ML20217C460

Reactivity Anomaly Events Methodology

ML20217C460
Person / Time
Site:	Comanche Peak
Issue date:	05/31/1991
From:	Husain A, Janne R, Maier S TEXAS UTILITIES ELECTRIC CO. (TU ELECTRIC)
To:
Shared Package
ML20217C457	List: ML20217C460 TXX-9118, Forwards Rept RXE-91-002, Reactivity Anomaly Events Methodology for Review & Approval,Per Re Plant Reload Analysis
References
RXE-91-002, RXE-91-2, NUDOCS 9107150303
Download: ML20217C460 (162)

Text

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l-l RXE-91-002 REACTIVITY ANOMALY EVENTS HETHODOLOGY MAY 1991 Dean W. Throckmorton Fred A. Monger Curtis V. DeVore David W. Hiltbrand REACTOR ENGINEERING d --

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REACTIVITY ANOMALY EVENTS METHODOLOGY l MAY 1991 1 Dean W. Throckmorton Fred A. Monger Curtis V. DeVore David W. Hiltbrand Reviewed: $*

Date: 8 !30 !9/

' Stephen M. Maier Supervisor, Transient Analysis Reviewed: 777aI2$/ M N((g/1 Date: 8 //

Micke y R. KilpJore ' '

Supervl r, Reac Physics

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Approved: / ./ Date: 5- 3 D *) )

i // ' Ra ~L. Janne Ma Nuclear Fuel I

Approved:

AAn AusafjHusain ex Date: If30f9l

1. l Director, Reactor Engineering j ^

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~.. ..s ACKNOWLEDCEMENTS This report is the culmination of many months of developmental effort in which numerous individuals participated. The Reactor Engineering department wishes to express its thanks to Ms. D. J. Edwards, and Messrs. R. M. Rubin, H. B. Giap, J. F. Harrison and C. L. Ritchey for their technical assistance in the development of the analysis-methodology and preparation of this report.

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DISCLAIMER The information contained in this report was prepared for the specific requirements of Texas Utilities Electric Company (TUEC),

and may not be appropriate for use in situations other than those j for which it was specifically prepared. TUEC PROVIDES NO WARRAN-TY HEREUNDER, EXPRESSED OR IMPLIED, OR STATUTORY, OF ANY KIND OR j NATURE WHATSOEVER, REGARDING THIS REPORT OR ITS USE, INCLUDING BUT NOT LIMITED TO ANY WARRANTIES ON MERCHANTABILITY OR FITNESS FOR A PARTICULAR PURPOSE.

By making this report available, TUEC does not authorize its use by others, and any such use is forbidden except with the prior written approval of TUEC. Any such written approval shall itself be deemed to incorporate the disclaimers of liability and dis-claimers of warranties provided herein. In no event shall TUEC have any liability for any incidental or consequential damages of cny type in connection with the use, authorized or unauthorized, of this report or the information in it.

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, .4 ABSTRACT This report describes the TU Electric methodology for performing the event analyses for the Comanche Peak Steam Electric Station (CPSES) Final Safety Analysis Report (FSAR) Chapter 15 evente characterized as reactivity anomalies. Specifically, this report addresses those events described in Section 15.4 of the CPSES FSAR. Both single- and multi-dimensional neutronic calculations are used to develop core physics parameters for the system and core thermal-hydraulic analyses. It is concluded that application of the TU Electric methodology will result in conservative predictions of the event consequences and that the applicabla acceptance criteria are met.

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Table of Contents

1.0 INTRODUCTION

. . . . . . . . . . . . . . . . . . . . . 1-1 1.1 Purpose . . . . . . . . . . . . . . . . . . . . 1-1 1.2 Report Organization . . . . . . . . . . . . . . . 1-1 2.0 ANALYSIS PHILOSOPHY . . . . . . . . . . . . . . . . . 2-1 2.1 Regulatory Bases . . . . . . . . . . . . . . . . 2-1 2.2 Event Classification . . . . . . . . . . . . . . 2-2 2.3 Acceptance Criteria . . . . . . . . . . . . .. . 2-5 2.4 Analytical Guidelines . . . . . . . . . .- . . . . 2-8 3.0 CALCULATIONAL TOOLS . . . . . . . . . . . . . . . . . 3-1 3.1 Overview . . . . . . . . . . . . . . . . . . . . 3-1

- 3.2 Cross Section Model . . . . . . . . . . . . . . . 3-1 3.3 Nodal Core Model . . . . . . . . . . . . . . . . 3-1 3.4 1-D Core Model . . -. . . . . . . . . . . . . . . 3-2 3.5 Core Thermal-Hydraulic Model . . . . . . . . . 3-2 3.6 System Thermal-Hydraulic Model . . . . . . . . . 3-3

, 3.7 Hot Spot Model . . . . . . . . . . . . . . . . . 3-7 3.9 Dilution Model . . . . . . . . . . . . . . . . . 3-9 4.0 ANALYSIS METHODOLOGY . . . . . . . . . . . . . . . . . 4-1

-4.1 Overview . .. . . . . . . . . . . . . . . . . . 4-1 4.2 Core Physics Methodology . . . . . . . . . . . . 4-1 4.3 DNB Methodology . .. . . . . . . . . . . . . . . 4-6 4.4 . Steady State Events Methodology . . . . . . . . . 4-8 4.5 Boron Dilution Methodology . . . . . . . . . . . 4-12 4.6 Control Rod Withdrawal Methodology . . . . . . . 4-22 4.7 Control Rod Drop Methodology . . . . . .. . . . 4-28 4.8 Control Rod Ejection Methodology . . . . . . . . 4-39 l

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Table of Contents (cont.)

5.0 EVENT ANALYSES . . . . . . . . . . . . . . . . . . . . 5-1 5.1 General . . . . . . . . . . . . . . . . . . . . . 5-1 5.2 Boron Dilution . . . . . . . . . . . . . . . . . 5-3 5.3 Control Bank Withdrawal at Power . . . . . . . . 5-11 5.4 Control Rod Drop . . . . . . . . . . . . . . . . 5-26 5.5 Control Rod Ejection . . . . . . . . . . . . . . 5-40

6.0 CONCLUSION

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7.0 REFERENCES

. . . . . . . . . . . . . . . . . . . . . . 7-1 Appendix A Statistical Combination of Uncertainties . . . . . . . . A-1 Appendix B Control Rod Ejection Sensitivity Studies . . . . . . . . . B-1 i

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List of Tables

_ Table 5.1-1 Common Input Parameters . . . . . . . . . . . . . 5-2 5.2-1 Mode Dependent RCS Conditions . . . . . . . . . . 5-7 5.2-2 Boron Dilution Input Assumptions . . . . . . . . 5-8 5.2-3 Boron Dilution Analysis Results . . . . . . . . . 5-9 5.3-1 Control Bank Withdrawal at Power Input Assumptions . . . . . . . . . . . . . . . . . . 5-15 5.3-2 Control Bank Withdrawal at Power Sequences of Events . . . . . . . . . . . . . . . . . . . . . 5-16 5.4-1 MTC and Inserted Control Bank Worth for the Generic Control Rod Drop Statopoints . . . . . . 5-31 5.4-2 Control Rod Drop System T-H Analysis Input Assumptions . . . . . . . . . . . . . . . . . . . 5-32 5.4-3 Control Rod Drop SCU Parameters . . . . . . . . . 5-33 5.5-1 Assumptions for CRE System T-H Analysis . . . . . 5-43 5.5-2 Assumptions for CRE Hot Spot Analysis . . . . . . 5-44 5.5-3 HFP CRE Sequence of Events . . . . . . . . . . . 5-45 5.5-4 CRE HFP Hot Spot Results . . . . . . . . . . . . 5-45

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5.5-5 HZP CRE Sequence of Events . . . . . . . . . . . 5-46 5.5-6 CRE HZP Hot Spot Results . . . . . . . . . . . . 5-46 B-1 HFP Neutronics Parameters Sensitivity Study Results . . . . . . . . . . . . . . . . . . . . . B-11 B-2 HZP Neutronics Parameters Sensitivity Study Results . . . . . . . . . . . . . . . . . . . . . B-12 a B-3 Thermal-Hydraulic Parameters Sensitivity Study Results . . . . . . . . . . . . . . . . . . . . . B-13 B-4 HFP Hot Spot Sensitivity Study Results . . . . . B-14 B-5 HZP Hot Spot Sensitivity Study Results . . . . . B-15 vii 3

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List of Figures Fiaure 3.6-1 System Thermal-Hydraulic Model . . . . . . . . . 3 3.7-1 Hot Spot Model . . . . . . . . . . . . . . . . . 3-11 4.7-1 Control Rod Drop Analysis Flow Chart . . . . . . 4-38 4.8-1 Control Rod Ejection Analysis Flow Che,rt . . . . 4-50 5.2-1 Limiting Boron Worth curves . . . . . . . . . . . 5-10 5.3-1 Core Average Power Response (75 pcm/sec, Minimum Feedback) . . . . . . . . . . . . . . . . . . . - . 5-17 5.3-2 Pressure Response (75 pcm/sec, Minimum Feedback) 5-17 5.3-3 Temperature Response (75 pcm/sec, Minimum Feedback) . . . . . . . . . . . . . . . . . . . . 5-18 5.3-4 DNBR Response (75 pcm/sec, Minimum Feedback) . . 5-18 5.3-5 Core Average Power Response (25 pcm/sec, Minimum Feedback) . . . . . . . . . . . . . . . . . . . . 5-19 5.3-6 Pressure Response (25 pcm/sec, Minimum Feedback) 5-19 5.3-7 Temperature Response (25 pcm/sec, Minimum Feedback) . . . . . . . . . . . . . . . . . . . . 5-20 5.3-8 DNBR Response (25 pcm/sec, Minimum Feedback) . . 5-20

-5.3-9 Core Average Power Response (1 pcm/sec, Minimum Feedback) . . . . . . . . . . . . . . . . . .. . 5-21 5.3-10 Pressure _ Response (1 pcm/sec, Minimum Feedback) . 5-21 5.3-11 Temperaturs Response (1 pcm/sec, Minimum Feedback) . . . . . . . . . . . . . . . . . . . . 5-22 5.3-12 DNBR Response (1 pcm/sec, Minimum Feedback) . . . 5-22 5.3-13 Control Bank Withdrawal from 102% RTP MDNBR vs. Reactivity Insertion Rate . . . . . . . 5-23 5.3 Control Bank Withdrawal from 62% RTP MDNBR vs. Reactivity Insertion Rate . . . . . . . 5-23 5.3-15 Control Bank Withdrawal from 12% RTP MDNBR vs. Reactivity Insertion Rate . . . . . . . 5-24

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I List of Figures (cont.)

Ficure 5.3-16 Pressure Transient for Control Bank Withdrawal . 5-24 5.3-17 Pressurizer Liquid Volume Response for Control Bank Withdrawal . . . . . . . . . . . . . . . . . 5-25 5.4-1 Bounding Excore Detector Tilt vs. Dropped Control Rod Worth . . . . . . . . . . . . . . . . . . . . 5-34 5.4-2 DNBR Limit Lines . . . . . . . . . . . . . . . . 5-34 5.4-3 Core Average Power Response at BOC . . . . . . . 5-35 5.4-4 Pressurizer Pressure Response at BOC . . . . . . 5-35 5.4-5 RCS Temperature Response at BOC . . . . . . . . . 5-36 5.4-6 ONBR Response at BOC . . . . . . . . . . . . . . 5-36 5.4-7 Core Average Power Response at EOC . . . . . . . 5-37 5.4-8 Pressurizer Pressure Respeide at EOC . . . . . . 5-37 5.4-9 RCS Temperature Response - C . . . . . . . . . 5-38 5.4-10 DNBR Response at EOC . . . . . . . . . . . . . 5-38 5.4-11 BOC Fui Conparison . . . . . . . . . . . . . . . 5-39 5.4-12 EOC Fui Comparison . , . . . . . . . . . . . . 5-39 5.5-1 BOC HFP Core Ave age Power Response . . . . . . . 5-47 5.5-2 BOC HFP Fuel Enthalpy Response . . . . . . . . . 5-47 5.5-3 BOC HFP Fuel Centerline Temperature Response . . 5-48 5.5-4 BOC HFP Average Fuel Temperature Response . . . . 5-48 5.5-5 EOC HFP Core Average Power Response . . . . . . . 5-49 5.5-6 EOC HFP Fuel Enthalpy Respense . . . . . . . . . 5-49 5.5-7 EOC HFP Fuel Centerline Temperature Response . . 5-50 a 5.5-8 EOC HFP Average Fuel Tempera?.ure Response . . . 5-50 5.5-9 BOC HZP Core Average Power Response . . . . . . . 5-51 5.5-10 BOC HZP Fuel Enthalpy Response . . . . . . .I . . 5-51 5.5-11 BOC HZP Centerline Puel Temperature Response . . 5-52 5.5-12 BOC HZP Average Fuel Temperature Responae . . . . 5-52 5.5-13 EOC HZP Core Average Power Response . . . . . . . 5-53 3

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List of Figures (cont.)

Floure 5.5-14 EOC HZP Fuel Enthalpy Response . . . . . . . . . 5-53 5.5-15 EOC HZP Fuel Centerline Temperature Response . . 5-54 5.5-16 EOC HZP Average Fuel Temperature Response . . . . 54 X

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s t. .o List of Abbreviations ANS American Nuclear Society ANSI American National Standards Institute ARI All Rods In l ARO All Roda Out LASME American Society of Mechanical Engineers BDMS Boren Dilution Mitigation System BOC Beginning of Cycle CPSES Comanche Peak Steam Electric Station CVCS Chemical and Volume control System DNB Departure from Nucleate Boiling DNBR Departure from Nucleate Boiling Ratio DTC Doppler Temperature Coefficient DWF Doppler Weighting Factor EOC End of Cycle EPRI Electric Power Research Institute Fa Hot channel peaking factor FPS Full-Power Seconds

. Fn Total peaking factor FSAR -Final S1fety Analysis Report HFP Hot Full Power HZP Hot Zero Power MDNBR Minimum Departure-from Nucleate Boiling Ratio MTC Moderator Temperature Coefficient NIS Nuclear Instrumentation System N-16 Nitrogen-16 NRC Nuclear Regulatory Commission OPN-16 Overpower N-16 OTN-16 Overtemperature N-16 PORV Power-operated Relief Valve a

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List of Abbreviations (cont.)

PZR Pressurizer RCP Reactor Coolant Pump RCS . Reactor Coolant System RHR Residual Heat Removal System RIL Rod Insertion Limits RMWS Reactor Makeup Water System RPS Reactor Protection System RTS Reactor Trip System RV Reactor Vessel l RWST Refueling Water Storage Tank l

SCU Statistical combination of Uncertainties l SDM Shutdown Margin SER Safety Evaluation Report SG Steam Generator SI Safety Injection SRSS Square Root Sum of the Squares T3yo RCS Average Temperature T-H Thermal-hydraulic Tuy Reference RCS Average Temperature TUEC Texas Utilities Electric Company UF Uncertainty Factor VCT Volume Control Tank xii 3:

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1.0 INTRODUCTION

1.1 Purpose This report describes the methods employed by TU Electric to perform the analyses of the FSAR Chapter 15.4 reactivity anomaly events. Additionally, this report describes the methods utilized to develop the physics parameters necessary to perform the safety analyses. The reactivity anomaly events include transients that f are initiated by movement of the control rods or changes in the Reactor Coolant System (RCS) boron concentration, and those events resulti..g from mispositioning of control rods or fuel assemblies. The methodology described in this report expands on

methods previously developed by TU Electric related to core reload design [1,2,3], core thermal-hydraulic analysis (4,5,6],

and system thermal-hydraulic analysis [7,8].

1.2 Report oraanization

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The main body of this report is organized into four sections:

the TU Electric analysis philosophy, the computer codes and models, the analysis methodology, and the specific event analyses.

A brief synopsis of the TU Electric analisis philosophy is

.A provided in Section 2.0 of this report. Subsections are provided to address the regulatory bases, event classification, acceptance criteria, and analysis guidelines used by TU Electric.

The computer codes and models used to develop input parameters and to perform the analyses of the reactivity anomaly avents are

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described in Section 3.0.

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The TU Electric' methodology employed to perform the analysis of the reactivity anomaly events is detailed in Section 4.0.

Four specific event analyses are provided in Section 5.0 to demonstrate the application of the TU Electric methodology.

These analyses include boron dilution, control bank withdrawal at power, control rod drop, and control rod ejecticn.

The conclusions of the report and the references utilized in the develr.pment of the methodology are provided in Section 6.0 and Section 7.0, respectively.

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~., .a 2.0 ANALYSIS PHILO?OPHY 2.1 Reaulatory Bases i 1

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Title 10 of the Code of Federal Regulations (9), Part 50 j (10CFR50), requires that a Final Safety Analyris Report (FSAR) be prepared for each reactor facility which includes information 1

that describes the facility, presents the design bases and limits I of operation, and presents a safety analysis of the facility as a whole. The purpose of the safety analysis is to demonstrate that the reactor facility can be operated without undue risk to the public hee.lth and safety. Appendix A of 10CFR50 establishes the minimum requirements for the principal design criteria for nuclear power plants (General Design Criteria or GDC',. The

American National Standard " Nuclear Safety Criteria for the Design of Stationary Pressurized Water Reactor Plants" (ANSI N18.2-1973) [10), was developed to amplify the guidance provided by the GDC. Additional requirements beyond those contained in the GDC have been established by the Nuclear Regulatory I

Commission (NRC) and committed to by TU Electric in establishing a

the safety analyses for CPSES. These requirements (industry standards, regulatory guides, NUREGs, etc.) are documented in the CPSES FSAR (11) and the Safety Evaluation Report (SER) [ 12 ).

issued by.the NRC and form the licensing basis for CPSES.

A defense-in-depth design philosophy is employed to safely a terminate and mitigate the consequences of accidents. Section II of the GDC provides the criteria for protection by use of multiple fission product barriers. The three physical barriers that constitute the defense-in-depth design are the fuel and cladding, the recctor coolant system, and the containment. The FSAR Chapter 15 events are analyzed in terms of their effect on

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these physical barriers because of the relationship between barrier integrity and dose. The NRC regulates the compliance 2-1 2

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l with the dose limits under the provisions of parts 20 and 100 to Title 10 of the Code of Federal Regulations, depending on the l

expected frequency of occurrence of the event. 10cFR20 provides the requirements for determining maximum acceptable lovels of radioactivi y in restricted and unrostricted arecs, while i

I 10CFR100 provides Reactor Site Criteria used in determining the maximum acceptable offsite doses.

2.2 Event C1 4ssification ANSI Standard N18.2-1973 is used to classify plant conditions or incidents in ac;ordance with their potential consequences and anticipated frequency of occurrence. The standard also provides general design requirements for each evant category. The basic principle applied in relating the genervi design requiremento to the plant condition is that the most frequent occurrences should yield little or no adverse consequences to the public, and the improbable extreme situations, having the potential for the greatest adverse consequence to the public, shall have a low probability of occurrence. The standard divides the spectrum of plant conditions into four categories. These categorios are:

1. Condition I Normal Operation; ,

2. Condition II: Incidents of Moderate Frequency;

3. Condition III: Infrequent Faults; and, l 4. Condition IV: Limiting Faults.

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Condition I: Normal Operation Condition I occurrences are operations that are expected frequently or regularly in the course of power operation, refueling, maintenance, or maneuvering of the plant. The only design requirement is that Condition I occurrences shall be 2-2 3

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~. .a accommodated with margin between any plant paramotor and the value of that paramotor which would require either automatic or manual protectivo action.

Condition II: Incidents o _r Moderate Frequency Condition II occurrences include incidents, any one of which may occur during a calendar year for a particular plant. Soveral design requiremon' .a associated with Condition II occurrences.

Those requiremonta a/9:

1. Condition II incidents shall be accommodated with, at must, a shutdown of the cactor with the plant being capable of retarning to operation after correctivo action;

2. A Condition II incident, by itself. cannot generato a more serious incident of the Condition III or IV typo without other incidents occurring indopondently;

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3. A singlo Condition II incident shall not cause consequential loss of function of any barrior to tho l

oscape of radioactivo products; and,

4. Any release of radioactive materials in offluent to unrestricted areas shall be in conformance with

.A 10CFR20.

Condition III: Infrecuent Faults Condition III occurrences include incidents, any one of which may occur during the lifetime of a particular plant. The following 2-3 J

I design requirements are associated with Condition III occurrencost

1. Condition III incidents shall not cause more than a l small fraction of the fuel elements in the reactor to be damaged, although sufficient fuel olomont damage might occur to preclude renumption of operation for a considerable outage time;

2. The release of rad [3 active material resulting from a Condition III incidant may exceed guidelines of 1 10CFR20, but shall not be sufficient to intorrupt or restrict public use of those areas beyond the exclusion I radius; and, 1

3. A condition III incident shall not, by itself, generato a condition IV fault or result in e consequential loss of function of the reactor coolant system or reactor containment barriers.

CQDdition IV: Limitina Faults condition IV occurrences are faults that are not expected to occur, but are postulated because their consequences would include the potential for the release of significant amounts of radioactive material. Two design requirements are associated with condition IV occurrences. These requiremonta ares l

1. Condition IV faults shall not cause a release of radioactive material that results in an undue risk to public health and safety excoading the guidelines c:'

10CFR100, and l

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2. A single Condition IV fault shall not cause a consequential loss of required functions of systems needed to cope with the fault including those of the reactor coolant system and the reactor containment system.

2.3 Acceptance Criteria To ensure adherence to the philosophy associated with the categorization of plant conditions or incidents, critoria are established to demonstrate that the integrity of each fission product barrier is maintalnud within the limits established fer the specific classification. Thec' criteria are generally referred to as acceptanca criteria. The acceptanco critoria provide a tangible means of evaluating the barrier performanco with respect to the more-generic criteria of the GDC. For example, General Design Criterion 25 requires that the reactor protection system be designed to assure that speciflod fuel design limits are not exceeded in the event of a single malfunction of the reactivity control system. This critorion can

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be satisfied by demonstrating that a reactor trip occurs in sofficient time to prevent centerline fuel mult or DNB. The events within the scope of this report are categorized below, with acceptance criteria specified for each classification.

2 Condition II: Incidents of Moderato Frequency The following reactivity anomaly events are classified as -

Condition II occurrences:

1. Control Bank Withdrawal from Suberitical;

2. Control Bank Withdrawal at power; 3.- Control Rod Drop; 2-5

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4. Statically Misaligned Control Rod; and,

5. -Boron Dilution.

The acceptance criteria for these events are established to preclude any fission product barrier degradation. The specific acceptance criteria are as follows:

1. The minimum Departure from Nucleate Doiling Ratio (DNBR) remains greater than the 95/95 DNBR limit, i.e.,

greater than a 95% probability at a 95't confidence level that DNB does not occur for the limiting tual pint

2. Pressure in the Reactor Coolant System (RCS) and main steam system is maintained less than 1101 (13) of their design limits; and,

3. Fuel centerline temperatures do not exceed the molting point.

An additional acceptance criterion is established for the Boron Dilution event. This criterion imposes a minimum time interval between the time when an alarm announces an unplanned boron dilution and the time at which operator intervention to terminate the event can be assumed. The required minimum time interval is 30 minutes during reftf, ling (Mode 6), and 15 minutes during all other operational modes (Modes 1 through 5).

Condition III: Infyeauent Paults The following reactivity anomaly events are classified as condition III occurrences:

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1. Single Control Rod Withdrawal, and  !

2. Hisloaded Fuel Assembly.

The acceptance criteria for these events are established to preclude any appreciable fission product barrier degradation.

These criteria are:

1. The RCS and main steam system pressures are maintained -

less than-110% of their design pressures, and

2. The offsite dose consequences must be a small fraction of the 10CFR100 guidelines of 25 REM whole body ar.d 300 REM to the thyroid. "Small fraction" is to be interpreted as less than 10% of the above values.

Condition IV: Limitina Faults The only event in FSAR Section 15.4 classified as a Condition IV occurrence is the control rod ejection-event. Because this event is a Limiting Fault, the associated acceptance criteria allow for a greater extent of fission product barrier degradation. The following acceptance criteria, consistent with Regulatory l Guide 1.77 (14),-are established for the control rod ejection i event: '

1. The radially averaged fuel pellet enthalpy shall not a

exceed 280 cal /gm at any axial location;

2. Offsite dose consequences must be "well within" the

-10CFR100 guidelines of 25 REM whole body and 300 REM to the thyroid. "Well within" is to be interpreted as less than 25% of the above values; and,

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3. The RCS pressure must remain less than the value that will cause stressen to exceed tite " Service Limit C" as defined in the ASME Code (13).

2.4 Analvtical Guidelines To achieve an overall conservative analysis, limiting analytical assumptions for system performance, initial conditions, and setpoint selection are employed. These assumptions are as follows:

1. The initial operating conditions assumed in the analyses are tMo most adverse with respect to the i

acceptance criterion of interest. These conditions are consistent with steady state operation, allowing for calibration and instrument errors and steady state fluctuations.

2. Only the plant systems and components classified as important to safoty are credited for event mitigation.

The single active failure which produces the most limiting results, relative to the event acceptance criterion of interest, is the worst single active failure and is included in the event analysis.

3. Control system operations are assumed only when action normally taken by the control system results in a more severe event.

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4. The setpoints and delay times for the Reactor l Protection System (RPS) are selected to provide I

sufficient operating margin to prevent spurious actuation, yet still provide adequate reactor -

protection.

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3.0 CALCULATIONAL TOOLS I

3.1 overvigW This section identifies the computer codes and models used to perform the various calculations required to analyze the reactivity anomaly events. A brief description of each code and model is provided.

3.2 Cross Section Model i

The fuel assembly cross sections used in the core physics models are gene ated by the CASMo-3 (15) computer code, using the TU Electric steady state physics methodology (1). CASMO-3 is a  ;

multi-group, two-dimensional transport theory code for burnup calculations on BWR and PWR assemblies of simple pin cells. The code analyzes a geometry consisting of cylindrical fuel rods of varying composition in a square pitch array with an allowance for

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fuel rods loaded with gadolinium, burnable absorber rods, cluster control rods, and incore instrument channels.

3.3 8947.1 Core Model The SIMULATE-3 (16) computer code, in conjunction with the 2

TU Electric steady state physics methodology (1,2), is used to determine reactivity coefficients, kinetics parameters, control rod worths, multi-dimensional power distributions, and event-specific axial power shapes. SIMULATE-3 is also used to calculate diacrete pin powers for a given core power distribution. TU Electric has performed verification

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calculations supporting the application of SIMULATE-3 to the calculation of discrete pin powers. In addition, this 3-1 i

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application of SIMULATE-3 has been approved for use in licensing analysis by another utility (17).

SIMULATE-3 is an advanced two-group nodal code used to perform '

steady state core physics calculations. The neutronics model in SIMULATE-3 solves a two energy group neutron diffusion equation using second-order polynomials to represent the tranavorce j leakage. These polynomials represent the intra-nodal flux distribution for both fast and thormal energy groups. The solution can be obtained in one , two , or three-dimensions.

SIMULATE-3 omploys discontinuity factors to permit discontinuous flux on the node boundary. The discontinuity factors aro l obtained from the same CASMO-3 casos which generate two-group assembly avertge cross sections. <

l 3.4 1-D Core Model l

The TU Electric 1-D core model and methodology [3] are used to generate axial power shapes for which event specific axial power shapes are not required. The 1-D core model employs a modified two-group, diffusion theory reactor simulator which is designed to execute with large mesh sizes. This model uses diffusion theory to solve for the faut group flux assuming, with minor corrections, the only source of thermal neutrons to be those slowing from the fast group and that no thermal leakaga occurs within each node.

3.5 Core Thermal-Hydraulic Model The VIPRE-01 (18) computer code is used in conjunction with the TU Electric core thermal-hydraulic (T-H) analysis methodology (4) to perform the reactor core T-H analyses. The CPSES core T-H l 3-2 i 3

v.. ..

model is consistent with the requiremonto specified in the VIPRE-01 SER (19).

VIPRE-01 was developed under the sponsorship of the Electric Power Research Institute (EPRI), to perform reactor coro thermal-hydraulic analysos involving the calculations of local heat flux, fluid conditions, and Departura from Hucloato Boiling Ratio (DNBR). VIPRE-01 allows tbo reactor coro to be modelod as a number of quasi one-dimensional flow paths-(channels) that communicate laterally by disorsion crossflow and turbulent mixing. Both steady stato and transient analysos can be performed with the VIPRE-01 code. '

3.6 System Thermal-Hydraulic Modal The RETRAN-02 (20) computer code la the principal analytical tool uced to perform the system T-H analysis portion of the evento described in this report. RETRAN-02 has the flexibility to model many different accident scenarios and plant conditions. As such, the NRC has stipulated in the SER for RETRAN-02 (21,22) that each user describe their application of the code. A description of TU Electric applications of RETRAN-02 is provided in other system T-H analysis reports (7,8).

The reeativity anomaly events are charactorized by upsets in the coro power distribution resulting from porturbations in the 2 ceactivity control mechanisms. As such, the ovents generally rcsult in a symmetric heat transfer loop response that is insensitive to secondary sido characteristics. The system T-H model is therefore selected as an equivalent single loop representation of the CPSES model (7). A noding diagram of the system T-H model is presented in Figure 3.6-1. Specific

^

components of the system T-H model which are important to the analysis of the reactivity anomaly events are described below.

3-3 1-l _

The RUTRAN-02 non-equilibrium pressurizer (PZR) model allows the simulation of specific situations where two distinct and separate ,

thermodynamic regions exist that ar1 not required to be in thermal equilibrium. The non-equilibrium model more accurately reproduces the effects of subcooled spray, electrical immersion heaters, liquid droplet rainout, and vapor rise within the pressurizer.

The c Te mver response is calculated with the point kinetics model 17 conjunction with thermal feedback effects and explicit reactivity forcing functions. Thermal feedback offects include Doppier reactivity feedback as a function of core average fuel temperature and moderator reactivity feedback as a function of core average fluid temperature (or density). The forcing functions represent -the reactivity contributions l' rom reactor scram and changes in control rod position. The point kinetics 4 model incorporates one prompt neutron group and six delayed neutron groups. Decay heat is calculated based on the 1979 American Nuclear Society (ANS) Standard (23). As recommended by the Standard, an uncertainty allowance of two sigma is applied to ,

the calculated decay heat.

The PPS functions included in the system T-H model are those functions related to the Reactor Trip System (RTS) which are important to the analysis of the reactivity anomaly events. The RETRAN-02 trip control logic is used to initiate an action when a monitored parameter exceeds a specified sotpoint and the specified delay time has passed. In addition to a " manual" reactcr trip signal, the following RTS functions are provided:

e

1. High neutron flux (high and low settings);

2. Overtemperature N-16;

3. Overpower N-16;

4. Low pressurizer pressure; 3-4 I

- . . . - . . _ _ _ _ _ __ _ __..,. _ . _ ., - _ _ . . ~ . _ , _ . . . . _ . . , _.

- . - _ _ . - . - . - - - - . . - _ - ~ _ - - - . _ . _ -

~... ..

5. High pressurizer pressure; and,

6. High pressurizer water level.

The flexibility of the RETRAN-02 control system models allows for the duplication of a wide variety of the CPSES control functions. l The system T-H model utilizes the RETRAN-02 control system capabilities to simulate the functions of the CPSES control systems important to the analysis of the reactivity anomaly events. In addition, the RETRAN-02 control system models are used to provide explicit calculations of parameters important to the analysis. These applications of the RETRAN-02 control system models are described below.  ;

l l

Pressurizer PreAsure Control SVbtam The Pressurizer Pressure Control System functions to maintain a relatively constant RCS pressure during normal plant operation and to assist in limiting pressure perturbations during an upset condition. Pressure regulation is achieved through use of the pressurizer sprays, heaters, and two power-operated relief valves (PORVs). The actions of the control system are based on two different signals. An uncompensated PZR pressure signal is used to govern the operation of one PORV. The remaining control system components utilize a compensated PZR pressure signal to direct their actions.

s Reactor Control System The Reactor Control System functions to maintain the RCS average temperature (Two) within a given tolerance of the programmed average temperature (Tur) by supplying rod drive demand signals to the Rod Control System. The Reactor Control System consists l

3-5 3

.. 4 ,

of two error signal channels, a Tuy-Two deviation signal and a mismatch between the turbine load and the nuclear power signal.

The sum of these two error signals is used by the rod speed programmer to produce a rod speed and direction demand signal for the Rod Control System.

Rod Control Syntqm The Rod Control System controls the motion of the control rod I banks within the core in response to signals from the Reactor Control System or the reactor operator. This system provides reactor power modulation by moving the full length control rod banke in a pre-selected sequence. The Rod Control System also allows manual operation of iridividual control banks.

I i

Nitrocen-16 Monitorina System 1

The Nitrogen-16 (N-16) activity in the primary coolant water is a process parameter used for continuous measurement of the reactor power level. The amount of N-16 activity in the primary coolant is directly proportional to the integrated fast neutron flux throughout the core.

By monitoring the gamma radiation resulting from the decay of the N-16, a means is provided to directly measure the total core power. CPSES uues this process parameter as an input to the Overtemperature N-16 and overpower N-16 protection system functions. The N-16 signal is also used in conjunction with the cold leg temperature to syntheeize an average RCS temperature.

3-6 I

~,. ..

I DNBR Evalua. tion _Model The RETRAN-02 control system models are used to assist in the evaluation of the Departure from Nucloato Boiling Ratio (DNDR).

A DNDR evaluation model is incorporated into tho system T-H model to monitor the DNBR trend during an ovent. This ovaluation model is based on parametric studies performed using the TU Electric core T-H methodology. Any analysis required to demonstrate compliance with the DNBR neceptance critorion is performed using the detailed TU Electric coro T-H methodology.

3.7 Hot Soot Mod 21 The RETRAN-02 computer code is used to predict the T-H response of the fuel at the hottest location within the reactor coro.

9 The hot spot model repreuents a stgmont of a single fuel pin at the location where the peak core power occurs during an ovent. As shown in Figure 3.7-1, a unit cell approach is applied to characterize the fuel pin and rurrounding fluid. The hot spot model used for the reactivity anomaly event analyses is consistent with the model described in the TU Electric T-H methodology report (7). Specific components of the hot spot model that are important to the analysis of the reactivity anomaly events are described below.

Three material regions are used to describe the fuel pellot, the a helium filled gap, and the zircaloy cladding. The cladding and fuel material properties are derived from MATPRO-Vornion 11 (24).

These property tables are consistent with those provided in tho

• system T-H model but cover a greater temperature range to allow for changes that occur as a result of fuel molting. The fuel pellet radial power density profile may be input to provide a l- conservative responso for the parameter of interest.

3-7 J

l l '

The RETRAN-02 option ta calculate the heat generated from an exothermic zircaloy-water (Zr-H io) reaction between the cladding and the coolant is included in the hot spot model. This option utilizes the parabolic rate law of Baker and Just (20).

The RETRAN-02 control system models are used to explicitly calculate the fuel pin surface heat transfer coefficient and the radially averaged fuel temperature and enthalpy. The Thom Subcooled Boiling Correlation (20) or the Bishop-Sandberg-Tong Film Boiling Correlation (25) is used to determine the heat transfer coefficient between the fuel pin and the coolant. The following methodology is used to determine the radially averaged fuel temperature and enthalpy.

Assume the fuel pellet is modelled in the radial direction by n equally spaced concentric rings bounded by n+1 mesh points or nodes. Let the temperature at node i be given oy To where i = 1, 2,..., n+1. The fuel centerline temperature in therefore T,i whereas T,.i is the pellet surface tamperature. It is further assumed that the temperature at any node represents the material temperature of the half region on either side of the node. The radially averaged fuel temperature, Tr , is given by the following equation:

9'

• T, = + Ti ( 21-2 ) + T,,i The fuel enthalpy at node i, hg is obtained for each value of To from a table of enthalpy versus temperature. A similar equation is then used to compute the radially averaged fuel enthalpy, h,:

'1' hi '4n-1 (Eq. 3.7-2) hr n2 ,4 I )* "' N 3-8 3

~ . . ..

I 3.8 Dilution Model The RETRAN-02 computer code in used to predict the system response to a boron dilution ovent while in Modus 3, 4, or S.

The dilution model is a one nodo representation of the RCS. The dilution model utilizes the RETRAN-02 general transport model to calculato the change in the RCS boron concentration as the event progresses.

The RETRAN-02 control system models are used to explicitly simulato portions of the Boron Dilution Hitigation System (DDMS).

The BDMS continually monitors the source range flux detectors to determine if the coro neutron population has increased by a factor of two or more within the previous ten minutes. Should the BDMS detect a flux doubling, a signal is generated to automatically avitch the charging pump suction from the VCT to the RWST.

During Modes 3, 4, and 5, the reactor core is subcritical. It in thereforo necessary to explicitly calculato the reactor kinetics

~

response associated with the addition of positive reactivity.

The RETRAN-02 control system models are utilized to perform these calculations.

s

=

3-9

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Figure 3.6-1 System Thermal-llydraulic Model

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Figure 3.7-1 Ilot Spot Model 1

- JUNCTION O cowrao' vo'uac

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4.0 ANALYSIS METHODOLOGY 4.1 overviev ,

The computer codes and models described in Section 3.0 are utilized in various manners to derive specific input parameters i and event responses for the reactivity anomaly events. This section discusses the TU Electric methodology for applying these computer codes and models to specific calculations.

4.2 Core Physics Mathodoloav i

i i

The core physics parameters relevant to the reactivity anomaly events include reactivity coefficients, kinetics parameters, ,

control rod worths, and power distributions. When an uncertainty is required, the minimum uncertainty applied is that calculated '

with the TU Electric Steady State Physics Methodology (1). i Proper accounting for changes to critical variables and the '

resultant impact on reactivity is important when calculating:the physics parameters.-- Xenon 1and samarium concentration, core average power,-boron concentration, moderator temperature, fuel

,,. temperature and control rod position all affect the reactivity of  :

.,.a given calculation. Therefore, calculations are performed by

, ' holding all parameters except the item of interest constant.

.The majority of the physics parameters are generically applicable to the analysis of reactor transients. The methodology used to calculate these generic parameters is discussed below.

Calculations for event-specific applications are discussed in the pertinent analytical methodology subsection.

4-1

]

., - - - . - . _ = . - . . . - - . . . . - . - . - - - - - - - - - .

Moderator Reactivity Feedback Moderator reactivity feedback is determined using the nodal iore model. The moderator temperature coefficient (MTC) is deterr.ined by performing a series of calculations in which the core inint tem,..rature is increased by an incremental amount and the resultant reactivity change is noted. The pertinent ranges for boron concentration, moderator temperature, core exposure, and control rod insertion are considered for these calculations.

The moderator reactivity feedback may also be expressed as a function of moderator density. The moderator density coefficient is determined by rewriting the MTC, including the uncertainty, in terms of density. The associated moderator density defect is calculated by integrating the moderator density coefficient over-a specific density range.

Doooler Reactivity Feedback The Doppler temperature coefficient (DTC) is also calculated using the nodal core model. These calculations are performed in a manner similar to the MTC calculations, except that the fuel temperature is uniformly increased in each fuel node. A range of

fuel temperatures anticipated to bound the temperatures for the event analyses is selected for the calculations.

Delayed Neutron Parameters t

The CASMO-3 model is used to calculate and importance weight the delayed neutron parameters for each fuel type. The delayed l neutron parameters are then appropriately adjoint flux weighted using the nodal core model. Further discussion of the 4-2 I l

calculation of these parameters can be found in the control Rod Worth Analysis report [2].

Control Rod Worth control rod worths are calculated with the nodal core model utilizing the TU Electric Control Rod Worth methodology. In l addition to the calculation of individual control rod worths,  !

three types of control bank worths are also provided.

i Inserted Worth l The inserted control bank worth is determined by calculating the reactivity difference between the condition of all control banks withdrawn ( ARO) .snd that with the control banks at the Rod Insertion Limits (RIL). Appropriate allowances are made for variations in the axial power distribution.

Scram Reactivity The scram reactivity corresponds to the amount of reactivity available for insertion upon a reactor trip. The available

. scram reactivity is determined by calculating the reactivity l a difference between the condition of the control banks at the l RIL and the condition with all control banks fully inserted (ARI). Typically the minimum scram worth is used in the safety analyses. A_ top-skewed axial power shape is utilized

! in the scram reactivity calculation to maximize the worth of the control banks at the RIL. The available scram l

l- reactivity is further reduced through application of a 10%

l calculational uncertainty and by conservatively assuming 4-3 2

that the highest worth control rod doer not fall into the Core.

Djfferential Worth calculations to determine the differential control bank vorth are performed by incrementally inserting the control banks and calculating the change in reactivity between each incremental insert ion, *

..n uncoctuinLy of 134 1: epplied to the calculated dif?erantial control bank worth.

Spram Reactivity Inser(ign Characterist$ng l

Normalized scram reactivir.r as a function of control rod position ist determined by incrementally inserting all control rods from the ARO position to the ARI position. Normalization is achieved by dividing the worth of all control rods at a specific position by the total worth at the ARI position. The worth of the control rodu during travel through the top half of the core is minimized by using a bottom skewed axial power distribution for the cels:ulation. The normalized scram reactivity as a function of

, control rod position is then converted to normalized scram 1

reactivity versus norcalized drop time using a conservative relationship of control rod position versus time.

i l

Core Power Distribution The core power distributions of interest include: relative axial power shape, hot channel peaking factor (Fw), and total peaking f actor (Fo) . The nodal core model is used to calculate all peaking factors. In addition, the nodal core model is used to 4-4 3

~.. ..

calculate discrete pin powers for the power distribution of interest.

Hot Channel Peaking Factor Fu,is defined as the ratio of the integral of linear power alcag the fuel rod with the highest integrated power to the average fuel rod power. For events in which the radial power distribution is anticipated to be more severe than for nominal operating conditions, an augmentation factor is applied to the best estimate Pui. This augmentation factor is defined as the ratio of the full power design limit for Fut to the maximum calculated Fu, at nominal full power conditions. In general, a single augmentation factor is i l_ calculated for each core design. This augmentation factor is also applied to the cniculated best estimate relative integral power of each fuel pin.

! I

( Total Peakina Factor

, l l

Fq is defined as the maximum local heat flux on the surface l of a fuel rod divided by the average fuel rod heat flux. A best estimate Fq is calculated using the nodal core model.

l This value is conservatively augmented as required on an

(.a event specific basis, A_xial Power Shane.

Axial power shapes are determined in accordance with the TU Electric methodology presented in the power distribution l- control analysis report (3). These axial power shapes are 4-5 3

supplemented, as needed, with axial power shapes calculated using the nodal core model. ,

1 I

4.3 DNB Methodoloov l

4.3.1 General The core T-H model is used to evaluate the event scenarios with respect to DNBR. For most events, a statepoint analysis is adequate, using boundary conditions for core power, system pressure, core inlet temperature, and core inlet flow rate from the system T-H analysis. The system boundary conditions selected for use in the statepoint analysis correspond to the most limiting event conditions with respect to DNB. Power distribution parameters such as the core axial power shape, radial peaking factor, and axial peaking factor are selected to be conservat1~$ with respect to those parameters obtained from the core physics calculations, l

For those situations where the fuel stored energy and thermal inertia effects are significant, the use of a transient calculation may be warranted, using time dependent syatem boundary conditions. The analytical approach employed for the DNBR analysis is determined on a case specific basis.

l 4.3.2 DNBR Limit Lines DNBR limit lines are generated to facilitate the analysis of l various events. The'DNBR limit lines represent a set of statepoints for which the MDNBR is equal to the DNBR design limit.

l 4-6 3:

. ~.. ..

l l

The intent of the CNBR limit lines in to specify the limiting core power as a function of core inlet temperature and pressure.

To accomplish this task, a matrix of core inlet temperatures and pressures is analyzed while the Fui, axial power shape, and core inlet flow rate are held constant. The range of core inlet temperatures and pressures analyzed is selected to bound the expected system responses for the event under consideration. The axial power shape and core inlet flow rate are selected to provide a limiting DNBR response.

DNBR limit lines can be used to calculate the limiting core power as a function of Fui, or vice versa, for a given statopoint.

This calculation utilizes v.he fact that sensitivity studies have demcnstrated that the MDNBR can be exp"essed as a function of the product of the Fui and the core power for a given combination of temperature, pressure, flow r.ste, and axial power shape. Thus, the following relationship may be used to define a limiting F 3s for a given statopoint.

(P

- ne) * ( F 3"y)-

Faw.uw = (Eq. 4.3-1)

- Pn where: Pn = Core power at a statopoint; Pma = DNBR limited core power at F5n, Ts7, and PRn; F$n = Design limit hot channel peaking factor; Tn = Core inlet temperature at e statepoint;

" = RCS pressure at a statopoint; and, PRn Fan.uu = DNBR limited hot channel peaking factor at Pn.

4-7 3

~ . .- - -_ __ -_- -- -- - . -

I 4.4 Steady State Events Methodoloav l

4.4.1 Event Description The events included in this section are those incidents associated with the misalignment of control rods or the misloading of fuel assemblies. These events are not actually transient events, but rather situations that may arise that could potentially degrade the fuel barrier performance.

Statien11v Misalianed Control Rod A misaligned control rod results when any single control rod is out of position with respect to the remaining control rods within the same bank. Misaligned control rods are detected by:

1. Asymmetric power distribution as seen on out-of-coro neutron detectors or core exit thermocouples;

2. Control rod deviation alarms; or,

3. Control rod position indicators.

Misloaded Fuel Assembiv Fuel and core loading errors such as those which can arise from the inadvertent loading of one or more fuel assemblies into improper positions, manufacturing of a fuel rod with one or more pellets of the wrong enrichment, or the manufacturing of a full fuel assembly with pellets of the wrong enrichment will lead to increased heat fluxes if the error results in placing thu fuel in core positions calling for fuel of lesser enrichment. Also included among possible

! 4-8 l

I _ . _ . . _ . _ _ _ _ . . _ . _ . . . , _ _ _ _ , - _ - _ . ,

~

1

~. ..

i core misloading errors is the inadvertent loading of one or more fuel assemblies with incorrect inserts.

To reduce the probability of core loading errors, each fuel assembly is marked with an identification number and loaded in accordance with a specified ccre loading pattern.

Following core loading, the identification number of each assembly loaded in the core will be checked against the desired core loading pattern. Serial numbers read during the core loading pattern verification are recorded on a loading diagram as a check on proper placement of each fuel assembly after the loading is completed.

The power distortion due to any combination of misplaced fuel assemblies which would significantly raise peaking factors should also be readily observable with the in-core flux monitors. In addition to the flux monitors, thermocouples are located at the outlet of about one third of the fuel assemblies in the core. There is a high probability that these thermocouples would also indicate any abnormally high coolant temperature rise. In-core flux

~

measurements are taken during the startup subsequent to every refueling operation.

4.4.2 Innut Assumotiong i The primary input parameters of interest for the statically l misaligned control rod are the hot channel peaking factor and the i-axial power shape. The primary parameter of interest for the l misloading of a fuel assembly is the increase in the hot channel i

peaking factor.

4-9 3

4.4.3 Analvtical Acoroach As stated previously, the mispositioning of control rods or fuel assemblies could potentially result in degradation of the fuel cladding integrity. These situations do not result in any type of transient condition for the RCS and therefore will not i

challenge the pressure acceptanes criterion. In addition, the resultant axial flux shapes and hot channel peaking factors are sufficiently small such that the acceptance criterion for fuel centerline melting is not violated. Thus, the acceptance L criterion of concern for these incidents is the m.dR acceptance criterion.

The basic analytical approach for each of these type events is essentially the same and consists of two separate calculations.

The first calculation involves the determination of the possible values for the parameters of interest for each event scenario.

The second calculation uses these values to perform a DNB analysis to determine the acceptability of the inputs with respect to the DNBR acceptance criterion.

Physics Calculationg Misa11aned Control Rod The most severe control rod misalignment situations, with respect to DNBR at significant power levels, arise-from cases in which one control rod is fully inserted or where a normally inserted control rod is fully withdrawn. The event can occur at any point in the cycle. Therefore, it is evaluated at all operating power levels at the time in the 4-10 I

l 1

[}..

i l

cycle at which the maximum full power Fw occurs for normal i

operation.

k Simplified screening calculations are used to determine 1 which of the single statically misaligned control rods 1 j

results in the greatest hot channel peaking factor. The nodal core model is then used to provide a more detailed f

l calculation of the hot channel peaking factor for the most l

~

limiting case identified by the screening calculations. The l hot channel peaking factor calculated with the nodal core model is conservatively increased using the augmentation factor described in Section 4.2.

l The most limiting normal operating axial power distribution, as identified for the power distribution control analysis

_ methodology (3), typically bounds the axiel power shapes for

[ the statically misaligned control rod scenarios. For those axial power shapes that are more limiting than described above, the core nodal model is used to determine the axial power shape. The effects of an adverse xenon distribution are considered when determining the auial power shape.

i Misloaded Fuel _A.ssembly  :

The types of fuel assembly misloadings considered are those l . which place the most reactive fuel types into the most A

reactive locations within the core, i.e. core locations surrounded by highly reactive fuel. Exchanges between fuel [

assemblies for each region of fuel. loaded into the core are considered.

Full corc screening calculations are used to evaluate the series of representative misloaded fuel assemblies. For l 4-11

each misloaded assembly configuration, the assembly relative powers et the incore detescor locations are examined to determine which misloaded assemblies would be detected.

Those misloaded fuel assemblies which would not be detected by the movable incore system are used to evaluate the DNBR.

Core T-H Calculations The core T-H model is used,'in conjunction with the pertinent physics parameters, to evaluate the core DNB response to the situations arising from the misalignment of a control rod or the misloading of a fuel assembly. These evaluations are performed utilizing the most limiting system conditions for full power steady state operation to ensure that the MDNBR calculated for each scenario remains greater than the DNBR design limit.

4.5 Boron Dilution Methodoloav 4.5.1 Event Lescription The boron dilution event is characterized by the reduction of the RCS boron concentration, resulting in the addition of positive reactivity to the core. The limiting boron dilution event is the addition of inborated, primary grade water from the Demineralized and Reactor dakeup Water System (RMWS) into the RCS through the reactor makeup portion of the chemical and Volume Control System (CVCS). To begin a boron dilution with these systems, the reactor operator must manually initiate the process. Strict

' administrative controls require the reactor operator to closely monitor the process and utilize procedures that restrict the rate and duration of the dilution. A boric acid blend system is 4-12 l

1

available to allow the operator to match the boron concentration of the makeup water to that of the RCS during normal charging.

The principal means of causing an inadvertent boron dilution are the opening of the primary water makeup control valve and failure of the blend system, either by controller or mechanical failure.

The CVCS and RMWS are designed to limit, even under various postulated failure nodes, the potential rate of dilution to values which, with indication by alarms and instrumentaticn, will allow sufficient time for automatic or operator response to terminate t.he dilution. An inadvertent dilution from the RMWS may be terminated by closing the primary water makeup control valve. All expected sources of dilution may be terminated by closing isolat.on valves in the CVCS. To regain the lost shutdown margia (SDM), the RWST isolation valves must be opened to allow the a.idition of 2000 ppm borated water to the RCS.

The status of the RCS is continuously available to the reactor operator in the form of parameter indicators, alarm annunciators, or reactor trip annunciators. Direct monitoring of the following parameters is-available to the reactor operators:

1. Boric acid flow rate;

2. Makeup flow rate;

3. CVCS pump status;

4. RMWS pump status;

5. Indicated Source Range Neutron Flux count rate; a 6. Audible Source Range Neutron Flux count rate; and,

7. OTN-16 turbine runback (at power).

The following alarms may also provide an indication to the reactor operator that a boron dilution event is in progress:

1. Boric acid flow rate deviation (>10% from set value);

2. Makeup flow rate deviation (>10% from set value);

4-13 3

_.._._m _ . _ . . . . _ . . . _~

l 1

l

.3. Highiflux' at shutdown - (Source Range) ; l

4. Neutron- flux doubling' (Source Range) ;.

5. 'Axiallflux difference (250% RTP); q

16. RIL-Low-

7. RIL Low Low; and,-

8. OTN-16=(at power).

The followh., automatic reactor trips may occur during an inadvertent boronLdilution to alert the' reactor operators that-an  ;

event: is in progress: ,

e

+

1. OTN-16; 2.. OPN-16; 3._ Power. Range High Neutron Flux - high; or, l; 4. -Power Range High Neutron Flux - low.

No: single active failure 11n any system or piece of equipment will adversely affect the. consequences of the event.

L

_ 4.5.2 Inout-Assumotigng

-:The : basic- parameters to consider in the performance cf the boron -

I dilution.eventLanalysis are described below.:

Critical Boron Concentrati2D The critical boron = concentration, Camr, is selected to be at the

' maximum anticipated concentration for the operational mode of concern. This selection results in the maximum reactivity insertion rate and thus-minimizes the time to. loss of SDM.

i..

II 4-14 3

~,. ~

The nodsl core model is used to calculate the mode-specific critical boron concentration. For Modes 3, 4, and 5, Ccarr is assumed to be the maximum critical boron concentration for any allowed condition. For Modes 1 and 2, Can is the maximLm ARO critical boron concentration at hot zero power conditions.

Initial Boron Concentration The initial boron concentrawlon, Co, is selected to be the concentration closest to critical based on the shutdown margin assumed for the operational mode of interest. This selection minimizes the time to loss of SDM.

Boron Worth The value of the boron worth used for the dilution event is a function of coolant density and boron concentration. Typically, the boron worth increases as the boron concentration decreases and/or the coolant density increases. The analyses must therefore consider the range of the temperature, pressure, and boron concentration changes expected during the event. The boron worth is selected to maximize the reactivity insertion rate and minimize the time to loss of SDM for the operational mode of interest.

" The nodal core model is used to calculate the boron worth for the range of temperatures, pressures, and boron concentrations anticipated for the fuel cycle design of interest. A series of calculations is performed in which the boron concentration is increased by an incremental amount and the resultant reactivity change is noted.

4-15 2

. . . . . . - _ ~ . . - - _ - . - - - . - . - . - - _ . _ - - . . . . - . . - . .

Shutdown Marain

.The minimum required SDM for the operational mode of. interest is used"for the analysis. . This assumption reduces the time to' Joss of SDM for the event.

3 Dilution Flow Rate The maximum dilution-flow rate is assumed for-eace of the

, analyses. Use of the maximum dilution: flow rate increoses the I

L reactivity insertion rate and thus minimizes the time to-loss of SDM.

Dilution Fluid Temoerature The minimum expected temperature is used for the dilution fluid during the event analysis. The minimum temperature provides the maximum dilution fluid density, thereby maximizing the mass addition-to the RCS and subsequent dilution rate.

I RCS Averace Temoerature The RCS average-temperature is selected to be at the maximum Lexpected value for each mode: of operation. This assumption minimizes.the RCS mass and thus enhances the effect of the

_ dilution.

Purae Volume and Purce Time L

The dilution event is terminated when the boron concentration of the RCS is no longer decreasing. Because the CVCS isolation 4-16 3

~.. ..

i valves from the RWST and VCT are some distance from the RCS, the amount of dilution liquid remaining in the lines and the time required to purge this liquid from the lines must be considered in the event analyses. The maximum purge volume is used in the analysis in conjunction with the dilution flow rate being analyzed to determine the corresponding purge time. The time to loss of SDM is minimized by assuming the maximum time required to purge the charging lines of dilution liquid.

Active RCS Volume The RCS volume is selected to be the minimum expected voluma for the operational mode of interest. The use of a minimum RCS volume provides a limiting system response by reducing the mass available to offset the dilution, thus enhancing the reactivity insertion characteristics.

Modes 1 and 2 During plant operation in Mode 1 or Mode 2, all four RCPs are in operation. The dilution nodel therefore includes the volume associated with the four RCS heat transfer loops as part of the active volume. The liquid volumes associated with the pressurizer, surge line, and RV upper head are conservatively excluded from the active system volume, as a these regions tend to be areas of low flow that are not immediately affected by the event.

Mode 3

- During Mode 3 operation, a minimum of one RCP is in operation. The total system volume is reduced to include 4-17 3

, _ _ _ _. _. . _ -. ~ .- _ . . _ . . _ __ . . _ _

l

- - 1 onlyfthe volume associated with the RV and the active RCS I heat transfer loop. As with the Mode 1 and Mode 2 analyses, the liquid volumes associated with the pressurizer, surge line, and RV upper head-are conservatively xcluded from the

• active volume.

Mode 4 During Mode 4 operation, a minimum of one train of the Residual Heat Removal (RHR) system is in operation. The total system volume is reduced to include only the volume associated;with the RV, the RCS flow loop established j between the inlet and exit of the RHR system, and the RHR volume itself. For the reasons previously stated, the l- liquid volumes associated with the pressurizer, surge line, and RV upper head are excluded from the active volume.

Mode 5 A minimum of one train of the RHR system is in operation during Mode 5 operation. During Mode 5 operation the RCS may be drained to allow for the removal of the RV upper head I

(head removal is actually classified as Mode 6). The total system volume used in the Mode 5 boron dilution analysis-therefore assumes that ths RCS has been drained to the level necessary for mid-loop operation. The total active volume includes the RV volume below the midplane of the inlet and outlet nozzles, the RHR volume, and the volume associated with'RCS flow loop established between the inlet and exit of the RHR system.

l l 4-18 3

~,. <

Mode 6 An uncontrolled boron dilution event cannot occur during this mode of operation. Inadvertent dilution is prevented by administrative controls which isolate the RCS from potential sources of unborated water. Any makeup which is required during refueling will be borated water supplied f rom ti.a RWST.

4 1

4.5.3 Analvtical ADoroach A boron dilution event occurring while the Rod Control System is in automatic does not result in an increase in the RCS pressure.

However, the initiation of a boron dilution event with the Rod control System in manual could result in a power excursion that causes a very slow increase in the RCS temperature and pressure.

The pressurization rate is sufficiently small so as not to challenge the RCS pressure acceptance criterion. Thus, the focus of the mode specific analyses is to demonstrate that the boron dilution event can be terminated prior to the onset of fuel failure. To limit the challenge to the acceptance criteria, it must be demonstrated that sufficient time exists for the reactor operators or BDMS to terminate the event prior to the loss of the SDM. The total time includes allowances for the CVCS purge time, electronic signal processing time, and valve stroke time.

s Mode 1 - Rod Control System in Automatic During Mode 1 operation, the reactor is critical with the control rods either partially or fully withdrawn from the core. During this mode of operation, the limit to which the control rods may a

be inserted into the core is dictated by the control rod insertion limits (RIL), thus ensuring that the minimum required 4-19

]

i SDM is'always available, Should a boron dilution event occur while operating in this configuration, the Rod Control System would slowly insert the control rods to compensate for the addition of-positive reactivity. Tin alarm would be activated when the controls rods approach the RIL, thus alerting the reactor operators to the condition. It can-therefore be concluded that the minimum time to the loss of SDM (and hence required operat'or action) occurs when the boron concentration is at the maximum-and the control-rods are at the RIL. The

"~llowing mathematical expression is used to determine the time required to erode the SDM.

f '

T=gnM'i Co (Eq. 4.5-1)

(cot 1T s v

l l

where: 7 = Time to loss the SDM, see M = Total Active RCS mass,_lbm m = Dilution flow rate, lbm/sec 4 = Initial boron concentration, ppm Camr = Critical boron concentration, ppm The insertion of the control rods to maintain a constant core power level results in:an increase to the core peaking factors.

._ A DNB-analysis is performed to demonstrate that the increased peaking does not result in fuel failures.

Modes 1 and 2 - Rod Control System in genual During operation in-this configuration, the reactor is critical with the-control rods either partially or fully withdrawn from L the core. Should a boron dilution event occur, the core power level, temperature, and pressure would slowly increase due to the addition of positive reactivity. The core power would eventually

~

4-20 i

~,, -

increase to the point that a RIL alarm would be activated or a reactor trip would occur to alert the reactor operator to the situation. In either case, the minimum SDM would still be available. It can therefore be concluded that the minimum time to the loss of SDM (and hence required operator action) occurs when the boron concentration is at the uaximum and the control rods are at the RIL. Equation 4.5-1 is used to determine the time required to erode the SDM.

operation at core power levels in excess of the limiting core power level established for the specific RIL could result in fuel failures. A DNB analysis is performed to demonstrate that the incretsed core power level does not result in fuel failures.

Modes 3. 4. and 5 During these modes of operation, the reactor core is subcritical with the control rods either partially or fully insei'ed into the core. Should a boron dilution event occur, the core neutron population would begin to slowly increase due to the addition of

^

positive reactivity. The neutrun flux would eventually increase to the point that a flux doubling signal would be generated, thus initiating the mitigative actions of the BDMS. The boron dilution model is utilized to determine the actuation time of the BDMS and the time required to erode the SDM.

e 4-21 1

_ - _ . . . - - . - _ . - ~ .... .. - - .

l

4. 6' Control Rod Withdrawal Methodoloav 4.6.1- Event Descriotion A control rod withdrawal event is defined as an addition of reactivity to the reactor core caused by the uncontrolled

' withdrawal of one or more control rods. Such an event could-be caused by operator action or by a malfunction of the Reactor Control System or Rod Control System.

The uncontrolled withdrawal of a control rod results in an immediate increase in.the neutron flux. The rate at which the neutron flux increases is proportional to the rate at which positive reactivity is added to the remator core. The thermal time constant of the fuel precludes-the core heat flux from increasing at the same rate as the neutron flux. Depending on the rate of reactivity addition, the core heat generation rate could exceed the heat removal rate through the SGs, resulting in an increase in the RCS temperature and pressure. In addition, the withdrawal of the control rod (s)-could cause a change to the core axial power profile.

Unless terminated by manual or automatic' action, the increase in core power, RCS temperature, and peaking could eventually lead to fuel-failures. The RPS is designed to terminate an uncontrolled control bank withdrawal, a Condition II occurrence, prior to the-onset _of DNB or fuel melt. The core power rise and increased peaking resulting from an uncontrolled witndrawal of a single control rod, a Condition III occurrence, may be sufficient to induce DNB. For those instances, the RPS functions to minimize-the extent of core damage.

4-22 l

l L

I

~,, .-

l The automatic features of the RPS which function to prevent fuel damage following the postulated event include the following reactor trips:

1. Power range neutron flux;

a. High neutron flux - high setting

b. High neutron flux - low setting

c. High neutron flux rate

~

2. Intermediate range high neutron flux;

3. Source range high neutron flux;

4. N-16 instrumentation;

a. Overtemperature N-16

b. Overpower N-16

5. High pressurizer pressure; and,

6. High pressurizer water level.

No single active failure in any system or piece of equipment will adversely affect the consequences of the event.

4.6.2 Inout Assumotions The basic parameters to consider in the performance of the control rod withdrawal event analysis are:

2 Initial RCS Pressure The initial RCS pressure is selected to maximize the challenge to the acceptance criterion of interest. For event scenarios used to demonstrate compliance with the RCS pressure acceptance

^

criterion, the initial pressure is selected to be the maximum pressure for nominal steady state operation. Conversely, the 4-23 I

l

i l

I minimum RCS pressure is selected for those event scenarios used to demonstrate compliance with the MDNBR criterion.

Initial RCS Core Inlet Temnerature The initial core inlet temperature is selected in a manner similar to that of the initial RCS pressure, i.e. to maximize the challenge to the acceptance criterion of interest. Typically, the initial core inlet temperature is selected to be at the maximum value for nominal steady-state operation.

Initial RCS Flow Rate l

The RCS flow rate is selected to be the minimum expected flow rate for.the number of operational RCPs assumed. During Moden 1 and 2, all four RCPs are in operation. For plausible events initiated from other modes of operation, the number of operating RCPs will be determined on an event specific basis. Use of a minimum flow rate reduces the ability of the coolant to remove the excess energy from the reactor core, thus resulting in a more severe event.

i i

Reactivity Insertion Rate The maximum reactivity insertion rate is typically selected for the event analysis. Those scenarios for which the limiting reactivity insertion rate cannot readily be discerned are analyzed for a spectrum of reactivity insertion rates.

l 4-24 l

I

Reactivity Feedback The reactivity feedback varies in accordance with the core exposure and boron concentration. To ensure adequate fuel integrity protection for all expected core exposures and boron concentrations, the event is analyzed for both the minimum and maximum expected reactivity feedback.

4.6.3 Analvtical Acorcach Physics Parametera Three items are calculated specifically for the Control Bank Withdrawal from Subcritical event. They are the differential control bank worth, the axial power shape, and the radial peaking factor. The calculations are performed at both the beginning and the end of the cycle at hot zero power conditions. The core is assumed to be xenon-free at the beginning of the cycle and to contain hot full power equilibrium xenon at the end of the cycle.

~

The calculation is performed by withdrawing the two sequential control banks with maximum combined worth in 100% overlap. All other RCCA banks are in the fully withdrawn or fully inserted postion depending on the normal withdrawal sequence. The axial power shape and corresponding Fa for a given rod position are evaluated in matched pairs. Combinations of axial power shapes a and radial peaking factors covering the full range of control rod motion associated with the simultaneous withdrawal of the two highest worth sequential control banks are evaluated. The hot channel peaking factors are augmented using the approach discussed in Section 4.2.

a 4-25

]

l The Control Bank Withdrawal at Power event is evaluated at both the beginning and the end of the cycle to bound all operating conditions. Calculations are performed by withdrawing the two sequential control banks with the maximum combined integral worth in normal overlap. All other control banks are positioned according to the normal withdrawal sequence. A control rod position dependent hot channel peaking factor is determined by recognizing that the majority of the increased peaking which occurs at reduced power is due to the control rod insertion. The hot channel peaking factor for this event is set equal to the Fa limit for the maximum normal operating power corresponding to the event control rod position.

System T-H Response Control Bank Withdrawal The system T-H model is used to predict the RCS response to a postulated control rod withdrawal event. The withdrawal of a control bank results in a symmetric insertion of positive reactivity into the reactor core. As such, the function of the secondary side is solely to provide a mechanism for heat removal. The level of model detail required to simulate the function of the SG is minimal.

Thus, the SG can be reduced in detail to a single node without affecting the event consequences.

A spectrum of cases encompassing the plausible range of reactivity insertion rates resulting from a single control bank withdrawal or a sequential withdrawal of two control banks is analyzed. To ensure adequate system protection, all combinations of reactivity feedback and power level are 4-26 1

~.. .

. considered. The limiting cases for reactivity feedback occur at-the extreme conditions corresponding to the most positive MTC in combination with the least negative DTC, and

! the most negative'MTC in combination with the most negative l DTC.

l \

l The OPN-16 reactor trip function provides primary protection against fuel centerline melting during postulated l

Condition II events. Therefore, compliance with the fuel melt acceptance criterion is demonstrated by ensuring the j OPN-16 trip setpoint utilized in the analyses is selected in accordance with the TU Electric setpoint methodology (3).

The event scenario used to demonstrate compliance with the system pressure acceptance criterion employs initial conditions and assumptions that maximize the peak RCS pressure. These conditions and assumptions include:

1. Maximum initial RCS pressure;

2. Maximum initial core inlet temperature; I

3. No credit for the Pressurizer Pressure Control System functioning to reduce the pressure.

l Compliance with the DNB acceptance criterion is demonstrated

[ by performing a detailed DNBR analysis of the limiting event scenario (s), to demonstrate that the MDNBR is greater than the associated DNBR design limit. The system responses for d pressure, core average power, and core inlet temperature for the limiting event scenario are used as boundary conditions to perform the detailed DNBR analysis. The effect of L control rod position on the hot channel peaking factors is considered when performing the detailed DNBR analysis.

l 4-27

I i

Sincle Control Rod Withdrawal The most limiting case for a control rod withdrawn with the control banks at the RIL is evaluated. The value of MTC corresponding to EOC conditions is used to analyze the event, as this tends to maximize the core power rise and minimize the tendency of the increased moderator temperature to flatten the core power distribution. Power distributions within the reactor core are calculated using the nodal core model.

The hot channel peaking factors are then used in the core T-H model to calculate the MDNBR for the event. The DNBR evaluation is performed at the power and coelant conditions at which the OTN-16 reactor trip function would be expected to trip the plant. An Fm limit that produces a MDNBR that is equal to the DNBR design limit is calculated.

A pin census is performed utilizing the nodal core model to determine the number of fuel pins with an integral power that exceeds the limiting Fa. Any fuel pin with an integral power greater than the limiting Fu is assumed to fail. The number of failed fuel pins is used as input to the offsite dose calculation to demonstrate compliance with the radiological consequences acceptance criterion.

4.7 pontrol Rod Dron Methodoloav 4.7.1 Event Description The Control Rod Drop event is initiated by an electrical or mechanical failure which results in one or more control rods from 4-28 2

~.. .

the same group of a given bank dropping to the bottom of the core. The negative reactivity insertion from the dropped control rod (s) causes a prompt reduction in nuclear power followed by a l decrease-in RCS pressure and hot leg temperature. Following a control rod drop event while operating with the Rod Control

{

System in manual, the plant will establish a new equilibrium power level at a value less than or equal to the pre-drop steady 1 state power level.

~

For a control rod drop event occurring while the Rod Control System is in automatic, the Rod Control System detects the drop in core power and initiates a withdrawal of the control banks.

The combination of control bank withdrawal and reactivity feedback causes the core power to increase and possibly overshoot the initial core power level. The core power distribution established following a control rod drop could result in an increase in the Fa. The increase in Fa coupled with the core power increase could potentially lead to fuel failure.

The automatic features of the RPS which function to prevent fuel damage following the postulated event include:

-o

1. Power range neutron flux;

a. High neutron flux - high setting

b. High neutron flux - low setting

c. High negative neutron flux rate

.4 2. N-16 instrumentation;

a. Overtemperature N-16

b. Overpower N-16

3. High pressurizer pressure;

4. High pressurizer water level; and, a 5. Low pressurizer pressure.

4-29 2

. ... - - _- . - - . ._ ..- . ~ . . _ -

No single active failure in any system or piece of equipment:will_-

adversely affect the consequences of the event. ,

l The control rod drop event does not result in a significant RCS pressure rise and, therefore, does not challenge the system pressure acceptance criterion. -Additionally, the increase in core power and hot channel-peaking factor does not challenge the fuel centerline melt acceptance criterion. Therefore, the focus of the analysis is directed on the DNBR acceptance criterion.

i 4.7.2 Inout-Assumotions The basic parameters to consider in the performance of the

-control' rod drop event analysis include:

l~

Initial RCS Pressure The initial RCS pressure is selected to be the minimum pressure consistent with nominal full power operation. This selection tends to maximize the approach to the DNBR acceptance criterion.

1 l

Initial RCS Temnerature The-initial-RCS temperature is selected to be the maximum value

~

consistent with nominal full power operation. Use of the maximum

' temperature tends to maximize the approach to the DNBR acceptance criterion.

l 4-30 J;

~,. ..

Initial RQS, Flow Rate The initial RCS flow rate is chosen to be the minimum expected flow rate consistent with operation of four RCPs. Use of the minimum flow rate reduces the ability of the coolant to remove the excess energy from the reactor core, thus resulting in a more severe event.

Initial Pressurizer Water Level The initial pressurizer water level is selected to be the minimum expected value for the power level being analyzed. Minimizing the initial pressurizer water level tends to result in a more limiting DNBR response for the event.

Doooler Reactivity Feedback The Doppler reactivity feedback is selected to provide the minimum amount of reactivity feedback during the post-drop power

~

increase. This selection produces a more severe event with respect to the DNBR acceptance criterion.

Eycore Detector Tilt 2 The excore detector tilt is a function of the dropped control rod worth and the detector location relative to the dropped control rod location. The excore detector tilt is defined as the ratio of the post-drop indicated power to the initial condition, with values of indicated power normalized to a core average power of one. The value of the excore detector tilt is selected to

• maximize the power mismatch, such that the power overshoot resulting from the withdrawal of the control bank is greatest.

4-31 l

1

..]

The limiting excore detector tilt corresponds to the minimum expected tilt value for the specific dropped control rod worth and core location.

4.7.3 Analvtical Anoroach A control rod drop event can occur at any time during the fuel cycle, at numerous locations throughout the core, and involve one or more control rods from a single bank. It is therefore difficult to characterize the event with a single set of bounding reactivity parameters. Because of this difficulty, individual i control rod drop cases are analyzed using appropriate physics parameters for each case. The steps involved in the analysis of the control rod drop event include:

1. Determination of the potential control rod drops and calculation of the associated physics parameters;

2. Calculation of the system T-H statepoints for each control rod drop scenario; and,

3. Evaluation of the core T-H response for each statepoint using DNBR limit lines appropriate for the event.

l The control rod drop analytical approach is shown schematically in Figure 4.7-1.

Physics Parameters The control rod drop event analysis requires a variety of event L specific physics parameters including dropped control rod worths, l hot channel peaking factors, excore detector tilts, MTCs, inserted control bank worths, and axial power shapes.

1 Each 4-32 3

parameter, with the exception of the axial power shapes, is determined for BOC, EOC, and the core exposure with the a.eatest normal operating Fui. A limiting axial power shape is determined based on the applicable shapes generated during the devt.lopment of the OTN-16 and OPN-16 reactor trip setpoints (3).

All potential control rod drops resulting from a single failure within the Rod Control System are evaluated. For the evaluation, screening calculations are performed to determine the possible sets of dropped control rod worth, hot channel peaking factor, and minimum excore detector tilt. These calculations are performed by assuming full insertion of each control rod combination into an otherwise ARO configuration. More detailed nodal calculations are performed, as required, for the limiting cases to confirm adequate conservatism.

The resulting hot channel peaking factor for each control rod drop scenario is expressed as a ratio of the post-drop peaking to the pre-dron peaking. This retio is multiplied by the maximum allowable Fui to generate a limiting post-drop hot channel peaking factor. The hot channel peaking factor ratios are parameterized as a function of the dropped control rod worth.

Precise calculations of excore detector tilt and RV downcomer density effects require the use of sophisticated methods, such as discrete ordinate transport methods. The use of such methodology is prohibitively expensive to employ on a cycle specific basis.

" Therefore, a spectrum of calculations was performed specifically for CPSES Unit 1 Cycle 1, and an algorithm was developed to address subsequent cycles. The excore detector tilt is approximated by calculating a parameter which is equal to the weighted sum of the relative powers of the eight assemblies nearest the excore detector. The relative powers of the three assemblies closest to the excore detector are weighted four times 4-33 2

4

, ,, E as heavily as the relative powers of the remaining five assemblies. The excore detector tilt is then calculated as the ratio of the post-drop parameter to the pre-drop parameter. A bounding curve of excore detector tilt versus dropped control rod worth is generated. Discrete ordinate transport methods are also used to evaluate the reduction in excore detector response due to the reduced temperature in the RV downcomer.

The inserted control bank worth is calculated in accordance with the methodology described in Se-tion 4.2. The inserted control bank worth is maximized by assuming a top skewed core power distribution.

System T-H Resoonse The system T-H model is used to calculate the system response to a postulated control rod drop event. Modifications are made to the model to incorporate features necessary to perform the control rod drop analysis.

The control rod drop event is analyzed assuming that the Rod Control System is in the automatic control mode. This mode of operation allows the control banks to be automatically withdrawn following the control rod drop, thus maximizing the post-drop power level and the approach to the DNBR acceptance criterion.

The extent to which the control banks are withdrawn following the control rod drop is governed by the duration and magnitude of the indicated power mismatch between reactor power and turbine load, and Tuy-Two. Operation of the Turbine Control System tends to prolong the power mismatch. To enhance the operation of the Turbine Control System, and thus prolong the indicated power mismatch, a conservatively large turbine control valve area is used.

4-34 2

The Nuclear Instrumentation System (NIS), as incorporated into the system T-H uodel, is modified to reflect a conservative excore detector response. As stated previously, increasing the duration and magnitude of the indicated power mismatch results in a more severe event. An undetectable failure of an auctioneering circuit within the Reactor Control System that results in the use of the lowest excore power indication to drive the Rod Control System is conservatively assumed for the analysis. In addition to the dropped control rod induced excore detector tilt, further reductions to the indicated excore power due to the increase in RV downcomer fluid density are also incorporated into the NIS simulation.

Operation of the Pressurizer Pressure Control System is assumed for the analysis. This system functions to suppress the pressure increase during the latter portion of the event, thereby producing more limiting results with respect to the DNBR acceptance criterion. Operation of the pressurizer heaters is not assumed so as to further inhibit the system pressure increase.

e The Steam Generator Water Level Control System is assumed to function normally during the event analysis. Operation of this system allows the main feedwater flow rate to follow the steam flow rate, thus prolonging the power mismatch.

The axial position of the control rod prior to dropping will impact the time required for the control rod to fall to the bottom of the core. To bound the drop time for all control roc positions, an instantaneous reactivity insertion is used.

The system T-H response following a control rod drop is sensitive to the rate of positive reactivity insertion resulting from the

~

withdrawal of control banks by the Rod Control System. The positive reactivity insertion rate is characterized by both the 4-35 I

J

differential worth and total bank worth. To maximize the post-drop power level, the total control bank worth is selected as the maximum expected worth of the Bank D control rods for a given core exposure. A conservative integral worth curve is used to maximize the differential worth of the control bank.

In place of a specific system T-H analysis for each dropped control rod scenario, a parametric approach to the system T-H analysis may be employed for the control rod drop event. This approach utilizes a matrix of control rod drop scenarios to generate a generic set of system statepoints for core average

! power, core inlet temperature, and system pressure. Each event scenario is defined by a different, yet consistent, set of physics parameters. The range of value.s for these parameters is selected to envelop all expected core exposures and cycle designs.

Ior each combination of physics parameters, a corresponding limiting system statepoint (core power, core inlet temperature, and RCS pressure) is determined to coincide with the time of MDNBR. The DNBR evaluation model incorporated into the systen T-H model is used to assist in determining the time of MDNBR.

The limiting control rod drop event occurs during full power operation, because the potential to violate the DNBR acceptance criterion is greatest due to the combination of power overshoot and core peaking.

An interpolation of the generic system T-H statepoints may be performed to determine the statepoint condition for a given control rod drop scenario. This interpolation utilizes the calculated specific physics parameters (dropped control rod worth, inserted control bank worth, and MTC) for a given dropped control rod scenario, as independent variables, to determine the set of system T-H statepoint conditions.

l l

l 4-36 1 l

3l 1

l T. . -- .

l l

l 1

-Core T-H Response The core T-H analysis for the control-rod drop event employs the TU Electric Statistical Combination of Uncertainties (SCU)-

methodology, as described in Appendix A, to derive a set of DNBR limit lines. The SCU methodology is applied to the uncertainties associated with core power, RCS pressure, core inlet temperature, and the hot channel peaking factor. The resultant Uncertainty Factor (UF) is used to reduce the calculated MDNBR-for each combination of core power, core inlet temperature, and RCS pressure.

The calculated system statepoint conditions for core inlet temperature and RCS pressure are used in conjunction with the DNBR limit lines to determine the core power, Pma, that would result in a DNBR equal to the DNBR design limit for the statepoint conditions. The values for Pme, the design limit hot channel f actor (F5H), and the system statepoint power (Pst) are used to derive a DNBR limited -hot channel peaking factor (Fuuiu) based on the relationship expressed in Equation 4.3-1. The calculated value of Fa for the control rod drop scenario (Faut) is then compared to the value of Fm.uu. If the value of Fa.uu is ~

the greater of the two values, it-is concluded that the MDNBR for the scenario analyzed does not violate the DNBR design limit.

1

=

4-37

.]

l i

Figure 4.7-1 Control Rod Drop Analysis Flow Chart CORE PHYSICS TRANSIENT DNBR LIMIT HOT CHANNEL PARAMETERS- STATEPOINT POWER

• MODERATOR -

Tyr -

PDNB(TsT, PR sT, F$n)

TEMPERATURE COEFFICIENT .

PRsT (Posg) * (F$n)

CONTROL BANK .

P ST REACTIVITY

--+. - -a.

• DROPPED CONTROL ROD WORTH

+

EXCORE TILT

+

FaH,sT DNBR DESIGN BASIS (P a)-(Ffu)

Fan.ux =

p FaH,si < F an uM i

4-38 l 3

l 4.8 gontrol Rod Eiection Methodolony 4.8.1 Event Descriotion The ejection of a con *.rol rod from the reactor core is postu.*4ted to be initiated by tua mechanical failure of a control rod mechanism pressure housing. The event is classified as a Condition IV occurrence because the failure of the pressure

~

barrier in the control rod mechanism housing is considered to be incredible. Should such a failure occur, however, the control rod and drive shaft would be rapidly expelled from the reactor crre due to the large pressure differential between the RCS and

.no containment. The reactivity increase due to the ejection of the rod causes a rapid power increase, which is limited by the inherent fuel temperature reactivity character 1ctics of the core.

The event is terminated when the RPS high neutron flux trip setpoint is exceeded and, after an appropriate delay, the remaining control rods fall into the core. The nuclear design of the reactor core and the limits on control rod insertion will limit the potential fuel damage resulting fram a postulated control rod ejection to acceptable levels. No single active failuro in any system or piece of equipment will adversely affect

the consequences of the event.

L I

4.8.2 Inout Assumotio.ng a

The basic parameters to consider in performing the trol rod ejection analysis are discussed below.

!4 4-39 1

i Initial RCS Fluid Conditigna l

I The enre average power response is insensitive to the values chosen for the initial RCS pressure and temperature. Therefore, these values are selected to be within the 1.3rmal steady state operational band consistent with the power level of interest. )

i l

Initial RCS Flow Rate The RCS flow rate is initialized to the minimum expected flow rate to maintain consistency with the hot spot and DNB analyses.

For the full power cases, the flow rate is consistent with the operation of four RCPs. The zero power cases are analyzcd l

I assuming only two of the four RCPs are operating.

Reactivity Feedbagh A minimum amount of reactivity feedback is used for the analysis to produce more severe event consequences. However, the extent of reactivity feedback varies as a function of core exposure and moderator boron concentration. Therefore, the values for the specific parameters affecting the reactivity feedback are selected on a case specific basis.

Effective Delayed Neutron FractioD

The value of 0.n is selected to maintain consistency with the j time-in-life being analyzed. Typically, the minimum value for a l

specified time-in-life is selected to provide a conservative prediction of the integrated full power seconds (FPS) of 4-40 3

. _ - _ , . ~ - _ . _ _ _ - . _ . .

~,. .

operation. This approach allows for the maximum amount of energy to.be produced by the reactor core.

Eiected Control Rod Worth The worth of the ejected control rod is determined based on tFf core exposure and location within the reactor core. The greate-u worth for the ejected control rod is selected far each case analyzed.

4.8.3 Analytical Approach Consistent with the current CPSES licensing basis, the control rod ejcction event is analyzed at four different core conditions for a given reload cycle:

1. BOC-HZP (Beginning of cycle - hot zero power);

2. BOC-HFP (Beginning of cycle - hot full powor);

3. EOC-HZP (End of cycle - hot zero power); and,

4. EOC-HFP (End of cycle - hot full power).

The analytical approach used to evaluate each of those cases involves several different but interdependent calculations. The approach can be divided into the following calculations:

a 1. Core physics input parameters;

2. System T-H response;

3. Hot spot fuel response;

4. Core T-H response;

5. Tuol pin census; and,

6. Offsite radiological dose consequences.

4-41 l

3

The control rod ejuction analytical approach is shown schematically in Figure 4.0-1.

Core Physics Parameters The majority of the core physics parameters are calculated using the methodology discussed in Section 4.2. However, special consideration must be given to the calculation of the following parameterdt l

3. Doppler reactivity feedback;

2. Total scram worth;  ;

3. Ejected control rod worth;

4. Total peaking factor; and,

5. Axial power shape.

Each of these parameters is calculated using the nodal core model. All cases assume that the post-ejected moderator temperature distribution remains unchanged from the pre-ejected distribution. The HZP cases assume no xenon at BOC, and HFP equilibrium xenon at EOC. The HFP cases account for non-equilibrium xenon by assuming a top skewed axial power distribution which maximizes the ejected control rod worth.

Doooler Rqaglivity Feedback During the event, the greatest fuel temperature reactivity feedback occurs at the location where the local power is greatest. Because the importance weighting of a region is a function of the neutron flux, these hot spots are the regions of higher importance. To correct for the increased Doppler feedback associated with the spatial effects of a non-uniform fuel temperature rise, a Doppler Weighting 4-42 3

_ = _ _ _ _ _ . . . . _ = . . _ _ _ _ . _ _ . . . _ _ . - _ _ _ . _ ___ _ _ . . . _ _

~.. .*

Factor (DWF) is calculated. The DWF is the ratio of the post-ejected spatially weighted DTc to the pre-ejected DTc. l The post-ejected spatially weighted DTC is calcalated by l allowing the fuel temperature of each node to change by an  ;

amount proportional to the relative power of the node. The pre-ejection DTC is calculated using the methodology ,

described in Section 4.2.

Scram Reactivity i The calculation of the total scram reactivity for the control rod ejection event is performed in accordance with the methodology described in Section 4.2. Ilowever, the ,

worth of the ejected control rod is excluded from the calculation. All other uncertainties, including the assumption of a stuck control rod, are considered.

Eiected Control Rod Worth

~

An ejected control rod worth is calculated for each control rod which is normally inserted under the conditions being evaluated. Potential misalignments of the ejected control rod are considered for this calculetion. The calculated worth of the ejected control rod is conservatively increased by 15%.

e Total Peakina Factor A best estimate total peaking fact 6r, Fn , is calculated utilizing the post-ejected core conditions, the best

. estimate Pg is increased through multiplicat..on by a' 4-43 2

l augmentation' factor. The augmentation factor for each case is calculated by dividing the maximum allowed total peaking i i

factor by the best estimate total peaking factor calculated in the HFP pre-ejection case.

Axial Power Shade The axial power shape for the control rod ejection analysis is selected to be the post-ejected hot channel power shape 1

j as calculated with the nodal core model. l System T-H ResDonse E

The main thrust of the system T-H analysis is the prediction of the core average power response to a postulated control rod -

ejection event. The system T-H model is used to predict this response using the point kinetics model in conjunction with Doppler weighting.

1 Input pare.neters for each case are selected to result in the production of the_ greatest amount of energy by the core, as measured in terms of full power seconds (FPS).- However, the selection of' input parameters is tempered by-also maintaining consistency with the time in life and initial core power level being evaluated.

The ejection of the control rod is simulated by introducing.the positive reactivity-into the reactor core linearly over a time interval of 0.1 second.

4-44 I

m

~,. .-

}{ot Soot Fuel Resnongg The positioning of the control rods within the reactor core i generally dictates which fuel assemblies will be the most reactive. However, the exposure and enrichment of the fuel assemblies are also contributors to defining the hot assemblies.

The hot spot analysis utilizes material properties and pellet radial power profiles that provide a conservative fuel response for the time in life being analyzed.

The initial radially-averaged fuel temperature is selected to be the maximum for the conditions analyzed in order to provide a conservative estimate of the initial fuel enthalpy. Maximizing the initial fuel enthalpy minimizes the amount of additional energy the fuel pellet can absorb without violating the peak enthalpy criterion.

The hot spot model utilizes the system T-H model 9: erated core average power response to predict the local power response for fuel at the most reactive location within the core. Ideally, the

~

core hot spot analysis would be performed uti , zing not only the core average power history, but also the total power peaking factor (Fq ) history for the control rod ejection scenario. In place of the transient specific Fq history, the hot spot ana]ysis utilizes the pro- and post-ejected Fn values to create a conservative Fq history. This history is_ generated by first assuming that the location. of the pre-ejected peak Fq and the a post-ejected peak Fn remains unchanged. The hot spot Fq history is then created by linearly increasing the value of P q from the pre-ejected value to the post-ejected value over the same time interval assumed for the ejection of the control rod. The local power response for the hot spot is then calculated by multiplying .

the core average _ power response by the assumed Fq history.

4-45 1

In addition to using conservative assumptions for the hot spot power response, conservative assumptions are also used to calculate thu amount of heat transfer from the fuel pin and to estimate the extent of the exothermic zircaloy-water reaction.

To conservatively decrease the ability of the fuel to transfer heat, the fuel pin is assumed to be in DNB shortly after the control rod is ejected from the core. This effect is simulated in the hot spot model by forcing the fuel rod-to-coolant heat. '

tra' 'er regime to switch from subcooled nucleato boiling to film bo l', t. Additional conservatism is applied to the calculation of th heat transfer coefficient through use of a safety factor to reduce the calculated heat flux and the assumption of a [

constant bulk fluid density.

The Zr-H 0 reaction is exothermic, and therefore tends to become self-sustaining if the proper conditions exist. The extent of the metal-water reaction between the_zircaloy cladding and the '

coolant is conservatively estimated by enhancing the probability that a reaction will occur, using the following techniques:

1. The bulk fluid enthalpy of the coolant is linearly increased to 1000 Btu /lbm during the first 0.2 second of the event to provide the steam environment necessary -

for the Zr-H 2 O reaction to occur, and

+

2. The cladding temperature is maximized by assuming ,

contact between the fuel pellet and the inner clad surface as a result of thermal expansion in the fuel.

To simulate this effect in the hot spot model, the fuel rod gap conductance is instantaneously increased to 10,000 Btu /ft*-hr *F at the time the control rod is fully ejected from the reactor core.

4-46 l

l 3

i

. - _ _ , _ . _ . _ _ . _ . _ . _ _ _ _ , . _ _ . . _ _ . _ . . _ . . . ~ . __

~, . .

l l

The application of the above assumptions to the hot spot analysis j results in a very conservative prediction of the fuel enthalpy and temperature distributions. The peak radially-avoraged fuol enthalpy can be directly compared to the acceptance critorion limit to dotonstrato complianco.

The probability of experiencing conter11no fuel molting is  ;

l significant for the control rod ejection ovent. Should fuel i melting occur for any of the casos analyzed, additional analysos are performed with the hot spot model to datormino the maximum

- value for the post-ojected Fn that does not result in fuel molting. This value, Foun3, is used as input to parform a fuel pin consus. Additionally, values for the post-ejected F n are calculated that restrict the extent of fuel molting to individual radial sogments.

1 m i Core T-H Response A core T-H analysis is performed to datormine the minimum hot channel peaking factor, Fouwn, for which DNB is expected to occur.. Inputs to the core T-H analysis include the limiting system T-H analysis conditions for core inlet temperature and RCS pressuro,.and the axial power shape. The limiting Fuuma is then used as input to perform a fuel pin census.

.A RCS Pressuro_RgEppngg i

overpressurization of the RCS is a possibility for this event due to the voiding that occurs in the vicinity of the ejected control rod. -As stated in the CPSES FSAR, this critorion has boon

- addressed generically (26) for CPSES. The generic analysis o a demonstrated that the peak RCS pressure arising from the ejection l

l 4-47

,2-

of a control rod remains less than 110% of the RCS design pressure. Further, analyses performed by other vendors (27) and utilities (28) have also demonstrated that the peak RCS pressure remains less than 110% of the RCS design pressure. Because the RCS pressure needed to produce sufficient stress to violate the AGME Stress Limit "C" is much greater than 110% of the design pressure, it can be concluded that the acceptance criterion for i system pressure will not be violated. Therefore, the overpressure analysis is not performed as part of the TU Electric methodology.

Puol Pin Census consistent with Regulatory Guide 1.77 (14), cladding failure is assumed to occur for any fuel pin experiencing a MDNDR equal to or less.than the DNBR design limit. Additional fission product releases may occur as a result of fuel melting. To assess the extent of fission product release resulting from cladding failure and fuel molting, a fuel pin census in performed. The pin census in divided into two portions: number of pins experiencing DNB and the fraction of the core experiencing fuel melting.  ;

The nodal core model is used to generate the discrete pin powers for the post-ejected core power distribution. The augmentation factor for the hot channel peaking factor is applied to this distribution to increase the discrete pin powers. The total number of' fuel pins experiencing D%B is then determined by totaling the pins having a integral power greater than or equal -

to Fm.om, as calculated in the core T-H analysis.

A separate calculation utilizing the augmentation factor for r n is used to conservatively increase the pointwise power produced in each fuel pin. The number of fuel segments with an Fn greater 4-48 1

than or equal to Fqwn is then datormined. An estimate of the total fraction of molted fuel is calculated by combining the Town consus with the hot spot data corrolating Fq to the radial

! fraction of fuel melt for a singlo pin.

Offsite Radioloalcal Doso Consecuences Offsite radiological dose calculations will be performed utilizing a source term calculated in accordance with the appropriate regulatory guidelinos. The extent of fuel damage will be conservatively estimated for incorporation into the source term calculation.

l l

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c M

O b

as t

Hot Spot w Analysis a o FqrRE Fqugy @

Fqrosi Qn Q o

w M

o 3 Core System T-H Fuel Offsite "

4 Physics Analysis Pin Dose M o Analysis '

Census Analysis o

p, DWF, AK,g n o

Orn N P >

F H Tg F33 m @

Fz o y Core T-H / b Analysis m

w O '

C l n

Dr M

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=

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5.0 EVENT ANALYSES S.1 general The methodology described in Section 4.0 is intended for use in analyzing the postulated reactivity anomaly events of FSAR Chapter 15.4. This chapter provides the analyses of four specific events to demonstrate the application of the methodology. Analyses of the following events are presented

1. Boron Dilution;

2. Control Bank Withdrawal at Power;

3. Control Rod Drop; and,

4. Control Rod Ejection.

A list of common input parametern used for the event analynes is presented in Table 5.1-1. Event specific input paramotors different from those in Table 5.1-1 are identified for each analysis, where applicable.

s i.=

5-1 2 ,

Table 5.1-1 Common Input Parameters l l

Parameter Value Core Power,mm gg, 3411 RCS Thermal Design Flow Rate, gpm/ loop 95,700 Pressurizer Pressure,m psia 2250 Pressurizer Level,m,m % span 60 ,

Pressurizer Sprays initiation pressure, psia 2275 full open pressure, paia 2325 Pressurizer PORV Setpoint, psia 2350 Pressurizor Safety Valve Setpoint, psia 2575 l Steam Generator Water Level,* % span 66.5 RCS Average Temperature,ma 'F 589.2 Scram Worth,* pcm 4000 Control Rod Drop Time to Dashpot, seconds 2.67 RCS Design Pressure, psig 2485 TUE-1 (5,6) DNBR Design Limitm 1,35 Reactor Trip Setpoints:

Power Range Neutron Flux (high), % RTP 118 Power Range Neutron Flux (low), % RTP 35 High Pressurizer Pressure, psia 2460 l

Low Pressurizer Pressure, psia 1860 High Pressurizer Level, % span 100 Overtemperature N-16 ki 1.370307 kg 0.014144 k3 0.000786 overpower N-16, % RTP 118

• nominal l

• 100% RTP l m For demonstration purposes only 5-2 3

~. .

~.. .

5.2 Boron Dilut12D 5.2.1 ADalysis To cover all phases of plant operation, a boren dilution during each oparational modo is considered in this analysis. As identified in Chapter 4, consor$rativo values for the rolovant paramotors aro used. These assumptions result in a conservativo determination of the tino available for opotator or syttom response after detection of a dilution event in progress. The HCS conditions for each modo of operation are provided in Tablo 5.2-1. A list of the input assumptions used 17 the analysis is provided in Table 5.2-2. A plot of the limiting boron worth versus concentration is provided in Figure 5.2-1. A timo of 260.1 seconds is assumed for the total time required to process the BDMS signal, reposit. ion the appropriate valves, and purgo the CVCS linos of unborated water.

5.2.2 Resulta The boron dilution event does not result in violation of any of the event acceptance critoria. The resultant coro power and peaking are within the limits of acceptability for DNB, i.e. the MDNBR is greater than the DNBR design limit. As shown in Table 5.2-3, the time available for automatic or operator a response to an unexpected boron dilution is sufficient to preclude the loss of SDM. Mode specific boron dilution results are discussed below.

5-3 I

)

3

[

Mode 5 - Cold Shutdown In the event of an inadvertent boron dilution while in Mode 5, the source range nuclear instrumentation detects a doubling of the neutron flux by comparing the current source range flux to that of approximately ten minutes earlier. Upon detection of the flux doubling, an alarm is sounded for the operator, and valve ,

movement is automatically initiated to terminate the dilution and >

start boration. The RWST isolation valves are opened such that 2000 ppm borated water is supplied to the suction of.the charging pumps. Subsequently, the cvCS isolation valves to the vcT close in order to terminate the dilution. These automatic actions are carried out in sufficient time to minimize the approach to y criticality and regain the lost SDM. No actions are required of the reactor operators to terminate the event.

Made 4 -

Hot Shutdown ,

The event scenario for Mode 4 operation is similar to that of Mode 5, but extended in time. The source range nuclear instrumentation detects a doubling of the neutron flux and

' initiates the sequence of actions to mitigate the event. These automatic actions are carried out in sufficient time to minimize the approach to criticality and regain the lost SDM. Again, no actions are required of the reactor operators to terminate the event. The difference in event timing is attributed to.the additional active RCS volume and increased SDM.

Mode 3 - Hot Standby The event scenario for Mode 3 operation is very similar to that

-of Mode 4, but of greater duration. The source range nuclear instrumentation detects a doubling of the neutron flux and 5-4 t

l $

L .--.-- _ -.- _ - -.

initiate the sequence of actions to mitigate the event. These automatic actions are carried out in sufficient time to minimize the approach to criticality and regain the lost SDM. Once ar an, no actions are required of the reactor operators to terminat the event. The difference in event timing is attributed to the additional active RCS volume and lower boron worth.

Mode 2 - startun In the event of an unplanned boron dilution while in the Startup mode, the plant status is such that minimal impact results. The i

plant slowly escalates in power until a RIL alarm is sounded and/or a reactor trip occurs on the Power Rango Neutron Flux -

High, low satroint. After receipt of a signal indicating a i

problem, at aast 27.6 minutes are available for the reactor operators terminate the event prior the loss of SDM. The actions the reactor operators involve opening the RWST isolat un valves nd c?.osing the CVCS isolation valves to the VCT.

• Mode 1 - Power With the Rod Control System in manual c.ontrol and no operator action taken to terminate the event, the resultant power and temperature increase cause a RIL alarm to be generated and/or a reactor-trip upon exceeding the OTN-16 trip setpoint or the Power Range Neutron Flux - High, high setpoint. After the receipt of a l

signal indicating a problem, at least 25.7 minutos are available for the reactor operators to terminate the event prior to the loss of SDM. The actions required of the reactor operators involve opening the RWST isolation valves and closing the CVCS

. isolation valves to the VCT.

5-5 J

With the reactor in automatic rod control, the pressurizer level controller limits the dilution flow rate to the maximum letdown flow rate by reducing the charging flow rate to match the letdown flow rate. An inadvertent boron dilution during automatic rod control results in a power and temperature increase such that the Rod Control System slowly inserts the control rods in an attempt to maintain a constant core power. The action of the rod control system results in the actuation of at least three alarms, thus alerting the reactor operators of the situation. These alarms aret

1. RIL - low level alarm;

2. RIL - low-low level alarm; and,

3. Axial flux difference alarm (AI outside target band).

I Given the many alarms, indications, and the inherent slow process of a boron dilution, the reactor operators have sufficient time to function to mitigate the event. For example, the reactor operators have at least 33.8 minutes from the RIL low-low alarm until 1.6% Ak/k is inserted.

5-6

)

~,, .

Table 5.2-1 Modo Dependent RCS Condit. ions Operational Mode Temper 31gtg 11 eat Transfer Los Mode 1 600'F All four RCS Loops (Power)

Mode 2 557'F All four RCS Loops (Startup) 1 Mode 3 557'F Ono RCS Loop i gliot Standby)

~

Mode 4 350'F one Train of RIIR (Hot Shutdown)

Mode 5- 200*F One Train of RilR (Cold Shutdown) 1 I

I 5-7 3

I l

Table 5.2-2 Boron Dilution Input Assumptions i

critical Boron Dilution Active operational Boron Worth SDM Flow Rate RCS Volume Mode _ oom nem/com 1 Ak/k com ft8 l 1 1600 (1) 1.6 127 9000 (Auto) 1 1600 (1) 1.6 167 9000 (Manual) 2 1600 (1) 1.6 167 9000 ,

3 1600 (1) 1.6 167 4800 4 1600 (1) 1.6 167 3950 5 1250 (1) 1.0 167 3550 .

I i-m See Figure 5.2-1 l 5-8 1

3

~,. .

Table 5.2-3 Boron Dilution Analysis Results Time of SDM Operational Alerting Signal, Lost, Mode Sianal seconds seconds 1 RIL 0.0 2029.4 (Auto) 1 RIL or 0.0 '543.7 1

(Manual) Rx trip

• l 2 RIL or 0.0 1659.6 1

Rx Trip

• 3 to* 442.1 878.2 4 tom 393,4 779,3 5 to m 287.5 550.5 l-

.s l

l

• at time of Rx trip the SDM 2 1.6% Ak/k

• Flux Doubling signal (does not include purge time, signal processing time, or valve stroke time) 5-9 3

- - - . . -. . . . . . . , . . . . . . - . _ _ , , - - - . ~ - . ,

Figure 5.2-1 Limiting Boron Worth curves b

I i

l I

R_

k_

A E

S N

O

~

, 8 i

=

8 l -

l l @

n n

• O G

l a w J t

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= D l , o - 8 O o

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k L 8 , , e o e "

n - U 5-10 I-

5.3 Control Bank Withdrawal at Power 5.3.1 Analysis Three cases at initial power levels of 102%, 62%, and 12% of RTP are analyzed using the methodology of Section 4.6. The initial plant conditions are provided in Table 5.3-1.

5.3.2 Essults Figures 5.3-1 through 5.3-4 provide the event response at full power to a control bank withdrawal incident having a large reactivity insertion rate. A reactor trip on high neutron flux occurs shortly after initiation of the event. Because the 4

neutron power response is rapid with respect to the thermal time constants of the plant, the event results in only a small change in RCS temperature and pressure, thereby maintaining the margin to DNB. A sequence of events is provided in Table 5.3-2.

The event responses to a control bank withdrawal incident from full power for smaller reactivity insertion rates are provided in

, Figures 5.3-5 through 5.3-12. Compared to the event response to a large reactivity insertion rate, a reactor trip occurs after a much longer period of time. The increases in RCS temperature and pressure are consequently larger for the smaller reactivity insertion rates. A reactor trip on high neutron flux will result unless the RCS temperature and power increase is sufficient to cause a reactor trip on OTN-16. The reactor trips occur in sufficient time to ensure the MDNBR remains greater than the DNBR l design limit. A sequence of events is provided in Tabic 5.3'2.

i . -The control bank withdrawal from full power requires only two reactor trip functions to provide DNB protection over the entire 5-11 3

.~ ._ _ __ . _. _ - - . - _ . _ _ _ _ _ . _ _ . . _ _ _ _

range of plausible reactivity insertion rates. These functions are the High Neutron Flux and OTN-16 reactor trips. MDNBR as a function of reactivity ir 7rtion rate, for an event initiated from full power operation utilizing either minimum or maximum reactivity feedback, in provided in Figure 5.3-13. In all cases, the MDNBR remains above the design limit.

The event responses to a control rod bank withdrawal initiated from 62% RTP and 12% RTP are provided in Figure 5.3-14 and Figure 5.3-15, respectively. The results are similar to the 102%

RTp case, with the exception that the range over which the OTN-16

] trip is effective is expanded. In each case, the MDNBR remains j above the DNDR design limit.

Many inflection points are visible on the curves of HDNDR versus  !

reactivity insertion rate in the referenced figures. These

! variations are due to the system conditions resulting from I protection and control system action during the event. For example, the curves presented in Figure 5.3-13 provide the following information:

l

1. For high reactivity insertion rates (i.e., between l

1 1.0 pcm/sec and 75.0 pcm/sec), the high neutron flux reactor trip provides DNB protection for the minimum reactivity feedback cases. For these cases, the i

neutron flux level in the core rises rapidly whi)e the core heat flux and RCS temperature lag behind due to i

the thermal capacity of the fuel and RCS fluid. Thus, high MDNBRs during the transient result when the reactor is tripped prior to a significant increase in the core heat flux or RCS temperature. As the reactivity insertion rate decreases, the core heat flux and RCS temperature remain essentially in equilibrium with the neutron flux. The MDNBR during the event 1

j 5-12

~, .

therefore decreases with diminishing reactivity insertion rates.

2. With further reductions in reactivity insertion rate, the OTH-16 and high neutron flux reactor trips become equally effective in terminating the transient (i.e.,

at =1.0 pcm/sec reacitvity insertion rate).

3. As the reactivity insertion rate decreases even further, the nuclear power escalates more slowly and RCS temperature reaches a greater value prior to initiation of a reactor trip. The increase in RCS temperature actually results in greater margin to the DNB limit due to the slope of the OTN-16 equation.

Thus, the MDNBR increases for lower insertion rates.

Because the withdrawal of a control bank is an overpower event, the fuel temperature continually rises until after the occurrence of a reactor trip. As stated previously, the nuclear power escalation resulting from a large reactivity insertion rate is more rapid than the increase in fuel temperature, due to the fuel

~

pin thermal time constant. The outcome is that the core heat flux lags behind the neutron flux response, thus preventing the peak core average heat flux from exceeding the thermal design limit. Even after considering the effects of the control bank withdrawal on the core power distribution, the peak fuel temperature still remains below the rolting temperature.

a For smaller reactivity insertion rates, the core heat flux remains more nearly in equilibrium with the neutron flux. The overpower condition is terminated by the OTN-16 reactor trip before a DNB condition is reached. The peak core average heat flux is maintained below the thermal design limit, even after

. considering the effects of the control rod position on the radial core power distribution.

5-13 2

i l

The reactor is tripped sufficiently fast during the control bank withdrawal event to ensure that the pressurizer safety valves can maintain the RCS pressure below the acceptance limit. The liquid volume remains below the top of the pressurizer, thus assuring '

the integrity of the pressurizer safety valves is maintained.

Figures 5.3-16 and 5.3-17 provide the limiting system responses  ;

for pressurizer pressure and volume, respectively. ,

I l

5-14 1

3 l-

,, . . . . , - - - , _ , , - - - - - , ~ , - . , . , - - - , . - - , - . - - - . , - - . . . , , - - - - - > - - - ,

l l

l Table 5.3-1 Control Bank Withdrawal at Power Input Assumptions

.102% RTP 62% RTP__ _12% RTD Doppler Defect, pcm Minimum' 973.5 973.5 973.5 Maximum 1189.8 1189.8 1189.8 Moderator Density Coefficient, Ak/g/cc Minimum 0.0 0.0 0.0 Maximum 0.43 0.43 0.43 RCS Average Temperature ('F) 594.6 581 7 S67.4 Pressurizer Pressure (psia)(D 2208 2208 2208 Pressurizer Level (% span) 65 51 33.5 RCS Flow Rate (gpm/ loop) 95700 95700 95700 t

l i 80 2280 psia used for overpressure analysis

.a -

I^

5-15

. , . . _ , , , _ . . , m . ~ . . - _ __ _ . _ _ _ - . . . , _ _ . _ . . _ . _ . . _ . _ _ . , _ . _ . , . _.. . _ , _ _ . . . _ . ,

l

. .. l I

Table 5.3-2 Control Bank Withdrawal at Power Sequences of Events Case At 75 pcm/s, Minimum Feedback Event Time, seconds, Initiation of withdrawal o.o Reactor trip (high neutron flux) 1.6 control rods begin to fall' 2.1 MDNBR occurs 3.0 Case B: 25 pcm/s, Minimum Feedback Event Time. seconds Initiation of withdrawal o.0 Reactor trip (high neutron flux) 5.27 Control rods begin to fall 5.77 MDNBR occurs 6.60 case C: 1 pcm/s Minimum Feedback Event Time, seconds Initiation of withdrawal o.0 Reactor trip (OTN-16) 144.5 Control rods begin to fall 146.5 MDNBR occurs 147.0 5-16 l

l 1 l _

Figure 5.3-1 Core Average Power Response (75 pcm/sec, Minimum Feedback)

POWER (fraction) 1.2 - 118. RTP

-~

~' ' ~ ~ ~

1.0 -

THERMAL 0.8 - N 0.6 -

NUCLEAR i 0.4 -

0.2 -

0.0 , , , , , -- --

il 0 1 2 3 4 5 TIME (seconds)

Figure 5.3-2 Pressure Response (75 pcm/ soc, Minimum Fe edbace PRESSURE (psla) 2500l 2400 -

l 2300 - g- -

a 2200 -

2100 -

l 2000 , , , , ,

0 1 2 3 4 5 6 TIME (seconds) l 5-17 1

Figure 5.3-3 Temperature Response (75 pcm/sec, Minimum Feedback)

TEMPERATURE (F) 610 600 -

590 -

580 , . . . i i

0 1 2 3 4 5 6 TIME (seconds)

Figure 5.3-4 DNBR P.esponse (75 pcm/sec, Minimum Feedback)

DNBR 4.0 3.0 -

2.0 -

1.0 . i i i i 0 1 2 3 4 5 6 TIME (seconds) 5-18 I

Figure 5.3-5 Core Average Power Response (25 pcm/sec, Minimum Feedback)

POWER (fraction) 1.2 - .....11.8"R..T..P............................_..

. .. m. ...............................

1.0 -

_ . - g TH8RMAL N

0.8 -

0.0 0.4 - NUCLEAR '

~

0.2 - \m 0.0 , , , , , , , , ,

0 1 2 3 4 5 6 7 8 9 10 TIME (seconds)

Figure 5.3-6 Pressure Response (25 pcm/sec, Minimum Feedbe.ck)

PRESSURE (psla) 2400 -

2300 -

[

a 2200 -

2100 -

2000 , , , , , , , ,

O 1 2 3 4 5 6 7 8 9 10

^

TIME (seconds) 5-19

.]

Figure 5.3-7 Temperature Response (25 pcm/sec, Minimum Feedback)

TEMPERATURE (F) 800 -

'590 -

580 , , , , , , , , ,

0 1 2 3 4 5 6 7 8 9 to TIME (seconds)

Figure 5.3-8 DNBR Response (25 pcm/sec, Minimum Feedback)

DNBR 4.0 3.0 -

2.0 -

- 1.0 , , , , , , , , ,

J 0 1 2 3 4 5 6 7 8 9 10 TIME (seconds) 5-20 4

l l

~,. .-

Figure 5.3-9 Core Average Power Response (1 pcm/sec, Minimu:n Feedback)

POWER (fraction) 1,2 -

1187. RTP 1.0 -

0.8 -

0.0 -

0.4 - NUCLEAR /

0.2 -

TIIERMAL [ \

0.0 , , , , , , ,

0 20 40 60 80 100 120 140 160 TIME (seconds)

Figure 5.3-10 Pressure Response (1 pcm/sec, Minimum Feedback)

PRESSURE (pata) 2400 -

2300 -

' a 2200 -

2100 -

2000 , , , , , , ,

0 20 40 80 80 100 120 140 160 l TIME (seconds) 5-21 l

'3

Figure 5.3-11 Temperature Response (1 pcm/sec, Minimum Feedback)

TEMPERATURE (F')

600 -

590 -

580 , , , , , , ,

.0 20 40 60 80 100 120 140 160 TIME (seconde)

Figure 5.3-12 DNBR Response (1 pcm/sec, Minimum Feedback)

DNBR 4.0 ,

i' 1

3.0 -

l L

2.0 -

l t

1.0 , . .

0 40 80 120 160 TIME (seconds) 5-22 2

1 7, ,.

b Figure-5.3-13 Control Bank. Withdrawal from 102%-RTP ,

MDNBR vs. Reactivity Insertion Rate MDNBR-2.2 4

1,8 -

MINIMUM FEEDBACK MAXIMUM i ~ FEEDBACK 1,4 - OTN-16 y HIGH FLUX OTN  : H1GH FLUX 1,0 , ,, , ,,,,, , , , , , , , , , , , , , , , , ,

-u. -0.1- 1 10 .

100 REACTIVITY INSERTION RATE (pem/sec)

Figure 5.3-14 Control Bank Withdrawal from 62% RTP MDNBR vs. Reactivity Insertion Rate -

MDNBR-2,6 .

2.24 MAXIMUM

-FEEDBACK 1'8 - MINIMUM FEEDBACK

/ >

a. __

\

g, OTN-lo w OTN-10 m HIGH FLUX i- 1.0 . , , , ,,,,,, , , , , ,,,,, , , , , , , , ,

^

0.1 1 10 100 REACTIVITY INSERTION RATE (pem/sec) 5-23

'3 -

l l

l ..

l Figure 5.3-15 Control Bank Withdrawal from 12% RTP MDNBR vs. Reactivity Insertion Rate MDNDR 3.0 -

2.6 - MAXIMUM FEEDDACK 2.2 - MINIMUM i FEEDUACK OTN-16 q 1.8 -

1.4 - C OTN-16 YU 1,0 i i , ,,iii i i i i ,,,,ii , i i ,,,,i 0.1 1 10 100 REACTIVITY INtERTION RATE (pcm/sec)

Figure 5.3-16 Pressure Transient for Control Bank Withdrawal PRESSURE (psia) 2400 --

2200 -

2000 . . i i O 4 8 12 16 20 TIME (seconds) 5-24 3

Figure 5.3-17 Pressurizer Liquid Volume Response for control Bank Withdrawal ~

PRESSURIZER LIQUID VOLUME (cuble feet)

~ ~~~

PRESSURIZER FULL 1500 -

, 1000 -

500 -

0 , , , ,

0 4 8 12 16 20 TIME (seconds)

A l

l 5-25 l

'2

5.4 Control Rod Dron 1

l 5.4.1 Analysis l

The system T-H methodology described in Section 4.7 is used to generate a series of generic statepoints representing the system ,

conditions at the limiting point for DNBR. Statepoints are determined for different dropped control rod worths, ranging from 50 pcm to 750 pcm, at each combination of MTC and control bank worth identified in Table 5.4-1. The excore detector tilt for each statepoint is determined as a function of the dropped rod worth, as shown in Figure 5.4-1. Additional input parameters, as provided in Table 5.4-2, are used for each statepoint l- calculation.

The core T-H methodology discussed in Section 4.7 is used to develop a set of DNBR limit lines for use in the control rod drop event analysis. These limit lines utilize the TU Electric SCU methodology to address the uncertainty associated with the core power, core inlet temperature, system pressure, and hot channel peaking factor. Table 5.4-3 provides the range of application dnd the uncertainty associated with each parameter used within the SCU methodology as applied to the control rod drop event analysis. The DNBR limit lines, Figure 5.4-2, are developed over the same range of applicability as the SCU uncertainty, utilizing a limiting axial power profile.

A set of physics parameters (MTC, dropped control rod worth, inserted control bank worth, and hot channel peaking factor) is generated for each specific control rod drop scenario. The generic statepoints and the DNBR limit lines are then used to evaluate each scenario by comparing the limiting hot channel peaking factor, Fm,uu , to the statepoint hot channel peaking 5-26 i I

l I

3l

~,. ;.'

factor, Fa.sr . I f F a .st is less than Fu.uu, it is concluded that the DNBR acceptance criterion is satisfied, i

5.4.2 Resulta System Response

~

The typical system response at BOC to a control rod drop event initiated from full power is provided in Figures 5.4-3 through

! 5.4-6. This analysis uses an MTC of -5 pcm/'F, a control-bank worth of 300 pcm, and a dropped rod worth of 350 pcm as input.

In. response to the control rod drop, the core average power decreases rapidly, leading to a decline in the hot leg temperature. The decrease in hot leg temperature consequently reduces the energy available for removal through the SGs, resulting in a reduction in SG pressure and temperature and RCS cold leg temperature. The cooling of the RCS produces a decrease in the RCS pressure and subsequent outsurge from the pressurizer.

The indicated NIS power is. influenced by the core nuclear power, excore detector tilt, and downcomer shielding. The combination of increased excore detector. tilt (control rod induced) and increased downcomer shielding (temperature induced) causes the

, -indicated NIS power to be significantly less than the actual core l a power. The mismatch between the indicated NIS power and the turbine power demand, in conjunction with a Tgu-Two mismatch, causes the Reactor Control System to generate a signal to

-withdraw the control bar.k. As the control bank is withdrawn, the core average power begins to increase. The combination of reactivity feedback, control bank worth, and dropped control rod i 4 worth is sufficient to cause the continued withdrawal of the 5-27 I

3 l

control bank such that the core power eventually overshoots the initial value.

The core thermal power response follows the nuclear power response, but because of the heat conduction characteristics of the fuel, the response is much slower. As the core thermal power gradually increases beyond the turbine load, the RCS pressure and temperature begin to increase.

The DNBR response to the event shows that the DNBR increases immediately after the event initiation due to the decrease in core power. As core power begins to recover, the DNBR peaks and then begins to decrease. The time of the MDNBR for the event, i.e., limiting statepoint condition, occurs at approximately the same time as the peak core thermal power. At the time of MDNBR, l the pressurizer pressure and core inlet temperature are 1

increasing but have not yet surpassed their initial values.

The matrix of generic statepoint case results indicate that the T-H statepoint conditions of the control rod drop event become less limiting as the dropped control rod worth is increased. The inability of the combined effects of MTC and control bank withdrawal to overcome the negatAve reactivity contribution-from the dropped centtol rod causes the statepoint power, temperature, and pressure to decrease in magnitude. If the dropped control rod worth is great enough, a reactor trip on low pressurizer pressure will occur.

L The matrix results also indicate that the T-H consequences of the control rod drop event become more limiting as the worth of the inserted control bank is increased. A greater inserted control bank worth allows the Rod Control System to more easily overcome the power reduction induced by the dropped control rod. The increased control bank worth also means a power overshoot is more likely to occur as a result of the control bank withdrawal.

5-28 3

i ,

~.. .

The use of a bounding excore detector tilt curve removes the location and exposure dependency from the calculation of the excore detector tilt. Because of the more limiting tilt at lessor dropped control rod worths, the combination of a low dropped control rod worth and high inserted control bank worth results in the most severe system T-H response. For these cases, the control bank worth is sufficient to compensate for the dropped control rod, but the increased tilt causes the Reactor Control System to generate a demand signal to withdraw the control banks for a longer period of time. By the time the withdrawal demand signal is terminated, the core thermal power has increased to a level significantly greater than the initial power level. The RCS pressure and temperature also tend to increase as the core thermal power increases.

At EOC, the more negative MTC tends to reduce the severity of the statepoint conditions due to the speed and degree with which the moderator feedback is provided. However, the control bank worth is typically greater at EOC, thus overshadowing the benefit of the more negative MTC. Figures 5.4-7 through 5.4-10 provide the typical system response at EOC to a control rod drop event

~

initiated from full power. The analysis utilizes an MTC of

-25 pcm/*F, a control bank worth of 400 pcm, and a dropped control rod worth of 350 pcm.

Cycle Specific Evaluation a

To demonstrate the analytical approach, an analysis for the CPSES Unit 1 Cycle 1 core design is performeo. The control rod drop combinations induced by a single initiating failure are determined along with consistent sets of MTC, control bank worth, dropped control rod worth, and hot channel peaking factor for

. each control rod drop scenario. The result of each scenario is 5-29 l

1 2

' evaluated by comparing the case specific hot channel peaking factor, Fa,s7, to the limiting hot channel peaking factor, Fm.uu.

The comparisons of the hot channel peaking factors for the single failure induced control rod drop combinations at BOC and EOC are provided in Figure 5.4-11 and Figure 5.4-12, respectively. For all cases, the Fws7 is less than Fa.uu, thus satisfying the DNBR acceptance criterion.

l 5-30 3

I Table 5.4-1 MTC and Inserted Control Bank Worth for the Generic Control Rod Drop Statepoints MTC Control Bank Worth (pem)

(pcm/*F) 2_qQ 300 LQQ 500 5_QQ 0 X X X

~5 X X X

-10 X X X

-15 X X X X

-20 X X X X

-25 X X X X

-30 X X X X

~

-35 X X X

-42 X X X "X" indicates combinations used to generate statepoints a

5-31 3

l

.)

Table 5.4-2 control Rod Drop System T-H Analysis Input Assumption...

Parameter _Value_

Core Power, MWni ,

Doppler Defect, pcm 756 RCS Average Temperature (*F) 590.7 Pressurizer Pressure (psia) 2238 Pressurizer Level (S span) 55 RCS Flow Rate (gpm/ loop) 95700 l-l l

l l

5-32 l

l 3 l

~ --

-l Table 5.4-3 i

Control Rod Drop SCU Parameters i

1 VARIABLE RANGE UNCERTAINTY i I

Core Power. 90 to 118 %RTP 12%

-Inlet Temperature 531.5 to 581.5 'F 14*F Pressure 1800 to 2350 psia 130 psi l

Hot Channel Factor 1.49 to 1.86 14%

U

=

5-33 l

3 l

Figure 5.4-1 Bounding Excore Detector Tilt vs. Dropped Control Rod Worth TILT 1.0 0.9 -

0.8 -

0.7 -

0.6 c

, 0.5 '

O 200 400 600 800 DROPPED CONTROL ROD WORTH (pem)

Figure 5.4-2 DNBR Limit Lines TEMPERATURE (F) 560 -

540 -

1800 1900 2000 2100 2200 2260 2350 pata psia psia psia psia psia psia 520 1.0 1.1 1.2 1.3 1.4 1.5 POWER (fraction) 5-34 3

~,. ,

Figure 5.4-3 Core Average Power Response at BOC POWER (fraction)

NUCLEAR 0.8

^

0.0 -

0.4 '

O 20 40 60 80 100 120 TIME (seconds)

Figure 5.4-4 Pressurizer Pressure Response at BOC PRESSURE (psia) 2300 -

2200 -

i a 2100 -

2000 -

1900 O 20 40 60 80 100 120

~

TIME (seconds) 5-35 1

3

+, .,

Figure 5.4-5 RCS Temperature Response at BOC TEMPERATURE (F) 620 HOT LEC 600 -

580 -

560 -

COLD LEG j- 540 -- - '

l O 20 40 60 80 100 120 TIME (seconds)

Figure 5 4-6 DNBR Respo7se at BOC DNBR 4.0 3.(, -

2.0 -

1.0 O 20 40 60 80 100 120 TIME (seconds) 5-36

II

l Figure 5.4-7 Core Average Power Response at EOC POWER (fraction)

\.

.O -

TilERMAL i

0.8 -

0.6 -

0,4 ' ' ' ' '

O 20 40 60 80 100 120 TIME (seconds)

Figure 5.4-8 Pressurizer Pressure Response at EOC PRESSURE (psia) 2300 -

2200 -

2 2100 -

2000 -

1900 O 20 40 60 80 100 120

^

TIME (seconds) 5-37 2

Figure-5.4-9 RCS Temperature Response at EOC TEMPERATURE (F)

HOT LEG 620 600 -

580 -

COLD LEG 560  %

540 O 20 40 60 80 100 120 TIME (seconds)

^

Figure 5.4-10 DNBR Response at EOC DNBR

.4.0 3.0 -

~

2.0 -

1.0 O 20 40 60 80 100 120 TIME (seconds) 5-38 3

c I

' Figure 5.4-11' ~ BOC Fw Comparison P(delta-H)-~

.= 2 A 2.2 -

LIMITING .

8o -

N + ++

4 9 STATEPOINT 1.8 - +g ao g/o

. + * + f*++ag BP -m CU O 0 M0h-O l.

1.6 --

0 B l C 1.4 - i ,

100- 200' 300 400 500 600' 700 m,

DROPPED CONTROL ROD WORTH (pcm)

Figure 5.'4-l'2 EOC Fw Comparison 2 A. F(delta-H) a

2.2 -

2.0 --

LIMITING H-+ + +

+

• -H.

~

-O o 0 0 - pcp 1.6 -

@ o O O-STATEP0lNT

, ' i , ,

- 1.4 i 0 -100 200- 300 400 500 600 700 DROPPED CONT 10L ROD WORTH (pem) i

! 5-39 L

l I

]

I

_ _ _- __ . . _ _ _ _ _ _ - . _ . . _ . _ _ _ _ _ _ _ _ _ _ _ . _ . . _ _ _ _ _ _ _ _ . . . ~ . _ _ ,

l 5.5 Control Rod _Eiection 5.5.1 Analysis Using the methodology of Section 4.8, four core operating configurations are analyzed. The input parameters used in the system T-H response analysis for each case are provided in Table 5.5-1. Sensitivity studies were performed to determine the conservative direction for the mor e important parameters. The results of the sensitivity studien are provided in Appendix B.

The case specific core average pawer response from the system T-H analysis is input to the hot spot model as a boundary condition.

The core average power response for each case is used in combination with the case specific input parameters as provided in Table 5.5-2. The values for the more important parameters are selected in accordance with the sensitivity study results, as presented in Appendix B.

5.5.2 Results EFP Analyses Upon ejection of the control rod, the core average power begins to increase due to the positive reactivity addition. The core

.; average power very quickly exceeds the high flux trip setpoint of 118% RTP, initiating a reactor trip. The delay associated with processing the reactor trip signal allows the core average power to continue increasing. The core average fuel temperature increases during this portion of the event due to the increased power production and the time required +.o transfer the energy from the fuel to the coolant. The core average power continues 5-40

\

l 3

l to increase until sufficient negative reactivity, due to Doppler feedback, is added to offset the ejected control rod worth.

After the appropriate trip signal delay, the shutdown banks fall into the core causing the core average power to decrease. The event is terminated upon insertion of the shutdown banks. The core average power responses for the BOC and EOC HFP cases are provided in Figure 5.5-1 and Figure 5.5-5, respectively.

The normalized core average power response r each HFP case is used to ascertain the corresponding hot spot fuel response. As with the core average power response, the hot spot power increases rapidly following the ejection of the control rod.

Coincident with the hot spot power increase, the coolant enthalpy is increased from its initial value to simulate void formation in the hot channel. In addition, the heat transfer correlation is switched to a film boiling correlation at the time the control rod is fully ejected from the core. These changes to the heat transfer regime and bulk fluid conditions diminish the ability of the fuel pin to transfer energy to the coolant. As a result, the fuel temperature and enthalpy begin to increase rapidly, and continue to increase until the energy removal rate of the coolant

~

exceeds the production rate of the fuej. As heat generation continues to decline, the fuel temperature and enuhalpy begin to decrease. The event responses for the average fuel enthalny, the fuel centerline temperature, and the average fuel temperatur( are provided in Figures 5.5-2, -3, and -4, respectively, for the HFP BOC case. The corresponding plots for the HFP EOC case are provided in Figures 5.5-6, -7, and -8, respectively. A sequence of events for both cases is provided in Table 5.5-3.

As indicated in Table 5.5-4, the peak radially averaged fuel enthalpy for both HFP cases remains well below the acceptance criterion value of 280 cal /gm. No fuel melt is expected for the

. BOC case, while the EOC case is limited to 12.25% by volume at the hot spot.

5-41 1

3

HZP Analyses The core average power response for the HZP cases is similar to that of the HFP cases, with the exception that the high neutron '

flux reactor trip signal is generated at 35% RTP. Figures 5.5-9 and 5.5-13 present the core average power response for the BOC HZP and EOC HZP cases, respectively.

As with tha im' cases, the normalized core average power response for each HZP case is used to ascertain the corresponding hot spot fuel response. The event responses for the average fuel enthalpy, the fuel centerline temperature, and the average fuel temperature are provided in Figures 5.5-10, -11, and -12,

respectively, for the HZP BOC case. The corresponding plots for

+ the HZP EOC case are provided in Figures 5.5-14, -15, and -16, respectively. A sequence of events for both cases is provided in Table 5.5-5.

As indicated in Table 5.5-6, the peak radially averaged fuel enthalpy for both HZP cases remains well below the acceptance criterion value of 280 cal /gm. No fuel melt is expected for either case.

l 5-42 l

l 3

y p L 1  ; .; q.4 r

l TABLE 5.5-1  !

Assumptions for CRE System T-H Analysis Assumption BOC HFP EOC HFP BOC HZP EOC HZP Power Level, fraction of RTP 1.02 1.02 1.0E-9 1.0E-9 Pressurizer. Pressure, psia 2280 2200 2280 2280 RCS Average Temperature, ' F- 593.7 593.7. 557.0 557.0 Average Fuel Temperature, 'F 1239 1239 557.0 557.0 RCS Flow Rate, % nominal 100 100 46 46 Ejected Rod Worth, pcm 200 265 700 950 ut 1 MTC, pcm/*F 0.0 -33.0 0.0 0.0 Trip Worth, pcm 4000 4000 2000 1000 Doppler Weighting Factor 1.0 1.0 2.4 4.5 Beta-Effective, Sg 0.0055 0.0044 0.0055 0.0044 Prompt Neutron Lifetime (t*) , psec 17.5 29.0 17.5 17.5 Time to the Dashpot, seconds 2.67 2.67 2.67 2.67 High Flux Trip Setpoint, % RTP 118 118 35 35 4

TABLE 5.5-2 t i

Assumptions for_CRE Hot Spot Analysis-Assumotion BOC HFP- EOC HFP BOC HZP EOC HZP-t Initial Power Level (fraction'of RTP) 1.02 1.02 1.0E-9 1.OE-9 Initial System Pressure (psia) i 2208 2208 2208 2208 i Initial Average' Fuel Temperature (*F)  ;

2480.4 2480.4 557.0 557.0 Initial-Clad Surface Temperature.(*F) 677.3 677.3 557.0 557.0 Channel Coolant Flow' Rate:(% nominal) 100 100 46 46 Post-Ejected Total Peaking Factor, Fo 4.43 5.41 13.00 17.44 Metal Water Reaction (On/Off) ON ON ON ON  !

r L

I

.i t

t

+

4 e

4

.__A - -

. - .. . - - - __ .._.e

Table 5.5-3 HFP CRE Sequence of Events

.. Time, seconds Event _ BOC EOC Initiation of Rod Ejection 0.00- 0.00 High Flux Trip Signal 0.05 0.04 control Rod FJlly Ejected 0.10 0.10 scram rods begin to drop 0.55 0.04 Peak Fuel Enthalpy 2.89 2.91 Peak Ct Fuel Temperature 3.02 3.07 e

Table 5.5-4 CRE HFP Hot Spot Rep'

~

Peak Fuel Parameter _ BOC ,,JL Radially Averaged Temperature, 'F 3tt4.2 3892.5 Radially Averaged Enthalpy, cal /gn 161.9 192.6

.A Centerline Temperature, 'F 4835.0 5278.9

% Fuel Holt 0.0 12.25 t

N 5-45 -

)

Table 5.5-5 HZP CR2 Sequence of Events Time. seconds _,_ .

Event BOC EOC Initiation of Rod Ejection 0.00 0 00 Control Rod Fully Ejected 0.10 0.10 High Flux Trip Signal 0.29 0.14 Scram rods begin to drop 0.79 0.64 Peak Fuel Enthalpy 3.40 3.33 Peak Co Fuel Temperature 3.91 4.18 i 1

1 Table 5.5-6 l 1

CRE HZP Hot Spot Results '

Peak Puel Parameter Boc EOc Radially Averaged Temperature, 'F 2588.9 3324.6 Radially Averaged Enthalpy, cal /gm 110.9 14 */ . 3 Centerline Temperature, 'F 3003.1 3889.9

% Fuel Melt 0.0 0.0 l-5-46 l

3

~.. .

Figure 5.5-1 80C llFP Core Avorage Power Responso P0h; .1 (fraction) 2.0 -

1.5 -

4 1.0 -

0.5 -

l 0.0 i , , i 0 1 2 4 5 T!WE (seconds)

Figure 5.5-2 BOC HFP Fuel Enthalpy Response ENTHALPY (cal /g)

._ . . . . . . . . . . . . . . . . - . . . . . . . . . . - - - - . . . . . . . . . . . . . ~ . . .

.. ..yg..g.

200 .

i 100 -

0 O 1 2 3 4 5 TlWE (seconds) l 1

5-47

)

i Figure 5.5-3 BOC liFP Fuel Centerline Temperature }tesponse l TEMPERATURE (F) l 6400 l

l l

5000 -

MELTING POINT l

4600 -

l 4200 -

3800 -

3400 '

O 1 2 3 4 6 TIME (seconds)

Figure 5.5-4 BOC HFP Average Fuel Temperature Response TEMPERATURE (F)

MELTING POINT 4400 -

3400 -

2400 O 1 2 3 4 b TIME (seconds) 5-48

[

l I

c. .

Figure 5.5-5 EOC liFP Core Averago Power Responso i POWEll (fraction) l 2.5 l

2.0 -

1.5 -

g,o . >

0.5 -

A 0.0 . i i i

0 1 2 3 4 5 TlWE (seconds)

Figure 5.5-6 EOC llFP Fuel Enthalpy Response

. ENTilALPY (cal /s) 300 ca

...'260.....f7g............................................~...~.....~..~..

200 -

I .h 100 -

l l

0 0 1 2 3 4 6

^

TlWE (seconds) l 5-49

\

i

-2

Figure 5.5-7 EOC HFP Fuel Conterline Temperature Responan TEMPERATURE (F) 5400 6000 -

N MEL

-.... ....T..I N...

G ...

P.O I N..T...............................

4600 -

4200 -

3000 '

3400 O 1 2 3 4 5 TIME (seconds)

Figure 5.5-8 EOC MFP Average Fuel Tomparature Responso TEMPERATURE (F)

MELTING POINT 4400 -

3400 -

2400 O 1 2 3 4 5 TIME (seconds) 5-50 I

)

Figuro 5.5-9 BOC liZP Coro Averago Power Rosponso POWER (fraction) 10 t i I It O.1 0.01 t 0.001 , i i i 0- 1 2 3 4 5 TIME (seconds)

Figure 5.5-10 BOC IIZP Fuel Enthalpy Response ENTilALPY (cal /s)

...g....,.g........................-............................-......-...

200 .

.a 100 -

0 O 1 2 3 4 5 TIME (seconds) 5-51 2

Figure 5.5-11 BOC HZP Centerline Fuel Temperature Response TEMPERATURE (F) 3000 -

2000 -

1000 -

]

O O 1 2 3 4 5 TIME (seconds)

Figure 5.5-12 BOC HZP Average Fuel Temperaturo Response TEMPERATURE (F) 3000 -

2000 -

1000 -

)

0 O 1 2 3 4 5 TIME (seconds) 5-52 3;

~.. ,

Figure 5.5-13 EOC llZP Core Average Power Responso power (fraction) 10 a 1s 0.1g 0.01 3 0.001 i i , ,

0 1 2 3 4 5

T!WE (seconds)

Figure 5.5-14 EOC llZP Fuel Entha'apy Responso l ENTilALPY (cal /g) l l .~f E0.....f7g ea 200 -

8 2 w 100 -

L 0

! O 2 3 1 4 5 TIME (seconds) 5-53 J'

i 1

1 l

1 l

Figure 5.5-15 EOC HZP Fuel Centerline Temperature Response TEMPERATURE (F) 3000 -

2000 -

1000 -

J l

l 0 >

0 1 2 3 4 5 TIME (seconds)

Figure 5.5-16 EOC HZP Average Fuel Temperature Response TEMPERATURE (F) 3000 -

2000 -

1000 -

a 0

O 1 2 3 4 S TIME (seconds) l 5-54 l

l 3

6.O CQ1LC11519]in TU Electric has described an analytical methodology for the reactivity anomaly events presented in Chapter 15 of the Final Safety Analysis Report. This methodology utilizes both neutronic and thermal-hydraulic codes to perform the analyscu. Acceptance critoria are specified for each of the evento discussed in the report, in accordance with applicable regulatory requirements.

Four specific event analyses are provided to demonstrate the application of the methodology. It can be concluded that the TU Electric methodology is an acceptabic means of analyzing the Reactivity Anomaly Events in Chapter 15 of the Final Safety Analysis Report.

A 6-1 J

i

7.0 REFERENCES

1. Edwards, D. J., Kostyniak, L. E., Monger, F. A.,

Rubin, R. M., anct Willingham, C. E., " Steady State Reactor Physics Methodology," RXE-89-003-P, TU Electric, July 1989.

2. Edwards, D. J., " Control Rod Worth Analysis," RXE-90-005, TU Electric, December 1990.

3. Bosma, J. T. and Grace, M. A., " Power Distribution Control Analysis and overtemperature N-16 and Overpower N-16 Trip Setpoint Methodology," RXE-90-006-P, TU Electric, February 1991.

4. Sung, Y. X. and Giap, H. B., "VIPRE-01 Core Thermal-hydraulic Analysis Methods for Comanche Peak Steam Electric Station Licensing Applications," RXE-89-002, June 1989.

5. Giap, H. B. and Sung, Y. X., "TUE-1 Departure from Nucleate Boiling Correlation," RXE-88-102-P, January 1989.

6. Giap, H. B. and Hiltbrand, D. W., "TUE-1 DNB Correlation, Supplement 1," RXE-88-102-P, Sup. 1, December 1990.

7. Lo, S. S., DeVore, C. V., and Boatwright, W. J., " Transient Analysis Methods for Comancho Peak Steam Electric Station 2 Licensing Applications," RXE-91-001, February 1991.

8. Boatwright, W. J., Maier, S. M., and Lo, S. S., " Design Basis Analysis of a Postulated Steam Generator Tube Rupture Event for Comanche Peak Steam Electric Station, Unit 1,"

RXE-88-101-P, March 1988, a

l 7-1 J

9. Code of Federal Regulations, Title 10 - Energy, Revised as of January 1, 1990.

10. ANSI N18.2-1973, " Nuclear Safety Criteria for the Design of Stationary Pressurized Water Reactor Plants."

11. " Comanche Peak Steam Electric Station Final Safety Analysis Report," Amendment 80, November 1990.

12. " Safety Evaluation Report Related to the Operation of Comanche Peak Steam Electric Station, Units 1 and 2,"

NUREG-0797, Supplement No. 24, USNRC, April 1990.

13. ASME Boiler and Pressure Vessel Code,Section III, " Nuclear Power Plant Components," Article NB-7000, " Protection Against overpressure," American Society Mechanical Engineers, 1971.

14. " Assumptions Used for Evaluating a Control Rod Ejection Accident for Pressurized Water Reactors," Regulatory Guide 1.77, USAEC, May 1974.

15. Edenius, M., Ahlin, A., and Forssen, B., "CASMO-3: A Funl Assembly Burnup Program User's Manual," Studsvik/NFA -

88/48, Studsvik of America, September 1988.

16. Umbarger, J. A. and DiGiovino, A. S., " SIMULATE-3: Advanced Three-Dimensional Two-Group Reactor Analysis Code User's Manual," Studsvik/SOA - 89/03, Studsvik of America, November 1989.

17. DiGiovine, A. S., Gorski, J. P., and Tremblay, M. A.,

" SIMULATE-3 Validation and Verification," YAEC-1659-A, Yankee Atomic Electric Company, September 1988.

7-2 1

~.. ,

18. Stewart, C. W., Cuta J. M., Montgomery, S. D., Kelly, J. M.,

Basehore, K. L., George, T. L., and Rowe, D. S., "VIPRE-01:

A Thermal Hydraulic Code for Reactor Cores,"

NP-2511-CCM, Revision 2, Electric Power Research Institute, July 1985.

19. Rossi, C. E., " Acceptance for Referencing of Licensing Topical Report, EPRI NP-2511-CCM, 'VIPRE-01: A Thermal-Hydraulic Analysis Code for Reactor Cores', Volume 1, 2, 3, and 4", NRC Letter to J. A. Blaisdell, UGRA Executive Committes,

Attachment:

Safety Evaluation Report on EPRI NP-2511-CCM VIPRE-01", May, 1986.

20. McFadden, J. H., Peterson, C. E., Paulsen, M. 2 1., and Gose, G. C., "RETRAN-02, A Program for Transient Thermal-Hydraulic Analysis of Complex Fluid Flow Systems,"

HP-1850-CCM-A, Revision 4, Electric Power Research Institute, November 1988.

21. Thomas, C. O., " Acceptance for Referencing of Licensing Topical Reports EPRI CCH-5, 'RETRAN - A Program for One Dimensional Transient Thermal-Hydraulic Analysis of Complex Fluid Flow Systems,' and EPRI NP-1850-CCM ;RETRAN A Program for Transient Thermal-Hydraulic s. 8.s of Complex Fluid Flow Systems,' NRC Letter to T. W. Lci..atz, UGRL Chairman,

Attachment:

" Safety Evaluation Repv*t on RETRAN -

A Program for Tran'sient Thermal-Hydraulic Analysis of a Complex Fluid Flov Systems," September 4, 1984.

22. Thadani, A. C., " Acceptance for Referencing Topical Report EPRI-NP-1850 CCM-A, Revisions 2 and 3 Regarding RETRAN02/ MOD 003 and MOD 004," NRC Letter to R. Furia, GPU Nuclear Corporation,

Enclosure:

" Safety Evaluation by the

. Office of Nuclear Reactor Regulation Relating to RETRAN02, Versions MOD 003 and MOD 004," October 19, 1988.

7-3 a

23. ANSI /ANS-5.1-1979, "American National Standard for Decay Heat Power in Light Water Reactors."

24. "MATPRO-Version 11 (Revision 2), A Handbook of Materials ,

Properties for Use in the Analysis of Light Water Reactor Fuel Rod Behavior," NUREG/CR-0497, TREE-1280, Rev. 2, USNRC, August 1981.

25. Tong, L.S. and Weisman, J., Thermal Analysis of Pressurized Water Reactors, Second Edition, American Nuclear Society, i 1979. l

26. Risher, D. H., "An Evaluation of the Rod Ejection Accident in Westinghouse Pressurized Water Reactors Using Spatial  ;

Kinetica Methodo," WCAP-7588, Revision 1-A, January 1975.

27. Herwig, W. M., " Control Rod Assembly Ejection - Analysis of the CRA Ejection Accident in B&W Pressurized Water Reactor,"

BAW-10150, wanuary 1982.

28. Duke Power Co., " Multidimensional Reactor Transients and Safety Analysis Physics Parameters Methodology,"

DPC-NE-3001, January 1990.

t 7-4 3~

l

.- . -- - ---. ._ - -- - - _ - . . - - - - - - - - - - - ~- - -

l

~..

l 1

Appendix A Statistical combination of Uncertaintien .

Historically, analysos have employed a deterministic approach to address the uncertainty associated with various input parameters.

The deterministic approach assumes that all applicabin paramotor  ;

uncertainties occur nimultaneously in the most adverso direction.

Although this approach is simple to understand and easy to implomont, the result is often a substantial penalty when applying the calculated uncertainty, one accepted method of reducing the excessivo conservatism is to treat the paruneter ,

uncertaintion in a statistical manner.

The TU Electric Statistical Combination of Uncertaintion (SCU) methodology treats the analysis input parameter uncertaintion in a statistical manner, This methodology utilizes a Square Root of the Sum of the Squares (SRSS) with Sensitivity Coefficients technique to statistically combine the effects of the individual ,

parameter uncertainties into a single Uncertainty Factor (UF).

The magnitude of the UF depends on the number of paramotors j considered, the statistical distribution of each paramotor, the _

~

magnitude of each uncertainty, and the significance of each parameter. Qualitatively, the significance of a parameter variable correlates to the sensitivity of the resultant t

calculation to a change in the parameter, e.g., a change of i

1% RTP may have a different significance than a prosauro change of 1%. The most notable difference between a SCO approach and a deterministic approach is in the use of each input paramotor and its uncertainty. The deterministic approach simultaneously combines the nominal value of each parameter with its maximum 4

uncertainty (applied in the most adverse manner) prior to calculating the result. The SCU approach statistically combines the calculated penalties associated with the parameter l . . uncertainties and then applies this uncertainty factor to the result calculated using the nominal paramotor values.

A-1 2

The SCU methodology is an approximate technique for estimating the variance of a multivariate function based on the expansion of a function as a Taylor's series. The technique la " approximate" because only the first order expansion of the Taylor's series is considered. Although including higher order terms can improve the results, the small gain in accuracy generally does not justify the additional computational effort. A brief discussion of the mathematical principle, as applied to the calculation of j the MDNBR uncertainty factor, is provided below. l

)

Assumotions  !

1. The independent variables are mutually exclusive. In other wc-ds, a change in one independent variable does not directly cause a change in another independent variable.

2. The sensitivity coefficient (S,) is constant over the ,

entire range of the independent variable. Typically, the most conservative value is selected for use.

3. The distribution of the variable Y can be represented as a normal distributi'.n The variable Y is defined as the ratio of MDNBRvAn to MDNBRuog.

a. MDNBRgou is defined as the calculated value of the MDNBR when the independent variables are at their respective nominal values, i.e. no uncertainty,

b. MDNBRvAn is defined as the calculated value of the MDNBR when the independent variables include their respective uncertainty.

l A-2 1

3

. -~. -.

l Plathematical Derivation The variable Y can be expressed as a function of n independent variables such that:

Y = f(X 1 , X p Xp ...,X,i ..., X) n (Eq. A-1)

It is further assumed that the distribution of X, can be expressed as a mean of 4,and a standard deviation of oi, where the mean for Xi is the nominal value. The sensitivity factor S, for the it independent variable can be defined as:

BY Si = 1SX i

= 8(In(Y))

6(In(X )) 1 (Eq. A-2)

Xi The value of S, can be interpreted as representing the percentage change in the MDl1BR value resulting from a change in X, with all other parameters being held constant. An estimate of the variance of Y (o8) can be expressed as a function of 44, oi, and S.i The equation for estimating oj is:

f 32 f 12 r 32 f 32 b = S[ b +S 8 b + ..... +S* b (Eq. A-3)

( Py, s 41; ( P2 3 ( Pn, For each independent variable X,, the sensitivity factor, mean, 2 and variance are known. Therefore, the Coefficient of Variation, o,/ yp , for the MD11BR uncertainty factor can be determined. The central limit theorem of statistics states that the probability distribution function for Y will approach a normal distribution even though the distribution of each independent variable X, may not be a normal distribution. The uncertainty factor for

.s A-3 l

l A

1 MDNBR (UPuusn) can be calculated as follows, assuming a 958 percontile for the normal distribution:

UF e33, a py - P(95%) oy (Eq. A-4)

By definition, the value of P(95%) for a normal distribution is 1.645, and the mean value of the variable Y is equal to unity.

The MDHDR uncertainty factor can therefore be expressed in the following manner:

UFamma = 1. 0 - 1. 6 4 5

• o y (Eq. A-5) l l

i A-4 l

, 3 t

l,,-, _ _ - . . . . - _ _

~.. .

Appendix B Control Rod Eiection Sensitivity _Studien B.1 Introduction A series of sensitivity studies were performed for both the system T-H response and hot spot response to quantify the impact of variations in core parameters and modelling assumptions on the predictions for the control rod ejection event. These studies are divided into three categoriest neutronics parameters, thermal-hydraulic parameters, and hot spot parameters. The first ,

-two categories investigate the sensitivity of the system T-H analysis. The third category investigates the sensitivity of the

hot spot analysis. The sensitivity of each parameter used in the study was determined for each of the four cycle conditions, i.e.,

BOC HZP, BOC HFP, EOC HZP, and EOC HFP.

B.2 Neutronics Parameters

~

A sensitivity study was performed on the point kinetics portion of:the system T-H analysis to assess the impact of various neutronics parameters on the predicted core average power response. For the control rod ejection event, the core average power response can be separated into two distinct power responses. The first of these power responses occurs during the a initial stage of the event and involves the prompt power

! response. During this segment of the event, the core average power response is dominated by the production of prompt neutrons resulting from the ejection of the control rod. core average power increases-rapidly until sufficient negative reactivity, due to Doppler feedback, is added to cause a reduction in the core a average power. The core average power decreases to a quasi steady state power level until insertion of the shutdown banks B-1 3

upon reactor trip. The event is sensitive to the core average power response during this time interval. An increase in the peak core average power or in the quasi steady state power level tends to increase the hot spot temperature prediction, thus producing a more severe event.

The latter segment of the event is dominated by the pcwor reduction due to the insertion of the shutdown banks and the effect of moderator reactivity feedback. During this time interval, the core average power level decreases in a manner matching the negative reactivity insertion characteristics of the shutdown banks.

Tablen B-1 and B-2 present a summary of the results of the sensitivity study for the neutronics parameters at ilFp and lizP operating conditions, respectively. The peak normalized core power level is presented for each case along with the number of equivalent full power seconds (FPS) of operation cecurring during the first five seconds of the event. This latter quantity relates to the integrated energy released during the event and is of particular importance for the HZP cases. The integrated FPS are used to demonstrate the event sensitivity to a given parameter.

The first entry in the table for each cycle condition is for a nominal set of input parameters, i.e. no perturbation to any of the neutronics parameters. Comparison of the values for the peak normalized power and energy release for each case with those of the nominal case provides a concise summary of the relative sensitivity of each case.

B-2 I

1 I

7 .

l D.anolor Reactivity Feedbagh The system T-H model incorporates a table of Doppler defect as a function of the coro average fuel temperature. Variations of ilot to the Doppler defect woro analyzed to datormine the event sensitivity.

The HZp casos tended to be more sensitive to the variations in the Doppler defect than the HrP cason. This selectivity is due mainly to the worth of the ejected rod for the HZP cases and the reliance on the Doppler reactivity foodback as the sole means of mitigating the power increase prior to shutdown bank insertion following reactor trip.

Mpilerator ReactivityJAndhAEh Moderator reactivity feedback offects are incorporated into the system T-H model through the use of a constant MTC. Variations of 13 pcm/*r for the MTC were analyzed for the sensitivity study.

The result was a negligible change in the core averago power

~

prediction. A noticeable delay is soon betwoon the time of peak power and the onset of the effect of the moderator foodback.

This delay is a result of the thermal timo constant of the fuel in transferring the onorgy to the moderator.

a Promnt Noutron Lifeting An increase in the value of the prompt neutron lifetime ( t*) is expected to slow the initial rate of increase in core averago power during the event. This offect is the result of the length of tino the average prompt neutron exists prior to absorption.

. Variations of 15 usoc in t' woro analyzed to determino the event B-3 2

sensitivity. Increasing the value of this parameter decreases the predicted peak normalized power although the total energy release shows little sensitivity.

Likewise, a decrease in the value of t' 'is expected to increase the initial rate of increase in the core average power during the  !

event. Because each neutron is absorbed sooner, an increase in '

the predicted peak normalized power results. However, the total energy release shows little sensitivity. The effects of these I changes is more pronounced for the HZP cases than the HFP cases due to the larger ejected control rod worth.

P I

Delaved Neutron Fraction.

The value of the delayed neutron fraction (#m) is a function of fuel exposure. As the number of fissions within the reactor core resulting from neutron absorption in-Uranium-235 decreases, the expected production of delayed neutrons changes.

Changing the value of Sm corresponds to a shift in the parcentage of delayed neutrons in the-core and thereby implies a different core response to the same reactivity perturbation.

Variations of 5% to $m were analyzed to determine the sensitivity of the event consequences. As expected, the lesser values for $m produced the more severe results.

Ejected Control Rod Worth Varying the ejected' control rod worth by 110% provides the greatest impact on the core average power response.. As expected, the increased positive reactivity insertion due to the ejection B-4

'l

~. .'

I l

of a greater control rod' worth results in a greater peak normalized power and integrated FPS.

Control Rod Eiection Time Sensitivity cases were performed to investigate variations of l

-50% and +100% to the ejection time of the control rod. The results show little sensitivity to either case. The principal outcome of increasing or-decreasing the ejection time is to accelerate-or retard, respectively, the time at which the peak core average power is attained. The extent of the sensitivity shown to an increase in the ajection-time is a result of having additional time for the negative feedback 1 effects to influence the core-average power response. 1 1

]

Reactor Trio Delav Time

, This sensitivity study investigated a variation of 10.5 second to

~

the reactor trip delay time._ Because the peak normalized power for the control rod ejection event typically occurs prior to shutdown bank insertion, increasing the reactor trip delay time J will have no impact on the peak normalized power response.

However,_a delay in negative reactivity insertion will slow the core average power rate of decrease, leading to aLgreater energy release throughout the duration of the event.

1 Scram Worth l: :For this sensitivity study, variations of 110% were applied to the scram. worth. As stated above,-the core average power L. response prior to insertion of the shutdown banks will not be-affected. The event sensitivity to the scram worth is minimal.

B-5 n2

~ .__

Time Steo S(Ag The time step sizes used to obtain the RETRAN-02 time-dependent numerical solution are selected according to the expected rate of change of core conditions during the event. The time stop size selected for use in the analysis is sufficiently small to provide numerical stability to the analytical results. An increase to the time step size, by an order of magnitude, was investigated to determine the effect on the analytical results. The results show little sensitivity to this change.

Delaved Neutron Parameters The RETRAN-02 point kinetics model contains a built-in standard set of beta yield fractions and decay constants for use in the reactor kinetics calculations. The sensitivity of this parameter was evaluated by utilizing a set of delayed neutron paramotors representative of expected CPSES reloads. The results show little sensitivity to changes in these parameters.

B.3- Thermal-Hydraulic Parameters Four thermal-hydraulic parameters-were evaluated in assessing the core average power response sensitivity of the system T-H analysis. Table B-3 presents a summary of the results of this sensitivity study.

Initial Core Averace Fuel Temperatur_e The system T-H model utilizes a constant fuel rod gap conductivity throughout the event analysis. The initial gap conductivity is selected to achieve a desired core average fuel B-6 I

~.. .i' temperature.- For the sensitivity study, the fuel rod gap conductivity was varied by 150% from_its nominal value to derive two new initial core average fuel temperatures. The results show little sensitivity to these changes.

Fuel Pin Model The system T-H model describes the fuel pin with six concentric fuel pin regions, a gap region and a cladding region. For the-purpese of the sensitivity study, the number of concentric fuel pin regions was increased from six to ten. As shown in-Table D-3, this change had little impact on the model predictions. In most cases, a slight reduction in the integrated '

FPS:Was observed. This result-is attributed to the increased Doppler' reactivity feedback due to a better prediction of the core average fuel temperature.

Core Inlet Temperatutg ,

The initial core inlet temperature was varied 15.5'F from the nominal value. 'A change of this magnitude showed little effect on the event results, f

System Pressure s

This sensitivity study evaluated a deviation of 130 psi from the nominal system pressure. As with the core inlet temperature, no noticeable-impact on the event response was observed.

P

-A B-7 2

a B.4 Hot Snot Parameters Sensitivity studies woro performed with the hot spot model to i evaluate-the impact of various thermal-hydraulic and neutronics parameters on the temperature and enthalpy response at the hot spot location. Tables B-4 and B-5 present a summary of the sensitivity study results for the hot spot model. The percentage of fuel molted and maximum fuel average enthalpy are compared for each sensitivity case with those predicted for a nominal case.

l Total Power Peakina Ppetor The core average power history predicted with the system T-H model is weighted by a total power peaking factor (F q) before use in the hot spot analysis. The post-ejected Fq for the sensitivity study cases was varied by lot from nominal. The results show a corresponding change in the radially averaged fuel enthalpy and extent of hot spot fuel melting.

Mass Plow Rate j The mass flow rate for the hot spot model was varied by 15%.

These changes produced a negligible difference in the event results.

Puol Pellet Radial Power Distribution The distribution of power radially within the fuel pellet varios l with initial Um enrichment'and fuel exposure. Sensitivity cases l were performed to investigate the response of the hot spot to

! variations in these parameters. The range of evaluation for each B-8

)

L _ . _ . . . _ . . . _ _ . _ _ . _ _ _ _ .. _ __

~,, .*

of those paramotors corresponds to the expected range of variation for the fuel cycle designs being considered for-CPSES.

The core exposure was varied from 0 MWD /MTU to 60 GWD/MTU, while H

the U ' onrichment was varied from 2.4 */o to 4.3 */o. As seen in Tabins B-4 and B-5, the sensitivity was observable but small.

Metal-Water Reaction The hot spot model utilizes the RETRAN-02 calculational model for the determination of the additional energy released due to the roaction betwoon the coolant and the zircaloy cladding. As expected, turning off this option reduces the radially averaged fuel enthalpy, t

'~

Initial Averace Fuel Temperature For the HFP casos, the conductivity of the fuel rod gap is adjusted to achieve a target initial radially averaged fuel temperature. For the sensitivity investigation, the value of the

~

fuel rod gap conductivity was varied to achieve a difference in the initial radially averaged fuel temperature of i100'F. As expected, a change to the initial fuel temperature causes a corresponding change to the peak enthalpy, with the greater fuel temperaturo resulting in a more severe event.

For the HZP cases, the initial fuel pin radially averaged fuel temperature. corresponds to that of the surrounding coolant. A change to the initial fuel temperature would therefore correspond >

to a change in the ir' ial coolant temperature. Because a separato coolant temperature sensitivity case is performed, no specific cases were analyzed for changes to tho fuel pin gap

. conductance at HZP.

B-9 3

Fuel Pin Msiel 1

I The hot spot fuel pin model is described with ten concentric fuel )

regions, a single gap region, and two concentric cladding regions. To test the sensitivity of this model, the number of fuel regions was reduced from ten to five. The results show little impact on the event results.

Coolant Temperature The initial coolant temperature surrounding the hot spot was-varied by 5.5'F for the sensitt ity study. As expected, the temperatures for the cases she. aa a corresponding deviation from the nominal values, but overall the effect was insignificant.

System Pressure The initial system pressure was varied by i30 psi for the sensitivj.y study. This produced a negligible change for the sensitivity cases, i

o i

l B-10

}

b k L i i . i' TABLE B HFP Neutronics Parameters Sensitivity Study'Results BOC HFP EOC HFP Peak Peak Parameter Variation Power __ FPS Power FPS Nominal 1.52 3.83 2.20 3.80 DTC -10% 1.52 3.87 2.21 .82

+10% 1.51 3.79 2.18 li , 7 'I MTC, pcm/*F +3 1.52 3.87 2.20 3.82

-3 1.52 3.79 2.19 3.77

(*, psec +5 1.52 3.83 2.19 3.80 m -5 1.52 3.83 2.21 3.79 b

H $,g -5% 1.56 3.83 2.20 3.80

+5% 1.48 3.83 2.20 3.80 Ejected Control Rod Worth +10% 1.60 3.92 2.48 3.88

-10% 1.44 3.74 1.97 3.72 Ejection Time Doubled 1.51 3.86 2.09 3.81 Halved 1.52 3.81 2.24 3.79 Reactor Trip Delay Time, seconds +0.5 1.52 4.38 2.20 4.29

-0.5 1.52 3.26 2.20 3.30 Scram Worth +1C1 1.52 3.77 2.20 3.74

-10% 1.52 3.89 2.20 3.85 Time Step Size, seconds 0.01 1.52 3.85 2.21 3.82 Reload Betas & Lamdas 1.52 3.84 2.20 3.80

i TABLE B-2 a y

HZPfNeutronics. Parameters Sensitivity Study'Results l BOC HZP EOC HZP 1 .

Peak' Peak-Parameter- Variation Power FPS Power FPS Nominal 8.67; 1.25 54.69. 1.31'  !

DTC -10%- '9.76 1.41 61.75 '1.47

+10%' 7.80 1.12 49.22 1.17 i

'MTC, pcm/*F

+3 '8. V I 14 54.98 ' 1.35

--3 D A9 ' 1.17 54.40 1.26'- 1 I  :

t*, psec +5 6.8 1. 2 <. 41.76 1.30 l m -5 12.C8 1.25 79.43 .1.32 s

$ctr -5% 12.35 . . "! 59.23- 1.32

+5% 5.70 1.17 50.32' 1.29 Ejected Control Rod Worth +10% 18.44 1.52 78.19- 1.49.

-10% 2.72 0.96 35.49 1.13 l Ejection Time Doubled 8.67 1.25 54.80 1.30-Halved 8.67 1.25 54.82 1.31 Reactor Trip; Delay Time, seconds  :+0.5 8.67 1.36 54.69 -1.37

-0.5 8.67 1.13 54.69' 1.25 Scram Worth- +10%' 8.67 1.24 54.69 1.30f l

-10% 8.67. 1.26 54.69 1.31 Time Step Size, seconds 0.01 11.40 1.29 39.76 '.1. 2 5 '

! Reload Betas & Landas 8.51- 1.27 .54.54 1.36 l

s' l

l l

L P L 4 . 4 4

TABLC B-3 Thermal-Hydraulic Pararoeters Sensitivity Study Results BOC HFP EOC HFP Peak Peak Parameter Variation Power FPS Power FPS Fuel Temperaturem +50% 1.52 3.83 2.19 3.78

-50% 1.52 3.83 2.21 3.86 Fuel Pellet, / mesh pts. 10 1.52 3.83 2.20 3.80 Inlet Temperature, *F +5.5 1.52 3.84 2.20 3.80

-5.5 1.52 3.84 2.20 3.80 RCS Pressure, psi +30 1.52 3.83 N/A N/A w -30 1.52 3.83 N/A N/A u

BOC HZP EOC HZP Fuel Temperature * +50% 8.71 1.30 54.77 1.37

-50% 8.63 1.15 54.61 1.21 Fuel Pellet, # mesh pts. 10 8.67 1.25 54.69 1.30 Inlet Temperature, 'F +5.5 8.68 1.25 54.86 1.31

-5.5 8.68 1.25 54.54 1.30 RCS Pressure, psi +30 8.67 1.25 54.69 1.31

-30 8.67 1.25 54.69 1.31

• Fuel rod gap conductance used for variance

TABLE B-4 HFP Hot Spot Sensitivity Study Results BOC HFP EOC HFP Peak Peak Enthalpy  % Fuel Enthalpy  % Fuel Parameter Variation cal /om Melt cal /am Melt Nominal 163.5 0.0 191.0 12.25 Post-ejected Fq +10% 172.9 2.25 208.1 20.25

-10% 154.3 0.0 175.5 6.25 Mass Flux +E% 162.7 0.0 190.0 12.25 to 'a % 164.4 0.0 192.1 12.25 8

5 Radial Power Distribution Uniforn 165.0 0.0 197.9 20.25

2. 4 */o 164.0 0.0 191.6 12.25 3.1 */o lh3.8 0.0 191.4 12.25 Metal-Water Reaction OFF 163.0 0.0 190.0 12.25

/ Radial Fuel Regions 5 162.2 0.0 187.2 9.00 Initial Coolant Temp, 'F +5.5 163.7 0.0 191.0 1~.25

-5.5 163.3 0.0 191.0 12.25 System Pressure, psi +30 163.2 0.0 190.6 12.25

-30 163.7 0.0 191.4 12.25 Initial Fuel Temperature, 'F +100 166.9 0.0 199.7 20.25

-100 160.2 0.0 187.6 12.25

, 4 g p L I 2 2

TABLE B-5 HZP Hot Spot Sensitivity Study Results LOC HZP EOC HZP Peak Peak Enthalpy  % Fuel Enthalpy  % Fuel Variation cal /cm Melt cal /cm Melt Parameter 110.0 0.0 142.2 0.0 Nominal

+10% 119.3 0.0 155.8 0.0 Post Ejected Fq 129.0 0.0

-101 100.8 0.0

+5% 109.4 0.0 141.3 0.0 Mass Flux 143.2 0.0

-5% 110.6 0.0 m

i b Uniform 112.3 0.0 145.5 0.0 Radial Power Distribution 141.8 0.0 2.4 */o 109.7 0.0 141.9 0.0 3.1 */o 109.8 0.0 OFF 109.5 0.0 139.8 0.0 Metal-Water Reaction 5 109.8 0.0 141.9 0.0

/ of Radial Fuel Regions

+5.5 110.2 0.0 142.5 0.0 Initial Coolant Temp, 'F

-5.5 109.9 0.0 142.0 0.0 109.7 0.0 141.7 0.0 System Pressure, psi +30 0.0 142.8 0.0

-30 110.3 Initial Fuel Temperature, *F +100 N/A N/A N/A N/A

-100 N/A N/A N/A N/A l

I

___