ML20078D967

Transient & Accident Methods & Verification

ML20078D967
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Site:	Fort Calhoun
Issue date:	09/30/1983
From:	OMAHA PUBLIC POWER DISTRICT
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OPPD-NA-8303-NP, NUDOCS 8310050252
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Omaha Public Power District Nuclear Analysis Reload Core Analysis Methodology Transient and Accident Methods and Verification OPPD-NA-8303- NP September 1983 s

e 8310050252 830926 PDR ADOCK 05000285 P. .PDR . _ .

f ABSTRACT This document is a Topical Report describing Omaha Public Power District's reload core transient and accident methods for application to Fort Calhoun Station Unit No.1.

The report addresses the District's transient and accident analysis method-ology and its application to the analysis of reload cores. In addition, comparisons of results using the NSSS simulation code to results from exper-imental measurements and independent calculations are provided.

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Proprietary Data Clause This document is the property of Omaha Public Power District (0 PPD) and con-tains proprietary information, indicated by brackets, developed by Combus-tion Engineering (CE). The CE information was purchased by 0 PPD under a proprietary information agreement.

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e-Table of Contents 1

Section Pace l

1.0 INTRODUCTION

AND

SUMMARY

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2.0 CHAPTER 14 EVENTS CONSIDERED IN THE RELOAD CORE ANALYSES 1 2.1 Criteria 1 2.2 USAR, Chapter 14, Safety Analysis Events Not Considered in Reload Core Analyses 3 2.2.1 Malpositioning of Part-Length CEA's 3 2.2.2 Idle-loop Startup Incident 3 2.2.3 Turbine Generator Overspeed Incident 4 2.2.4 Loss of Load 4 2.2.5 Malfunctions of the Feedwater System 5 2.2.6 Steam Generator Tube Rupture Incident S 2.2.7 Loss of Coolant Accident 6 2.2.8 Containment Pressure Analysis 6 2.2.9 Generation of Hydrogen in Containnent 6 2.2.10 Fuel Handling Accident 7 2.2.11 Gas Decay Tank Rupture 7 2.2.12 Waste Liquid Incident 7 2.3 USAR, Section 14, Events Considered in a Reload Core Analysis 7 3.0 TRANSIENT AND ACCIDENT ANALYSIS AND TECHNICAL SPECIFICATIONS 8 4.0 TRANSIENT AND ACCIDENT ANALYSIS MODELS 9 4.1 Plant Simulation Model 9 4.2 DNBR Analysis Models 10 4.3 Application of Uncertainties 11 5.0 TRANSIENT AND ACCIDENT METHODS 12 5.1 CEA Withdrawal 13 l 5.2 Boron Dilution Incident 21 .

5.3 Control Element Assembly Drop Incident 26 5.4 Four-Pump Loss of Flow Event 30 5.5 Asymmetric Steam Generator Event 34 5.6 Excess Load Incident 37 5.7 RCS Depressurization 44 5.8 Main Steam Line Break Accident 47 iii

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i Table of Contents (Continued) l i

Section _Pa ge 5.0 TRANSIENT AND ACCIDENT METHODS (Continued) 5.9 Seized Rotor Event 55 5.10 CEA Ejection Accident 60 5.11 Loss of Coolant Accident 63 6.0 TRANSIENT ANALYSIS CODE VERIFICATION 6.1 Introduction 64 6.2 Comparison to Plant Data 64 6.2.1 Turbine-Reactor Trip 65 6.2.2 Four-Pump Loss of Coolant Flow 67 6.3 Comparison Between OPPD Analyses and Independent Analyses Previously Performed by the Fuel Vendors 69 6.3.1 Dropped CEA 70 6.3.2 Hot Zero Power Main Steamline Break 71 6.3.3 Hot Full Power Main Steamline Break 73 6.3.4 RCS Depressurization 75 6.4 Summanf 77

7.0 REFERENCES

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I 0maha Public Power District Reload Core Analysis Methodology Transient and Accident Methods and Verification

1.0 INTRODUCTION

AND

SUMMARY

This report discusses the methodology the Omaha Public Power District I utilizes to analyze transients and accidents for reload cores. In addition, the report discusses the District's verification of the CE NSSS simulator CESEC for Fort Calhoun Station transients. The pur-pose of this verification is to demonstrate the District's ability to properly utilize the CESEC code.

The District's transient and accident analysis methodology for reload cores is based upon the reanalysis of those Updated Safety Analysis Report (USAR), Chapter 14 events whose consequences may be adversely affected by changes in parameters associated with any reload core.

The USAR Chapter 14 events which must be considered during a reload core analysis are discussed in Section 2.0. Section 3.0 discusses the transient analyses which detennine certain parameters specified in the Technical Specifications. The District's transient analysis models are discussed in Section 4.0. The District's application of these transient analysis models to the various Chapter 14 events is -

discussed in Section 5.0. The verification of the NSSS simulator model used by the District is discussed in Section 6.0. References are provided in Section 7.0.

2.0 CHAPTER 14 EVENTS CONSIDERED IN THE REL0AD CORE ANALYSES This section discusses the criteria utilized to determine if a Chapter 14 event need be considered in reload core analyses. Each event which is not fomally considered in a reload core analysis is discussed and the reasons given for not normally including the event in the reload core analyses. The methodology applied to these events will not be discussed in this report.

2.1 Criteria The criterion used to detemine the events considered in reload core analyses is that changes in various neutronics parameters

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2.0 CHAPTER 14 EVENTS CONSIDERED IN THE RELOAD CORE ANALYSES (Continued) 2.1 Criteria (Continued) adversely effect the safety analyses of these events. The core parameters considered are the pin peaking factors, FR a nd Fxy >

the Moderator Temperature Coefficient (MTC), the Fuel Tempera-ture Coefficient (FTC) or Doppler Coefficient, the baron concen-tration, the inverse boron worth, the neutron kinetics parame-ters, the CEA reactivity worth and the cooldown reactivity asso-ciated with a steam line break. If these parameters change such that the previously reported results for a Chapter 14 event are no longer conservative, then this event must be re-analyzed. If these parameters are conservative with respect to the values assumed in the referenced safety analyses, the cri-teria of 10 CFR 50.59 are met and this event is not reanalyzed.

If a change in some of the parameters may cause the results of a safety analyses to be nonconservative, the event is reanal-yzed. If the criteria for the event are still met, then the requirements of 10 CFR 50.59 are satisfied. The event is re-ported as being reanalyzed and that it has been detemined that no unreviewed safety question exists for the event. In some cases it may be possible that an event is reanalyzed and it is detemined that an unreviewed safety question exists. In these cases the analysis for these events are submitted. In addi-tion, any safety analyses which are perfomed as a result of a change in the_ Technical Specifications are reported as part of the supporting documentation for a Facility License Change.

Criteria not directly associated with the reload core but which may be considered in a reload analysis are changes to plant sys-tems which would take place during a refueling and would first 's be utilized during the operation of the subsequent core. In cases, where either physical modifications or modifications in operating procedures are made such that they do impact the safety analyses, the results of the revised safety analyses are

! reported in a reload core analysis. This methodology report does not consider the methodology that is required to analyze i all events which could be affected by this criteria, rather, if 2 l I

2.0 CHAPTER 14 EVENTS CONSIDERED IN THE RELOAD CORE ANALYSES (Continued) 2.1 Criteria (Continued) submittals are made which require analyses of events other than those discussed in this report, revisions to this methodology report will be made to incorporate the methodology used for these events.

2.2 USAR, Chapter 14, Safety Analysis Events Not Considered in Re-load Core Analyses This section discusses the USAR, Section 14, safety analyses which are not normally considered in a reload core analysis.

The USAR section is discussed and the reasons for not including it in the scope of these analyses is discussed. Typically, the reasons for not analyzing these events are that the operating modes considered in the events are no longer allowable at Fort Calhoun Station, the event is not associated with any core parameters or the event is analyzed by a fuel vendor for the District. 1 2.2.1 Malpositioning of Part-Length CEAs This event is not analyzed in the reload core analysis because the use of the part-length CEAs is prohibited by the Technical Specifications. In addition, the drop of a part-length CEA is less severe than the drop of a full-length CEA.

2.2.2 Idle-Loop Startuo Incident This event is not analyzed because part-loop operation is not penaitted by the Fort Calhoun Technical Specifications.

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2.0 CHAPTER 14 EVENTS CONSIDERED IN THE RELOAD CORE ANALYSES (Continued) 2.2 USAR, Chapter 14, Safety Analysis Events Not Considered in Re-load Core Analyses (Continued) 2.2.3 Turbine Generator Overspeed Incident This event is an analysis of the consequences of a tur-bine wheel failure and is unrelated to any reload core changes.

2.2.4 Loss of Load The loss of load to both generators is assessed to determine if:

A. The pressurizer safety valves limit the reactor coolant system pressure to a value below 110% of design pressure (2750 psia) in accordance with Section III of the ASME Boiler and Pressure Ves-sel Code, and sufficient thennal margin is main-tained in the hot fuel assembly to assure that Departure from Nucleate Boiling (DNB) does not occur throughout the transient. This event is not analyzed with respect to the first criteria since the relief capacity of the pressurizer safety valves does not change and the initial energy contained in the reactor coolant system will not change unless power level is raised above 1500 MW or the reactor coolant system inlet tenperature is significantly increased. Section 14.9 of the USAR reports that the DNBR for the ,

loss of load transient never decreases below the initial value considered in the analysis. There-fore, it is concluded that any change in a para-meter which could Effect the DNBR for this event would much more significantly effect other events

and that it is not necessary to analyze this

event with respect to DNBR criteria.

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2.0 CHAPTER 14 EVENTS CONSIDERED IN THE REL0AD CORE ANALYSES (Continued) 2.2 USAR, Chapter 14, Safety Analysis Events Not Considered in Re-load Core Analyses (Continued) ,

2.2.4 Loss of Load (Continued)

The loss of load to one steam generator is discussed in this methodology report as one of the asymmetric steam generator transients.

2.2.5 Malfunctions of the Feedwater System The analyses which are reported in USAR, Section 14.10 Malfunctions of the Feedwater Systen, are the total loss of feedwater flow and the loss of feedwater heating. The results of the total loss of feedwater flow show that the minimum DNBR does not decrease below its initial steady state value and that no safety limits are approached during the event. Therefore, this event is not reanalyzed in a reload core analysis.

The loss of feedwater heating is the most adverse feedwater malfunction in tenns of cooling on the RCS.

This event, like the excess load event, is more limit-ing at EOC. This event has the same effect on the pri-mary system as a small increase in turbine demand which is not matched by an increase in core power. As a re-sult, the DNBR degradation associated with this event is less severe than that for the excess load where a large effective increase in turbine denand is analyzed.

The excess load event analysis is reported elsewhere in this document.

2.2.6 Steam Generator Tube Ruoture Incident i

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The steam generator tube rupture incident is analyzed to detennine if the off site dose acceptance criteria of 10 CFR Part 100 is met. The analysis is a radio-5

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2.0 CHAPTER 14 EVENTS CONSIDERED IN THE RELOAD CORE ANALYSES (Continued) l 2.2 USAR, Chapter 14, Safety Analysis Events Not Considered in Re-l 1oad Core Analyses (Continued) i l 2.2.6 Steam Generator Tube Rupture Incident active material release analysis based upon 17. failed fuel within the core. It is not dependent upon any reload core analysis related parameters, therefore, it is not analyzed in the reload core analysis. In the future, the steam generator tube rupture incident anal-ysis may be verified for high burnup fuel.

2.2.7 Loss of Coolant Accident The loss of coolant accident as reported in USAR, Sec-tion 14.15, is analyzed for the District by ENC and CE.

The large break analysis was perfomed by EdC, the small break analysis was perfomed by CE. The District confims the assumptions used in these analyses are valid for each reload core. If reanalysis is required, the reanalysis is done by a nuclear fuel vendor. The District does not perfom any loss of coolant accident analyses.

2.2.8 Containment Pressure Analysis Containment pressure analysis is dependent upon the ini-tial liquid mass and energy contained in the primary or secondary system. Since these parameters do not change when the core is refueled, the contaiment pressure analysis is not done in a reload core analysis.

• 2.2.9 Generation of Hydrogen in Containment The generation of hydrogen in containment analysis is independent of any reload core parameters, therefore, the analysis is not perfomed during the course of a reload core analysis.

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r 2.0 CHAPTER 14 EVENTS CONSIDERED TN THE RELOAD CORE ANALYSES (Continued) 2.2 USAR, Chapter 14, Safety Analysis Events Not Considered in Re-load Core Analyses (Continued) 2.2.10 . Fuel Handlina Accident The fuel handling accident is a function of the isotopic inventory contained in the fuel pins. This is not nomally considered in a reload core analysis, however, it may be necessary to reconsider this analyses for high burnup fuel.

2.2.11 Gas Decay Tank Rupture The gas decay tank rupture is independent of any para-meters associated with refueling the core. Therefo re, the analysis is not perfomed during a nomal reload core analysis.

2.2.12 Waste Liouid Incident The waste liquid incident analysis is not affected by refueling the core. Therefore, the waste liquid inci-dent analysis is not perfomed in the course of a nomal reload core analysis.

2.3 USAR, Section 14, Events Considered in a Reload Core Analysis The reload core analysis consists of analyzing several events which are considered in the USAR and two events which previous--

i ly were not analyzed in the USAR. These events are analyzed in ,

accordance with the criteria discussed in this report and to l detemine if an unreviewed safety question exists for a reload j core. The USAR Chapter 14 events considered in a reload core l analysis are the Control Element Assembly Withdrawal (CEAW) in-cident, the boron dilution incident, the Control Element Assem-bly (CEA) drop incident, the loss of coolant flow incident, the excess load incident, the steam line break accident, the CEA 7

t 2.0 CHAPTER 14 EVENTS CONSIDERED IN THE RELOAD CORE ANALYSES (Continued) 2.3 USAR, Section 14, Events Considered in a Reload Core Aralysis (Conti nued) ejection accident and the seized rotor accident. In addition, analyses are perfomed for incidents resulting from the malfunc-tion of one steam generator and for the RCS depressurization incident. The analysis for each of these events will be dis-  !

cussed in detail in Section 5.0 of this report.

3.0 TRANSIENT AND ACCIDENT ANALYSIS AND TECHNICAL SPECIFICATTONS Results of transient and accident analyses are used in the Technical Specifications in two ways. The first way is that values from the Technical Specifications are included in the initial conditions of the transient analyses. These Technical Specifications guarantee that the various transient and accident analysis acceptance criteria will not be exceeded if the reactor is operated within the bounds of these Technical Specifications. Technical Specifications of this type include the limits on F R Exy, the PDIL and the Moderator Temper-ature Coefficient.

The second type of values factored into the Technical Specifications are t' ose that are detemined by transient analysis. These parame-ters consist of the transient response term applied to the TM/LP equa-tion, the minimum required shutdown margin, the linear heat rate LC0 and the DNBR LC0. The transient response tenn applied to the TM/LP equation in the Technical Specifications is a result of the analysis of the RCS depressurization event or excess load event. The minimum required shutdown margin at hot shutdown conditions is determined by the steam line break event. This value is also confirmed for the boron dilution event. The minimum required shutdown margin for cold

• shutdown and refueling shutdown conditions is determined by the baron dilution event or the five percent subcriticality requirement for ref uel i ng. The values used in the linear heat rate LC0 are typically detennined by the loss of coolant accident. These values are also confirmed for the dropped CEA event. The LC0 on DNBR margin is calculated based on results from the dropped CEA event, the loss of four pump flow analysis or. the CEA withdrawal analysis.

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I 4.0 TRANSIENT AND ACCIDENT ANALYSIS MODELS The District utilizes the latest version of the CESEC code (CESEC-III and hereafter referred to as CESEC) in the simulation of plant response to non-LOCA initiating events. The District utilizes the CETOP and TORC conpeter codes for calculation of DNBR during these events.

4.1 Plant Simulation Model The District utilizes the CESEC digital conputer code, Refer-ences 4-1 through 4-10, to provide the simulation of the Fort Calhoun Station nuclear stean supply systen. The progran cal-culates the plant response to non-LOCA initiating events for a wide range of operating conditions. The infomation presented in Reference 4-9 supercedes infomation provided in References 4-1 through 4-8. Additional infomation on the model is pro-vided in Reference 4-10. The CESEC program, which numerically integrates one dimensional mass and energy conservation equa-tions, assumes a node / flow-path network to model the NSSS. The primary system canponents considered in the code include the re-actor vessel, the reactor core, the primary coolant loops, the pressurizer, the steam generators and the reactor coolant pumps. The secondary system components include the secondary side of the steam generators, the main steam systen, the feed-water system and the various steam control valves. In addi-tion, the progran models some of the control and plant protec-tion systems.

The code self initializes for any given, but constant, set of reactor power level, reactor coolant flow rate and steam gener-ator power sharing. During the transient calculations, the 5 time rate of change in the system pressure and enthalpy are obtained from solution of the conservation equations. These derivatives are then numerically integrated in time under the assumption of ther.nal equilibrium to give the systen pressure and nodal enthalpies. The fluid states recognized by the code are subcooled and saturated; superheating is allowed in the pressurizer. Fluid in the reactor coolant system is assumed to ,

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4.0 TRANSIENT AND ACCIDENT ANALYSIS MODELS (Continued) ,

I 4.1 Plant Simulation Model (Continued) be homogenous. Reference 4-9 provides a description of the CESEC code, including the major models, and the input, output and plot packages.

The pressurizer model is described in Reference 4-9 and further discussed in Reference 4-10. The District utilizes the wall heat transfer model to pemit simulation of voiding in any node in which steam fomation occurs. Voiding may occur in events such as a steam line break or steam generator tube rupture.

Nodalization of the closure head, described in Reference 4-9 and further discussed in Reference 4-10, allows for the foma-tion of a void in the upper head region when the pressurizer empties. Flow to the closure head is teminated in simulations of those events in which natural circulation occurs and in those events such as the steam line break where this action delays salety injection.

The capabilities and limitations of the CESEC code are dis-cussed in References 4-9 and 4-10. The District's CESEC model of Fort Calhoun Station is valid for the transients discussed in Section 5 of this report, with the exception ~of the CEA Ejec-tion Analysis and LOCA Analysis. The CESEC model is also valid for analysis of the loss of load, malfunctions of the feedwater system and the steam generator tube rupture incidents.

The CESEC code is maintained by CE on the CE computer system in -

Windsor, Connecticut. The District accesses the code through a time sharing system. CE maintains all documentation and qual-ity assurance programs related to this code.

i 4.2 DNBR Analysis Models l

l The DNBR analysis is currently perfomed using either the TORC code, Reference 4-11, or both the TORC and CETOP codes, Refer-ence 4-12. The TORC code is used as a benchmark for the CETOP 10

r 4.0 TRANSIENT AND ACCIDENT ANALYSIS MODELS (Continued) 4.2 DNBR Analysis Models (Continued) code model. TORC solves the conservation equations, as applied to a three-dimensional representation of the open lattice core, to determine the local coolant conditions at all points in the core. Lateral transfer of mass and energy between neighboring flow channels (open core effects) are accounted for in the cal-culation of local coolant conditions. These coolant conditions are then used with a Critical Heat Flux (CHF) correlation sup-plied as a code subroutine to determine the minimum value of DNBR for the reactor core. The CE-1 CliF correlation (Refer-ences 4-13 and 4-14) is used for the Fort Calhoun reactor as approved in Reference 4-15. The Detailed TORC code is used directly in the seized rotor analysis.

The CETOP code has been developed to reduce the computer time needed for thennal hydraulic analyses while retaining all of the capabilities of the TORC design model. The CETOP model provides an additienal simplification to the conservation aqua-tions due to the specific geometry of the model. A complete description of the CETOP code is contained in Reference 4-12 and a description of the District's application of the CETOP code is contained in Reference 4-16.

The fraction of inlet flow to the hot assembly in the CETOP model is adjusted such that the model yields appropriate MDNBR results when compared to the results of the TORC analysis for a specified range of operating conditions.

The CETOP code is used to calculate DNBR for all transient anal-

• yses discussed in Section 5 with the exception of the seized rotor analysis.

4.3 Application of Uncertainties

! Uncertainties are taken into account either by deteruinistic or statistical methods. The detenninistic method applies all un-l 11

4.0 TRANSIENT AND ACCIDENT ANALYSIS MODELS (Continued) 4.3 Application of Uncertainties (Continued) certainties adversely and simultaneously when calculating the approach to a limit.

Uncertainties in DNBR calculations are taken into account by statistical methods. The statistical method takes into account the likelihood that the uncertainties will all be adverse. The statistical method is discussed in Reference 4-17. In this method the impact' of component uncertainties on DPR is assessed and the DNBR SAFDL is increased to include the effects of the uncertainties. Since the uncertainties are accanodated by the increased DNBR SAFDL in the statistical method, engin-eering factors are not applied to the DNBR analysis model. The statistical method of applying uncertainties is applied to the CEA withdrawal, CEA drop, loss of RCS flow, excess load, seized rotor and asymmetric steam generator event DNBR calculations.

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS This section addresses the evaluation of the various transients and accidents that are perfonned during a reload core analysis. Specific methods are described for each transient and accident. For each acci-dent or transient the following material is described:

A. Definition of the Event - A brief description of the causes, consequences, and RPS trips involved in the incident.

B. Analysis Criteria - A brief description of the classification of the event and the Specified Acceptable Fuel Design Limit (SAFDL) or the offsite. dose criteria which must be met.

C. Objectives of the Analysis - A brief description of the methods that are used to assure that the criteria of the analysis are met.

D. Key Parameters and Analysis Assumptions - A description of the key parameters and assumptions used in the analysis.

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5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued)

E. Analysis Method - A description of the methodology employed by the District to analyze the event.

F. Analysis Results and 10 CFR 50.59 Criteria - The expected re-sults of the analysis and a discussion of the methods used to determine if the event meets the criteria of 10 CFR 50.59.

G. Conservatism of Results - A description of the conservatism of the analysis.

The values of the trip setpoints and trip delay times used in these analyses are shown in Table 5.0-1.

5.1 CEA Withdrawal 5.1.1 Definition of the Event A sequential CEA Group Withdrawal Event is assumed to occur as a result of a failure of the control element assembly drive mechanism control system or by operator error. The CEA Block System eliminates the possibility of an out of sequence bank withdrawal or single CEA withdrawal due to a single failure.

Any controlled or unplanned withdrawals of the CEA's re-sults in a positive reactivity addition which causes the core power, core average heat flux and reactor cool-ant system temperature and pressure to rise and in turn l decrease the DNB and Linear Heat Rate (LHR) margins.

The pressure increase, if large enough, activates the 5 pressurizer sprays which mitigate the pressure rise.

In the presence of a positive Moderator Temperature Coefficient (MTC) of reactivity, the temperature in-crease results in an additional positive reactivity addition further decreasing the margin to the DNB and LHR limits.

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Table 5.0-1 REACTOR PROTECTIVE SYSTEM TRIPS AND SAFETY INJECTION Used in Analysis Trip Setpoint Uncertainty Delay Time (Sec) Setpoint High Rate-of-Change of Power 2.6 dec/ min 10.5 dec/ min 0.4 2.1 dec/ min High Power Level 107% 5.0% 0.4 112%

Variable High Power Level 9.1% above set 0.9% 0.4 10% above initial power level to power level a low of 19.1%

Low Reactor Coolant Flow 95% 12% 0.65 93%

High Pressurizer Pressure 2400 psia 122 psi 0.9 2422 psia 3: Thermal Margin / Low Pressure (l) 1750 psia 122 psi 0.9 1728 psia Low Steam Generator Pressure 500 psia 122 psi 0.9 478 psia low Stean Generator Water 31.2% of narrow 110 in. (5.7% of narrow 0.9 25.5% of span Level range span range span)

Containment Pressure High 5 psig 10.4 psi 0.1 5.4 psig liigh Pressure Safety Injection 1600 psia 122 psi 12(2) 1578 psia (1) Values represent the low limit of the thermal margin / low pressure trip. The setpoint of this trip is discussed in Reference 5-2.

(2) Pump start - loop valve opening time.

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.1 CEA Withdrawal (Continued) 5.1.1 Definition of the Event (Continued)

Withdrawal of the CEA's causes the axial power distribu-tion to shift to the top of the core. The associated increase in the axial peak is partially compensated by the corresponding decrease in the integrated radial peaking factor. The magnitude of the 3-D peak change depends primarily on the initial CEA configuration and axial power distribution.

The withdrawal of the CEA's causes the neutron flux as measured by the excore detectors to be decalibrated due to CEA motion, i.e., rod shadowing effects. This decal-ibration of excore detectors, however, is partially com-pensated by neutron attenuation rising fran moderator density changes (i.e. , tenperature shadowing ef fects).

As the core power and heat flux increase, a reactor trip on high power, variable high power, or Thennal Mar-gin / Low Pressure may occur to tenninate the event de-pending on the initial operating conditions and rate of reactivity addition. Other potential trips include the axial power distribution and high pressurizer pressure trips. If a trip occurs, the CEA's drop into the core and. insert negative reactivity which quickly terminates further margin degradation. If no trip occurs and cor-rective action is not taken by the operators, the CEA's fully withdraw and the NSSS achieves a new steady state $

equilibrium with higher power, tenperature, peak linear heat rate and lower hot channel DNBR value.

5.1.2 Analysis Criteria The CEA Withdrawal (CEAU) event is classified as an Anticipated Operational Occurrence (A00) for which the 15

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.1 CEA Withdrawal (Continued) 5.1.2 Analysis Criteria (Continued) following criteria must be met:

a. The transient minimum DNBR is greater than the )

95/95 confidence interval limit for the CE-1 correlation, and

b. the Peak Linear Heat Generation Rate (PLHGR) does not exceed 21 kw/ft.

5.1.3 Objectives of the Analysis The objectives of the analysis performed for the "at power" CEAW event is to calculate the Required Over-power Margin (ROPM) which must be factored into the setpoint analysis.

The objective of the analysis for the hot zero power CEAW event is to demonstrate that the Variable High Power Trip (VHPT) is initiated in time to insure that the analysis criteria are met.

5.1.4 Key Parameters and Analysis Assumotions The initial conditions assumed in the CEAW analysis are shown in Table 5.1-1. The reactor state parameters of primary importance in calculating the margin degradation are:

1. CEA withdrawal rate * (i.e., reactivity insertion rate),

2. Gap thermal conductivity (HGAP),

• NOTE: The term CEA withdrawal rate and CEA reactivity insertion rate are used interchangably in this report.

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5.0 TRANSIENT AND. ACCIDENT ANALYSIS METHODS (Continued) 5.1 CEA Withdrawal _(Continued) 5.1.4 Key Parameters and Analysis Assumotions (Continued)

3. Initial power level,

4. Flux power level deter:ained' from the excore detec-tor response during the transient,

5. The moderator tenperature coefficient reactivity, and

6. Changes in the axial power distribution and plan-ar and integrated radial peaking factor during the transient.

The excore responses for each initial power level anal-yzed are based on the CEA insertions allowed by the Power Dependent Insertion Limit (PDIL) at the selected power level,'the cha,nges in CEA position prior to trip, and the corresponding rod i shadowing and temperature attenuation (shadowing) _ factors.

For the CEAW cases where combinations of parameters re-sult in a reactor trip, the scram reactivity versus in-sertion characteristics are assumed to be those associ-ated with the core average axial power distribution peaked at the bottom of the core. The scram reactivity versus 'asertion characteristics associated with this bottom peak shape minimize tbe amount of negative reac-E tivity inserted during initial portions of the scram following a reactor trip.

j All control systems except the pressurizer pressure con-i trol system and the pressurizer level control system i

l s are assumed to be in a manual' mode. These are the most advede operating nodes for this event. The pressuri-l r <

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Y, 17 L ,

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Table 5.1-1 Initial Conditions Assumed in CEAW Event Analysis Parameter Units Value Initial Core Power Level MWt 1-1500*

Initial Core Inlet Coolant ,

• F 532-545*

Temperature Moderator Temperature Coefficient X10-4 Ap/ F Tech. Spec. Range Initial RCS Pressure psia Minimum allowed by*

Tech. Specs.

Fuel Temperature Coefficient X10-4 ap/ F Least Negative i

Predicted During A Cycle Initial Core Mass Velocity 106 lbm/hr Minimum allowed by*

Tech. Specs.

Fuel Temp. Coeff. Uncertainty  % -15.0 Gap Thennal Conductivity BTU /hr-f t2 [ ]

CEA Differential Worth *10-4 / inch C 3 CEA Withdrawal Speed i n/ min 46.0 Radial Peaks \ Maximum Allowed by Tech. Specs. for a Given Initial Power Level Scram Reactivity

• Minimum Predicted During a Cycle High Power Trip Analysis 'A. Vint  % of 1500 MWt 112.0 Variable High Power f:;,b ! sis

% Above Initial 10.0 Setpoint Power Level Temperature Shadowing Factor  % Power /*F [ ] *

• For DNBR'calcu'lations, effects of ur, certainties are combined statistically, i

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?, '1 4 p ,.

.e. s - 4 m .I +

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.1 CEA Withdrawal (Continued) 5.1.4 Key Parameters and Analysis Assumotions (Continued) l zer pressure control system and pressurizer level con-trol system are assumed to be in the automatic mode since the actuation of these systems minimizes a rise in the coolant system pressure. The net effect, is to delay a reactor trip until a high power trip is initi-ated. This allows the transient increases in power, heat flux and coolant temperature to proceed for a long-er period of time. In addition, minimizing the pres-sure increase is conservative in the margin degradation calculations since increases in pressure would offset some of the DNB margin degradation caused by increases in the core heat flux and coolant temperatures.

5.1.5 Analysis Methodoloa3

. The methodology used for analysis of the CEAW event is described in CEN-121(B)-P, Reference 5-1. The District does not perfonn all parametric analyses discussed in Reference 5-1 for Fort Calhoun Station. Rather, the District utilizes the analyses performed in Reference 5-1 to linit the number of analyses necessary for Fort Calhoun Station. Specifically, the District utilizes the result that [

] In addithn, the result from a Reference 5-1 that [

.] when combined with [

] can be used to perfonn sensitivity analyses on the CEA withdrawal rate to achieve [

'] i s ut ili zed.

19

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.1 CEA Withdrawal (Continued) 5.1.5 Analysis Methodology (Continued)

The rod shadowing factors for the Fort Calhoun Station full power case with Bank 4 inserted are the inverse of the rod shadowing factors used in Reference 5-1 (The rod shadowing factors for Fort Calhoun Station are such that the excore detectors see more flux when the rods are withdrawn than when they are inserted. Therefore, the [

] during a full power CEA withdrawal event). Be-cause of this effect, it may be necessary to assume a

[ ~ ] in order to ach'ieve [

]

The analysis at intemediate power levels is the same as documented in Reference 5-1.

The hot ero power CEAW event is analyzed assuming the variable high power trip is initiated at 29.1% (19.1%

plus 10% uncertainty) of rated thennal power. In addi-tion, the analysis assumes that the maximum CEA with-drawal rate is conbined with the maximum differential rod worth. This case is analyzed using CESEC and the minimum DNBR is calculated using CETOP_ using the assump-tions discussed in Reference 5-1.

The CEAW event analyzed to detennine the closest approach to the fuel centerline melt SAFDL assumes those values of the CEAW rate and Hg ap' discussed in Reference 5-1. This combination of CEAW rate and H gap was used to detennine the PLHGR at all power levels.

5.1.6 Typical Analysis Results and 10 CFR 50.59 Criteria l

The results of the analyses of the CEAW event for Fort Calhoun Station at full power and at intennediate power

\

20

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued)  !

5.1 CEA Withdrawal (Continued) 5.1.6 Typical Analysis Results and 10 CFR 50.59 Criteria (Continued) levels are expected to be similar to those presented in Reference 5-1. The results of the hot zero power CEA withdrawal analysis are expected to be similar to those discussed in the Cycle 8 reload submittal and the 1983 update of the USAR. The 10 CFR 50.59 criteria are met if the analysis for the full power and intermediate power level CEAW events shows that the required over-power margin for these events is less than the avail-able overpower margin required by the current Technical Specification DNB and PLHGR LC0's. The 10 CFR 50.59 criteria is satisfied for the hot zero power CEAW event if the minimum DNBR is greater than that reported in the latest submitted analysis.

5.1.7 Conservatism of Results Conservatism of the results of the CEAW incident analyses is discussed in Reference 5-1 for the full power, intermediate power level and hot zero power Cases.

5.2 Baron Dilution Incident 5.2.1. Definition of Event Boron dilution is a manual operation, conducted under 'I strict procedural controls which specify permissible limits on the rate and magnitude of any required change in boron concentration. Boron concentration in the l reactor coolant system can be ' decreased by either I controlled addition of unborated makeup water with a corresponding removal of reactor coolant or by using 21

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.2 Boron Dilution Incident (Continued) 5.2.1. Definition of Event (Continued) the deborating fon exchangers. To effect boron dilution the makeup controller mode selector of the chemical and volume control system (CVCS) must be set to " dilute" and then the demineralized water batch quantity selector set for the desired quantity. When the specific amount has been injected, the demineralized water control valve is shut automatic-al ly. An inadvertent boron dilution can occur only if there is a combination of operator error and a CVCS mal-function occurring at the same time. No RPS trips are assumed to teminate this iacident.

5.2.2 Analysis Criteria The boron dilution event is classified as an A00 for which the following criteria cannot be exceeded:

A. DNBR greater than the 95/95 confidence interval limit using the CE-1 correlation, and B. The PLHGR less than 21 kw/ft.

5.2.3 Objectives of the Analysis The DNBR and PLHGR criteria are met by showing that suf-ficient time exists for the operator to take corrective action to teminate the event prior to exceeding the SAFDLs. This is acceplished by calculating the time interval in which the minimum T( :hnical Specification shutdown margin is 1ost. The acceptable time interval for the operator to take corrective actions before shut-down margin is lost are 15 minutes for Modes 2, 3 and 4 and 30 minutes in t1 ode 5.

I 22 l

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.2 Baron Dilution Incident (Continued) 5.2.4 Key Parameters and Analysis Assumotions The boron dilution event at power (Mode 1) is bounded by the faster reactivity insertion rate of the CEA withdrawal event and it lacks the local power peaking associated with the withdrawn CEA. For the boron dilu-tion event in Modes 2 through 5, it is assumed that all three charging pumps are operating at their maximum cap-acity for a total charging rate of 120 gpm. For the dilution at hot standby (Mode 2) the event is assumed to be initiated at the Technical Specification hot shut-down margin requirement at 532 F. The reactor coolant system is 5,506 cubic feet.

The baron dilution incident at hot shutdown (Mode 3) is assumed to be initiated from the Technical Specifica-tion shutdown margin requirement at 210 F. The boron dilution incident cold shutdown (Mode 4) is initiated fran the Technical Specification minimum shutdown mar-gin requirement at 68 F. The analysis is conducted for two RCS volumes, one of 5,506 cubic feet and the other of 2,036 cubic feet, which corresponds to the volume for a refueling operation condition. The analysis for the lower volume cold shutdown condition assumes that shutdown groups A and B are withdrawn fraa the core and all regulating groups are inserted in the core with the

. exception of the most reactive rod which is assumed to be stuck in its fully withdrawn position. These assump- ,

tions are consistent with the Technical Specifications for cold shutdown conditior.s. The baron dilution event during refueling is analyzed assuming that reactor re-fueling has just been completed and the head is in place but the coolant volume is sufficient to only fill the reactor vessel to the bottom of the piping nozzles i

23

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.2 Boron Dilution Incident (Continued) 5.2.4 Key Parameters and Analysis Assumptions (Continued)

(2,036 cubic feet) and the minimum permissible boron concentration allowed by Technical Specification for refueling exists. All CEA's are withdrawn from the core.

These assumptions represent shutdown conditions for the various modes wherein the core reactivity is greatest, the water volume and total boron content is at a mini-mum, and the rate of dilution is as large as possible.

Hence, these conditions represent the minimum time to achieve inadvertent criticality in the event of an un-controlled boron dilution.

5.2.5 Analysis Methods The method used to calculate the dilution time to criticality fran Modes 2 through 5 is through the use of the following equation:

atcrit

• TBD . In l-Where TBD = baron dilution time constant, which is a function of RCS volume and temperature (sec)

CB = critical boron concentration (ppm)

SDM = shutdown margin (%Ap)

• IBW = inverse boron worth (ppm /%ap)

As can be seen from this equation, the dilution time to criticality is minimized with a greater critical boron.

concentration, a smaller inverse boron worth, or a smaller T BD.

24-

5.0 TRANSIENT AND ACCIDEiiT ANALYSIS METHODS (Continued) 5.2 Boron Dilution Incident (Continued) 5.2.6 Analysis Results and 10 CFR 50.59 Criteria The analysis results are similar to those reported in the Cycle 8 safety analysis report and in the 1983 update of the USAR. The criteria of 10 CFR 50.59 are satisfied if the Technical Specification requirements on shutdown margin and the refueling boron concentra-tion is unchanged as a result of this analysis.

5.2.7 Conservatism of Results Because of the procedures involved in the boron dilu-tion and the numerous alann indications available to the operator, the probability of a sustained or erron-eous boron dilution is very low. There is usually a large interval between the calculated time and the time limit for the boron dilution at hot standby and hot shutdown modes. Therefore, the results show consid-erable margin to the limit. The calculated time to critical for the boron dilution at cold shutdown with the minimum RCS volume is reasonably close to the acceptance criteria; however, the event is analyzed with only shutdown groups A and B being fully withdrawn from the core. Cold shutdown is nomally achieved with the shutdown groups A and B fully inserted in the core and, therefore, the core has a much lower keff than assumed in the analysis. The boron dilution at refueling is conservative since it is improbable that '

more than a few CEA's will be removed at any one time during a refueling and the approach to critical following refueling is done under strict administrative control with only one bank of CEA's reaoved at a time.-

The analysis assumes that all CEA's are withdrawn from the core.

25

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.3 Control Element Assembly Drop Incident 5.3.1 Definition of Event The control element assembly (CEA) drop incident is de-fined as the inadvertent release of a CEA causing it to drop into the reactor core. The CEA drive is of the rack and pinion type with the drive shaft running paral-lel to and driving the rack through a pinion gear and a set of bevel gears. The drive mechanism is equipped with a mechanical brake which maintains the position of the CEA. The CEA drop may occur due to an inadvertent interruption of power to the CEA drive magnetic clutch or an electrical or mechanical failure of the mechani -

cal brake in the CEA drive mechanism when the CEA is being moved.

The full-length CEA drop event is classified as an A00

'which does not require an RPS trip to provide protec-tion against exceeding the SAFDLs. The CEA drop re-sults in a redistribution of the core radial power dis-tribution and an increase in the radial peaks which are not directly monitored by the RPS and which are not among those analyzed in detennining the DNB and LHR LCOs and LSSSs. As such, initial steady state margin must be built into the Technical Specification LCOs to allow the reactor to " ride out" the event without exceeding the DNBR and LHR SAFDLs.

5.3.2 Analysis Criteria s

The full-length CEA drop event is classified as an Anti-cipated Operational Occurrence for which the followinj criteria must be met: a) The transient minimum DNBR must be greater than or equal to the 95/95 confidence interval limit, using the CE-1 correlation, b) The Peak Linear Heat Rate (PLHR) must be less than or equal to 21 kw/ft.

26

f 5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) i l

l 5.3 Control Element Assembly Drop Incident (Continued) l l 5.3.3 Objectives of the Analysis The objective of the analysis is to determine the Re-quired Overpower Margin (ROPit) which must be built into the LCOs to assure the DN3R and LHR SAFDLs are not ex-ceeded for the CEA drop which produces the highest dis-tortion in the hot channel power distribution. Since the R0PM is dependent upon initial power level, rod con-figuration and axial shape index, an analysis parametric in these variables is performed.

5.3.4 Key Parameters and Analysis Assumptions Table 5.3.4-1 contains a list of the key parameters assumed in the full-length CEA drop analysis. Assump-tions used in the analysis include:

1.

2. The rod block system is assumed to prevent any other rod motion during the transient.

3. The turbine admission valves are maintained at a constant position during the transient. This is because the turbine admission valve position is set manually at Fort Calhoun Station and, there-fore, the turbine admission valves will not auto-matically open in response to a reduced electri-

. cal generation output.

27 l

Table 5.3.4-1

, KEY PARAMETERS ASSUMED IN THE FULL LENGTH CEA DROP ANALYSIS Parameter Units Value 2 i Initial Core Power Level MW t

750 to 1500' O

Initial Core Inlet F Maximum allowed t Temperature by Tech. Specs.

Initial RCS Pressure psia Minimum allowed

• by Tech. Specs.

Initial Core Itass Flow Rate *10 lLm/hr Minimum allowed

• by Tech. Specs.

Moderator Temperature *10-4ap/0F Most negative Coefficient allowed by Tech.

Specs.

CEA Insertion  % Insertion Maximum allowed by Tech. Specs.

Radial Peaking Distortion Maximum value predicted Factor during core life Dropped CEA Worth %Ap Core Average H gap BTU /hr-Ft - F

~

Fuel Temperature *10-4Ap/ 0F - ~

Coefficient t

For DNBR calculations, the effects.of uncertainties on these parameters are combined statistically.

28

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) {

5.3 Control "lement Assembly Drop Incident (Continued) 5.3.5 Analysis Method The analysis methods utilized by the District to anal-yze the CEA drop incident are discussed in Section 8 of Reference 5-2.

5.3.6 Analysis Results and 10 CFR 50.59 Criteria Typical analysis results are contained in Section 8 of I Reference 5-2 and in the 1983 update of the Fort Ca'-

houn Station Unit No. 1 USAR. The criteria of 10 CFR 50.59 are met if the required overpower margin calcu-lated for this incident is less than the overpower mar-gin being maintained by the current Technical Specifica-l tions.

l 5.3.7 Conservatism of Results The following areas of conservatism are included in the j analysi s.

1. The most negative moderator [

] coefficients of reactivity are utilized be-cause these coefficients produce the minimum RCS coolant temperature decrease.

2. The [ ] distortion factor at any time dur-ing core life is combined with the [ ] CEA worth at any time during core life. *

3. The moderator temperature coefficient assumed in the analysis is the most negative value allowed by the Technical Specifications. The actual end of life value, including measurement uncertainty, is less negative.

29

l 5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.3 Control Element Assembly Drop Incident (Continued) 5.3.7 Conservatism of Results (Continued) 4.

5.4 Four-Pump Loss of Flow Event 5.4.1 Definition of the Event The four-pump loss of coolant flow event is initiated by the simultaneous loss of electrical power to all ,

four reactor coolant pumps. The loss of AC power to reactor coolant pumps may result from either the complete loss of AC power to the plant, or the failuire of the fast transfer breakers to close after a loss of offsite power.

Reactor trip for the loss of coolant flow is initiated by a low coolant flow rate as determined by a reduction in the sum of the steam generator hot to cold leg pres-sure drop. This signal is conpared to a setpoint which is a function of the number of reactor coolant pumps in operation (which current Technical Specificaticns re-quire to be four). A reactor trip would be initiated when the flow rate drops to 93% of full flow (95% minus 2% uncertainty).

i 5.4.2 Analysis Criteria The four-pump loss of flow-event is classified as an A00 for which the transient minimum DNBR must be 30

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.4 Four-Pump Loss of Flow Event (Continued) 5.4.2 Analysis Criteria (Continued) greater than the 95/95 percent confidence interval limit using the CE-1 correlation.

5.4.3 Objectives of the Analysis The objective of the analysis is to determine the re-quired overpower margin that must be built into the DNB LCOs such that in conjunction with the low flow trip the DNBR SAFDL is not exceeded. Since the required overpower margin is dependent upon both axial shape index and the CEA rod configuration, an analysis para-metric in these parameters is performed.

5.4.4 Key Parameters and Analysis Assumptions The closest approach to the DNBR SAFDL occurs for a loss of flow event initiated from the full power conditions. Table 5.4.4-1 gives the key parameters used in this analysis. The flow coast down is calculated in the CESEC code.

5.4.5 Analysis Method The analysis method used by the District to analyze the four-pump loss of coolant flow is discussed in Section 7 of Reference 5-2. The District utilizes the CESEC-TORC method to analyze axial power distributions '

characterized by both negative and positive shape indices. The STRIKIN-TORC method is not utilized by the District because of the high rotational energy of the pumps (N = 1185 rpm, I = 71,000 lb-ft /2pump). The District also utilizes the [ static reactivity insertion rate rather than the space time reactivity insertion rate.]

31

l l

Table 5.4.4-1 KEY PARAMETERS ASSUMED IN THE LOSS OF COOLANT FLOW ANALYSIS Parameter Uni ts Value Initial Core Power Level MW t 1500*

Initial Core Inlet o Temperature F Initial RCS Pressure psia 6

Initial Core Mass *10 lbm/hr.

Flow Rate Moderator Temperature *10 Ap/ F Maximum allowed Coefficient by Tech. Specs.

Fuel Temperature *10~4Ap/UF ' .:ast neq. ative Coefficient predicted during core life.

Low Flow Trip Delay Time sec. Maximum CEA Drop Time sec. Maximum allowed by Tech. Specs.

Scram Reactivity Worth no Minimum predicted during core lifetime Scram Reactivity Curve Consistent with axial shape of interest 2

Core Average H gap BTU /hr-Ft oF k

t For DNBR calculations, effects.of uncertainties on these parameters were a combined statistically.

4 32 L

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.4 Four-Pump Loss of Flow Event (Continued) 5.4.6 Analysis Results and 10 CFR 50.59 Criteria Expected analysis results are presented in Section 7.1 of Reference 5-2. The main difference between these re-sults and the results for Fort Calhoun Station is that the R0PM will be significantly reduced for Fort Calhoun Station. This is because of the higher rotational energy of the Fort Calhoun reactor coolant pumps.

The criteria of 10 CFR 50.59 are met if the required overpower margin calculated for the four-pump loss of coolant flow event is less than the overpower margin being maintained by the current Technical Specifica- '

tions.

5.4.7 Conservatism of Results The conservative nature of the DNBR R0PM values calcu-lated for the four-pump loss of flow event is demon-strated by the following conservative assumptions.

1. Field measurenents of the CEA magnetic clutch de-cay is more rapid than assumed in the safety anal-ysis.

l l

2. The available scram worth is higher than assumed in the safety analysis.

l '3. The MTC at full power is more negative than the value assumed in the safety analysis.

4. The actual CEA drop time to 90% inserted is faster than that assumed in the safety analysis.

5. The conservatism of the CETOP calculations is discussed in Section 7 of Reference 5-2.

33

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.5 Asymmetric Steam Generator Event 5.5.1 Definition of the Event The asymmetric transients arising from a secondary sys-tem malfunction in one steam generator result in changes in core power distribution which are not inher-ently covered by the TM/LP or APD LSSS. Consequently, these events must be analyzed to detemine the initial steady state themal margin which is built into and maintained by the Technical Specification LC0 such that assurance is provided that the DNBR and peak linear heat rate SAFDLs are not exceeded for these transients.

The four events which effect the steam generator are:

1. Loss of load to one steam generator.

2. Loss of feedwater to one steam generator.

3. Excess feedwater to one steam generator.

4. Excess load to one steam generator.

The possible RPS trips which can occur to mitigate the consequences of these events include the low steam gen-erator level, TM/LP, low steam generator pressure, and the asymmetric steam generator transient protection trip function (ASGTPTF). The particular trip which in-tervenes is dependent upon the event initiator and the initial operating conditions.

• The ASGTPTF trip will be installed in the Fort Calhoun Station RPS prior to operation of Cycle 9 to reduce the margin requirements associated with these asymmetric

! events and to insure that these events do not become a limiting A00 for establishing initial margin which must 34 i

^

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.5 Asymmetric Steam Generator Event (Continued) 5.5.1 Definition of the Event (Continued) be maintained by the LCO. A system description of the ASGTPTF is presented in Appendix B of Reference 5-2.

5.5.2 Analysis Criteria The asymmetric steam generator events are classified as 4

A00s for which the following criteria must be met, a) the transient minimum DNBR must be greater than or equal to the 95/95 confidence interval limit using the CE-1 correlation, and b) the peak linear heat must be less than or equal to 21 kw/ft.

5.5.3 Objectives of the Analysis The objectives of the analysis are to detemine the required overpower margin that must be built into the LC0's such that in conjunction with the ASGTPTF the DNBR and PLHGR SAFDL's is not exceeded.

5.5.4 Key Parameters and Analysis Assumotions Section 7 of Reference 5-2 demonstrates that the loss of load to one steam generator (LL/ISG) is the limiting asymmetric steam generator transient for establishing initial steady state themal margin which must be main-tained by the Technical Specification LC0. Therefore, infomation is only provided for this asymmetric steam generator event. The key parameters used in the analy-sis of the LL/ISG event are given in Table 5.4.4-1.

The charging pumps and proportional heater systems are assumed to be inoperable during the transient. This maximizes the pressure drop during the event. The tur-35

Table 5.5.4-1 KEY PARAMETERS ASSUMED IN.THE LL/1SG EVENT

'Pa rameter Units Value Initial Core Power fM t 700 to 1500*

Initial Core Inlet OF Maximum allowed t Temperature by Tech. Specs.

Initial Reactor Coolant psia Minimum allowedt System Pressure by Tech. Specs.

~ _

Moderator Temperature -4

• 10 ap/ F Coefficient Fuel Temperature *10-4ap/0F Coefficient Core Average H gap BTU /hr-f 2t ,of Maximum value predicted during core life.

6 Initial Core Mass *10 lbm/hr. Best estimate flow.t Flow Rate Scram Reactivity Worth %Ap Minimum oredicted during core life.

t For DNBR calculations, effects of uncertainties on these parameters were combined statistically.

s 36

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.5 Asymmetric Steam Generator Event (Continued) 5.5.4 Key Parameters and Analysis Assumptions (Continued) bine admission valves are assumed to maintain a con-stant position throughout the event since the turbine control system at Fort Calhoun utilizes manual setting of the turbine admission valves.

5.5.5 Analysis Method The method utilized by the District to analyze the LL/1SG is discussed in Section 7 of Reference 5-2.

5.5.6 Analysis Results and 10 CFR 50.59 Criteria The results of the analysis for the LL/1SG event are discussed in Section 7 of Reference 5-2.

The results for Fort Calhoun Station are expected to be similar. The criteria of 10 CFR 50.59 are satisfied if the required overpower margin calculated for the LL/1SG event is less than the overpower margin being main-tained by the current Technical Specifications.

5.6 Excess Load Incident i 5.6.1 Definitica of Event i

l An excess load transient is defined as any rapid in-

! s crease in the steam generator steam flow other than a i steam line break. Such a rapid increase in steam flow results in a power mismatch between the reactor core and the st.aam generator load demand. In addition, there is a decrease in the reactor coolant temperature and pressure. Under these conditions the negative mod-erator temperature coefficient reactivity causes an increase in core power.

37

.l l

5.0l TRANSIENT AND ACCIDENT ANALYSIS METHODS '(Continued) 4 5.6_ Excess Load Incident (Continued) l l

I 5.6.1 Definition of Event (Continued) l The rapid opening of the turbine admission valves or

, the steam dump bypass to the condenser causes an excess 3 load event. Turbine valves are not sized to accomodate steam flow for powers much in excess of 1500 MWt. The

- steam dump valves and steam bypass valves to the conden-I ser are sized to accomodate 33% and 5%, respectively, i of the steam flow at 1500 MW.- Therefore, the following load increase incidents .are examined:

1 A. Rapid opening of the turbine control valves at power: The maximum increase in the steam flow due to the turbine control valves opening is ,

limited by the turbine load limit control. The load limit control function is used to maintain load, so unless valve failure occurs, _the control

valves will remain where positioned.

j B. Opening of all dump and bypass valves at power due to steam dump control interlock failure: The circuit between the steam dump controller and the i

dump valves is open when the turbine generator is 3

on line. Accidental closing of the steam dump control interlock under full load conditions, ac-

, cording to the temperature program of the control-

} 'ler, causes full opening of the dump and bypass-s

- val ves. Since the reactor coolant. temperature de-- ,

creases during the event, these valves will be closed again after the av_erage reactor coolant-

- temperature . decreases to 535 F.

C. Opening of the dump and bypass valves ~at hot .

standby conditions due to low reference temper-

' ature setting-in the steam dump contrclleri :When -

38

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

I 5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.6 Excess Load Incident (Continued) 5.6.1 Definition of Event (Continued)

C. (Continued) the plant is in hot standby conditions the dump valve controller is operative but does not act because the hot standby temperature is lower than the lowest value required to open the valves. At hot standby the reactor coolant temperature is 532 F, which is 8 F below the minimum tenperature required to open the dump and bypass valves (540 F). The maximum error that can be intro-duced in the referenced temperature setting is limited to 70 F since a narrow range instrument is used for this purpose. Reducing the dump valve controller reference setting from 532 to 515 would result in a partial opening of the valves but as soon as the reactor coolant tenper-ature dropped to 518 F the valves would again be completely closed.

D. Opening the dump and bypass valves at hot standby due to stean dump controller malfunction: The most severe incident at hot standby would occur in the event the steam dump valve controller yields an incorrect signal and causes the steam dump and bypass valves to open completely. This case is considered to be much less probable than i case C above but represents the most limiting event under hot standby conditions.

The possible RPS trips that might be encountered during this event are:

1. Variable high power trip (VHPT).

39

i 5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.6 Excess Load Incident (Continued) 5.6.1 Definition of Event (Continued)

D. (Continued)

2. TM/LP trip.

3. Low steam generator water level trip.

4. Low steam generator pressure trip.

The RPS trip initiated to mitigate the conse-quences of the event will depend upon the initial conditions and the rate of reactivity insertion due to moderator feedback effects.

5.6.2 Analysis Criteria The excess load event is classified as a A00 for which the following criteria must be met.

A. The transient minimum DNBR must be greater than or equal to the 95/95 confidence interval limit using the CE-1 correlation.

B. The peak linear heat rate (PLHR) must be less than or equal to 21' kw/ft.

5.6.3 Objectives of the Analysis The objectives of the analysis are to calculate a [

-] which, when incorporated in the TM/LP i

equation will ensure that the SAFDL's are exceeded for those excess load events which require a TM/LP trip for protection and to ensure that the DNBR and LHR SAFDL's are not exceeded for excess load events for which the TM/LP does not provide protection.

40 l

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.6 Excess Load Incident (Continued) 5.6.4 Key Parameters and Analysis Assumotions As discussed in Section 5 of CENPD-199-P (Reference 5-2), sensitivity studies performed by CE have demen-strated that the maximum calculated [

.] for the excess load event occurs for the [

] at hot full power conditions. District sensitivity stud-ies show similar results. Therefore, only the hot full power case is analyzed. The key parameters used in the analysis of the excess load event are given in Table 5.6.4-1. The remaining assumptions are the same as those discussed in Reference 5-2.

5.6.5 Analysis Method The steps used for determining the [ ]

value and calculating the largest [ ]

for all excess load events which rely on the TM/LP trip for DNBR protection are given in Section 5 of CENPD-199-P (Reference 5-2). The minimum transient DNBR value for excess load events protected by the Variable High Power Trip is calculated using the procedure dis-cussed in the same Section.

The PLHR is calculated by obtaining the core average linear heat rate at time of peak core power and multi-plying it by the appropriate peaking factors and asso- '

ciated uncertainties.

5.6.6 Analysis Results and 10 CFR 50.59 Criteria The results of the excess load analysis are similar to those presented in Section 5 of CENPD-199-P (Reference 41

Table 5.6.4-1 KEY PARAMETERS ASSUMED IN THE EXCESS LOAD EVENT ANALYSIS Parameter Units Value Initial Core Power

• MW t 1500t Initial Core In-let Temperature OF At Power Maximum allowedt by Tech. Specs.

Initial Reactor Coolant psia Minimum allowed t System Pressure by Tech. Specs.

Initial Core Mass 6

• 10 1bm/hr. Minimum allowed

• Flow Rate by Tech. Specs.

CEA Drop Time sec. Maximum allowed by Tech. Specs.

Scram Reactivity %ap t11nimum predicted Worth during core life.

, Moderator Temperature *10-4ap/ F Negative values up to the Coefficient most negative value allowed by Tech.

i Specs.

tFor DNBR calculations, effects of uncertainties on these parameters were combined statistically.

s ,

42

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.6 Excess Load Incident (Continued) 5.6.6 Analysis Results and 10 CFR 50.59 Criteria (Continued) 5-?). The criteria of 10 CFR 50.59 are met if the

[ pressure bias term] is less than or equal to the value used in the current TM/LP trip equation.

5.6.7 Conservatism of Results The following points demonstrate the conservatism of the overall results for the excess load event:

1. Field measurements demonstrate that the CEA magnetic clutch decay time is less than that assumed in the analysis.

2. The actual scram worths are higher than those in the analysis.

3. Where the most negative MTC is used, the value is more negative than that measured during plant operation.

4. The actual Doppler reactivity is more negative than assumed in the analysis.

5.

6. Field data demonstrates that the actual CEA drop time is less than that assumed in the analysis.

7. The conservatis:n of the [ ] is discussed in Section 5 of CENPD-199-P (Reference 5-2).

43

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.7 RCS Depressurization 5.7.1 Definition of Eve.g The RCS depressurization event is characterized by a rapid decrease in the primary system pressure caused by either the inadvertent opening of both power operated relief valves (PORVs) or the inadvertent opening of a f single primary safety valve operating at rated thermal l power. Following the initiation of the event, steam is discharged from the pressurizer steam space to the quench tank where it is condensed and stored. To com-pensate for the decreasing pressure the water in the pressurizer flashes to steam and the proportional heat-ers increase the heat added to the water in the pressur-izer in an attempt to maintain pressure. During this time the pressurizer level also begins to decrease caus-ing the letdown control valves to close and additional charging pumps to start so as to maintain level. As pressure continues to drop, the backup heaters energize to further assist in maintaining primary pressure. A reactor trip is initiated by the TM/LP trip to prevent exceeding the DNBR SAFDL.

5.7.2 Analysis Criteria The RCS depressurization event is classified as an A00 for which the transient minimum DNBR must be greater than or equal to the 95/95 percent confidence interval limit using the CE-1 correlation. ,

5.7.3 Objectives nf the Analysis This event is classified as an A00 for which there must be sufficient margin built into the TM/LP trip such that the DNBR SAFDL is not exceeded. The objective of this analysis is to calculate a conservative [.

] for incorporation into the TM/LP equation.

44

I 5.0 - TRANSIENT AND. ACCIDENT ANALYSIS METHODS (Continued) l 5.7 RCS Depressurization (Continued) 5.7.4 Key Parameters and Analysis Assumptions The key parameters for the RCS depressurization event analysis are given in Table 5.7.4-1. Additional assump-tions are discussed in Section 5 of CENPD-199-P (Refer-ence 5-2).

5.7.5 Analysis Method The methods used by the District to analyze the RCS depressurization event are contained in Section 5 of CENPD-199-P (Reference 5-2).

5.7.6 Analysis Results and 10 CFR 50.59 1

i s Results of the RCS depressurization transient are discussed in Reference 5-2 and in the 1983 update of the Fort Calhoun Station Unit No.1 USAR. The criteria of 10 CFR 50.59 are satisfied if the [

] is less than or equal to the value used in the current TM/LP trip equation.

5.7.7 Conservat' ism of Results The conservatism of the calculated pr: ssure bias term is obtained by using the combination of the following conservative key parameters:

1. Conservative . scram reactivity characteristics are used in the analysis.

2. . Conservatively slow RPS response times are used.

3. Conservatively high 'pr"imary relief or safety valve areas are used.

(- .

45

\ 1.

e

/

.t -1 <

)

- Tabl e 5.7.4-1 2

i KEY PARAMETERS ASSUMED IN THE RCS DEPRESSURIZATION EVEflT ANALYSIS Parameter '

Units Value Initial Core Power Level IW t

1530 Initial Core In- 3 F Maximum allowed by let Temperature . Tech. Specs.

a Initial Reactor Coolant 8 Psia Upper limit of normal System Pressure , ,

operating range Moderator Temperature *1b"Ap/ F Most negative allowed Coefficient by Tech. Specs.

Fuel Temperature *10-4as/0F Most negative predicted Coefficient during core life.

Core Average H BTU /hr.-Ft.2-F Minimum predicted gap during core life.

Total Trip Delay sec. ' 1.4 Time 46 j

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.7 RCS Depressurization (Continued) 5.7.7 Conservatism of Results (Continued)

4. The RCS pressure is initially assumed to be in its upper limit as opposed to the nomal oper-ating pressure.

5.8 Main Steam Line Break Accident 5.8.1 Definition of the Event A large break of a pipe in the main steam system causes a rapid depletion of steam generator inventory and an increased rate of heat extraction from the primary sys-tem. The resultant cooldown of the reactor coolant, in the presence of a negative moderator temperature coeffi-cient of reactivity, will cause an increase in nuclear power and trip the reactor. A severe decrease in main steam pressure will also initiate reactor trip and cause the main steam isolation valves to close. If the steam line rupture occurs between the isolation valve and the steam generator outlet nozzle, blowdown of the affected steam generator will continue. (However, clo-sure of the check valve in the ruptured steam line, as well as closure of the isolation valves in both steam lines, will teminate blowdown from the intact steam generator) . The fastest blowdown, and therefore, the most rapid reactivity addition, occurs when the break ,

is at a steam generator nozzle. This break location is  ;

assumed for the cases analyzed. l Both full power and no-load (hot standby) initial con-dition cases were ' considered for two-loop operation l (i.e., four reactor coolant pumps).

47

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.8 Main Steam Line Break Accident (Continued) 5.8.1 Definition of the Event (Continued)

Since the steam generators are designed to withstand reactor coolant system operating pressure on the tube side with atmospheric pressure on the shell side, the continued integrity of the reactor coolant systen barrier is assured.

The most probable trip signals resulting fran an MSLB include low steam generator pressure, high power, low steam generator water level, TM/LP, and high rate-of-change of power (for the no-load case).

5.8.2 Analysis Criteria The steam line break accident event is classified as a postulated accident for which the site boundary doses must be within the 10 CFR 100 criteria. Acceptable site boundary doses are demonstrated by showing that the critical heat flux is not exceeded.

5.8.3 Objectives of the Analysis The objectives of the analysis are to demonstrate that

{

the margins to DNB for the reload core no-load two-loop and full-load two-loop main steam line break cases are greater than that for the Cycle 1 cases given in the original FSAR. This is accanplished by demonstrating that the return to power during the event for the reload core is less than the return to power calculated for Cycle 1.

5.8.4 Key Parameter and Analysis Assumptions The tiSLB accident is assumed to start from steady state conditions with the initial power. being 1530 MWt (102%)

48

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.8 Main Steam 1.ine Break Accident.

5.8.4 Key Parameter and Analysis Assumptions (Continued) for the full power case and 1 MWt for the no-load case.

The reactor coolant system cooldown causes the greatest positive reactivity insertion into the core when the Moderator Temperature Coefficient (MTC) is the most negative. For this reason the Technical Specification negative MTC limit corresponding to the end-of-cycle is assumed in the analysis. Since the reactivity change associated with moderator feedback varies significantly over the temperature range covered in the analysis, a curve of reactivity insertion versus temperature rather than a single value of MTC is assumed. This curve is derived on the basis that upon reactor trip the most reactive CEA is stuck in the fully withdrawn position thus yielding the most adverse cambination of scram worth and reactivity insertion. Although no single value of MTC is assumed in the analysis, the moderator cooldown reactivity function is calculated assuming an initial MTC equal to the most negative Technical Specification limit.

Reactivity feedback effects fraa the variation of fuel temperture (i.e., Doppler) are included in the analysi s. The most negative Doppler defect function, when used in conjunction with the decreasing fuel temperature causes' the greatest positive reactivity insertion during the MSLB event. In addition to ,

assuming the most negative Doppler defect function, an additional 157, uncertainty is assumed, i.e. , a 1.15 multiplier. This multiplier conservatively increases the subcritical multiplication and results in a larger return-to-power.

The delayed neutron precursor fraction, S, assumed is the maximum absolute value including uncertainties for.

49

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.8 Main Steam Line Break Accident (Continued) 5.8.4 Key Parameter and Analysis Assumptions (Continued) end of cycle conditions. This is conservative since it also maximizes subcritical multiplication and thus, enhances the potential for a return-to-power.

The steam generator low pressure trip, which occurs at 478 psia (including a 22 psia uncertainty below the nom-inal trip setting of 500 psia), is the trip assumed in the analysis. No credit is taken for the high power trip which occurs at approximately the same time for the full power case. For the cases analyzed, it is assumed that the most reactive CEA is stuck in the fully withdrawn position. If all CEA's insert (no stuck CEA's), there is no return-to-critical and no power transient following trip.

The cold edge temperatures are used to calculate mod-erator reactivity insertion during the cooldown, thus maximizing the return-to-critical and return-to-power potentials.

The Emenjency Operating Procedures are to be revised prior to the operation of Cycle 9 to eliminate the requirement of manually tripping the reactor coolant pumps (RCP's) as discussed in Reference 5-3. However, if tripping of the RCP's is required, the following discussion is applicable. ,

The MSLB case with the RCP's tripped is similar to the MSLB case with a loss of offsite power since the RCP's coastdown -in both events. As discussed in Reference

~ 5-4, the loss of offsite power delays safety injection due to the time' delay for the emergency diesel genera-50 e

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued)

S.8 Main Steam Line Break Accident (Continued) 5.8.4 Key Parameter and Analysis Assumptions (Continued) tors to restore power to the safety injection pumps and causes a coastdown of the RCP's. The coastdown affects the degree of overcooling and increases the time for safety injection borated water to reach the core mid-plane. Because manual tripping of the RCP's results in a later coastdown of the RCP's and because safety irdec-tion is not delayed since offsite power is available (i.e., the diesel generator startup and pump loading de-lays are not present), the injected boron will arrive at the core midplane sooner for a MSLB with the RCP's tripped than for a MSLB with a loss of offsite power.

Therefore, the reactivity effects of a MSLB with the RCP's tripped are less severe than for the MSLB with a loss of offsite power.

Reference 5-4 states that the MSLB case with a loss of offsite power results in the injected boron being dom-inant over the RCS cooldown and concludes that the reac-tivity effects of a MSLB accident would be reduced in severity with a concurrent loss of offsite power when conpared to the same event with offsite power available and the RCP's operating. Because the reactivity effects of a MSLB with the RCP's tripped after SIAS are less severe than a MSLB with a concurrent loss of off-site power, it is concluded that the reactivity effects for the MSLB case with the RCP's tripped after SIAS are i less severe than for a MSLB with offsite power avail-able and RCP's operating.

The reactor coolant volumetric flow rate is assumed to be constant during the incident. . . The LCO flow rate (197,000 gpm) was used in order to obtain the most ad-l

-51

l 5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) i l

5.8 Main steam Line Break Accident (Continued) 5.8.4 Key Parameter and Analysis Assumptions (Continued)

' arse results. A lower flow rate increases the initial fuel and average primary coolant tenperatures and conse-quently results in a higher steam generator pressure and a greater stean generator mass inventory. These effects cause a longer blowdown, a greater blowdown rate and a greater decrease in average primary coolant tenperature. After MSIV closure the lower flow rate decreases the rate of reverse heat transfer fran the intact steam generator, thereby increasing the heat extracted fran the primary steam by the ruptured stean generator. The overall effect is that the potential for a return-to-power is maximized.

Maximum values for the heat transfer coefficient across the steam generator are used for the no-load initial condition case, while noninal values are used for the full-load initial condition. These heat transfer coef-ficients result in the most severe conditions during the incident because of the shape of the reactivity versus moderator tenperature function and the differ-ence in average moderator tenperature for the maximum and minimum values of the steam generator heat transfer coef fici ents.

The fast cooldown following a MSLB results in a rapid shrinking of the reactor coolant. After the pressur-t izer is emptied, the reactor coolant pressure is assumed to be equal to the saturation pressure corres-ponding to the highest tenperature in the systen.

Safety injection actuation occurs at 1578 psia (i.e.,

1600 psia minus the 22 psia uncertainty) after the 52 1

- . i

5.0 TRANSIENT AND ACCIDENT ANALYSIS METiiOD (Continund) 5.8 Main Steam Line Break Accident. (Continued) 5.8.4 Key Parameter and Analysis Assumptions (Continued) pressurizer empties. Additional time is required for pump acceleration, valve opening, and flushing of the unborated part of the safety injection piping along with the requirenent that the RCS pressure decrease below the shutoff nead of the safety injection pumps (1376 psia for high pressure safety injection (HPSI) pumps and 201 psia for 1ow pressure safety injection pumps (LPSI) pumps). The analysis takes credit for one HPSI pump, one LPSI pump, and the safety injection tanks.

The boric acid is assumed to mix homogeneously with the reactor coolant at the points of injection into the cold legs. Slug flow is assumed for movement of the mixture through the piping, plena, and core. After the boron reaches the core midplane, the. concentration with-in the core is assumed to increase as a step function after each loop transit interval.

The boron concentration of the safety injection water is assumed to be at the Technical Specification minimum limit. The values of the inverse boron worth are con-servatively chosen to be large to minimize the negative reactivity insertion fron safety injection.

Since the rate of temperature reduction in the reactor ,

coolant systen increases with rupture size and with steam pressure at the point of rupture, it is assumed that a circumferential rupture of a 25-inch (inside diameter) steam line occurs at the steam generator main steam line nozzle, with unrestricted blowdown. Criti-cal flow is assumed at the point of rupture, and all of i

53

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.8 Main Steam Line Break Accident (Continued) i 5.8.4 Key Parameter and Analysis Assumptions (Continued) the mass leaving the break is assumed to be in the steam phase. This assumption results in the maximum heat removal from the reactor coolant per pound of secondary water, since the latent heat of vaporization is included in the net heat removal. A single failure of the reverse flow check valve in the ruptured steam generator is assumed; so that the intact steam gener-ator will have steam flow through the unaffected steem line and back through and out the ruptured line. Based on sensitivity analyses perfonned by the District, this is the most severe single failure for the steam line break event. The analysis credits a choke which is in-1 stalled in each steam line immediately above the steam generator and assumes the steam flow fran the intact steam generator is through a 507, area reduction choke I installed in a 24 inch steam line. This flow will be terminated upon MSIV closure.

The feedwater flow at the start of the MSLB corresponds to the initial steady state operation. For the full

load initial condition, it is autonatically reduced in accordance' with the program used in the valve control-l e r. For the no load initial condition, feedwater flow is assumed to match energy input by the reactor coolant pumps and the 1 MWt core power. Feedwater isolation upon the receipt of a low steam generator pressure (at 478 psia) is credited for both the full load and no load cases. A valve closure time of 30 seconds was used.

(

54

I 5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.8 Main Steam Line Break Accident (Continued) 5.8.5 Analysis Method The analysis of the main steam line break accident is performed using CESEC which models neutron kinetics with fuel and moderator tenperature feedback, the reactor protective system, the reactor coolant system, the steam generators and the main steam and feedwater systens.

5.8.6 Analysis Results and 10 CFR 50.59 Criteria The results of the analysis for the Fort Calhoun steam line break event are discussed in Section 14.12 of the 1983 update of the Fort Calhoun Station Unit No.1 USAR. The criteria of 10 CFR 50.59 are met if the calculated return-to-power is less than the return-to-power reported for the Cycle 1 analysis, using the cur-rent Technical Specification limit on shutdown margin and moderator tenperature coefficient.

5.8.7 Conservatism of Results Conservatism is added to the analysis by inclusion of uncertainties in moderator and fuel temperature coeffi-cients of reactivity, by taking no credit for void reac-tivity feedback, by taking credit for only 1 HPSI pump and by taking no credit for the stuck CEA worth, s

5.9 Seized Rotor Event 5.9.1 Definition of Event The seized rotor event is assumed to be caused by a me- -

chanical failure of a single reactor coolant pump. It 55

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.9 Seized Rotor Event (Continued) 5.9.1 Definition of Event (Continued) is assumed that the rotor shears instantaneously, leav-ing a low inertia impeller attached to a bent shaft.

This latter combination comes to a halt immediately causing a sharp drop in the flow rate. The rapid reduc-tion in core flow will initiate a reactor trip on low flow within the first few seconds of the transient.

5.9.2 Analysis Criteria A single reactor coolant pump shaft seisure is classified as a postulated accident for which the dose rates must be within 10 CFR 100 guidelines.

5.9.3 Objective of the Analysis The objective of the analysis is to demonstrate that the radiological releases are within a small fraction of 10 CFR 100 guidelines. This objective is met if it can be shown that less than 1% of the pins fail during the event.

5.9.4 Key Parameters and Analysis Assumotions The key parameters used in the analysis of the seized rotor event are given in Table 5.9.4-1. The seized rotor is conservatively assumed to result in a 0.1

• second rampdown of the core flow from its initial value -

to the 3 pump value. For CETOP calculations, [

l 3 i

56

Table 5.9.4-1 l KEY PARAMETERS ASSUt1ED Ifl THE SEIZED ROTOR ANALYSIS I

Parameter Units Value l i

Initial Core Power Level MW t 1500 i

Initial' Core Inlet O F Maximum allowed

• Tempera ture by Tech. Specs.

4 Initial Reactor Coolant psia Minimum allowed' System Pressure by Tech. Specs.

Initial Core Mass *10 lbm/hr. Minimum allowed +

Flow Rate by Tech. Specs.

Moderator Temperature *10 ap/ UF fiost positive allowed Coefficient by Tech. Specs.

Fuel Temperature *10-4ap/0F Least negative pre-Coefficient dicted during core life.

Core Average H BTU /hr-Ft. - F Minimum predicted gap a

during core life.

CEA Drop Time sec. Maximum allowed by Tech. Specs.

Scram Reactivity %Ap Minimum predicted Worth during core life.

I Urcertainties on these parameters are combined statistically.

5

, ~57

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.9 Seized Rotor Event (C?ntinued) 5.9.5 Analysis Method Two methods of analyzing the seized rotor event are dis-cussed in this section. Section 5.9.5.1 discusses a method which does not require transient analysis input.

Section 5.9.5.2 discusses a method which utilizes trans-ient analysis input.

5.9.5.1 Analysis Method Without Transient Analysis Response Input This method calculates the number of pin failures assuming that the core flow instan-taneously decreases to the 3-pump flow rate. This method utilizes the TORC anal-ysis with a 3-pump inlet flow distribution.

The initial RCS pressure and core inlet tem-perature are used as input to TORC and the core average heat flux is conservatively assumed to remain at its initial value.

The maximum value of F R T is combined with a conservatively flat power distribution.

The TORC calculation [

] the num-ber of pins that have failed is calculated.-

5.9.5.2 Analysis Methods Using Transient Analysis This method utilizes the CESEC code to cal-culate the transient response for the 58

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.9 Seized Rotor Event (Continued) 5.9.5 Analysis Method (Continued) 5.9.5.2 Analysis Methods Usina Transient Analysis (Continued) seized rotor event. The CETOP code is then used to determine the time of minimum DNBR.

The TORC code utilizes the 3-pump inlet flow distribution, 3-pump core flow rate, and the RCS pressure, core inlet tenpera-ture and core heat flux calculated at the time of minimum DNBR by CESEC. The steps to determine the number of pin failures is then performed [ ] as discussed in Section 5.9.5.1.

5.9.6 Analysis Results and 10 CFR 50.59 Criteria The results of the seized rotor analysis are contained in the 1983 update of the Fort Calhoun Station Unit No.

1 USAR. The criteria of 10 CFR 50.59 are met, if the number of pin failures is less than one percent.

5.9.7 Conservatism of Results Conservatism in the calculated number of fuel pins pre-dicted to experience DNBR is added through the use of the following assumptions:

s

1. The most positive MTC is assumed in the analysis. -

The actual MTC is more negative and would limit core power and heat flux rise.

59

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.9 Seized Rotor Event (Continued) 5.9.7 Conservatism of Results (Continued)

2. A relatively flat pin census is assumed in the an-alysi s. A more peaked pin census distribution would lower the number of pins predicted to exper-1ence DNB.

3. For the case without transient analysis, no cred-

} it is taken for the pressure increase during the transient and calculating the minimum transient DNBR.

5.10 CEA Ejection Accident 5.10.1 Definition of Event A CEA ejection accident is defined as a mechanical failure of a control rod mechanical pressure housing such that the coolant system pressure would eject the CEA and the drive shaft to a fully withdrawn position.

The consequences of this mechanical failure is a rapid reactivity insertion which when combined with an adverse core power distribution potentially leads to localized fuel damage. The CEA ejection accident is the most rapid reactivity insertion that can be reasonably postulated. The resultant core and thennal power excursion is limited primarily by the Doppler reactivity effect of the increased fuel teaperatures and is tenainated by reactor trip of the remaining CEA's activated by the high power trip or variable high power trip.

60

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.10 CEA Ejection Accident (Continued) 5.10.2 Analysis Criteria The CEA ejection event is classified as a postulated accident. The design and limiting criteria are:

1. The average fuel pellet enthalpy at the hot spot will be equal to or less than 280 calories / gram.

2. The peak reactor pressure during a portion of the transient will be less than the value that will cause stress to exceed the emergency conditions stress limits as defined in Section 3 of the ASME Boiler and Pressure Vessel Code.

3. Fuel melting will-be limited to keep the offsite dose consequences well within the guidelines of 10 CFR 100.

These limiting criteria are taken fron the NRC Regula-tory Guide 1.77 " Assumptions Used for Evaluating a Con-trol Rod Injection Accident for Pressurized Water Reac-to rs" .

5.10.3 Objectives of the Analysis The objective of the analysis is to demonstrate that the. total average enthalpy of thr. hottest fuel pellet for the hot full power and hot 2ero power cases is less than ~that reported for the reference cycle.

~

l 5.10.4' Analysis Method The Olstrict utilizes the CEA Ejection Accident Analysis Methodology of our current fuel vendor, Exxon l

1 61

1 5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.10 CEA Ejection Accident (Continued) 5.10.4 Analysis Method (Continued)

Nuclear Corporation. This analysis methodology is

, documented in Reference 5-5. This methodology utilizes physics parameters, calculated by the District in accordance with the methods outlined in Reference 5-6.

The power peaking factor, F q T, is defined as the post ejected 3-D fuel rod power peak.

5.10.5 Analysis Results and 10 CFR 50.59 Criteria The results of the CEA Ejection Analysis are reported in Section 14.13 of the Fort Calhoun Unit 1 USAR.

Criteria of 10 CFR 50.59 are satisfied if the total average enthalpy of the hottest fuel pellet is less than or equal to the values reported in the reference cycle.

5.10.6 Conservatism of Results The major area of conservatism is the calculation method used to obtain the ejected CEA worth and the ejected radial peak. The ejected worth and the ejected radial peak are calculated without any credit for Doppler or Xenon feedback. In addition, the hot full power ejected worth and ejected peak are calculated assuming the no-load temperature of 532 F. The lower temperature-is more adverse since this .causes a power role to the core periphery which also happens to be the location of the ejected CEA. Also, the ejected worth is calculated assuming the CEA'S are fully inserted for hot full power case regardless of PDIL. Thus, the ejected worth is conservative.

62

5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.11 Loss of Coolant Accident The District does not perform the loss of Coolant Accident Analysi s. The large break loss of coolant analysis was performed by Exxon Nuclear Corporation (ENC) and the small break analysis was performed by Combustion Engineering. The large break analysis shows the closest approach to the Appendix K criteria for ECCS analysis. The District verifies that the physics input assumptions and the maximum rod burnup are within the bounds assumed in the ENC large break analysis.

1 k

S l

1 1

63

6.0 TRANSIENT ANALYSIS CODE VERIFICATION 6.1 Introduction The District utilizes the CESEC-III coaputer code to calculate the transient response of the NSSS during events discussed in this document. Canbustion Engineering has provided overall verification of the CESEC-III code in Reference 6-1. The purpose of the work reported here is to demonstrate the Dis-trict's ability to correctly utilize the CESEC-III code.

In order to demonstrate Omaha Public Power District's ability to correctly use the CESEC-III canputer code, verification work has been perfomed by benchmarking both actual plant transient data and independent safety analyses previously accepted by the NRC. The plant transients which were bench-marked were the Turbine-Reactor trip and Four-Pump Loss of Coolant Flow events. The ind9 pendent safety analyses which ~

were benchmarked were the Dropped CEA, Main Steamline Break, and RCS Depressurization events. Each of the conparisons will be addressed below.

6.2 Comparison to Plant Data A prerequisite for beginning perfornance of transient analyses is verification that the code will stabilize with the correct systen parameters when simulating steady state operation.

This step was perfomed following setup of the CESEC-III code and correct results were obtained.

For plant transient benchmarking, the type of transients that have occurred and both the quality and quantity of data exist-ing for each is very limited. In nearly all cases, operators ,

take actions which reduce the consequences of the event, intro-ducing ccmplicated perturbations in system response which can-not be easily modeled, because the actions taken and the time at which they are perfomed are not recorded. Strip chart re-cordings on an extrenely caapressed time scale are generally i

l 64

6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued) 6.2 Comparison to Plant Data (Continued) the only fonn of data available. This canpressed time scale (with graduations typically of 10 minutes) do not permit ade-quate comparisons to CESEC-III modeling in which seconds are of major concern. The only source of plant transient data in which system parameters were measured with high speed strip chart recorders and no operator action taken, was during the Cycle 1 startup testing. Good data existed for a nominal full power turbine-reactor trip and a 35*. power total loss of RCS flow event. The CESEC-III computer code was set up to model Cycle 1 in a best estimate mode to pemit accurate conparisons to the actual measured plant responses for both of the above cases. A summary of each of these comparisons follows.

~

6.2.1 Turbine-Reactor Trip For the turbine-reactor trip case, the plant compar-ison data were obtained fran the Cycle 1.startup test-ing perfomed May 10, 1974. The event was initiated fran 97% of full power, all-rods-out, and equilibrium xenon. The plant response data used in the CESEC-III comparisons were obtained fran vendor test recorders.

No operator action was taken following the manual gen-erator-turbine trip (which provided the RPS " loss of load" trip). Prior to the trip the main feedwater, the pressurizer pressure, and pressurizer level con--

trol systems were all in the automatic mode, and the letdown backpressure control valve was in the manual mode. With the exception of adjusting the letdown '

backpressure control valve at 20 seconds, no operator action was taken for 60 seconds following the trip.

Figures 1-1 through 1-7 show plots of the conparisons between the measured plant responses and the CESEC-III predicted responses. It should be noted that '

l 65 1

6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued) 6.2 Comparison to Plant Data (Continued) 6.2.1 Turbine-Reactor Trip (Continued) this test was perfonned based on a rated power level of 1420 MWt rather than the current limit of 1500 MWt (the design power for which licensing was obtained in Cycle 6).

Figure 1-1 shows the nuclear power response following the turbine-reactor trip. The CESEC-III prediction follows the same power decay rate, however, the end-point residual power is slightly higher, i.e., conser-vative. It should be noted that trip delays included in the CESEC-III modeling prevent the immediate power drop observed in the plant data; again this is conser-vative. The pressurizer pressure response predicted by CESEC-III and shown in Figure 1-2 shows very good agreenent with the plant response. The CESEC-III case was initiated 10 psia above the plant data and remained slightly above the plant response for the duration of the transient. The difference-between the predicted and measured pressurizer pressures in-creased slightly due to the higher residual power after trip as shown in Figure 1-1. This difference between the predicted and measured pressurizer pres-sures at 60 seconds is only 19 psia, a value which is less than the pressure measurenent uncertainty. Fig-ure 1-3 shows the pressurizer level response. The comparison between the measured and predicted values- .

shows excellent agreement. Figure ~1-4 shows the RCS cold-leg 'and hot-leg tenperature responses for each steam generator loop for the plant data and the CESEC-III predicted average cold-leg and hot-leg ten--

peratures. The differences in the transient response . -1 of the two steam generator loops.for the plant data -

66

i 6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued) l 6.2 Comparison to Plant Data (Continued) 6.2.1 Turbine-Reactor Trip (Continued) is attributable to the differences in the main feed-water flow rate rampdown after trip (see Figure 1-5).

The CESEC-III responses lead the loop measurements because of the measurement delays associated with the response time of the RTDs (resistance temperature de-vices) providing the temperature signals. Figure 1-5 shows the measured and predicted steam generator pres-sure responses. These two plots show very good agree-ment with each other with only minor differences.

The predicted pressure is slightly higher early in the event due to a combination of the greater heat residual as shown in Figure 1-1, a quicker turbine stop valve closure, and quicker steam dump-bypass .

~

operation assumed in the CESEC-III analysis. The latter two. effects, which are shown in the steam flow of Figure 1-7, would show better agreenent if the CESEC-III input were modifled, however, the overall differences are small enough not to warrant the rean-alysis.

In conclus son, the CESEC-III predicted parameters for the turbine-reactor trip show very good agreement with those measured in the Cycle 1 startup testing perfonned at nominal full power conditions.

6.2.2 Four-Pump loss of Coolant Flow For the four-pump loss of coolant flow case, the plant comparison data were obtained fran the Cycle 1 startup test perfomed March 6,.1974. This event was initiated fran 35', power by inanually and simulatane-ously tripping all four reactor coolant pumps. At i

67

r i

l l

6.0 TRANSIENT ANALYSIS CODE VERIFTCATION (Continued) i 6.2 Comparison to Plant Data (Continued) 6.2.2 Four-Pumo loss of Coolant Flow (Continued) the time of trip the pressurizer pressure, pressuri-zer level, main feedwater, and steam dump and bypass controllers were in the automatic mode. At approxi-mately 20 seconds after the trip, the operators took manual control of feedwater in order to preclude over-feeding of the steam generators and too rapid of a cooldown for the following natural circulation test.

The behavior of the various RCS and secondary parame-ters that were measured and the CESEC-III predictions for the first 30 seconds following the RCP trips are shown in Figures 2-1 through 2-8. These comparisons show excellent agreement. The minor differences that exist are discussed below.

Figure 2-1 shows a plot of the measured total RCS flow versus time and that predicted by the CESEC-III code which incorporates explicit modeling of the reac-tor coolant pumps. These data show excellent agree-ment with the predicted flow being slightly conserva-tive. Figures 2-2 and 2-3 show the pressurizer pres-sure and level response comparisons which also show excellent agreement. Figure 2-4 shows plots of core nuclear power versus time. As in the turbine-reactor.

trip case, CESEC-III shows a slightly higher residual power after trip. The predicted and measured steam

• generator pressure responses as plotted in Figure 2-5, also show very good. agreement. The response of the hot-leg and cold-leg temperatures, as shown in l Figure 2-6, is consistent with the data obtained from the turbine-reactor trip case. Again the delay asso- I ciated with the RTD- response causes the predicted tem-68 l

I 6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued) 6.2 Comparisor,10 Plant Data (Continued) 6.2.2 Four-Pumo Loss of Coolant Flow (Continued) peratures to lead those that were measured. Figure 2-7 shows that the main feedwater input function used in CESEC-III was acceptable in tenns of the actual feedwater system response. It should be noted that the operator action of assuming manual control of the main feedwater system at approximately 20 seconds had little effect on any of the other system parameters examined, and that following a several second reduc-tion in flow the previous flow rate was reestab-lished. Figure 2-8 shows that turbine stop valve closure rate assumed in the CESEC-III analysis was quicker than the actual valve response. The figure also shows a steam flow rate mismatch between the two steam generators for the plant data. This is some-thing one would not expect and raises the question of the validity of the measurenent or its uncertainty for this steam generator steam rate flow, because the two corresponding feedwater flow rates (in Figure 2-7)areconsistent.

In conclusion, the CESEC-III predicted parameters for the 35% power total loss of coolant flow show very good agreement with those measured during Cycle 1 startup testing.

6.3 Comparisons Between OPPD Analyses and Indeoendent Analyses Previously Performed by the Fuel Vendors Of the transients analyzed by 0 PPD for reload core licensing (using CE methodology) no plant data existed, so conparison of the limiting events to previous independent analyses perfonned by either Exxon Nuclear Company (ENC) or Combustion Engineer-69

6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued) 6.3 Comparisons Between OPPD Analyses and Independent Analyses Previously Performed by the Fuel Vendors (Continued) ing (CE) was done. For the conparison cases, the assumptions used in the analyses were similar to those used by the Dis-trict, i.e., the core physics parameters did not vary signifi-cantly between fuel cycles. The events chosen for comparison were:

(1) The Dropped CEA event is dependent upon the initial available overpower margin to prevent exceeding the SAFDL's. The goal of the analysis is to detennine the DNBR required overpower margin (R0PM).

(2) The Hot Zero Power (HZP) Hain Steamline Break which detemines the minimum required shutdown margin.

(3) The Hot Full Power (HFP) Main Steamline Break which detennines the most negative moderator temperature coefficient of reactivity allowed.

(4) . The RCS Depressurization event which is used in the detennination of the [ ]. The [.

] accounts for DNBR margin degradation in the thennal margin / low pressure (TM/LP) trip [

]

6.3.1 Dropped CEA B

The Cycle 8 Dropped CEA anclysis perfonned by 0 PPD was conpared to the previous analysis, contained in the Updated Safety Analysis Report (USAR). The USAR analysis was. perfonned by ENC for Cycle 6. Table 1 summarizes the parameters and their values for Cycles 70

l 6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued) 6.3 Comparisons Between OPPD Analyses and Independent Analyses l Previously Performed by the Fuel Vendors (Continued) l 6.3.1 Dropped CEA (Continued) 6 and 8. Plots of core power versus time for the OPPD (Cycle 8) and ENC (Cycle 6) analyses are found in Figure 3-1. The curves show a very similar prompt drop, to 69% versus 70%, respectively, and both cases show a return to a nominal 100% power. Both cases assumed that the turbine admission valves opened to their full open position in an attempt to maintain full load during the event (i.e., the turbine control system was placed in the load set mode which is not used at Fort Calhoun Station). The core heat flux plots are contained in Figure 3-2. Both are very sim-ilar, as was the case in the core power cases. Fig-ure 3-3 contains c'.ats of the coolant average tempera-ture versus time. Both figures are in good agreenent showing a drop in average coolant temperature to 567 F. Plots of the inlet and outlet temperatures for Cycle 8 are also included. Figure 3-4 shows plots of the pressurizer pressure versus time. The minimum pressures predicted at 160 seconds are 1957 psia and 1945 psia for Cycle 8 and Cycle 6, respect-ively. This difference is small enough to be less than the pressure measurement uncertainty.

In summary, the primary system responses between the ENC and OPPD analyses show excellent agreement with each other which is consistent with reload cores

• having similar core physics parameters.

6.3.2 Hot Zero Power Main Steamline Break The hot zero power (HZP) Main Steamline Break, which is the basis for detennination of the required shut-71

6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued) 6.3 Comparisons Between OPPD Analyses and Independent Analyses Previously Performed by the Fuel Vendors (Continued) 6.3.2 Hot Zero Power Main Steamline Break (Continued) down margin, was analyzed by 0 PPD for Cycle 8. The results of this analysis have been compared to those of ENC in their Cycle 6 analysis and to those ob-tained by CE in their Cycle 6 control grade auxiliary feedwater (AFW) system analysis. Table 2 shows com-parisons of the pertinent input values for each of the analyses.

Figure 4-1 shows plots of core power for the Cycle 8 OPPD analysis and Cycle 6 ENC analysis, respectively.

The maximum return-to-power is less for Cycle 8 than for Cycle 6 and occurs later due to the use of a high-er shutdown margin. The Cycle 6 CE AFW analysis power is not included because there was no return-to-critical and no return-to-power. Figure 4-2 shows plots of the core average heat flux for OPPD, ENC and d

CE, respectively. Both the OPPD and CE analyses, which were performed using CESEC-III and CESEC-I, res-pectively, show a slight heat flux increase at approx-inately 12 seconds. This is due to suberitical multi-plication. Otherwise, the heat flux curves within the specific analyses are essentially the sams as the core power curves with a slight decay. Figure 4-3 shows the total reactivity versus time for each of the analyses. With very similar moderator cooldown 5 curves, the peak reactivities occur chronologically with increasing shutdown margin as expected; i.e.,

for increased shutdown margin (CEAs) it takes longer to be offset by the positive moderator cooldown reac-tivity insertion.

72

6.0 TRANSIENT ANALYSIS CODE VERTFICATION (Continued) l 6.3 Comparisons Between OPPD Analyses and Independent Analyses Previously Performed by the Fuel Vendors (Continued) 6.3.2 Hot Zero Power Main Steamline Break (Continued)

Figure 4-4 shows plots of RCS pressure versus time for Cycle 8 (0 PPD) and Cycle 6 AFW (CE); Also in-cluded in Figure 4-4 is the Cycle 1 (CE) results.

All three of these curves show excellent agreenent.

The Cycle 6 AFW (CE) analysis shows a lower endpoint pressure than the Cycle 1 (CE) and Cycle 8 (OPPD) analyses due to the assumption of auxiliary feedwater addition. The ENC data available did not include the RCS pressure response.

Figure 4-5 shows plots of the steam generator pres-sures for Cycle 8 (OPPD) and Cycle 6 AFW (CE), res-pectively. These plots show reasonable agreenent between pressures and times. The increase in the intact steam generator's pressure is due to MSIV closure; i.e., failure of the reverse flow check valve onsthe _ intact steam generator was chosen as the most adverse single failure. Following dryout of the runtured steam generator, the pressure drops to atmos-pheric. The times of dryout are slightly different due to the increased nonnal water level value used in the Cycle 8 analysis.

In summary, the HZP Main Steamline Break analysis for Cycle 8 shows trends si.nilar to those in Cycle 6 as ,

analyzed by both CE 'and EfC.

6.3.3 Hot Full Power Main Steamline Break The hot full power (liFP) Main Steamline Break pro-  !

vides an acceptance criteria for the most negative s u l 73  ;

/ ,

6, s

6.0 TRANSIENT ANALYSIS CODE VERIFICATI0ft (Continued) 6.3 Comparisons Between OPPD Analyses and Independent Analyses PreviouslLPerformed by the Fuel Vendors (Continued) 6.3.3 Hot Full Power Main Steamline Break (Continued) moderator tanperature coefficient (MTC) of reactiv-ity. If a return-to-critical occurs, the goal of the reload analysis is to show that the return-to-power is bounded by the most limiting case which, for the Fort Calhoun Station, is the Cycle 1 analysis. The Cycle 8 HFP analysis of this event was conpared to the previous analyses perfonned by ENC in Cycle 6 and by CE in-their Cycle 6 control grade AFW systen anal-ysis. Table 3 shows a conparison of the important input parameters for each of the analyses.

Figures 5-1, 5-2, and 5-3 show plots of core power, core average heat flux, and total reactivity for Cycle 8 (0 PPD), Cycle 6 (ENC), and Cycle 6 AFW (CE).

Within each cycle's analysis, the core average heat flux slightly lags the core power which peaks at a time several seconds after the peak reactivity is reached (for the return-to-critical cases). The return-to-power peaks occur at different times due to the different scram worths used, as explained for the shutdown margin in the HZP Steamline Break analysis 4

section. y 9

Figure 5-4 shows plots of the RCS pressure versus time for the Cycle 8, Cycle 6 AFW, and Cycle 1 anal- i yses. These plots are very similar and show excel-lent av eement. Figures 5-A and 5-B show plots of the RCS temperatures for Cycle 8 and Cycle 6 AFW.

Again good agreement exists to approximately 180 seconds. .At this time, the Cycle 6 AFW analysis assumed runout' F1,0w from both AFW pumps to the rup- i 74

l 6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued) 6.3 Comparisons Between OPPD Analyses and Indeoendent Analyses Previously Performed by the Fuel Vendors (Continued) 6.3.3 Hot Full Power Main Steamline Break (Continued) tured steam generator which resumed the RCS cooldown.

This additional cooldown caused by the AFW system is prevented from occurring in Cycle 8 by the logic of the neuer safety grade AFW system.

Figure 5-6 shows plots of steam generator pressures versus time for Cycle 8 and Cycle 6 AFW (CE). These results are very similar except that the intact steam generator pressure, in the CE analysis, begins to drop after 180 seconds due to the AFW induced RCS cooldown.

6.3.4 RCS Depressurization The RCS Depressurization analysis is perfonned to cal-culate a [ ] for the TM/LP trip which accounts for the DNBR margin degradation

['

]

Because no figures fran previous cycle analyses ex-ist, comparison was made between the transient anal-ysis training manual sample analysis and the figures generated by 0 PPD for Cycle 8. Pertinent input para-meters are summarized in Table 4.

Figure 6-1 shows the plots of RCS pressure versus time for the initial case run without a trip which is used to detemine the time manual trip is to be used.

75

6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued) 6.3 Comoarisons Between OPPD Analyses and Indeoendent Analyses Previously Performed by the Fuel Vendors (Continued) 6.3.4 RCS Depressurization (Continued)

A manual trip is next simulated at the time of maxi-mun margin degradation; i.e., at the time the maximum RCS Depressurization rate occurs. The maximum RCS Depressurization rate occurs in approximately the first 20 seconds and is constant. Therefore, the time at which a manual trip should occur is arbitrary but must be in the first 20 seconds. A trip time corresponding to a 100 psia drop is adequate to perform the analysis.

Figure 6-2 shows plots of core power versus time for the Cycle 8 analysis and CE's example. The core aver-age heat flux curves are found in Figure 6-3. The RCS pressure versus time plots are shown in Figure 6 -4 . In the CE example, the initial pressure was 2300 psia, a value which corresponds to the maximum pressure before which tne pressurizer sprays will be activated in a 2700 MW(th) class plant (whose normal RCS pressure is 2250 psia). In the Cycle 8 analysis, a value of 2172 psia was used for the initial RCS l pressure, since the normal operating RCS pressure at Fort Calhoun is 2100 psia. The Fort Calhoun pressur-izer sprays are fully closed at 2175 psia and fully open at 2225 psia.

5 s The comparison of the figures show good agreement in the trends for the core power, core average heat flux, and RCS pressure. The [

]

76

6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued) l 6.4 Summary Initial setup and operation of the CESEC-III code was per-

formed by showing that the code stabilized for steady state plant operation. Benchmarking against Cycle 1 plant data for the Turbine-Reactor Trip and the Four-Pump Loss of Coolant Flow was performed and excellent agreenent between the pre-dicted and observed responses was obtained.

For transients in which plant data were not available, conpar-isons were performed between the OPPD Cycle 8 analyses of the limiting transients and the Cycle 6 anlayses of the fuel ven-dors (CE and ENC) and, in one case, the transient analysis training manual example. In all cases, these benchmarking comparisons showed very good agreenent.

4 5

77

i TABLE 1-

' COMPARISON OF PARAMETERS INCLUDIrlG UNCERTAINTIES USED IN THE CEA DROP ANALYSES FOR CYCLES 6 AND 8 Parameter Units Cycle 6 Cycle 8 Initial Core Power Level MWt 102% of 1500 102% of 1500 Core Inlet Temperature *F 547 547 Pressurizer Pressure- psia 2053 2053

' RCS Flow Rate gpra 190,000 197,000 Moderator Temperature Coeff. 10-4 ap/*F -2.3' -2.7 Doppler Coeff. Multiplier 1.20 1.15 CEA Insertion at Full Power  % Insertion 0.0 25.0 Dropped CEA Worth %Ap -0.34 -0.28 l

a

$ g 76

TABLE 2 COMPARISON OF PARAMETERS INCLUDING UNCERTAINTIES USED IN THE HZP MAIN STEAMLINE BREAK ANALYSIS FOR CYCLES 6 AND 8 Cycle 6 Parameter Units Cycle 6 AFW Cycle 8 Initial Core Power Level MWt 0.0 1.0 1.0 Core Inlet Temperature *F 532 532 532 Pressurizer Pressure psia 2053 2175 2172 RCS Flow Rate gpm 190,000 190,000 197,000 Effective Moderator Temperature 10-4 ap/ F -2.3 -2 . 3 -2.5 Coefficient Doppler Coeff. Multiplier 0.8 1.15 1.15 Minimum CEA Scram Worth  % ap -3.0 -4.2 -4.0 (Shutdown Margin)

Initial Steam Generator Pressure psia N/A 90 0 895 Initial Steam Generator Mass  % Narrow 63 63 70 Inventory (Level) Range Scale l

79

TABLE 3 COMPARISON OF PARAMETERS INCLUDING UNCERTAINTIES USED IN THE HFP MAIN STEAMLINE BREAK ANALYSES FOR CYCLES 6 AND 8 Cycle 6 Parameter Units Cycle 6 AFW Cycle 8 Initial Core Power Level MWt 102% of 1500 102% of 1500 102% of 1500 Core Inlet Temperature 'F 547 547 547 4

Pressurizer Pressure psia 2078 2175 2172 RCS Flow Rate gpm 190,000 190,000 197,000 Moderator Temperature 10-4 ap/*F -2.3 -2.3 -2.5 Coefficient-Doppler Coef f. Multiplier 0.8 1.15 1.15 Minimum CEA Scram Worth  % ap -5.81 -5.81 -6.68*

Initial Steam Generator psia N/A 880.5 890 Pressure l

Initial Steam Generator  % Narrow 63 63 70 Mass Inventory (Level) Range Scale

,

• Reduced to -6.57 to account for axial shape.

4 e

6 1

i 80 l l

TABLE 4 i

COMPARISON OF PARAMETERS INCLUDING UNCERTAINTIES USED IN THE RCS DEPRESSURIZATION ANALYSES FOR CYCLES 6 AND 8 Parameter Units Example Case

• Cycle 8 Initial Core Power Level MWt 102% of 1500 102% of 1500

, Core Inlet Temperature *F 547 547 Pressurizer Pressure psia 2300 2172 RCS Flow Rate gpm N/A 209,796 Moderator Temperature 10-4 ap/ F -2.5 -2.7 Coefficient Doppler Coeff. Multiplier 1.15 1.15

• Example case input data consistent with 2700 MWt plant operating characteristics.

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PLANT DATA: TEST PERFORMED MAY 10, 1974 FullPowerTurbineTrip DaahaPublicPowerDistrict Figure SteamFlowvsTime FortCalhounStation-UnitNo.i 1-7 .

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l .4-Pu:pLossOfFlow CmahaPublicPowerDistrict Figure L

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CEADropIncident OmahaPublicPowerDistrict' Figure CoolantTemperaturevsTime FortCalhounStation-UnitNo.i 3-3

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, CEACropIncident 0:laha Public Power District figure i PressurizerPressurevsTime FortCalhounStation-UnitNo.i 3-4 100

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'1 '- - - - - -

0 40- 80 120 160 200 TIME, SECONDS NOTE :

CYCLE - 6: ENC ANALYSIS WITH SDM= 3.0% y CYCLE 8: OPPD' ANALYSIS WITH SDM= 4.0% 9 ZeroPowerSteamLineBreakIncident OmahaPublicPowerDistrict Figure CorePowervsTime FortCalhounStation-UnitNo.i 4-1

ini

i<

l ,

\,

I O

in 80 i i  ; i M

d 60 -

,g -

ti.

O y 40 _ /

.i s. ," CYCLE 6 4MB i

Q 20 -

y ,/

'N cLE 6 Afil e

h ,:0

/ '-

'N

@ 0 40 80 120 160 200 tu g > TIME, SECONDS 4 +

a

~

, j

/

~

NOTE : i t -

CYCLE 6.' ENC ANALYSIS WITH SDM= 3.0%as CYCLE 6 AFW:

CEANALYSISWITHSDM=4.h%af CYCLE 8: OPPD ANALYSIS-WITH SDM= 4.0%sp L ,

r

.e,

'ZeroPowerSteasLineBreakIncident .0mahaPublicPowerDistrict Figure '

CoreAverageHeatFluxvs. Time FortCalhounStation-UnitNo.i 4-2 7 ,

,g idd .

8~<'

4 i i i i 3 -

2 -

4 x CYCLE 6

. 1 -

N

- [ CYCLE 8 g

0 ,-

9 -

1

/ y' ,.

d -1 -i j' ' ' -

-3

$ !i / '

H ii /

/

-2 '['

l/

CYCLE 6 AFW

i. f I se

-3

-f I

-4 I I __

0 40 80 120 130 200 TINE, SECONDS.

NOTE: '

CYCLE 6: ENCANALYSISWITHSDM=3.0%sp CYCLE 6 AFW: CE ANALYSIS WITH SDM= 4.2% 6,g CYCLE 8:

OPPDANALYSISWITHSDM=4.0%af ZeroPowerSteamLineBreakIncident CaanaPublicPowerDistrict F'qure TotalReactivityvsTite FortCalhounStation-UnitNo.1 3 g___

1M

2400 i i i i I

n 2000 -

5 '

!2

, CYCLE i w 1600 -

5 \

e \\

bE \,

a.

s 1200 -

i -

3 \ cLE 6 AFX o

8 g 800 -

i, \ -

t; 4,\

5 a-

- jj ,,,,,,,,,,,,,

-.~.- _. _ - - .

0 1 ' ' '

0 40 80 120 160 200 TIME, SECONDS NOTE :

CYCLE 1: CE ANALYSIS CYCLE 6 AFW: CE ANALYSIS CYCLE 8: OPPD ANALYSIS ZeroPowerSteamLineBreakIncident OmahaPublicPowerDistrict Figure CoolantSystemPressurevsTime FortCalhounStation-UnitNo.i 4-4 ,

1000 i i t i M

a.

800 - -

5 600 - m s.s. m a a h . _

E g ---

m ACT S.G. CYCLE 8 y 460 - } ' s, -

5 e

i t

$ \,,

'\

g -

r"J 200 -

\ s.s. CYCLE 6 AFW cn g' . . CYCLE 8 0 I ' ' i_. 1 '

0 40 80 120 160 200 TIME, SECONDS NOTE :

CYCLE 6 AFW: CE ANALYSIS i CYCLE 8: OPPD ANALYSIS ZeroPowerSteamLineBreakIncident OmahaPublicPowerDistrict- Figure SteamGeneratorPressurevsTime FortCalhounStation-lJnit?!o.i 4-5 ins

140 i i i 1 120 -

N I

i 100'g -

8 i

'E g 80 -

a

[ 60 - -

5

a. f y

a 40 -

[" 8 " -

! CLE B 20 --

---

Y'8

, ,'-_' :~ry= __.-- -  ; - ...q.

0 40 80 120 160 200 TIME, SECONDS NOTE :

CYCLE 6: ENC ANALYSIS (CEA WORTH = 5.81Lf )

CYCLE 6 AFW: CE ANALYSIS (CEA WORTH = 5.81hf)

CYCLE 8: OPPD ANALYSIS (CEA WORTH = 6.57hp)

FullPowerSteamLineBreakIncident Om6haPublicPowerDistrict Figure CorePowervsTime FortCalhounStation-Unit 110.i 5-1

"'a

120 i i i i

s A

S 100 E -

S 4

7"8 F-1 /

5 I

80 --

e il Cd 6 AFM

" (( l n

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$ 40 - \.\ -

w I.

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CLE 6

< 20 -

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o '\'::z---

'-~~~~

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: 2 .- . =N.

0 ' _I 0 40 80 120 160 200 TIME, SECONDS NOTE :

CYCLE 6: ENC ANALYSIS CYCLE 6 AFW: CE ANALYSIS CYCLE 8: OPPD ANALYSIS FullPowerSteamLineBreakIncident OmahaPublicPowerDistrict figure '

CoreAveraceHeatFluxvsTime FortCalhouaStation-Unit.No.i 5-2 107  !

1 2 i i i i CE 6 AFW

' ~

y Y P,~. . -_

n i

,/ \

~

% -2

/ cm a w D

/ -

u //

e

" .li I

-4 -

d pf ce s B

-6 -

-8 ' ' ' '

0 40 80 120 160 200 TIME, SECONDS NOTE :

CYCLE 6: ENC ANALYSIS (CEA WORTH = 5.81Lf ) '

CYCLE S AFW: CE ANALYSIS (CEA WORTH = 5.81hf )

CYCLE 8: OPPD ANALYSIS (CEA WORTH = 6.57h ) f l

L i

l FullPowerSteamLineBreakIncident OmahaPublicPowerDistrict Figure TotalReactivityvsTime FortCalhounStation-UnitNo.1 5-3 10A

2400 , , , ,

\ LYcLE 6 M 2000 g 1600 -

u3 -

E a -

a 1200 -

83 e

N .6 0

cr 800 -

\ , -

i ,

'N 400 -

q ,_____

0 I ' ' '

0 40 80 120 160 200 TIME, SECONDS NOTE : '

CYCLE 1: CE ANALYSIS CYCLE 6 AFW: CE ANALYSIS CYCLE 8: OPPD ANALYSIS FullPowerSteamLineBreakIncident OmahaPublicPowerDistrict Figure CcolantSystemPressurevsTime FortCalhounStation-UnitNo.-1 5-4 100

I 700 , , , ,

T out' u- 600 T avg -

c.o E T in us E 500 - -

E

& 'x ,

[

W 400 -

N~ -

5 8

a 300 -

I 200 I ' I  !

0 40 80 120 160 200 TIME, SECONDS NOTE:

CYCLE 6 AFW: CE ANALYSIS .

j Full Power Steam Line Break Incident 0 alla Public Pa' der District Ficure l Reactor Ceolant Temperatures vs Ti.te ! Fcrt Calhoun Station-Unit No. i s er i "a .. . . . ..

700 l i i i i T out

u. 600 -

g T avg tu o .

T in en E 500 -

?

E N E 400 -

~

N 5

a 8 300 -

200 ' ' ' '

0 40 80 120 160 200 TIME, SECONDS NOTE:

CYCLE 8: OPPD ANALYSIS '

l l

Full Power Steam Line Break Incident DaahaPublicPowerDistrict Figure i i ReactorCoolantTemaeraturesvsTir.e FortCalhounStation-Unittia.i s-es i.

1000 i i i I M

c.

800 -

CLE 6 H INTACT S.G. -

W g s g i 's m 600 - -

. m a m m S.G. _

l E l

~\'x m

?

y 400 -

5 e ,

{ CLE 6 H RPTURED 5.6.

W 200 -

\ g 7CLE B RElrnED S.G.

1 I

0 0 40 80 120 160 200 TIME, SECONDS NOTE :

CYCLE 6 AFW: CE ANALYSIS '

CYCLE 8: OPPD ANALYSIS FullPowerSteamLineBreakIncident OmahaPublicPowerDistrict Figure SteamGeneratorPressurevsTite FortCalhounStation-UnitNo.i 5-6 12

~ -

2400 l

l 2300 3

, 2200 a

in E 2100

!'s cc 2000 1900 __ _

0 4 8 12 16 TIME, SECONDS NOTE :

EXAMPLE CASE: CE ANALYSIS FOR 2700 MWt UNIT CYCLE 8: OPPD ANALYSIS HCSDepressurizationEvent D.tanaPublicPowerDistrict Figure RCSPressurevsTime FortCalhounStation-UnitNo.i 6-1 111

120 110 100 90 e

g 80 C

e 70 60 E

5 50 E 40 8

30 20 10 0,_ _ . , _ _ ,

0 2 4 6 8 10 12 14 16 TIME, SECONDS NOTE :

EXAMPLE CASE: CE ANALYSIS FOR 2700 MWt UNIT CYCLE 8: OPPD ANALYSIS RCSDepressurizationIncident OmahaPublicPowerDistrict Figure CorePowervsTime FortCalhounStation-UnitNo.i 6-2 m _ - _ _ - - _ - - _ - - - - - - - - - - -.

110 100 5

cd D 90 E

B Q

80 Ei a

>i y 70 -

E 60 50,_ _.

0 2 4 6 8 10 12 14 16 TIME, SECONDS NOTE :

EXAMPLE CASE: CE ANALYSIS FOR 2700 MWt UNIT CYCLE 8: OPPD ANALYSIS ACSDepressurizationIncident OmahaPublicPowerDistrict Figure CoreAverageHeatFluxvsTime FortCalhounStation-UnitNo.i 6-3 115

~ ~

2300 2250 y 2200 ,

E  !

l2150 a

E 2100 0

cc 2050 2000 1950 1900 _ _

0 2 4 6 8 10 12 14 16 TIME, SECONDS NOTE : '

EXAMPLE CASE: CE ANALYSIS FOR 2700 MWt UNIT CYCLE 8: OPPD ANALYSIS ACSDepressurizationIncident OmahaPublicPowerDistrict Figure

.__RCSPressurevsTime FortCalhounStation-UnitNo.i 6-4 16

7.0 REFERENCES

)

, Section 4 References 4-1 CENPD-107, "CESEC-Digital Simulation of a Combustion Engineering Nuclear Steam Supply Systen", C-E Proprietary Report ( April ,1974).

4-2 CENPD-107-P, Supplement 1, "ATUS Model Modifications to CESEC", C-E Proprietary Report (September,1974).

4-3 CENPD-107, Supplement 2-P, "ATUS Models for Reactivity Feedback and Effect of Pressure on Fuel", C-E Proprietary Report,(September,1974).

4-4 CENPD-107, Supplement 3, "ATWS Model Modifications to CESEC",

C-E Non-Proprietary Report (August,1975).

4-5 CENPD-107, Supplement 1, Amendment 1-P. "ATWS Model Modifications to CESEC", C-E Proprietary Report (November, 1975).

4-6 CENPD-107, Supplement 4-P, "ATWS Model Modification to CESEC",

C-E Proprietary Report (December,1975).

4-7 CENPD-107, Supplement 5-P, "ATWS Model Modifications to CESEC, C-E proprietary Report (June,1976).

4-8 CENPD-107, Supplement 6, "CESEC - Digital Simulation of a Combustion Engineering Nuclear Steam Supply Systen", C-E Non-Proprietary Report (August,1973).

4-9 CESEC, Digital Simulation of a Combustion Engineering Nuclear Steam Supply System, December,1981, transmitted as Enclosure 1-P to LD-82-001, January 6,1982.

4-10 CEN-234(C)-P, Louisiana Power and Light Company, Waterford Unit 3, Docket 50-382, Response to Questions on CESEC, Decunber,1982.

4-11 CENDD-161-P, TORC Code, A Computer Code for Detennining the Thennal Margin of a Reactor Core," July,1975.

4-12 CEN-191(B)-P, "CETOP-D Code Structure and Modeling Methods for Calvert Cliffs 1 and 2," December,1981.

4-13 CENPD-162-P-A, "CE Critical Heat Flux, Critical Heat Flux i Correlation for CE Fuel Assemblies with Standard Spacer Grids Part 1 Unifonn Axial Power Distributions," Septenber,1976.

4-14 CENPD-207-P, "CE Critical Heat Flux, Critical Heat Flux Correlation for CE Fuel Assemblies with' Standard Spacer Grids Part 2 Nonunifonn Axial Power Distributions," June,1978.

4-15 Letter fron E. G. Tourigny (NRC) to W. C. Jones (0 PPD) dated March 15, 1983.

4-16 OPPDNA-8301, " Reload Core Analysis Overview", Septonber,1983.

, 117

7.0 REFERENCES

(Continued)

Section 4 References 4-17 CEN-124(0)-P, " Statistical Combination of Uncertainties Methodology Analyses for Fort Calhoun Station Unit 1",

Octobe r, 1983.

Section 5 References 5 -1 CEN-121(B)-P, "CEAW, Method of Analyzing Sequential Control Element Assembly Group Withdrawal Event for Analog Protected Systems", November,1979.

5-2 CENPD-199-P, Revision 1-P, "CE Setpoint Methodology", April, 1982.

5-3 Letter from W. C. Jones to R. A. Clark, LIC-83-157, dated June 30, 1983.

5-4 Fort Calhoun SER on Automatic Initiation of Auxiliary Feedwater, contained in the letter to W. C. Jone from Robert A. Clark, dated February 20, 1981.

5-5 XN-NF-78-44, "A Generic Analysis of the Control Rod Ejection Transient for Pressurized Water Reactors", January,1979.

5-6 OPPD-NA-8302, " Nuclear Design Methods and Verifications",

Septenber,1983.

Section 6 References 6-1 "CESEC - Digital Simulation of a CE NSSS," Enclosure 1-P to LO 001, January 6,1982.

t 118

_ _ _ . _ _ .