boelter or Jens-Lottes correlation to i determine the film heat transfer before DNB, and the Bishop-Sandburg-Tong correlationfO to determine the film boiling coefficient after DNB. The Dittus-Boetter correlation is used to compute the breed convection heat transfer coefficient before DNB. At the same time, the Jens-Lottes correlation is used to compute the surface temperature to support local boiling. The code automatically switches to this mode of heat transfer when the surface temperature reaches the local boiling surface temperature. The 14.2-35

BVPS-1-UPDATED FSAR Rev. 0 (1/82) Moderator and Doppler coefficient The critical boron concentrations at the beginning of life and end of life were adjusted in the nuclear code in order to obtain moderator density coefficient curves which are conservative compared to actual design conditions for the plant. As discussed above, no wzighting factor is applied to these results. The Doppler reactivity defect is determined as a function of power level using the one dimensional steady state computer code with a Doppler weighting factcr of 1.0. The resulting curve is conservative compared to design predictions for this plant. The Doppler weighting factor should be larger than 1.0, just to make the present calculation agree with design predictions before ejection. This weighting factor will increase under accident conditions, as discussed above. The transient weighting factor used in the analysis is presented in Table 14.2-3. Delayed Neutron Fraction Calculations of the effective delayed neutron fraction (S ) typically yield values of .70 percent at beginning of life *Nd 0.50 percent at and of life for the first cycle. The accident is o sensitive toA if the ejected rod worth is nearly equal to or Pet greater thanias in zero power transients. In ord +n - allow for 20.5'5 ruture fuel cycles, pessimistic estimates of g of percent at l

                 <V beginning of cycle and .4/ percent at end of cyc                                                                                                               were used in          ,

l the analysis. 7 Trip Reactivity Insertion lThetripreactivityinsertionisassumedtobe percent from hot REV full power and 2 percent from hot zero power including the effect of one stuck rod. These values are reduced by the ejected rod reactivity. The shutdown reactivity was simulated by dropping a rod of the required depth into the core. The start of rod motion occurred 0.5 seconds after the high neutron flux trip point is reached. This delay is assumed to consist of 0.2 second for the , instrument channel to produce a signal, 0.15 second for the trip ' breaker to open and 0.15 second for the coil to release the rods. 2.7 analyses presented are applicable for a rod insertion time of

                          .      seconds from _ coi { release to entrance to the dash pot.                                                        ,                                   ,,
                        . . = - _ , . . _ _ _ _ _ _ . . . . . . . . . . . . . ==.                                             . . . .....                                     ...w u u        u. uv..t
                        *e-1,4-seconde..                                                 The choice of such a conservative insertion rate ROJ      means that there is over 1 second after the trip point is reached                                                                                                                         l before significant shutdown reactivity is inserted into the core.

This is. a particularly significant conservatism for hot full power accidents. The rod insertion versus time is described in Section 14D.5. 14.2-38

                                                                             ~
                                                                                                                                                                       )

l BVPS-1-UPDATED FSAR Rev. 0 (1/82) 1 14.2.6.2.2 Results The values of the parameters used in the analysis, as well as the results of the analysis, are presented in Table 14.2-3 and discussed below. Beginning of Cycle, Full Power Control bank D was assumed to be inserted to its insertion limit. garThe worst ejected rod worth and hot channel factor were .20 tak and 7.pFl respecH u=1v. The peak i-t ;;;; clad average 255"1 '_ temperature was F. The peak hot spot fuel center ature excseu.0 e beginning of life melt temperature of qqoo F. However, melting was restricted to less than 10% of the et. FEV Beginning of Cycle, Eero Power and - For this t n, ,,po.n. trol bank D was assumed to be fully

            ' inserted an                    at E W insertion limits. The wor =* ajected rod 39 is located in control b                         D and has a worth of                                                            tak and a hot channel                      of                 The peak :.et ;;;; cla                                                   temperature REV reached only w p at 4uen 2?

1818 o 60.0 u4empwahre wu INS *F.

ag End of Cycle, Full Power 0.2 % Control bank D was assumed to be inserted to its inse limit. g T ejected rod worth and hot channel factors were tok and

# 8

_ _ ._ ."_ 5."_ .Y', .I *_

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The peak hot ' g resulted in a peak eiggemperature of F. g b ge spot fuel 4 temperature excI'e'ded the end ife melt 4,800 F. However, melting was restricted to less than 10% of the pellet. The variation in melt temperature with burnup is discussed in Section 3.4.1.2. EEV End of Cycle, Zero Power iO Ebrd The ejected rod worth and hot channel factor for t is case we ed and ban C obt at in& insertiond assuming limits. control Thebank D to bewere results fully i tak l and a r 21.'75 respectively. "e .. ce- , t!.e .;11 eie -ee J u. i.u -e

_::tir: 7:12 :, ----ly 0.00 U.% e.. r ir I t ^- i. . The peak clad g and fuel center temperatures were and .

RE\l aveny. '291I oni49 A summary of the cases presented above is given in Table 14.2-3. The nuclear power and hot spot fuel and clad temperature transients for the worst cases (BOL full power and EOL zero l power) are presented in Figures 14.2-11 through 14.2-)d. is on Product Release 7Ev It is assumed that fission products are released from the gaps of all rods entering DNB. In all cases considered, less than 14.2-39 ____-____________- _ _ _ _ a

BVPS-1-UPDATED FSAR Rev. 0 (1/82) TABLE 14.2-2 (CONT'D) TIME SEQUENCE OF EVENTS FOR CONDITION IV EVENTS Time (Sec.)

                                                                      -Accident                                               Event                 3-Loop Rupture of main                     Feedline rupture occurs                                                    O feedwater pipe High pressure reactor                                                      9.5 trip setpoint reached (This trip was not considered in the analysis).

Affected steem generator 17.5 liquid discharger low level coincident with feed / steam flow mismatch in other steam generators; reactor trip setpoints reached. Reactor trip occurs 19.5 Peak steam relief from 20.0 pressurizer safety valves Presurizer fills 416 Bulk boiling begins in 744 reactor coolant fluid Core decay heat decreases 2055 to auxiliary'feedwater heat removal capacity IMSEIRT l Rev l 2 of 2

BVPS-1 UPDATED FSAR 14.2.6 Spectrum of RCCA Ejection Accidents Inserts for Zire Grids and Increased Peaking Factors Insert A: Effect of Rod Travel Housina Longitudinal Failgrgi Insert B: Effect of Rod Travel Housino Circumferential failures Insert C: Possible Consequences Insert D: This event is classified as an ANS Condition IV incident. Insert E: The calculated sequence of events for the rod ejection accidents, as shown in Figures 14.2-11 through 14.2-14, are presented in Table 14.2-2. For all cases, reactor trip occurs very early in the transient, after which the nuclear power excursion is terminated. The reactor will remain subcritical following reactor trip. The ejection of an RCCA constitutes a break in the RCS, located in the reactor pressure vessel head. The effects and consequences of loss-of-coolant accidents (LOCAs) are discussed in FSAR Section 14.3. Following the RCCA ejection, the operator would follow the same emergency instruction as for any other LOCA to recover from the event. i . . . .

BVPS-1 UPDATED FSAR 14.2.6 Spectrum of RCCA Ejection Accidents Inserts for Zirc Grids and Increased Peaking Factors Insert F: TABLE 14.2-2(cont.) TIME SE00ENCE OF EVENTS Accident Eygni Time (sec) ' Rod Cluster Control Assembly Ejection

1. Beginning-of-Life, Initiation of Rod Ejection 0.0 Full Power Power Range High Neutron Flux Setpoint Reached 0.05 Peak Nuclear Power Occurs 0.14 Rods Begin to Fall 0.55 Peak Fuel Average Temperature Occurs 2.38 Peak Clad Average Temperature Occurs 2.44 Peak Heat Flux Occurs 2.47

2. End-of-Life, Initiation of Rod Ejection 0.0 Zero Power Power Range High Neutron Flux Low Setpoint Reached 0.15 Peak Nuclear Power Occurs 0.17 Rods Begin to Fall 0.65 Peak Clad Average Temperature Occurs 1.04 Peak Heat Flux Occurs 1.04 Peak Fuel Average Temperature Occurs 1.59

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g80 - w 8 v 60 - 20 - l I I I i o 0 0.5 1.0 1.5 2.0 .5 10.0 TIME (SECONDS) FIGURE 14 2-11 NUCLEAR POWER TRANSIENT BOL HFP ROD EJECTION ACCIDENT SE AVER VALLEY POWER STATION UNIT NO. 1 UPDATED FINAL SAF(TV ANALYS15 REPORT

                                                                                                                                    . _ _ _ _ _ _ _ _ _ - -_ _ _ _ - _ .      __-_=__=__._ _ _ _ -

4 l 4 Figure 14 2.-11 j Nuclear Power Transient BOL-HFP Ejection Accident l 3.0 2.5-2.0-5 g i.s. b 8 z 1.0-4

                          .S0-0M u i 2   3       4      $      6 7           6                6    to i                                            TIWE    (SECONDS)

                                                             ,                                                            REV. 0 (1/82) 6000 MELTING EL CENTER TEMPERATURE I

4000 - I FUEL AVG PERATURE 5 $ w l g 3000 - ( l 2 Og I CLAD TEMPERATURE 2000 - L 1000 I I I 0 O l.0 2.0 3.0 4.0 10.0 TIE (SECONDS) FI6URE 14 2-12 HOT SPOT FUEL AND CLAD TEMPERATURE VS TIME, BOL HFP R0D EJECTION ACCIDENT BEAvf4 V ALLE Y POW E R 5f Afl0N UNIT NO. 1 UPDATED FINAL SAFETY ANAlf5th REPOAT

k ( Figure 14.2.- 12. Peak Fuel and Clad Average Temperature verpus Time BOL-HFP Rod E}ection Accident sees. I ELTING TEMPERATURE 4900 Fl s... . _ _ _ _ k____________ l FUEL CENTER TEMPERATURE-4994. - i A w seee. . E IFUEL AVERACE TEMPERATURE l af 2996. - 180@a ' l CLAD OUTER TEMPERATURE 8.

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io 0 .5 1.0 1.5 2.0 2.5 10.0 Tl* (SECONOS) FIGURE 14 2-13 NUCLEAR POWER TRANSIENT E0L HIP ROD EJECTION ACCIDENT BE AVER VALLEY POWER STAT!0h UNIT NO. 1 UPDATED FINAL 5AFETY ANALY5IS REP 0at _-________-____-___?_____-___-_______--_____-

Figure 14.2.-15 Nuclear PowerTransient EOL-HZP Rod Ejection Accident IC' ; 4 101 -- 5 8 c g 100 . b 8 z 10-i-

                   'IDo     .60  1.0      1.5      2.0     2.5 3.0 3.5 4.0 TIME    (SECONDS) e mes

l REV. 0 (1/82) 6000 M ~ MELil_N G 48000F e0 - FUEL TER TEMPERATURE FUEL AVC TEMPER UAE s

                                                                                                              *a                                                                                              N i

CLAD TEMPERATURE 2000 . - i (>b ' 1000 - i 0 0 1.0 2.0 3.0 0 10.0 TIME (SECONDS) FIGURE 14 2-14 HOT SPOT FUEL AND CLAD TEMPERATURE VS. TIME, E0L HZP R0D EJECTION ACCIDENT BE AVER V ALLEY POW ER ST ATION UNIT NO. 1 UPDATED FINAL $ AFETY AN ALYSIS REPORT _._.-__m_ __. - . . _ _ -

                                                                                                                - . _ _ _ _ _ _ _ _ _ - _ . _ . _ _ . _ - _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _                            _J

l Figure m.2.-M Peak Fuel and Clad Average Temperature versus Tlrne EOL-HZP Rod Ejec4on Accident 6000. l ElilNG TEMPERATURE 4800 F1 ! 5000. _ _________________d _ _ _ _ _ _ _ _ _ l l FUEL CENTER TEMPERATURE 4000. - d a W 3000. - f 2000. - 1000. 1 FUEL AVERAGE TEMPERATURE l [ d ' lCLAD OUTER TEWPERATUREl 0. O. 1. 2. 3. 4. 5. 6. 7. 3. 9. 10, TIME (SECORDS)

M 1 BVPS-1-UPDATED FSAR Rsv. 0 (1/82) 4 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 primary circuit. The analyses have demonstrated that upper limit on fission product release as a result of a number of fuel rods entering DNB amounts to 10 percent. 14.2.7 Single Reactor t, ' ant Pump Locked Rotor 14.2.7.1 Identification of Causes and Accident Description The accident postulated is an instantaneous seizure of a reactor coolant pump rotor. Flow through the a'ffected 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 l 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, 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 reduced heat transfer in the steam generators causes an insurge into the pressurizer and a pressure increase throughout the Reactor Coolant System. 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 three 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 5.n the analysis, efff3gg7 g 14&2.7.2 Analysis of Effects and Consequences 14.2.7.2.1 Method of Analysis digital-computer codes are used to analyze this trans . The PH "I"I code is used to calculate the resulti oop and core coolant following the pump seizure. LOFTRAN illI code is then used to ulate the time o ctor trip, based on the flow calculated by , nuclear power followir. reactor trip, and to determi ak pressure. The thermal behavior of the fuel lo at the core ot is investigated using the FACT code, using the core alculated by PHOENIX and nuclear power calculated by LOFTRAN. ACTRAN udes the use of a film boiling heat tran code ficient. RC4L ke un1/ Dsser B 14.2-41

                                                                               .LVPS-1-UPDATED FSAR                                                               R2v. 0 (1/82) ases are analyzed:

[

1. All 1 erating, one locked ro 3

w@ 2. Two out of three operating with no isolated r> loops, locked r a3 j 3. ut of three loops operating with solated

                                ,,                                  loop, one locked rotor.

At the beginning of the postulated locked rotor accident, i.e., at the time the shaft in one of the reactor coolant pumps is assumed to seize, the plant is assumed to be in operation under the most adverse steady state operating conditions zith :::;::t madam,,t; th; ;;;in te ;;;, i.e., maximum steady state power level, e6m6mnam- steady state pressure and maximum steady state coolant average temperature. With all but ene lwp eimreting, the

                                                       ==4   x pes;; 1Erel (including errere) elle..J in e.i. wa. vi
                                                       ;;;;;ti;; i; :.;;T- M.

msv When the peak pressure is evaluated, the initial pressure is conservatively estimated as 30 psi above nominal pressure (2,250 psia) to allow for errors in the pressurizer pressure measurement and control channels. This is done to obtain the highest possible rise in the coolant pressure during the transient. To obtain the maximum pressure in the primary side, conservatively l pressure. high loop pressure drops are The pressure responseh added toshown the calculatedin Figures pressurizer 14.2-J8% csythrough 14.2-17 are the responses at the point in the Reactor Coolant System having the. maximum pressure. The maximum pressure in the reactor coolant system following a locked rotor accident occurs at the pump discharge of pumps which have not experienced the locked rotor. Evaluation of the Pressure Transient After pump seizure, the neutron flux is rapidly reduced by control rod insertion effect. Rod motion is assumed to begin one second after the flow in the affected loop reaches 87 percent of nominal flow. The time delay of 1.0 second used in connection with the low flow reactor trip is a very conservative allowance for the total time delay between the time the flow reaches 87 percent of nominal and the time the rods begin moving into the core. This total includes individual delays associated with the following: flow sensors / transmitters, solid state protection system input relays, solid state protection system, voltage drop on reactor trip breaker undervoltage and control rod gripper release. No credit is taken for the pressure reducing effect of the pressurizer relief values, pressurizer spray, steam dump or controlled feedwater flow after plant trip. 14.2-42

BVPS-1-UPDATED FSAR Rev. 0 (1/82) Although these operations are expected to occur and would result in a lower peak pressure, an additional degree of conservatism is provided by ignoring their effect. The pressurizer safety valves are actuated at 2,575 psia and their capacity for steam relief is as stated. Evaluation of the Effects of DNB in the Core During the Accident For this accident, DNB is assumed to occur in the core and, therefore, an evaluation of the consequences with respect to fuel rod t,hermal transients is performed. Results obtained from analysis of this " hot spot" condition represent the upper limit with respect to clad temperature and zirconium water reaction. 2.544 In the evaluation, the r power at the hot spot is conservatively assumed to be times the average rod power (i.e., F(Q) = at the initial core power level. 2.5 % N Film Boiling Coefficient The film boiling coefficient is calculated in the FACTRAN code using the Bishop-Sandberg-Tong film boiling correlation. The fluid properties are evaluated at film temperature (average between wall and bulk temperatures). The program calculates the film coefficient at every time step based upon the actual heat transfer conditions at the time. The neutron flux, system pressure, bulk density and mass flow rate as a function of time are used as program input. For this analysis, the initial values of the pressure and the bulk density are used throughout the transient since they are the most conservative with respect to clad temperature response. For conservatism, DNB was assumed to start at the beginning of the accident. Fuel Clad Gap Coefficient The magnitude and time dependence of the heat transfer coefficient between fuel and cladding (gap coefficient) has a pronounced influence on the thermal results. The larger the value of the gap coefficient, the more heat is transferred between pellet and clad. Based on investigations on the effect of the gap coefficient upon the maximum clad temperature during the transient, the gay coefficient was assumed to increase from steady state value consistent with the initial fuel temperature to 10,000 Btu /hr ft 8 F at the initiation of the transient. This assumption causes energy stored in the fuel te be released to the clad of the initiation of the transient and maximizes the clad

                                                                                                                ' ~ -

t'empe'YhTUFt t ring the transient. ) i 14.2-43

I BVPS-1-UPDATED FSAR rov. 0 (1/82) Zi'rconium-Steam Reaction i The zirconium-steam reaction can become significant above 1,300 F (clad temperature). The Baker-Just parabolic rate equation shown below is used to define the rate of the zirconium steam reaction. d - = 33.3 x 10 8 exp (14.2-3) where: w = amount reacted (mg/cm 8) t = time (sec) T = temperature (K) The reaction heat is 1,510 cal /gm gev 14.2.7.2.2 Locked Rotor Results sient values of reactor vessel flow coastdown, neutron fl , and channel heat flux are shown in the following figur .

                                                                          @                                                Number of In operating loops                          Figure Nos e                                              3                                14     8, 14.2-19
                                                                              ~

2 2 (no isolated loop) .2-20, 14.2-21 b 3 2 (isolated loop) 14.2-22, .2-23 W Maximum rea coolant system pressure, maximum c temperature and amo of zirconium-water reaction are containe Table

14. . Figure 14.2-24 shows the clad temperature transis or N e worst case.

REV Conclusions

1. Since the peak Reactor Coolant System pressure reached during any of the transients is less than that which would cause stresses to exceed the faulted condition stress limits, the integrity of the primary coolant system is not endangered.

2. Since the peak clad surface temperature calculated for the hot spot during the worst transient remains considerably less than. 2,700 F and the - amount of Zirconium-water reaction is small and the core will remain in place and intact with no consequential loss of core cooling capability.

14.2-44

l 1\! g p n o i o t~ L ay rld ele 5 , pat 2

      )                 Oia      7,       *

2 tl 8 sio pns 2

1 oII ( o L e 0 n ' O g

         .                                                  n v         S                                          i e         T                                          t R          N                                           a E                                          r I       g     s                             e S       n     p                             p N      i                                   o A      t R       ay                                 s T       rl                                  p eld     0                           o R       pae     8                           o O      Oit      6,       *

l R T ta A O sil l S R pno l F oIs a 4 D o I D - E L r E 2 K o o T . C 2 N f 1 A 4 O D 1 L n f P a o U E R h

        -    L   O                                          t          1 1      B   F          g
        -    A              n                                r S      T   S         i                                 e P          T         t                                 w V          L          ay                               o B          U          rl                      6       l S          el E          pa   5         1        0        y R          Oi   7         3                l t 6,        0,               l F          si pg a

O 2 2 i o t Y o n R L a A t M s M b U u S S 0 r ) ) ao osa tei t F (t Hp S cts asp d o aep Eth r ey( lrS og RS Cu fHi e tt o e ntr u sao t ew n nu urH tr aas ils i pe me noy uCb x e aar xmr aeo o mt4 MCP MTC Aa( V E R Q ! f3 OjNg llllll lli'

BVPS-1 UPDATED FSAR 14.2.7 Single Reactor Coolant Pump Locked Rotor Inserts for Zirc Grids and Increased Peaking Factors Insert A: This event is classified as an ANS Condition IV incident. Insert B: Two digital computer codes are used to analyze this transient. The LOFTRAN Code (Reference 11) is used to calculate the resulting loop and core flow transients following the pump seizure, the time of reactor trip based on the loop flow transients, the nuclear power following reactor trip, and to determine the peak pressure. The thermal behavior of the fuel located at the core hot spot is investigated using the FACTRAN Code (Reference 8), which uses the core flow and nuclear power calculated by LOFTRAN. The FACTRAN Code includes a film boling heat transfer coefficient, insert C: The effect of zirconium-steam reaction is included in the calculation of the " hot spot" cladding temperature transient. Insert D: The transient results with and without offsite power available are shown in Figures 14.2-15 through 14.2-18. The results of these calculations are also summarized in Table 14.2-4. The peak Reactor Coolant System pressure reached during the transient is less than that which would cause stresses to exceed the faulted condition stress limits. Also, the peak clad surface temperature is considerably less than 2700 degrees F. It should also be noted that the clad temperature was conservatively calculated assuming that DNB occurs at the initiation of the transient. The calculated sequence of events is shown on Table 14.2-4. Figure 14.2-15 shows that the core flow rapidly reaches a new equilibrium value for the case with offsite power available. With the reactor tripped, a stable plant will eventually be attained. Normal plant shutdown may then proceed. I -m__._ ._. _ _ _ _ . ._._.____ _ . . _ _ _ _ . _ _ . . - _ _ _ _ _ _ _ _

           .                                             BVPS-1 UPDATED FSAR 14.2.7 Single Reactor Coolant Pump. Locked Rotor Inserts for Zire Grids and Increased Peaking Factors l                            Insert E:

TABLE 14.2-4

SUMMARY

OF RESULTS FOR LOCKED ROTOR TRANSIENTS 3 Loops Operating 3LoopsOperat!ing (with offsite power) (without offsite oower) Maximum Reactor Coolant Pressure (psia) 2597 2642 l Maximum Cladding Temp-erature at Core Hot Spot, (*F) 1795 1870 Zirc-Water Reaction at Core Hot Spot (weight %) 0.269 0.415 4

BVPS-1 UPDATED FSAR 14.2.7 Single Reactor Coolant Pump Locked Rotor Inserts for Zirc Grids and Increased Peaking Factors Insert E (cont.): TABLE 14.2-4 (cont.) TIME SE00ENCE OF EVENTS l Accident Event Time (sec) Locked Rotor 3 Loops Operating (with offsite power)

                                                                               -w 1 RCP Rotor Locks                                0.0~

Low Flow Trip Point Reached 0.06 Rods Begin to Drop 1.06 Maximum RCS Pressure Occurs 2.8 Peak Clad Temperature Occurs 3.6 Locked Rotor-3 Loops Operating (without offsite power) 1 RCP Rotor Locks 0.0 Low Flow Trip Point Reached 0.06 Rods Begin to Drop 1.06 RCPs Lose Power, Coastdown Begins 1.06 Maximum RCS Pressure Occurs 3.2 Peak Clad Temperature Occurs 3.9

    ~                ^

l REV. 0 (1/82) 2s00 2700 - s 2 00 _ .c I, Y - _ 2s00 I o s e - g2400 2 00 - 1 1 I I 2 oo O 2.0 4.0 6.0 8.0 10.0 TIE (SECONOS) i O h . vP SG FI6URE 14 2 15 ALL LOOPS OPERATIN6, ONE LOCKED ROTOR , PRESSURE VS TIME l EEAVER VALLEY POWER STATION UNIT 20. 1

UPDATED FINI4L 5AFiff ANAlf515 REPORT

REV. o (1/82) t 2eco I 2700 - ! 2600 - 7 z; t w A

                                  = 2500    -

E E E 2soo - 2300 - I I I I 2200 0 2 4 6 5 10 TIE ($EcoNDS) y- FIGURE 19 2-16 AV TWO OUT OF THAEE 1,00F3 CPERATING, GNE o bV/ LOCKED ROTOR (NO LOOPS ISOLATED) PRESSURE VS TIME BEAVER VALLEY POWER STATION UN!T NO, 1

                                                              .                  UPDATLD FINAL 5AFEff ANALYSIS REP 0kT

l. i '

REV. 0 (1/82) seco I m - l l

soo -

E g - l l 5 aco - I 5 1 nao - - t 2soo - I I i 1

                                                'O                                         2     4               6               8           10 TIME (SECOS8)

I&

                                                       '7                                                           FIGUR8 14 2-17 TVO 00T OF THREE LOOPS OPERATlW6, ONE Or                                                                                              LOCKED ROTOR PRESSURE VS TIME (ONE LOOP ISOLATED)

BE AY A VALLf r POWEA STATICW UN!T WO.1 HPDATED Fll.AL SAFEff ANAlf5!$ AEP0AT

[4 m REV. 0 (1/82)

1.2 1.0 -

                                   **7 5

b 5 .8 l 5 G .6 - W E g .. - w a .2 - l i l 0 0 2 4 6 8 10 TIME (SECONOS) 1 l l h' C/ J l FIGURE 162*18 ALL LOOPS OPERA!!NG, ONE LOCKED ROTOR CORE FLOW VS TIME BF. AVER VALLEY POWER STATION UNIT N0. 1 UPDATED FINAL SAF(TV ANALY5t$ aEPORT l

1 l i REV. 0 (1/82) i i l I l.2 s.o 9

                             .8    -

HEAT FLUX 7 l 5 5 .s - W E l l t E.. - NEUTRON FLUX

                             .2    -

I I ! I o 0 2 s 6 8 10 l TIE (SECONOS) b p FIGilRE 14 2-19 Q ALL LOOPS OPFAATlii6, 06lE LOCKED ROTOR FLUX TRAMSIENTS Y$ TIME SE AVEK V ALLEY POWER STATION UNIT NO. 1

                                           .                  UPDATED FINAL SAFETV ANALYS!5 REPORT
                                                                                                       ]

REV. 0 (1/82) l.2 , 1 _ l.0 - t i: m

                                     .s    -

m s j.s - e l 5

                                     .2  -

I I I I I 0 0 2.5 5.0 7.5 10.0 12.5 15.0 TIE (SEcou05) l gv' v F16URE 14 2-20 TWO QUT OF THREE LOOPS OPEPATlWG, g OuE LGCKED LOTOR (N0 ISOLATED 1.00P3) CORE FLOW VS ilME SE AFER V ALLEJ PGEER STAT!0N UNIT NO., 1

                                                          .                                                   UPDATED f!hAL SAFETF AKALV5f5 tir0LT

___________ _ _ _____ _ _ _ __ - - \

fREV. 0 (1/ t

                                                                                                                     'l . 2 1.0       -
                                                                                                       '.I g 0.8 m

50.6 EAT FLUX w [ 0.4 -- NEUT20e FLUX

                   .                                                                                                0.2          -

o l I I I 0 2 4 6 s 10 TIME (SECONOS)

j p 1 Q FIGURE 14 2-21 TWO OUT OF THREE LOOPS OPERAil#G, t OeE t.0CKED ROTOR (WO IS01).TED t03P) I Fl.UX TRANSIENT VS TIME BEAVER VALLEY POWik 17ATlom UNIT h0. 1

                                                                                                                                                                         .                             UPDATED FINAL $AFEff ANALYS!$ REPORT l

   -                                                                                                         J l
     ,                                                                                   REV. 0 (1/82) i t                                                                                  .

1.2 , 7 1.0 - - 5 t

       =                                                                                                     I

{ '0.s - 5 g0.6 - 5. h0.4 - t 0.2 - l I I I o 0 2 4 6 8' 10 TIE (SECONOS) 6 h FIGURE 14 2-22 TWO OUT OF THREE LOOPS OPERAilb6, OuE l LOCKEB ROTOR (1 150 LATED LOOP) CORE FLOW V5 TIME BEi.VER f ALLEY P'W ER STATION UNIT NO. 1

                             .                                    UPDATED r!NAL SAFETY ANALY315 REPORT

REV. 0 (1/82) 0.8 0.7 - L . 7 0.6 - E_ l 0.5 - s NEAT FLUX 30.4 - atuTaon Flux G 0.3 - 6

                                                  =
                                                                    ~
                                                  $ 0.2    -

0.1 - 0' I ! I I 0 2.0 4.0 6.0 8.0 10.0 Tile (SECONOS) I FIGURE 14 2 23 TWO DUT 47 1HREE LOOPS OPEkATING, ONE i LrX. KCL ROTOR (ONE LOOP ISOLATED r FLi!X TRANSIENTS VS TIME BE AVER VALLEY POWER $7Afl0N UNIT No.1

                                                                           .                                                                                                               UPDATED FINAL SAFETY I.NALY$15 REPJ9T

i REV. 0 (1/82) I l 2000 - 1800 - 1600 C f1400 4 i d 1200 - . 1000 . . 800 - i I i 1 0 2.0 40 6.0 8.0 10.0 TIE (SEcoll0S) FIGURE 14 2-24

                                                                                         /v                                                ALL LOOPS OPERAfil!6, ONE LOCKED ROTOR                                                             :

CLAD TEMPERATURE VS TIME BE AVER V ALLEY POW ER ST ATION UNIT NO. 1

                                                                                                              .                            UPDATED FINAL SAFETY ANALY$!$ REPORT v

Figure 14.E-15 Flow Transients for Three Loops in Operation, One Locked Rotor

                                                        ,.4 5.2-OFF l */O"gEeTg*pgwgm
                                                                                                                                  * * - --       -I

_1 . o - E5-3 . so. tM

w - - - - - - - - - - - - - - - - - - - - - ~ ~ ~ - - - - - - -

E E . co-E_E 40-I

                                                        . 2 0-
                                                          *%     ,            a       >     <       a                 a       1      .         ,      ,a TIWC      (SECONDS)
                                                          ...                                                                                            .~

I F8,FETie'5e:2: 5" _ _ _'l ag , eE k=.. l \ DE g- _N l , l

                                                           ** o  ,            .       .     .        .                 .       7               .

YlWE (SECONDB)

I j 4 Figure t%-16 l Reactor Coolant System Pressure Transient for Three Loops in Operation One Locked Rotor 2700 W/0 0FFSITE POWER W OFFSITE POWER - - - - 2600- , ~,,

               /               %
                                                         \

2500- s

                                                             \
                                                               \

m g I \ ,

5 c 2400- %

                                                                       \
                                                                                                               )

- s - LLI g W 2300- g

M N e o.

                                                                                     's ---

2200 ___,, M (.) Ct: ( 2100-W 2000-1900-u i i 3 4 6 6 1 s e 10 TIME (SECONDS)

Figure 14.f.-17 Nuclear Power Transient, Average and Hot Channel Host Flux Transients for Three Loops in Operation, One Locked Rotor 1.. 1.2-

1. O <

g E . .O. E S . eo-E . .O.

                                                         .20<

t_

                                                         .CO   W                                  3               *
                                                                                             '           3             *

7 e 10 YIME CsEcONOs) 1..

1. 2 a g 1. 0 -

m EE E . .O. EE..O< ..O.

                                                        .30-
                                                        .OO  W                             '     3      3        *    *

w 7 e O YsME CsEcONOs) 1..

1. 2 -

1. O -

5 ,u = o - m m E...O. .O.

                                                       .30
                                                       .00,                              ,      ,      ,       ,     ,
                                                                                                                             ,   ,       ,   y  o TfME    CsEc@NCS)

Figure 14.2-2 { Maximum Clad and Fuel Centerline Temperatures at Hot Spot for Three Loops in Operation One Locked Rotor s... L 30 0 3..O. %g*

                 ? ,....
                              ,,,,.                                                                                                                             I"'8,M7Fe'5o:e:55 _ _ _I
                              ' *a o       .            .        .                             .                                                .                    . 7     .        .      ,.

T1ME (SECONOS) 3000

                                                                                                                                                                   ~~~.

g.. ' g, - . s g .. . g..

                              ,,,.                                                                                                                             IV8,KTie'5e:t:55 - - _l t

ib 1 3 3 e . . 7 . . to T ut (sccONO3)

BVPS-1-UPDATED FSAR Rev. 0 (1/82) confirm that resulting power distribution effects will either be readily detected by the in-core moveable detector system or will cause a. sufficiently small perturbation to be acceptable within the uncertainties allowed between nominal and design power shapes. i 14.2.9 Complete Loss of Forced Reactor Coolant Flow 14.2.9.1 Accident Description A complete loss of forced reactor coolant flow 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 ofJoct of loss-of-coolant flow is a rapid increase g in the coolant temperature. This increase could result in DNB , with subsequent fuel damage if the reactor were not tripped promptly.@ The following provide necessary protection against a loss-of-coolant flow accident:

1. Undervoltage or underfrequency on reactor coolant pump power supply busses ;

2. Low reactor coolant loop flow. !

The reactor trip on reactor coolant pump bus undervoltage is provided to protect against conditions whic h can cause a loss of l voltage to all reactor coolant pumps, i.e., station blackout. This function is blocked below approximately 10 percent power (Permissive P-7) . I The reactor trip on reactor coolant pump underfrequency is provided to open the reactor coolant pump breaker and trip the reactor for an underfrequency condition, resulting from frequency disturbances on the major power grid. The trip disengages the reactor coolant pumps from the power grid so that the pump kinetic energy is available for full coastdown. mE*T @ tuv The reactor trip on low primary coolant loop flow is provided to

                 -                                                                        protect against loss of flow conditions which affect only one reactor coolant loop.      It also serves as a backup to the undervoltage   and underfrequency                                                           trips.      This function is generated by two out of three low flow signals per reactor coolant loop. Above approximately 30 percent power (Permissive 8), low flow in any loop will actuate a reactor trip. Between approximately    10   percent             power                                                    and 30   percent    power (Permissive 7 and Permissive 8), low flow in any two loops will actuate a reactor trip.

Normal power for the reactor coolant pumps is supplied through busses from a transformer connected to the generator. Each pump 14.2-47 i

s BVPS-1-UPDATED FSAR Rsv. 0 (1/82) is on a separate bus. When generator trip occurs, the busses are automatically' transferred to a transformer supplied from external power lines, and the pumps will continue to supply coolant flow to the core. Following any turbine trip, where there are no electrical faults which require tripping the generator from the network, the generator remains connected to the network for approximately 30 seconds. The reactor coolant pumps remain connected to the generator thus ensuring full flow for 30 seconds after the reactor trip before any transfer is made. 14.2.9.2 Method of Analysis s transient is analyzed by four digital-computer codes. Fir the HOENIXf1'l code is used to calculate the loop and core ow durin the transient. The LOFTRAN fill code is then d to the time of reactor trip, based on the flows culated calculat by PHOENIX, and Ithe nuclear power transient foll ategthe reactor trip. The F RAN 'I code is then used to cal heat flux transient ed on the nuclear power from FTRAN and flow from PHOENIX. Fin ly the THINC code is to calculate the minimum DNBR during a transient based the heat flux from FACTRAN and flow from HOENIX. The grid spacer factor is @ applied to the W-3 cor ations. The transients presented represent the minimum of the ic or thimble cell. h 3 The following 3 cases have an zed:

%               Loop                                   P                Loop y           initially                                coastin            stop 5           operating                                  down           valves 3                                     3               e g

d 2 opn 2 2 , 2 closed esvThe method of analysis and the assumptions made regarding initial operating conditions and reactivity coefficients are identical to those discussed in Section 14.1, except that following the loss of supply to all pumps at power, a reactor trip is actuated by either bus undervoltage or bus underfrequency. ated sequence of events is shown on Table 14.2-b the three analyzed. Figures 14.2~30 thro . - 8 show reactor vesse owns, the nuclear Q 3y the loop coastdown , power coastdowns, and the and hot channel heat flux e reactor is assumed three cas 5g coastdowns for each._ The DNB for each of gg to trip g M dervoltage signal. sa M cases is not less than 1.30. Rev 14.2-48

g' l BVPS-1-UPDATED FSAR Rev. 0 (1/82) l b 14.2.9.3 Conclusions sh- r -'_ymis parfonned has demonstrated that for *ha :: .-l.se loss of forced re L:r -lant flow- *M LA- ces not decrease or 1Fbelow 1.30 durina +h t.=---TantT--I S"- there is no clad damage r:1=-.Yfission products to the Reacthoy-+==. ~ nsv 14.2.10 Single accA Withdrawal at Full Power 14.2.10.1 Accident Description No single electrical or mechanical failure in the rod control system could cause the accidental withdrawal of a single rod cluster control assembly from the inserted bank at full power operation. N operator could deliberately withdraw a single rod cluster control assembly in the control bank. This feature is necessary in order to retrieve an assembly should one ha accidentally dropped. In the extremely unlikely event of simultaneous electrical failures which could result in single rod cluster control assembly withdrawal, rod deviation and rod control urgent failure would be displayed on the plant annunciator, and the rod position indicators would indicate the relative positions of the assemblies in the bank. N urgent failure alarm also inhibits automatic rod motion in the group in which is occurs. Withdrawal of a single rod cluster control assembly by operator action, whether deliberate or by a combination of errors, would result in activation of the same alarm and the same visual indications. Each bank of rod cluster control assemblies in the system is divided into two groups of four mechanisms each. h rods comprising a group operate in parallel through multiplexing thyristors. N two 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 and deactuation of the stationary gripper, movable gripper and lift coils of a mechanism is required to withdraw the rod cluster control assembly attached to the mechanism. Since the four stationary gripper, FGVable gripper and lift ooils associated with the four rod cluster control assemblies of a rod group are driven in ' parallel, any single failure which would cause rod withdrawal would affect a minimum of one group, or four rod cluster control assemblies. Mechanical failurer are in the direction of insertion, or immobility. In the unlikely event of multiple failures which result in continuous withdrawk1 of a single rod cluster control assembly, it is not possible, in all cases, to provide assurances of automatic reactor trip such that core safety limits are not violated. Withdrawal of a single rod cluster control assembly results in both positive reactivity insertion tending to increase core power, and an increase in local power density in the core area " covered" by the rod cluster control assembly. I 14.2-49 l._-...._ - -- -

m a-,----,--,.- - - , - - , - ,- e-,,-,.- 34 BVPS-1-UPDATED FSAR Rev. 0 (1/82) 14.2.10.2 Method of Analysis Power distributions within the core are calculated by the TURTLE (12I code based on macroscopic cross section generated by LEOPARD II O . The peaking factors calculated by TURTLE are then used by THINC to calculate the minimum DNB for the event. The case analyzed was the worst rod withdrawn from bank D inserted at the insertion limit, with the reactor initially at full power. F(6H) for this case was 1.71 including appropriate allowances for calculational uncertainties. 14.2.10.3 Results Two cases have been considered as follows:

1. If the reactor is in the manual control mode, continuous withdrawal of a single rod cluster con _ trol assembly 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 failed rod cluster control assembly. In terms of the overall system response, this case is similar to those presented in Section 14.1.7; however, the increased local power peaking in the area of the withdrawn rod cluster control assembly results in lower minimum DNBR's than for the withdrawn bank cases. Depending on initial bank insertion and location of the withdrawn rod cluster control assembly, automatic reactor trip may not occur sufficiently fast to prevent _ the minimum core DNB ratio from falling below 1. 30. ' r; valuation o this case at the power and coolant conditions at which the overtemperature AT trip would be expected to trip the plant shows that an upper limit for the number of rods within a DNBR less than +r-SO- 5 er n. ,

2. If the reactor is in automatic control mode, w thdrawal of a single rod cluster control assembly will result in the immobility of the other rod cluster control assemblies in the controlling bank. The transient will then proceed in the same manner as Case 1 delscribed above. For such cases as above, a trip will ultimately ensue, although not sufficiently fast in al"1 cases to prevent : minim- ^

ratic in 'h ccrc

                                                   -ef 1:::  t.'.= 1. 3 0 . vs/ots oF.A=hvor sa Q caetyds 14.2.10.4           Conclusions sccefuce,UIlenits For the case of one rod cluster control assembly fully withdrawn with the reactor in the automatic or manual control taode and initially operating at full power with Bank D at the insertion limit, an upper bound of the number of fuel rods experiencing DNBR < 1.0 is 5 percent of the total fuel rods in the core.

fl

                                            % des & IM

14. 2-:50

k BVPS-1-UPDATED FSAR Rev. 0 (1/82) For both cases discussed, the indicators and alarms mentioned would function to alert the operator to the malfunction before DNB could occur. For case 1 discussed above the insertion limit alarms (low and low-low alarms) would also serve in this regard. 14.2.11 Minor Secondary System Pipe Breaks 14.2.11.1 Identification of causes and Accident Description l Included in this grouping are ruptures of secondary system lines which would result in steam release rates- equivalent to a 6 inch diameter break or smaller. 14.2.11.2 Analysis of Effects and Consequences Minor secondary system pipe breaks must be accommodated with the failure of only a small fraction of the fuel elements in the reactor. Since the results of analysis presented in Section 14.2.5 for a major secondary system pipe rupture also meet this criteria, separate analysis for minor secondary system pipe breaks is not required. The analysis of the more probable accidental opening of a secondary system steam dump, relief or safety valve is presented , in Section 14.1.13. These analyses are illustrative of a pipe ' break equivalent in size to a single valve opening. 14.2.11.3 Conclusions The analysis presented in Section 14.2.5 demonstrates that the consequences of a minor secondary system pipe break are acceptable since a DNBR of less than -1rS- does not occur even for a more critical major secondary system ipe break. f/ce /es Ya /4 E-14.2-51

~ _ _ _ _ _ _ .

BVPS-1-UPDATED FSAR Rev. 0 (1/82) References to Section 14.2 (Cont'd)

12. S. Altomare and R. F. Barry, "The TURTLE 24.0 Diffusion Depletion Code," WCAP-7758, Westinghouse Electric Corporation (September 1971).

13. R. F. Barry, LEOPARD - A Spectrum Dependent Non-Spatial Depletion Code for the IBM-7094." WCAP-3269-26, Westinghouse Electric Corporation (September 1963) .

14. " RADIOISOTOPE, A Computer Program for Calculating Residual Activities in a Closed System After One or More Decay Periods," RP-1, Stone & Webster Engineering Corporation (November 1972).

15. " ACTIVITY A computer Program for Calculating Fission Product Activity in Fuel, Coolant, and Selected Tanks for a Nuclear Power Plant," RP-3, Stone & Webster Engineering Corporation (January 1973).

16. "IONEXCHANGER, A Computer Program for Determining Gamma Activities in Ion Exchangers or Tanks as a Function of Time for Constant Feed Activity," RP-2, Stone & Webster Engineering Corporation (December 1972).

NSEltT F REV i i 14.2-53

                                                                                                      ~      ~

- - _ _ _ _ _ - - _ _ _ _ _ _ - - _ - - - - _ - _ - __________-__~~

BVPS-1-UPDATED FSAR Rev. 0 (1/82) TABLE 14.2-5 T EQUENCE OF EVENTS FOR CONDITION III EVENTS T (seconds) Accident Event 3 Loop

1. Complete Loss of Coastdown beg s 0 Forced Reactor Qh Coolant Flow All Loops Operat: ng, Moti Begins 1.50 pr All Pumps Coasting Min DNBR Occurs 2.85 W Down

2. Two out of Three Coastdown Beg 0 3

 !          Pumps Operating, Two Pumps Coast            Rod Motion Begins                  1.50 m           Down, Icop St ja          valves Ope                 Min 4= = DNBR Occurs               2.

g 3. Two of Three Coastdown Begins 0 g Operating, o Pumps Coasting Rod Motion Begins 1.50 Down, Loop Stop Valves Closed Minimum DNBR Occurs 2.35 REN 1 of 1

BVPS-1 UPDATED FSAR 14.2.9 Complete Loss of Forced Reactor Coolant Flow Inserts for Zirc Grids and Increased Peaking Factors Insert A: This event is classified as an ANS Condition III incident. Insert B: If the maximum grid frequency decay rate is less than approximately 5 Hz/sec, this trip function will protect the core from underfrequency events without requiring tripping of the RCP breakers. Reference 17 provides analyses of grid frequency disturbances and the resulting nuclear steam supply system protection requirements which are generally applicable. Insert C: 14.2.9.2 Method of Analysis The complete loss of flow transient has been analyzed for a loss of all three reactor coolant pumps with three loops in operation. Insert D: Figures 14.2-30 through 14.2-33 show the transient response for the loss of power to all reactor coolant pumps with three loops in operation. The reactor is assumed to be tripped on an undervoltage signal. Figure 14.2-33 shows that the minimum DNBR is always greater than the safety analysis limit. 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 do not increase significantly above their respective initial values. The calculated sequence of events for the case anlayzed is shown in Table 14.2-5. The reactor coolant pumps will continue to coast down and natural circulation flow will eventually be established. With the reactor tripped, a stable plant condition will eventually be attained. Normal plant shutdown may then proceed. Insert E: 14.2.9.3 Conclusions The analysis performed has demonstrated that for the complete loss of forced reactor coolant flow, the DNBR does not decrease below the limit value at any time during the transient. Thus, the DNB design i basis as described in Section 3 is met. There is no fuel clad damage l or release of fission products to the Reactor Coolant System. Insert F:

17. Baldwin, M.S., Merrian. M.M., Schenkel, H.S., and Van De Waile, D.J., "An Evaluation of Loss of Flow Accidents Caused by Power System Frequency iransients in Westinghouse PWRr," WCAP-8424, Revision 1, June 1975.

                       ,,   e          -+         o     e *      -w~

BVPS-1 UPDATED FSAR 14.2.9 Complete Loss of Forced Reactor Coolant Flow

            . Inserts for Zirc Grids and Increased Peaking Factors Insert G:

TABLE 14.2-5 TIME SE00ENCE OF EVENTS FOR CONDITION III EVENTS Accident Event Time (sec) Complete Loss of Forced Reactor Coolant Flow (3 Loops Operating, 3 RCPs Coasting Down) Coastdown Begins 0.0 RCP Undervoltage Trip Point Reached 0.0 Rods Begin to Drop 1.5 Minimum DNBR Occurs 3.4

                                     ,                , , . .      %   . .-4

                        .                                                                  REV. 0 (1/82) i 1.2 n  l.0 a

0.8 - s 8 g0.6 - 5 f 0.4 - 8 5 0.2 - l l l l 0 2.0 g,o 6.0 3,o 10.0 TIME (SECONDS) l l 1 l FIGURE 14 4-30 ALL 3 LOOPS OPERATING, ALL 3 PUFPS COASTING DOWN, CollE FLOW VS.' TIME , 3(Ay!R fiAl.LEt POWER STATICM UNIT Wo. t j UPDATED r!aAL 5AF(Tr ANALY3!$ REPORT 'j _ )

                                          ~           --                . - . . .                         , _

REV. 0 (1/82) 1 l.2 l.0 -

                                                 '2 1

l 0.s AvEnAct cNANNEL NEAT Flux AND j b NOT CNANNEL NEAT FLUX 3 0.6 - g NUCLEAn FLUX m 0.4 - 0.2 - i I I l n 0 2.0 s.0 6.0 8.0 10.0 TIE (SECONDS) 0 e 4f i c FIGURE 14 2-31 l l ALL 3 LOOPS OPERATING, ALL 3 PUMPS l 40ASTIM DOW, FLUX TRANSIENT VS TIME l sEAvEn VALLEY kW Et ".TATion uNtf no. 1 UPP ATFD FINAL SAFETY AKALYSIS REPORT

<a-am_------------___,__-.-.-----,--aa-----           - - - - - - - - - - - _ - - -

t REV. 0 (1/82) 2.0 l.9 - l.8 - g 1.7 1.C - l.5 -

                                                 !                              I 1.4 O    l .0 .        2.0               3.0             4.0 i                                                       TIE (SEcol105) l sV F160t2 14 2-32 ALL 3 LOOPS OPERATING, ALL 3 PttMPS COASTING DOWN, DMBR VS TIME 5( AVER VALLtf POWER STATIDM Un!T h0. I UPDATED FINAL $AFEff ANtLf!!5 AEPORT

_ = rums. .

REV. 0 (1/82) l.2 l.0 W ^ o.s - C1 C0aE FLOW E~ LOOP FLOW gg 0.e- - gg0 .s - t:. r_. B3 d d 0.2 - 5$

                                                                                                                                               ,        I        I            I              I 0    2.0       s.0         6.0            8.0          10.0 TIE (SECONOS)
                                                                                                                                                    #l FIGURE 14 2-33 TWO MIT OF THREE L30P3 OPERAfirL TWO PURPS COASTINE DNN, Fl.0W CDASTDOWWS VS. TIME (N0 isotATED LOOM 4E AVf!t VALLEY POWER $TATIOR 0N!" N0.1 UPDATED FINAL 5AFiff ANALY11$ REPORT e

g j 1 REV. 0 (1/82) i1 l l.2 l

                                     ' l.0     -

l0.8 i 50.s - g NOT CNANNEL MEAT FLUE w 20.4 - NUCLEAR FLUX AVERACE CMAnnEL 0.2 - NEAT FLUK 0 I I I l 0 2.0 4.0 6.0 g.0 go,o TIME (SEC001DS) 9 M l I FIGURE 14 2-34 TWO OUT OF in2Cf LOOR OPERAililG, TWO PtMPS COASTIF2 DOWN, (LUX TRAMSIENTS VS TIAE (No ISULATED LocrS) EE AVfL VAf. LEY P0utR STATIDN unit no.1 UPDATED 7!P.AL $AFETT AFAlf535 RE7087

REV. 0 (1/82) 2.5 2.4 - - 1 2.3 - 2.2 - 1.1 2.0 - i i

                                   ...                               i 0     1.0        2.0        3.0          4. 0 TIE (SEC0 LIDS)

I h O

FIGURE 14 2-35 TWO OUT OF THREE LOOPS OPERATING, TWO PUMPS COASTING DOWN, (NO ISOLATED LOOPS) DNBR VS TIME BE AVER VALLET POWER STAT!DN UNIT NO. 1 UPDATED FINAL $AFETY ANAlf515 REPORT

                           ~ ~ -                                        --

l , REY. 0 (1/82) i l.2 i 1.0

           ;3a 0.s     -
               -                             LOOP FLOW 3I                   CORE TLOW I
           'o ,* 0. s  -

8s-53

          -m.E 0. 4 88 d d 0.2    -

5$ v s 1 I I l o 0 2.0 4.0 6.0 8.0 10.0 TIE (SECONO3) 8 r F* ""tE 14 2-36 li.0 Oui 0F THREE LOOPS OPERATING, TWO PUMPS COASTING DOWN, FLOW COASTDOWNS VS TIME (I LOOP ISOLATED) BE AVER VALLEV POWER STATION UNIT NO.1 UPDATED FINAL 5AFETY ANALYS15 REPORT

f REY. 0 (1/82) I 1.2 1.0 -

         '3 if
         ! .a n                                                                                                                                                                    :

NOT f.dANNEL g NEAT Flux '

f

             .4    -

g NEUTADN flux 1 m AVERAGE CNANNEL

       .     .2    -

NEAT Flux - 0 I l l l 0 2.0 4.0 6.0 8.0 10.0 TIE (SECONOS) 6h 9 FIGURE 14 2-3, TWO OUT OF THn E LOOPS OPERATING, TWO PURPS COASTING DOWN, FLUX TRANSIENTS VS TIRE (1 LOOP ISOLATED) SE AVER VALLEY POWER STATION UNIT No.1 UPDATED FINAL SAFETT ANALT515 REPORT

t REV. 0 (1/82) 2.5 2.4 - 2.3 - g 2.2 2.1 - 2.0 - I I I i.s 2.0 3.0 0 1.0 4.0 TIE (SECONDS) 66 o FI6dRE142-38 - TWO OUT OF THREE LOOPS OPCPATim6, TWO PURPS C0ASTING DOWN (1 LOOP ISOLATED) DN8R VS. TIME BE AVER VALLET POWER STATION UNIT NO. I UPDATED FINAL $AFETY ANALYSIS REPORT

i Figure W.2.-3o , Core Flow Coastdown versus Time for Three Loops in Operation, Three Pumps Consting Down, Complete Loss of Flow 1.4 1.2-1.0- )

 .m o<

az 3 Jo .30-w z l

M D LL. wo z o .60-NM wm "O

       .40-1
       .20-u    1    2      3     4     6      6          7                                     6 9   to TIWE   (SECONDS)

Figure 94.2.-31 Nuclear Power Transient and Pressurizer Pressure Transient For Ttwee Loops in Operation, ; Three Loops Coasting Down, Complete Loss of Flow 1.4

1. 2 -

1. 0-5 g *B . so.

s g 5. co-E 40-

    . 2 0-
    .00 g           3    g      a      4      3     g      7           g y     go TIME    (SECONDS) 8400 3300*

8800 A 3 ,0c. 12 2000 1500< iOOO i > 1 3 3 4 5 5 7 5 5 to TaMc (sccONOS) l

Figure N.2.-32 Average and Hot Channel Heat Flux Transients for Three Loops in Operation, Three Pumps Consting Down, Complete Loss of Flow 1.4

1. 2 -

1. 0-

_a E R .so-im5.co-BE 40

                                                     .20-
                                                     .00 g    3     2       a      4     o     e    7                                      e           a 10 TIME -(SECONDS) i 1.4

1. 2 -

1. 0-WY gg. O.

m

                                                  $ . sO-al      40-
                                                      . 2 0-
                                                      .OO g    i    a       a       . o     e     f                                         a       w 10 TIME   (SCCONOS)

Figure 114 . 2. - 3 3 l DNBR vs Time for Three Loops in Operation, Three Pumps Consting Down, i Complete Loss of Flow 2,4-2.2-2.0-x 1.8- . E 1.6-1.4-1.2-1.0 0 i 2 3 4 L TlWE (SECONDS)

I I BVPS-1-UPDATED FSAR Rev. 0 (1/82) fuel rod model l21 which has been reviewed and approved by the NRC. 14D.3.3 Power Distribution The transient response of the reactor system is dependent on the initial power distribution. The nuclear design of the reactor core minimizes adverse power distribution through the placement of control rods and operation instruction. The power distribution may be characterized by the radial factor F (AH) and the total peaking factor F (Q) . The peaking factor limits are given in the Technical Specifications. For transients which may be DNB limited the radial peaking factor is of importance. The radial peaking factor increases with decreasing power level due to rod insertion. This increase in F(AH) is included in the core limits illustrated in Figure 14D-1. All transients that may be DNB limited are assumed to begin with a F(Q) consistent with the initial power level defined in the Technical Specifications. The axial power shape used in the DNB calculation is th; 1.55 ch pp:d : in 22 discussed in Section 3.4.3.2.2. REY For transients which may be over power limited the total peaking factor F(Q) is of importance. The value of F (Q) may increase with decreasing power level such that full power hot spot heat flux is not exceeded, i.e. F(Q) times Power = design hot spot heat flux. All transients that may be overpower limited are assumed to begin with a value of F(Q) consistent with the initial power level as defined in th'e. Technical Specifications. The value of peak kW/ft can be directly related to fuel temperature as illustrated on Figures 3.4-1 and 3.4-2. For transients which are slow with respect to the fuel rod thermal time constant: the fuel temperatures are illustrated on Figures 3.4-1 and 3.4-2. For transients which are fast with respect to the fuel rod thermal time constant, for example, rod ejection, a detailed heat transfer calculation is made. 14D.4 TRIP POINTS AND TIME DELAYS TO TRIP ASSUMED IN ACCIDENT ANALYSES A reactor trip signal acts to open two trip breakers connected in series feeding power to the control rod drive mechanisms. The loss of power to the mechanism coils causes the mechanisms to release the RCCA which then fall by gravity into the core. There are various instrumentation delays associated with each trip function, including delays in signal actuation, in opening the trip breakers, and in the release of the rods by the mechanisms. The total delay to trip is defined as the time delay from the time that trip conditions are reached to the time the rods are free and begin to fall. Limiting trip setpoints assumed in 14D-4

l BVPS-1-UPDATED FSAR Rev. 0 (1/82) accident analyses and the time delay assumed for each trip function are given in Table 14D-3. Reference is made in that table to overtemperature and overpower AT trip shown in Figure 14D-1. Thc c;;rtcmpercturc'a?-cctpcint: 4chcen in rigu;; 140 1.along ;;ith all othcr ;;;1uated OMOR'O ^,icze ;&lcul& Led eseumiug appihiiuaLuly

                                                                        --lM mu 3 u .u Lhu w L uul Loat fl u valvulaLivu, ao di=vu==ed in Cectica 3.4.2.1.

The difference between the limiting trip point assumed for the analysis and the nominal trip point represents an allowance for instrumentation channel error and setpoint error. During preliminary startup tests, it was demonstrated that actual instrument errors and time delays are equal to or less than the assumed values. 14D.5 INSTRUMENTATION DRIFT AND CALORIMETRIC ERRORS I TOLTE RANCE NEUTRON TLUX g instrumentation drift and calorimetric errors used ' est hing the maximum overpower setpoint are prese in Table 14 The calorimetric r is the error assumed the determination of core thermal po as obtained om secondary plant measurements. The total 1 chambe urrent (sum of the top and bottom sections) is calibrated equal) to this measured power on a periodic basis. The secondary pow is obtained from me'as ent of feedwater flow, feedwat inlet temperature to the steam erators, and steam pre re. High accuracy instrumentation is p ded for these measurements with accuracy tolerances much tighte han e which would be required to control feedwater flow. N 14D.6 ROD CLUSTER CONTROL ASSEMBLY INSERTION CHARACTERISTIC ,, The negative reactivity insertion following a reactor trip is a$ function of the acceleration of the RCCA, and rod worth as a function of rod position. the variation With respect into*E accident analyses, the critical parameter is the time of 1 insertion up to the dashpot entry or approximately 85 percent off the rod cluster travel. For accident analyses dit is g *g ', con tively assumed that the insertion time to dashpot entry is seconds. The RCCA position versus time assumed in acci ent analyses is shown in Figure 14D-2. REV Figure 14D-3 shows the fraction of total negative reactivity insortion for a core where the axial distribution is skewed to the lower region of the core. An axial distribution which is skewed to the lower region of the core can arise from a xenon oscillation or can be considered as representing a transient 14D-5

h. . . - _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ _ _ _ _

BVPS-1-UPDATED FSAR Rev. 0 (1/82) i axial distribution which would exist after the RCCA bank had , already traveled some distance after trip. This curve is used me inp ut. to :.11 poi .t hinctico cm. mudolo used .u L onmouL mmly.;cc. *m h analqses and evahra+iews of h trans14n4s whic.h uMit w. point kineAicta com mode 4. REV There is inherent conservatism in the use of this curve in that it is based on a skewed distribution which would exist relatively infrequently. For cases other than those associated with xenon oscillations significant negative reactivity would have been inserted due to the more favorable axial distribution existing prior to trip. The normalized RCCA negative reactivity insertion versus time is shown in Figure 14D-4. The curves shown in this figure were obtained from Figure 14D-2 and 14D-3. 14D.7 REACTIVITY COEFFICIENTS The transient response of the reactor system is dependent on reactivity feedback effects, in particular the moderator temperature coefficient and the Doppler power coefficient. These reactivity coefficients and their values are discussed in detail in CccLica 4. F 2 ": p a.L Se chh RW In the analysis of certain events, conservatism requires the use of large reactivity coefficient values, whereas, in the analysis , of other events, conservatism requires the use of small reactivity coefficient values. Some analyses, such as loss of reactor coolant from cracks or ruptures in the RCS, do not depend on reactivity feedback effects. The values used are given in Table 14D-2. Reference is made in that table to Figure 14D-5 which shows the upper and lower Doppler power coefficients as a function of power used in the transient analysis. The justification for use of conservatively large versus small reactivity coefficient values are treated on an event by event basis. To facilitate comparison, individual sections in which justification for the use of large or small reactivity coefficient values is te be found are referenced below. Condition II Events Section

1. Uncontrolled RCCA Bank Withdrawal 14.1.1 from a Subcritical Condition

2. Uncontrolled RCCA Bank Withdrawal at 14.1.2 Power

3. RCCA Misalignment 14.1.3

4. Uncontrolled Boron Dilution 14.1.4

5. Partial Loss of Forced Reactor Coolant 14.1.5 Flow 14D-6

BVPS-1-UPDATED FSAR- Rev. 0 (1/ 82 ).

1. A'sufficiently large number of radial space increments to handle fast transients such as rod ejection accidents.

2. Material-properties which are functions of temperature and_ a sophisticated fuel-to-clad gap heat transfer calculation.

3. The- necessary calculations to handle post-DNB-transients: film boiling heat transfer correlations, Ziracloy-water reaction and partial ' melting of the materials.

The gap heat transfer coefficient is calculated according to an elastic pellet model '(refer to Figure 14D-7). The thermal expansion of the pellet is calculated as the sum of - the radial (one-dimensional) expansions of the rings. Each ring is assumed to expand-freely. The cladding diamdter is calculated based on thermal expansion and internal and external pressures.

                                 -If the 'outside radius of the expanded pellet is smaller than the inside radius of the expanded clad, there is'no fuel-clad contact and the gap conductance is calculated on the basis of the thermal conductivity of the gas contained in the gap.                                                                                                                                                                                     If ' the ' pellet outside radius: so calculated is larger - than the clad' inside radius (negative gap), the pellet and the clad are pictured as exerting upon each other a pressure sufficiently important' to reduce the gap to zero by elastic deformation of both.                                                                                                                                                                                             This contact pressure determines the gap . heat transfer coefficient.

FACTRAN is further discussed in Reference 7. ,

                                             .10.2      BLKOUT                                                                                                                                                                                                                                  ,

The B UT code ral is used to analyze specifically the ong term

                                    -(slow) t             sient behavior of the RCS process variabl . The "two loop" anal             cal model employed permits studyin                                                                                                                                                                              ffects, caused by anomalous              nditions                                      in                  one                         loop                               using                   i                      ed                          parameters     to describe          the    e     itions of all remaini                                                                                                                                    loops.                                             The plant components simulate                         are the entire RCS                                                                                                                    luding the pressurizer and      the      associated                                      essurizer co rol systems, the steam generator, chemical and                                                           ume con                                                                       system, and the steam dump system.

The code is generally ppli le for studying transients following perturbation in reacto coolant loop flow, nuclear power, feedwater fl or enthalpy, a flow and charging or letdown flow. er the reactor trips, the decay heat curve, shown in Figur 4D-6, is used for the comp tion of the nuclear power input r BLKOUT. The effect of transie nuclear power is accounte for in the station blackout and 1 s-of-feedwater analy * .- The heat generation rate from the nuc1 power and re ual fissions (shown in Figure 14D-8) are added to e decay at curve as input to the BLKOUT code. A very cons vative 2 second delay time is assumed from a trip signal to ip. 14D-11 l l

BVPS-1-UPDATED FSAR Rev. 0 (1/82) l the minimum shutdown margin of 1 percent Ak , assumed ximizes the residual fissions, ease of astored energy in t is accounted f sauming eight full power. seconds of stored one ng a trip. The eight full

     <4                              '

power seconds of one e fuel ted for by assuming fthis valu thermal time constant of explained in Reference 8.

                                                                                                              .                     This is REV    14D.10.3    MARVEL The MARVEL codel'I is used to determine the detailed transient behavior of multi-loop pressurized water reactor systems caused by prescribed initial perturbations in process parameters.                                                                      The code is useful in predicting plant behavior when different conditions are present in the loops.          For

analytical purposes, the physical, thermal and hydraulic characteristics of a multi-loop plant are represented by two " equivalent" loops. The perturbation is considered to occur in one of the equivalent loops which may represent one or more physical loops. The other equivalent loop thus represents in lumped form, the remaining loops in the plant.

The code simulates the coolant flow through the reactor vessel, hot leg, cold leg, steam generator plus the pressurizer surge line. Neutron kinetics, fuel-clad heat transfer and the rod control system characteristics are modeled. Simulation of the , reactor trip system, ESF (safety injection) and chemical and volume control system is provided. MARVEL also has the capability of calculating the transient value of DNB ratio' based on the input from the core limits illustrated on Figure 14D-1. The core limits represent the minimum value of DNBR as calculated for a typical or thimble cell. MARVEL is further discussed in Reference 9. 14D.10.4 LOFTRAN The LOFTRAN programII'I is used for studies of transient response of a pressurized water reactor system to specified perturbations in process parameters. LOFTRAN simulates a multi-loop system by a lumped parameter single loop model containing reactor vessel, hot and cold leg piping, steam generator (tube and shell sides) and the pressurizer. The pressurizer heaters, spray, relief and safety valves are also considered in the program. Point model 1 neutron kinetics, and reactivity effects of the moderator, fuel, boron and rods are included. The secondary side of the steam generator utilizes a homogeneous, saturated mixture for the thermal transients and a water level correlation for indication and control. The reactor protection system is simulated to include reactor trips on neutron flux, overpower and over-temperature reactor coolant AT, high and low pressure, low flow, and high pressurizer level. Control systems are also simulated 14D-12

BVPS-1-UPDATED FSAR Rev. 0 (1/82) including rod control, steam dump, feedwater control, and pressurisur pressure control. The safety injection system including the accumulators are also modeled. LOFTRAN is a versatile program which is suited to both accident ' evaluation and control studies as well as parameter sizing. LOFTRAN also has the capability of calculating the transient value of DNB ratio based on the input from the core limits ~ illustrated on Figure 14D-1. The core limits represent the 'f minimum value'of DNBR as calculated for typical or Thimble cell. ' LOFTRAN is further discussed in Reference 10. 14D.10.5 LEOPARD The LEOPARD computer programl11I determines fast and thermal spectra, using only basic geometry and temperature data. The code optionally computes fuel depletion effects for a dimensionless reactor and recomputes the spectra before each discrete burnup step.

LEOPARD is further described in Reference 11. I 14D.10.6 TURTLE TURTLE (12) is a two-group, two-dimensional neutron diffusion code featuring a direct treatment of the nonlinear effects of xenon, enthalpy, and Doppler. Fuel depletion is allowed. TURTLE was written for the study of azimuthal xenon oscillations, but the code is useful for general analysis. The input is simple, fuel management is handled directly, and a boron I 1 criticality search is allowed. TURTLE is further described in Reference 12.

                     ~+ INSET.T @

14D.10.7 TWINKLE REV The TWINKLE programI10 is a multi-dimensional spatial neutron kinetics code, which was patterned after steady-state codes presently used for reactor core design. The code uses an implicit finite-difference method to solve the two-group transient neutron diffusion equations in one, two and three dimensions. The code uses six delayed neutron groups' and contains a detailed multi-region fuel-clad-coolant heat transfer model for calculating pointwise doppler and moderator feedback effects. The code handles up to 2000 spatial points, and performs its own steady state initialization. Aside from basic l cross-section data and thermal-hydraulic parameters, the code accepts as input basic driving functions such as inlet temperature, pressure, flow, boron concentration, control rod motion, and others. Various edits provide channelwise power, 1 14D-13 _ _ _ _ _ _ _ _ _ . . ._ . 1

i l BVPS-1-UPDATED FSAR Rev. 0 (1/82) { axial of fset, enthalpy, volumetric surge pointwise power, fuel temperatures, and so on. I The TWINKLE code is used to predict the kinetic behavior of a I reactor for. transients which cause a major perturbation in the , spatial neutron flux distribution. 1 TWINKLE is further described in Reference 13. !

       .10.8  WIT WIT il 4     a one-region neutron kinetics program with a                                                                                               gle   ;

axial lump cription of thermal kinetics making i seful in j the analysis o transients in a heterogeneo reactor core consisting of fuel s, fuel rod cladding, water moderator and coolant. The code basically a model and; therefore, generally useful for fast etiv transients which terminate before significant effects oc rom the remainder of the plant, i.e. transients shorter t . the lo transit time. WIT 'is used i fety analysis of ' react 1 accidents from a subcritica ndition. is further described in Reference 14. ~ t .9 PHOENIX The PHOENIX c (2 a calculates the individual flows, core flow and pump speeds function of time equent to failure of any number of the rea coola umps. The analysis is based on a momentum balance aro h reactor coolant loop and across the reactor core, momentum nee is combined with the continuity equat , a pump momentum ba and the pump characteristics y number of reactor coolant s are accommoda p to a maximum of 6. ENIX is further described in Reference 15. 14D.10.10 THINC The THINC code is described in Section 3.4.3.1. 1 14D-14 _______.___________-______________m_ _ _ _ _ _ _ _ _ __ _ _ _ _ _ _ _ _ _ _ _ ____m_____ -

BVPS-1-UPDATED FSAR Rev. 2 (1/84) l References to Appendix 14D

1. Supplemental information on fuel design transmitted from R.'Salvatori, Westinghouse, to D. Knuth, Atomic Energy

_., Commission, as attachments to letters NS-SL-518 (12/22/72), NS-SL-521. (1/4/73), NS-SL-524 (1/4/73) and NS-SL-543 (1/12/73), (Westinghouse Proprietary); and supplemental

                 'information on fuel design transmitted from R. Salvatori, to'D. Knuth,.as attachments to letters NS-SL-527 (1/2/73) and NS-SL-544 (1/12/73).

I

2. K. Shure, " Fission Product Decay Energy" WAPD-BT-24, pp.

1-17, Westinghouse Bettis Atomic Power Laboratory (Decer.ber 1961).

3. K. Shure and D. J. Dudziak, " Calculating Energy Released by Fission Products," Trans. American Nuclear Society, 4 (1)

p. 30 (1961).

4. Teake, United Kingdom Atomic Energy Authority Decay Heat-Standard. (Private Communication)

5. J. R. .Stehn and E. F. Clancy, " Fission-Product . Radio-activity and Heat Generation" in " Proceedings of the Second United Nations International Conference on the Peaceful

                'Uses of Atomic Energy, Geneva, 1958," Volume 13, pp. 49-54, United Nations, Geneva, 1958.                                                             .

6.' F. E. Obenshain and A. H . Foderaro, " Energy from Fission Product. Decay," WAPD-P-652, Westinghouse Bettis Atomic Power Laboratory (1955.) .

7. C. Gunin,~"FACTRAN, a FORTRAN IV Code for Thermal Transient in a UO Fuel Rod," WCAP-7908, Westinghouse Electric Corporati$n (June 1972).

6. J. M. .m . , E. Oel ;teri, "L;ng T;;- T :n;i:nt ..' :ly;;r (jy Imp -.. fe. 7" 0 'OL"0* T Cede;," t'O'." 7 0 00 , 'c';;tingh; = :

g a ":lectai; C;.g;;;ti;n Mun: 1972).

9. J. M. Gutz, " MARVEL (A Digital Computer Code for Transient Analysis of a Multi-loop PWR System)," WCAP-7909, Westinghouse Electric Corporation (June 1972).

10. T. W. T. Burnett, C. J. McIntyre, J. C. Baker, R. P. Rose, "LOFTRAN Code Description," WCAP-7907, Westinghouse Electric Corporation (June 1972).

11. R. F. Barry, " LEOPARD, a Spectrum Dependent Ncn-Spatial Depletion Code for the IBM-7094," WCAP-3269-26, Westinghouse Electric Corporation (September 1963) .

14D-15 l l

                                                                                                                                                        ^
                                      . _ - _ _____-________ _ ____--_                      _l____-____-___                    _-                           L__ _ _ L

BVPS-1-UPDATED FSAR Rev. 2 (1/84) References to Appendix 14D (Cont'd)

12. R. F. Barry and S. Altomare, "The TURTLE 24.0 Diffusion Depletion Code": WCAP-7213, (Westinghouse Proprietary)

(June 1968) ; WCAP-7758 (September 1971).

13. D. H. Risher, Hr., R. F. Barry, " TWINKLE -

A Multi-Dimensional Neutron Kinetics Computer Code," WCAP-7979, Westinghouse Electric Corporation (November 1972).

14. O. ". T:irir:th::, H. C. Errgr: 72 ""IT-f ?!!?ter TriNSi?rt -

Tumly.i. Ce yo.ez Tres;;; 0:Oripti:n," "C.'."- ? ? ? O , IU W .Liuy m e. 21.; 1; C:spcz;t.icn '0:::-icr 1972'. r4 Y M. T. ". "crd 1:n "C:10:1:ti:r

                     .                                                      f F10L Ce r rtdO /r After 10r? Of
                 ":::t:r     Cr:1:nt                 ?" ;                '"                    Tv                                                        ad a i - "   WF A D- 7 4 7 't ,

p Weetinghc ;; 21 :tri: C::p::: tier 'Erptr-her l??2' -

16. F. M. Bordelon, et al., " SATAN-VI ' Program: Comprehensive Space-Time Dependent Analysis of Loss-of-Coolant",

WCAP-8306, Westinghouse Electric Corporation (June 1974). lQSET.T h e.sv I

                                                                    ,                                                                                                                                                               1 i

l l i 14D-16 I 1 [ ._ _ . . _ _ _ . . _ . . _ __ _ _ _ .__ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ - - _ _ _ _ _ _ _ _ - - - -

l~ BVPS-1 UPDATED FSAR APPENDIX 14D Inserts for Zire Grids and Increased Peaking Factors Insert A: 140.10.6.1 Advanced Nodal Code (ANC) ANC is a two-group, multi-dimensional nodal diffusion theory code. ANC.can be used for all nuclear core design calculations, including critical boron concentrations, control rod worths, and reactivity coefficients. l ANC is further described in Reference 17. 14D.10.6.2 PHOENIX-P PHOENIX-P is a two-dimensional multigroup transport theory code used to calculate lattice physics parameters for pressurized water reactors. This code generates cross section and feedback parameters that are consistent with dimensional code requirements. PHOENIX-P is further described in Reference 18. Insert B:

17. Liu, Y. S. , et. al . , "ANC: A Westinghouse Advanced Nodal Computer Code," WCAP-10965-P-A, September 1986.

18. Nguyen, T. Q., et. al., " Qualification of the PH0ENIX-P/ANC Nuclear Design System for Pressurized Water Reactor Cores",

WCAP-11596-P-A, June 1988. i i i

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I I BVPS-1-UPDATED FSAR Rev. 0 (1/82) TABLE 14D-4 j DETERMINATION OF MAXIMUM OVERPOWER TRIP POINT - PO ER RANGE NEUTRON FLUX - BASED ON NOMINAL SETPOINT/ CONSIDERING INHERENT INSTRUMENTATION ERRORS /
\
~~A. Nomi al Setpoint (% of rated power) 109~~
\
B. Calorim tric Errors in the Measurement of Se ndary System hermal Power: stimated Accuracy of ffect on' Measurement Thermal Power of Variable Determination Variable (% error) (% error) Feedwater temperat re +0.5 -- Feedwater pressure 15 0.3 (Small correction on enthalpy) Steam pressure _,2 -- (Small correction q) on enthalpy) Feedwater' flow (A i .25 1.25 l l C. Assumed Calorimetric E ror (% of ated power) 2 D. Axial power distribu ion effects on total ion chamb r current: Estimated ror (% of rated powe , 3 Assumed ror (% of rated power) 5 E. Instrumental n channel drift and setpoint r reducibility: Est' ated Error (% of rated power) 1 A sumed Error (% of rated power) 2 F. Maxi overpower trip point assuming all in vidual errors are simultaneously in the mo t adverse direction (% of rated power) 18 1 of 1
~~. - - - - - - - - _ - . - ----_______.--______________________j~~
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ML20246K454

Contents

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1.0 INTRODUCTION

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6.0 REFERENCES

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ML20246K454

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1.0 INTRODUCTION

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