ML20154H655
| ML20154H655 | |
| Person / Time | |
|---|---|
| Site: | Fort Calhoun  |
| Issue date: | 04/30/1988 |
| From: | OMAHA PUBLIC POWER DISTRICT |
| To: | |
| Shared Package | |
| ML20153E168 | List: |
| References | |
| OPPD-NA-8303-NP-R02, OPPD-NA-8303-NP-R2, NUDOCS 8809220087 | |
| Download: ML20154H655 (137) | |
Text
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Omah.s Public Power District Nuclear Analysis Reloat. Core Analysis Methodology Transient anA Accident Methods and Verification OPPD NA 8303 NP Rev. 02 April 1988 Copy No.
8009220007 000030 PDR ADOCK 05000205 p PDR
ABSTRACT This document is a Topical Report describing Omaha Public Power District's re-load core transient and accident methods for application to Fort Calhoun Station Unit No. 1.
The report addresses the District's transient and accident analysis methodology and its application to the analysis of reload cores. In addition, comparisons of results using e.he NSSS simulation code to results from experimental measure-ments and independent calculations are provided.
OPPD.NA 8303 NP. Rev. 02 L
e Tpble of Contents 4
1.0 INTRODUCTION
AND
SUMMARY
l 2.0 CHAPTER 14 EVENTS CONSIDERED IN THE RELOAD CORE ANALYSES 1 2.1 Criteria 2 2.2 USAR, Chapter 14, Safety Analysis Events Not Considered in Reload Core Analyses 3 2.2.1 Malpositioning of Group N CEAs l (formerly Part Length CEAs) 3 2.2.2 Idle Loop Startup Event 4 2.2.3 Turbine Generator overspeed Event 4 2.2.4 Loss of Load Event 4 2.2.5 Malfunctions of the Feedwater System 5 2.2.6 Steam Generator Tube Rupture Accident ,
6 2.2.7 Loss of Coolant Accident 7 2.2.8 Containment Pressure Analysis 7 2.2.9 Cenaration of Hydroben in Containment 8 2.2.10 Fuel Handling Accident 8 2.2.11 Cas Decay Tank Rupture 8 2.2.12 Waste Liquid Event 8 2.3 USAR, Section 14 Events Considered in a Reload Core Analysis 8 3.0 TRANSIENT AND ACCIDENT ANALYSIS AND TECHNICAL SPECIFICATIONS 9 4.0 TRANSIENT AND ACCIDENT ANALYSIS MODELS 10 4.1 Plant Simulation Model 10 !
4.2- DNBR Analysis Models 12 I, 4.3 Application of Uncertainties 13 !
4' 5.0 TRANSIENT AND ACCIDENT METHODS 14 5.1 CEA Withdrawal Event 16 5.2 Boron Dilution Event 23 l 5.3 Control Element Assembly Drop Event 28 i 5.4 Four Pump Loss of Flow Event 32 l
5.5 Asymmetric Steam Generator Event 36 ;
5.6 Excess Load Event -
40 l 5.7 RCS Depressurization Event 47 5.8 Main Steam Line Break Accident 50 l OPPD NA 8303 NP. Rev. 02 11
Table of Contents (Cantinued)
Section f_aEft 5.0 TRANSIENT AND ACCIDENT METHODS (Continued) 5.9 Seized Rotor Accident 60 5.10 CEA Ej ection Accident 64 5.11 Loss of Coolant Accident 67 5.12 Loss of Load to Both Steam Generators Event 68 5.13 Loss of Feedwater Flow Event 73 6.0 TRANSIENT ANALYSIS CODE VERIFICATION 79 6.1 Introduction 79 6.2 Comparison to Plant Data 80 6.2.1 Turbine Reactor Trip 80 6.2.2 Four Pump Loss of Coolant Flow 83 6.3 Comparison Between OPPD Analyses and Independent Analyses Previously Performed by the Fuel Vendors 85 6.3.1 Dropped CEA 86 6.3.2 Hot Zero Power Main Steamline Break 87 6.3.3 Hot Full Power Main Steamline Break 88 6.3.4 RCS Depressurization 90 6.4 Sum nary 92
7.0 REFERENCES
97
(
OPPD.NA 8303 NP. Rev. 02 111
LIST OF TABLES TABLE TITLE EAff 5.0-1 Reactor Protective System and Safety Injection. 15 5.1-1 Initial Conditions Assumed in CEAU Event Analysis 20 5.3.4 1 Key Parameters Assumed in the Full Length CEA Drop 31 Analysis 5.4.4 1 Key Parameters Assumed in the Loss of Coolant Flow 34 Analysis 5.5.4 1 Key Parameters Assumed in the LL/lSG Event 39 5.6.4 1 Key Parameters Assumed in the Excess Load Event Analysis 45 5.7.4 1 Key Parameters Assumed in the RCS Depre.ssurization Event 49 Analysis 5.9.4 1 Key Parameters Assumed in the Seized Rotor Analysis 62 5.12.4 1 Key Parameters Assumed in the Loss of Load to Both Steam 71 Generators Analysis 5.13.4 1 Key Parameters Assumed in the Loss of Feedwater Flow 77 Analysis 6.3 1 Comparison of Parameters Including Uncertainties Used in 93
- te CEA Drop Analysis for Cycles 6 and 8 6.3 2 Comparison of Parameters Including Uncertainties Used in 94 the HZP Main Steamline Break Analysis for Cycles 6 and 8 6.3 3 Comparison of Parameters Including Uncertainties Used in 95 the HFP Main Steamline Break Analysis for Cycles 6 and 8 6.3 4 Comparison of Parameters Including Uncertainties Used in 96 the RCS Depressurization Analysis for Cycles 6 and 8 OPPD NA 8303 NP, Rev. 02 iv
LIST OF FIGURES FIGURE TITLE EAGE 1-1 Full Power Turbine Trip Nuclear Power vs Time 99 1-2 Full Power Turbine Trip Pressuriser Pressure vs Time 100 13 Full Power Turbine Trip Pressurizer Level vs Time 101 14 Full Power Turbine Trip RCS Temperatures vs Time 102 15 Full Power Turbine Trip Steam Cenerator Pressure vs Time 103 16 Full Power Turbine Trip Main Feedsater Flow vs Time 104 17 Full Power Turbine Trip Steam Flov vs Time 105
, 21 4 Pump Loss of Flow Total RCS Flow vs Time 106 22 4 Pump Loss of Flow Pressuri:er Pressure vs Time 107 23 4- Pump Loss of Flow Pressurizer Level vs Time 108 24 4- Pump Loss of Flow Core Power vs Time 109 25 4 Pump Loss of Flow Steam Generator Pressure vs Time 110 26 4 Pump Loss of Flow RCS Temperatures vs Time 111 27 4. Pump Loss of Flow Main Feedvater Flow vs Time 112 28 4 Pump Loss of Flow Steam Flow vs Time 113 31 CEA Drop Incident Core Power vs Time 114 32 CEA Drop Incident Core Average Heat Flux vs Time 115 33 CEA Drop Incident Coolant Temperature vs Time 116 34 CEA Drop Incident Pressurizer Pressure vs Time 117 41 Zero Power Steam Line Break Incident Core Power vs Time 118 42 Zero Power Steam Line Break Incident Core Average Heat 119 Flux vs Time 43 Zero Power Steam Line Break Incident Total Reactivity 120 vs Time
- OPPD NA 8303 NP, Rev. 02 v
r-LIST OF FIGURES (Continued)
FICL'P E TITLE fl01 44 Zero Power Steam Line Break Incident Coolant System Pres- 121 sure vs Time 45 Zero Power Steam Line Break Incident Steam Generator 122 Pressure vs Time 51 Full Power Steam line Break Incident Core Power vs Time 123 52 Full Power Steam Line Break Incident Core Average Heat 124 Flux vs Time 53 Full Power Steam Line Break Incident Total Reactivity Heat 125 Flux vs Time 54 Full Power Steam Line Break Incident Coolant System Pressure 126 vs Time 5 5A Full Power Steam Line Break Incident Reactor Coolant Temper- 127 atures vs Time: Cycle 6 5 5B Full Power Steam Line Break Incident Reactor Coolant Temper- 128 atures vs Time: Cycle 8 56 Full Power Steam Line Break Incident Steam Generator Pres- l?9 sure vs. Time 1
OPPD NA 8303 NP. Rev. 02 vi
OHAHA PUBLIC PO'JER DISTRICT RELOAD CORE ANALYSIS METHODOLOGY TRANSIENT AND ACCIDENT METHODS AND VERIFICATION REVISION DATE 00 September 1983 01 November 1986 02 April 1988 l OPPD NA 8303 NP, Rev. 02 vii n
~l l
Omaha Public Power District Reload Core Analysis Methodology Transient and Accident Methods and Verification ;
n
1.0 INTRODUCTION
AND
SUMMARY
This report discusses the methodology the Omaha Public Power. District utill:es to analyze transients and accidents for reload cores. In addi-tion, the report discusses the District's verification of the combustion Engineering System Excursion Code (CESEC) for Fort Calhoun Station transients. The purpose 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 trans.
Lent analyses which determine certain parameters specified in the Tech.
nical Specifications. The District's transient analysis models are dis-cussed 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 RELOAD CORE ANALYSLS 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 formally 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 dis-cussed in this report.
OPPD NA 8303 NP. Rev. 02 Page 1 of 129
2.0 CllAPTER 14 EVEllTS CONSIDERED IN THE RELOAD CORE ANALYSES (Continued) 2.1 Criteria The criterion used to determine the events considered in reload core analyses is that changes in various neutronics parameters ad-versely affect the safety cnalyses of these events. The core para-meters considered are the pin peaking factors, FR and Fxy, the Moderator Temperature Coefficient (MTC), the Fuel Temperature Coefficient (FTC) or Doppler Coefficient, the boron concentration, the inverse boron worth, the neutron kinetics parameters, the CEA reactivity worth and the cooldown reactivity associated with a steam line break. If these parameters change such that the previ-ously reported results for a Chapter 14 event are no longer conser-vative, then this event must be reanalyzed. If these parameters are conservative with respect to the values assumed in the refer-enced safety analyses, the criteria of 10 CFR 50.59 are set and this event is not reanalyzed. If a change in some of the para-meters may cause the results of a safety analyses to be nonconser-vative, the event is reanalyzed. If the criteria for the event are still met, then the requirements of 10 CFR 50.59 are satis-fled. The event is reported as being reanalyzed and that it has been determined that no unreviewed safety question exists for the event. In some cases it may be possible that an event is reanal-yzed and it is determined that an unreviewed safety question exists. In these cases the analyses for these events are submit-ted. In addition, any safety analyses which are performed as a re-sult of a change in the Technical Specifications are reported as part of the supporting documentation for a Facility 1.icense 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 be utili:ed during the operation of the subsequent core. In cases where either physical modifications or modifications in operating OPPD NA 8303 NP, Rev. 02 Page 2 of 129
1 2.0 CilAPTER 14 EVENTS CONSIDERED IN THE RELOAD CORE ANA1.YSES (Continued) 2.1 Criteria (Continued) procedures are made that'do impact the safety analyses, the re-suits of the revised safety analyses are reported in a reload core analysis. This methodology report does not consider the methodo1 ogy that is required to analyze all events which could be affected by this criteria, rather, if submittals are made which require analyses of events other than those discussed in this report, revi.
sions to this methodology report will be made to incorporate the methodology used for those events.
2.2 USAR. Chaeter 14. Safety Analysis Events Not considered in Reload Core Analyses This section discusses the USAR, Section 14, safety analyses which are not normally consi'.ered 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.
2.2.1 Malcositionine of Croue N CE/ _ (former1v Part Lenzth Sf2LL1 This event is not analyzed in the reload core analysis because the Group N part length CEAs are replaced with full length CEAs starting with Cycle 11. The use of Group N during power operations will still be prohibited by the Technical Specifications. The drop of a Group N CEA is considered in the full length CEA analysis.
4 OPPD.NA 8303.NP, Rev. 02 Page 3 of 129
2.0 Cl! APTER 14 EVENTS CONSIDERED IN THE RELOAD CORE ANALYSES (Continued) 2.2 USAR. Chaoter 14 Safety Analysis Events Not Considered in Reload Core Analyses (Continued) 2.2.2 Idle-Loon Startuo Event Thi9 event is not analyzed because part loop operation is not permitted by the Fort Calhoun Technical Specifications.
2.2.3 Turbine Generator Oversoeed Event This event is an analysis of the consequences of a turbine wheel failure and is unrelated to any reload core changes.
2.2.4 Loss Of Load Event A. The loss of load to both generators is assessed to determine if:
The pressuri:er safety valves limit the reactor cool-ant system pressure to a value below 110% of design pressure (2750 psia) in accordance with Section III of the ASME Boiler and Pressure Vessel Code, and suffi-cient thermal margin is maintained in the hot fuel assembly to assure that Departure from Nucleate Boil-ing (DNB) does not occur throughout the transient.
This event is not analyzed with respect to the first criteria since the relief capacity of the pressuri:er safety valves does not change and the initial energy I
contained in the reactor coolant system will not change unless power level is raised above 1500 M*J or the reactor coolant system inlet temperature is sig-nificantly increased. Section 14.9 of the USAR re-ports that the DNBR for the loss of load transient ,
never decreases below the initial value considered in OPPD NA.8303 NP, Rev 02 ,
Page 4 of 129
h 2.0 Cl! APTER 14 EVENTS CONSIDERED IN THE RELOAD CORE ANALYSES (Continued) 2.2 USAR. Chaeter 14 Safety Analysis Events Not Considered in Reload
, Core Analyses (Continued) 2.2.4 Loss of Load Event (Continued) the analysis. Therefore, it is concluded that any change in a parameter 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.
Steam generator tube plugging performed during a re.
fueling outage has the potential for altering the heat transfer characteristics assumed in Section 14.9.1 of the USAR. Section 5.12 of this document addresses the methodology to be employed should the see i generator tube plugging exceed or be expected to exceed the cur.
rent USAR analysis assumptions.
B. 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 reedvater System The analyses which are reported in USAR. Section 14.10 Mal.
functions of the Teodwater System, are the total loss of feedvater flow and the loss of feedwater heating. The re.
sults of the total loss of feedvater flow show that the min.
inum DNBR does not decrease below its initial steady state value and : hat no safety limits are approached during the event. Therefore, this event is not reanalyzed in a reload core analysis -
OPPD.NA 8303.NP. Rev. 02 Page 5 of 129
W 2.0 CHAPTER 14 EVENTS CONSIDERED IN THE RELOAD CORE ANALYSES (Continued) 2.2 USAR. Chaeter 14 Safety Annivsis Events Not Considered in Reload Core Analyses (Continued) 2.2.5 Malfunctions of the Feedvater System (Continued)
The loss of feedwater heating is the most adverse feedwater malfunction in terms of cooling on the RCS. This event, like the excess load event, is more limiting at EOC. This event has the s,me effect on the primary system as a small increase in turbine demand which is not matched by an in-crease in core power. As a result, the DNBR degradation associated with this event is less severe than that for the excess load where a large effective increase in turbine de- l mand is analyzed. The excess load event analysts is report-ed in Section 5.6 in this document. '
Steam generator tube pluggi y performed during a refuelin6 outage has the potential for degrading the heat transfer characteristics assumed in Section 14.10.1 of the USAR for 4
the Loss of Feedwater Flow Event. Section 5.13 addresses the methodology to be employed should steam generator tube plugging exceed or be expected to exceed the assumptions of
- the current Less of Feedwater Flow Event. Reduced heat transfer for the Loss of Feedwater Heating Event does not [
1 require reanalysis, since it is an oNercooling event and the !
increase in plugged tubes reduces the consequences of the event. ,
2.2.6 Steam Cenerator Tube P.ueture Accideng The steam generator tube rupture accident is analyzed to determine if the offsite dose acceptance f.riteria of 10 CFR Part 100 is met. The analysis is a radioactive material i
t OPPD NA 8303 NP, Rev. 0 Page 6 of 129 !
I 2.0 CllAPTER 14 EVENTS CONSIDERED IN THE RELOAD C03E ANALYSES (Continued) I i
2.2 USAR. Chaeter 14. Gafety Analysis Events Not Considered in Reload Cora Analyses (Continued) ,
2.2.6 Steam Generator Tube Rtinture Accident (Continued) i release analysis based upon 14 failed fuel within the core. I It is not dependent upon any reload core analysis related 'l t
parameters, therefore, it is not analyzed in the reload core !
a
. analysis. In the future, the steam generator tuba rupture t t
i 1
accident analysis may be verified for high burnup fuel and/ i or a change in heat transfer characteristics for an increase t
] in the number of plugged tubes in the generators.
! l j 2.2.7 Loss of Coolant Accident I !
l ;
i The loss of coolant accident as reported in USAR, Section [
- I4.15, is analyzed for the District by CE. The large and f j ,
c.nall break anal ses f were performed by CE. The District !
4 confirms the assumptions used in these analyses are valid [
i for each reload core. If reanalysis is required, the reanal- [
, ysis is done by a nuclear fuel vendor. The District does not perform any loss of coolant accident analyses. t 2.2.8 Containment Pressure Analysis 1
j Containment pressure analysis is dependent upon the initial ,
liquid mass and energy contained in the primary or secondary system. Since these paraceters do not change when the core is refueled, the containment pressure analysis is not done in a reload core analysis, i I
! I i
I
! l 4
6 l
4 OPPD NA 8303.Nr. Rev. 02 [
Page 7 of 129
! t 1
ii
2.0 CilAPTER 14 EVENTS CON,sIDERED IN THE RELOAD CORE ANALtSES (Continued) 2.2 USAR. Chaeter 14. Safety Analysis Events Not Considered in Reload Core Analyses (Continued) 2.2.9 Ceneration of Hydroren in Contairgglit.
The generation of hydrogen in containment analysis is inde-pendent of any reload core parameters, therefore, the anal.
ysis is not performed during the course of a reload core analysis.
2.2.10 Fuel Handline Accident The fuel handling accident is a function of the isotopic inventory contained in the fuel pins. This is not normally considered in a reload core analysis, however, it may be }
necessary to reconsider thi6 analyses for high burnup fuel.
2.2.11 Cas Deca?,jank Rueture 1
The gas decay tank rupture is independent of any parameter l associated with refuelin5 the core. Therefore, the analysis is not performed during a normal reload core analysis.
l 2.2.12 Vaste Licuid Event I
The vaste liquid event analysis is not affected by refueling '
the core. Therefore, the waste liquid event analysis is not performed in the course of a normal reload core analysis.
i I i USAS. Section 14. Events Considered in a Reload Core Ansivsis 2.3 .
The reload core analysis consists of analyzing several events which ;
are considered in the USAR and two events which previously were not 4
J i
OPPD.NA 8303.NP, Rev. 02 .
Page 8 of 129 '
l
2.0 Cl! APTER 14 EVENTS CONSIDERED IN THE RELOAD CORE ANALYSES (Continued) 2.3 l'S AR . Section 14 Events Considered in a Reload Core Analysis (Continued) '
analyzed in the USAR. These events are analyzed in accordance with the criteria discussed in this report and to determine if an unre-viewed safety question would exist for a reload core. The USAR Chapter 14 events considered in a reload core analysis are the Control Element Assembly Withdrawal (CEAW) event, the boron dilution '
event, the Control Element Assembly (CEA) drop event, the loss of coolant flow event, the excess load e mne, the steam line break accident, the CEA ejection accident a W. the seized rotor accident.
In addition, analyses are performed for incidents resulting from the malfunction of one steam generator and for the RCS depressuri:ation event. The analysis for each of these events will be discussed in detail in Section 5.0 of this report.
310 TRANSI'.NT AND ACCIDENT ANALYSIS AND TECHNICAL SPECIFICATIONS Results of transient and accident analyses are used in the Technical Spec-ifications in two ways. The first way is that values from the Technical Specifications are included in the initial conoitions of the transient analyses. These Technical Specifications Suarantee that tha various trans- l 1ent and accident analysis acceptance criteria vill not be exceeded if the l reactor is operated within .he bounds of these Technical Specifications.
Technical Specifications of this type include the limits on TR , F xy, l the PDIL and the Moderator Temperature Coefficient. i The second type of values factored into the Technical Specifications are those that are determined by transient analysis. These parameters consist of the transient response term applied to the TM/LP equation, the minimum required shutdovr. margin, the linear heat rate LCO and the DNBR LCO. The l trsnsient response term applied to the TM/LP equation in the Technical Spec-ifications is a result of the analysis of the RCS depressuri:ation event or I i
OPPD NA 8303 NP. Rev. 02 Page 9 of 129
3,0 TPANSIENT AND ACCIDENT ANALYSIS AND TECHNICAL SPECIFICATIONS (Continued) excess load event. The minimum required shutdown margin at hot shutdown conditions is determined by the steam line break accident. 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 boron dilution event or the five percent suberiticality requirement for refueling. The values used in the linear heat rate LCO are typically deter.
P mined by the loss of coolant accident. These values are also confirmed for the dropped CEA event. The LCO on DNBR margin is calculated based on re-sults from the dropped CEA event, the loss of four pump flow analysis or the CEA withdrawal analysis.
4.0 TRANSIENT AND ACCIDENT ANALYSIS MODELS The District utilizes the latest vursion 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 utili:es the CETOP and TORC computer codes for calculation of DNBR during these events. ,
4.1 Plane Simulation Model The District utili:es the CESEC digital computer code, References 4 1 l and 4 2, to provide the simulation of the Fort Calhoun Station nu.
clear steam supply system. The program calculates the plant response l to non LOCA initiating events for a vide range of operating condi-tions. Additional information on the model is provided in Reference j 4 3. The CESEC program, which numerically integrates one dimensional !
mass and energy conservation equations, assumes a node / flow path net-work to model the NSSS. The primary system components considered in the code include the resetor vessel, the reactor core, the primary I coolant loops, the pressuri:er, the steam generators and the reactor coolant pumps. The secondary system components include the secondary side of the steam generators, the main stess system, the feedvater system and the various steam control valves. In addition, the pro-grau models some of the control and plant protection systems.
t OPPD NA 8303 NP. Rev. 02 Page 10 of 129
4.0 TRASSIENT AND ACCIDENT ANA1.YSIS MODELS (Continued) 4.1 Plant simulation Model (Continued)
The code self initializes for any given, but constant, set of reactor power level, reactor coolant flow rate and steam generator power shar-ing. During the transient calculations, the time rate of change in the system pressure and enthalpy are obtained from solution of the conservation equations. These derivatives are then numerically inte-grated in time under the assumption of thermal equilibrium to give the system pressure and nodal enthalpies. The fluid states recog- ,
ni:ed by the code are subcooled and saturated; superheating is al. I loved in the pressurizer. Fluid in the reactor coolant system is assumed to be homogenous. Reference 4 1 provides a description of the CESEC code, including the major models, and the input, output and -
plot packages.
The. pressuri:er model is described in Reference cd and further dis-cussed in Reference 4 2. The District utilizes the wall heat trans.
fer model to permit simulation of voiding in any node in which steam formation 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 1 and further discussed in Reference
[
4 2, allows for the formation of a void in the upper head region when '
the pressurizer empties. Flow to the closure head is terminated in
- simulations of those events in which natural circulation occurs and in those ovents such as the steam line break where this action delays (
safety inj ection. l The capabilities and limitations of the CESEC code are discussed in References 4 1 and 4 2. The District's CESEC model of Fort Calhoun Station is valid as indicated in Reference 4 3 for the transients discussed in Section 5 of this report, with the exception of the CEA Ejection Analysis and LOCA Analysis. The CESEC model is also valid
. for analysis of the loss of load, malfunctions of the feedwater sys-tem and the steam generator tube rupture incidents.
t OPPD NA 8303 NP, Rev. 02 Page 11 of 129
4.0 TRM;SIENT AND ACCIDENT ANALYSIS MODELS (Continued) 4.1 Plant Simulation Model (Continued)
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 quality assurance programs related to this code.
4.2 DNBR Ansivsis Models The DNBR analysis is currently performed using either the TORC code.
Reference 4 4. ar both the TORC and CETOP codes, Reference 4 5. The TORC code is used as a benchmark for the CETOP code model. TORC solves the conservation equations, as applied to a three dimensional representation of the open lattice core, to determine the local cool-ant 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 calculation of local coolant conditions. These coolant conditions are then used with a critical Heat Flux (CHT) cor. ,
relation supplied as a code subroutine to determine the minimum value of DNBR for the reactor core. The CE 1 CHF correlation (References 4 6 and 4 7) is uaed for the Fort Calhoun reactor as approved in Reference 4 8. The Detailed TORC code is used directly in the sei:ed i rotor analysis. I The CETOP code has been developed to reduce the computer time needed for thermal hydraulic analyses while retaining all of the capabili-ties of the TORC design model. The CETOP model provides an addi-tional simplification to the conservation equations due to the spec-ific geometry of the model. A complete description of the CETOP code ;
is contained in Reference 4 5 and a description of the District's application of the CETOP code is contained in Reference 4 9.
OPPD NA 8303 NP. Rev. 02 i Page 12 of 129 [
4.0 TRANSIENT AND ACCIDENT ANALYSIS MODELS (Continued) 4.2 DNBR Analysis Models (Continued)
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 analyses discussed in Section 5 with the exception of the seized rotor analy-sis.
4.3 Aeoliention of Uncerealities Uncertainties are taken into account either by deterministic or stat-istical methods. The deterministic method applies all uncertainties adversely and simultaneously when calculating the approach to a limit.
Uncertainties in DNBR calculations are taken into account by statis-tical methods. The statistical method takes into account the like.
11 hood that the uncertainties will all be adverse. The statistical method is discussed in Reference 4 10. In this method the impact of component uncertainties on DNBR is assessed and the DNBR SAFDL is in-creased to include the effects of the uncertainties. Since the uncer.
tainties are accommodated by the increased DNBR SAFDL in the statis-tical method, engineering factors are not applied to the DNBR an41y-wis model. The statistical method of applying uncertainties is ap- !
plied to the CEA withdrawal. CEA drop, loss of RCS flow, excess load, sei:ed rotor and asymmetric steam generator event DNBR calculations.
4 OPPD NA 8303 NP, Rev. 02 Page 13 of 129
I 5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS This section addresses the evaluation of the various transients and acci.
dents that are performed during a reload core analysis. Specific methods are described for each transient and accident. For each accident or trans.
ient the following material is described:
A. Definition of the Event . A brief description of the causes, conse.
quences, and RPS trips involved in the incident.
B. Analysis Criteria /. brief description of the classification' of the event and the Specifisd 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.
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 results of the analysis and a discussion of the methods used to determine if the event meets the criteria of 10 CFR 50.59.
C. Conservatism of Results . A description of the conservatism of the analysis.
The values of the trip setpoints and trip delay times used in these anal.
yses are shown in Table 5.0 1.
OPPD.NA.8303 NP. Rev. 02 Page 14 of 129
1 Table 5.0-1 REACIIR HUIECTIVE SYSIIM 'IRIPS AND SMTIY TK7BLTION Used in Analysis Trip Setnoint Unoettain_ty D!alav Time (sec) E9.t_roint Iligh Rate-of-Onnrp of Power 2.6 dec/ min 10.5 dec/ min O; 2.1 dec/ min
'- liigh Power IcVel 107% 5.0% 0.4 112%
, Variable High Ibwer Imvel 9.1% above set 0.9% 0.4 10% above i power le/el to initial power
] a low of 19.1% level
, Im Reactor Coolant Flow 95% f2% 0.65 93%
1 liigh Pressurizer Pressure 2400 psia f22 psi 0.9 2422 psia
- Ihermal Margirg/Im PressureIII 1750 psia f22 psi 0.9 1728 psia Im Steam Generator Pressure 500 psia 122 poi 0.9 478 psia low Steam Generator Water Ievel 3L2% of narrow 110 in. (5.7% of narrut 0.9 25.5% of span rarty span range span)
Steam Generator Differential
- Pressure 135 psid 140 psi 0.9 175 psid I
Caltairment Pressure Hit #1 5 psig 10.4 psi 0.1 5.4 psig fligh Pressure Safety Injection 1600 psia 122 psi 12(2) 1578 psia I
l II) Values represent the low limit of the thermal margirVlow pressure trip. 'the setpoint of this trip is diewM 1 in Reference 5-3.
l (2) Rap start - loop ulve opening time.
I OPH)-NA-8303-P, Rev. 02
5.0 TPANSIENT ASO ACCIDENT ANALYSIS METHODS (Continued) 5.1 CEA Vithdrawal Event 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 cont st system or by operator error. The CEA Block Syseta eliminates the possibility of an out of sequence bank withdrawal or single CEA withdrawal due to a r single failure.
Any controlled or unplanned withdrawals of the CEAs results in a positive reactivity addition which causes the core power, core average heat flux and reactor coolant system temperature and pressure to rise and in turn decrease the DNB and Linear Heat Rate (LHR) margins, The pressure in-crease, if larga enough, activates the pressurizer sprays which mitigate the pressure rise. In the presence of a positive Moderator Temperature Coefficient (MTC) of reactiv-
- ty, the temperature increase results in an additional pos-itive reactivity addition further decreasing the margin to the DNB and LJiR limits.
Withdrawal of the CEAs causes the axial power distribution 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 mag-nitude of the 3 D peak change depends primarily on the ini-tial CEA configuration and sxial power distribution.
The withdrawal of the CEAs causes the neutron flux as mea-sured by the excore detectors to be decalibrated due to CEA motion, i.e., rod shadowing effects. This decalibration of OPPD NA 8303 NP, Rev. 02 Page 16 of 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.1 G A_V.1t.hdrawal Event (Continued) ,
5.1.1 Definition of the Event (Continued)
[
r excore detectors, however, is partially compensated by ;
[
neutron attenuation rising from moderator density changes l (i.e., temperature shadowing effects). (
l As the core power and heat flux increase, a reactor trip on high power, variable high power, or Thermal Margin / Low Pres- f sure may occur to terminate the event depending on the ini-tial operating conditions and rate of reactivity addition, j Other potential trips include the axial power distribution f and high pressurizer pressure trips. If a trip occurs, the CEAs drop into the core and insert negative reactivity which i
- quickly terminates futtbar margin degradation. If no trip l occurs and corrective action is not taken by the operators, I t
I the CEAs fully withdraw and the NSSS achieves a new steady j
- neate equilibrium with higher power, tempareture, peak l 4
linear heat rate and lower hot channel DNBR value. ,
O 3
5.1.2 Analysis Criteria ;
I l
The CEA Vithdrawal (CEAV) event is classified as an Antici- i paced Operational Occurrence (AOO) for which the following i 1
criteria must be n.ot: [
i
[
A. The transient minimum DNBR is greater than the 95/95 f confidence interval limit for the CE 1 correlation, f t
and [
i 1
B. the Peak Linear Heat Cen'eration Rate (PMCR) does not i e
, exceed 22 kw/ft (Reference 5-1). ',
i I i
OPPD NA 8303 NP, Rev. 02 l
Page 17 of 129 -
t
5.0 TRANSIE:,'T AND ACCIDENT ANALYSIS METHODS (Continued) 5.1 CEA Withdrawal Event (Continued) 5.1.3 Obiectives of the Ansivsis The objectives of the analysis performed for the "at power" CEAV event is to calculate the Required Overpower Margin (ROPM) which must be factored into the setpoint analysis.
The objective of the analysis for the hot tero power CEAV 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 Kev Paraceters and Analysis Assumotioni The initial conditions assumed in the CEAV analysis are shown in Table 5.1 1. The reactor state parameters of primary importance in calculating the margin degradation are:
A. CEA withdrawal rate * (i.e., reactivity insertion rate). ,
- 8. Cap thermal conductivity (H5ap)*
C. Initial power level, D. Flux power level determined from the excore detector response during the transient, ;
I E. The moderator temperature coefficient reactivity. and i
- NOTE: The term CEA withdrawal rate and CEA reactivity insertion rate are used I interchangeably in this report. l
[
OPPD NA 8303 NP. Rev. 02 {
Page 18 of 129
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5.0 TRANSIENT AND ACCIDENT ANALYSIS METliODS (Continued) l 5.1 CEA Withdrawal (Zing (Continued) i 5.1.4 Kev Parameters and Analysis Assuretions (Continued) l f
F. Changes in the axial power distribution and planar and [
, integrated radial peaking factor during the transient.
The excore responses for each initial power level analyzed are based on the CEA insertions allowed by the Power Depend- l ent Insertion Limit (PDIL) at the selected power level, the f changes in CEA position prior to trip, and the corresponding f rod shadowing and temperature attenuation (shadowing) fac-tors.
i' For the CEAV cases where combinations of parameters result f in a reactor trip, the scram reactivity versus insertion 1
characteristics are assumed to be those associated with the i I
core average axial power distribution peaked at,the bottom !
l of the core. The scram reactivity versus insertion charac. j
, teristics associated with this bottom peak shape minimize [
- the amount of negative reactivity inserted during initisi j portions of the seras following a reactor trip.
All control systems except the pressurizer pressure control system and the pressurizer level control system are assumed to be in a manual mode. These are the most adverse operat.
l ing modes for this event. The pressurizer pressure control system and pressurizer level control system are assumed to j be in the automatic mode since the actuation of these sys- [
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Table 5.1 1 Initial Conditions Assumed in CEAW Event Analysis Pareeter Units 2. alm Initial Core Power MWt 1(HZP)/1530(HFP)t i Initial Core Inlet coolant 'T 532(HZP)t Temperature Maximum allowed by Tech. Specs.
Moderator Temperature Coefficient x10*'ap/'F Tech. Spec. Range Initial RCS Ptessure psia Minimum allowed byt Tech. Specs.
Fuel Temperature Coefficient x10ap/'F Least Negative Predicted During A Cycle Initial Core Mass Velocity x106 lbm/hr Minimum allowed byt Tech. Specs.
Fuel Temp. Coeff. Uncertainty n 15.0 CEA Vithdrawal Speed in/ min 46.0 Radial Peaks Maximum Allowed by Tech. Specs. for a Given Initial Power Level Scram Reactivity 4 Minimum Predicted During a Cycle High Power Trip Analysis Setpoint t of 1500 MVt 112.0 Variable High Power Trip Analysis 4 Above Initial 10.0 Setpoint Power Level t For DNBR calculations, effects of uncertainties are combined statistically.
OPPD NA 8303 NP. Rev. 02 Page 20 of 129 i
5.0 TPANSIENT AND ACCIDE!Tt ANA1.YSIS METHODS (Continued) 5.1 CEA Vithdrawal Event (Continued) f 5.1.4 Kev Parameters and Analysis Assumotions (Continued) tems minimizes a rise in the coolant system pressure. The not effect. is to delay a reactor trip until a high power trip is initiated. This allows the transient increases in power, heat flux and coolant temperature to proceed for a longer period of time. In addition, minimizing the pressure increase is conservative in the margin degradation calcula-tions 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 Methodology The methodology used for analysis of the CEAV event is de.
Scribed in CEN.121(8) P, Reference 5 2. The District does not perform all parametric analyses discussed in Reference 5 2 for Fort Calhoun Station. Rather, the District utill:es the analyses performed in Reference 5 2 to limit the number of analyses necessary for Fort Calhoun Station.
The rod shadowins 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 2 (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.) The analysis at intermediate power levels is the same as documented in Reference 5 2.
1 OPPD NA 8303 NP. Rev. 02 Page 21 of 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.1 CEA Withdrawal Event (Continued) l 5.1.5 Analysis,Methodolory (Continued)
The hot zero power CEAV event is analyzed assuming the vari-able high power trip is initiated at 29.1% (19.1% plus 10%
uncertainty) of rated thermal power. I.1 addition, the anal-ysis assumes that the maximum CEA withdrawal rate is com.
bined with the maximum differential rod worth. This case is analyzed using CESEC and the minimum DNBR is calculated us-ing CETOP using the assumptions discussed in Reference 5 2. l The CEAW event analyzed to determine the closest approach to the fuel centerline melt SAFDL assumes those values of the CEAV rate and H g ,p discussed in Reference 5 2. This com- l bination of CEAV rate tnd Hg,p was used to determine the PUICR at all power levels.
5.1.6 Tveical Analysis Results and 10 CFR 50.59 Criteria The results of the analyses of the CEAV event for Fort Cal-houn Station at full power and at inter:sediate power levals are expected to be similar to those presented in Reference 5 2. The results of the hot :ero power CEA withdrawal anal-ysis 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 '.f the analysis for the full power and intermediate power level CEAV events shows that the required overpower inargin for these events is less than the available overpower margin required by the current Technical Specification DNS and PUiCR LCOs. The 10 CFR 50.59 criteria is satisfied for the hot :ero power CFAV event if the minimum DNSR is greater than that reported in the latest submitted analysis.
OFFD NA 8303 NP. Rev. 02 Page 22 of 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.1 CEA Vithdrawal Event (Continued) 5.1.7 Conservatism of Resul_ts Conservatism of the results of the CEAU incident analyses is discussed in Reference 5 2 for the full power, intermediate power level and hot zero power cases.
5.2 Boron Dilutten Event 5.2.1 Definition of Everie Boron dilution is a manual operation, conducted under strict procedural controls which specify permissible limits on the rate and magnitude of any~ required change in boron concentration. Boron concentration in the reactor coolant system can be decreased by either controlled addition of unbor,ated makeup water with a corresponding removal of reactor coolant or by using the deborating ion 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 domineralized water batch quantity !
selector set for the desired quantity. When the specific !
amount has been injected, the domineralized water control l
.alve is shut automatically. An inadvertent boron dilution can occur only if there is a combination of operator error I and a CVCS malfunction occurring at the same time. No RP3 trips are assumed to terminate this incident.
5.2.2 Analysis Criteria l
t The boron dilution event is classified as an A00 for which the following criteria cannot be exceeded:
OPPD NA-8303 NP. Rev. 02 Page 23 of 129
5.0 TPASSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.2 Boren Dilutten Event (Continued) 5.2.2 Analysis Criteria A. DNER greater than the 95/95 confidence interval limit using the CE 1 correlation, and B. The PLHCR less than 22 kv/ft.
5.2.3 Obicetives of the Ansivsis The DNER and PLHCR criteria are met by showing that suffi-cient time exists for the operator to take corrective action to terminate the event prior to exceeding the SAFDLs. This is accomplished by calculating the time interval in which the minimum Technical Specification shutdown margin is lost.
The acceptable time interval for the operator to take correc-tive actions before shutdown margin is lost are 15 minutes for Modes 2. 3 and 4 and 30 minutes in Mode 5.
5.2.4 Kev Parameters and Analysis Assuretions 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 dilution event in Modes 2 through 5. it is assumed that all three charging pumps are 1
I operating at their maximum capacity for a total charging rate of 120 gpa. For the dilution at hot standby (Mode 2) the event is assumed to be initiated at the Technical Spec-ification hot shutdown margin requirement at 532'T. The reactor coolant system is 5.506 cubic feet.
OPPD NA 8303 NP. Rev. 02 Page 24 of 129 l
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5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.2 Poron Dilution Event (Continued) 5.2.4 Mev Parameters and Analysis Assumotions (Continued)
The boron dilution event at hot shutdown (Mode 3) is assumed to be initiated from the Technical Specification shutdown margin requirement at 210'T. The boron dilution event at cold shutdown (Mode 4) is initiated from the Technical Spec-ification minimum shutdown margin requirement at 68'T. The analysis is conducted for two RCS volumes, one of 5,506 cubic feet and the other of 2,036 cubic feet, which corre.
sponds 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 from 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 assumptions f
, are consistent with the Technical Specifications for cold shutdown conditions. The boron dilution event during refuel-ing is analyzed assuming that reactor refueling has just j been completed and the head is in ple:e but the coolant vol-use is sufficient to only fi).1 the reactor vessel to the bot-tom of the piping nozales (2,036 cubic feet) and the minimun permissible boron concentration allowed by Technical Specifi- !
cation for refueling exists. All CEAs are withdrawn from the core.
These assumptions represent shutdown conditions for the var-ious modes wherein the core reactivity is greatest. the water voluse and total boron content is at a minimum, and ,
the rate of dilution is as large as possible. Hence, these conditions represent the minimum time to achieve inadvertent l
criticality in the event of an uncontrolled boron dilution. I i
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OPPD NA 8303 NP. Rev. 02 Page 25 of 120 I
5.0 TRANSIENT AND ACCIDENT ANA1.YSIS METHODS (Continued) 5.2 Bnron Dilutten Event (Continued) 5.2.5 Analysis Methods The method used to calculate the dilution time to critical.
l icy from Modes 2 through 5 is throten the use of the follow-ing equation: l l
l Atterit " 'BD In C3 + SDM*IEW C
5 Where r5D - boron dilution time constant, which is a function of RCS volume and temperature (sec)
C 3 - critical boron concentration (ppm)
SDM - shutdown margin (top)
IBW - inverse boron vorth (pps/ap) l As can be seen from this equation, the dilution time to criticality is minimized with a greater critical boron concentration, a sa:11er inverse boron worth, or a smaller
'SD' 5.2.6 Analysis Results and 10 CFR 50.So criteria The analysis results are similar to those reported in the Cycle 11 safety analysis report and in the 1987 update of l the USAR. The criteria of 10 CTR 50.59 are satisfied if the Technical Specification requirements on shutdovn margin and l l the refueling boron concentration are unchanged as a result of this analysis
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Fage 26 of 129 j i
5.0 TPR;SIE!;T AND ACCIDE!iT A!!ALYSIS METHODS (Continued) 5,2 Poron Dilution Event (Continued) 5.2.7 Conservatism of Results Because of the procedures involved in the boron dilution and the numerous alarm indications available to the operator, the probability of a sustained or erroneous boron dilution is very lov. 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 considerable 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 normally achieved with the shutdown groups A and 5 fully inserted in the core and, therefore, the core has a
.. much lower k,gg than assumed in the analysis. The boron dilution at refueling is conservative since it is improbable that more than a few CEAs 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 CEAs removed at a time, The analysis assumes that all CEA: are withdrawn from the core, OFFD NA 8303 NP, Rev, 02 Page 27 of 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) '
i 5.3 control Element Afsembiv Droo Event l 5.3.1 Definition of Event The control element asserably (CEA) drop event is defined as l' the inadvertent release of a CEA causing it to drop into the i reactor core. The CEA drive is of the rack and pinion type with the drive shaf t running parallel 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 mechanical brake in the CEA drive mechanism when the CEA l is being moved.
The full length CEA drop event is classified as an ADO which '
does not require an RPS trip to provide protection against exceeding the !AFDLs. The CEA drop results in a redistri-bution of the core radial power distribution and an increase in the radial peaks which are not directly monitored by the l RPS and which are not among those analysed in determining the DNB and MR LCOs and LSS$s. 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 WR SAFDLs. l i
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OPPD NA 8303 NP, Rev. 02 i Page 2S of 129 l l
5.0 TPW;SIENT AND ACCIDENT ANALYSIS HETHODS (Continued) 5.3 Cnntrol Element Assembly Dron Event (Continued) l 5.3.2 Analysis Criteria The full length CEA drop event is classified as an Antici-l pated Operational Occurrence for which the following cri-teria must be met:
1 A. The transient minimum DNBR must be greater than or l equal to the 95/95 confidence interval limit, using i the CE 1 correlation, and B. The Peak Linear Heat Race (PLHR) must be less than or equal to 22 kv/ft.
5.3.3 Obiectives of the Analysis
't The objective of the analysis is to determine the Required Overpower Margin (ROPM) which must be built into the LCOs to assure the DNBR and LRR SAFDLs are not exceeded for the CEA i drop which produces the highest distortion in the hot chan. I nel power distribution. Since the R0PM is dependent upon initial power level, rod configuration and axial shape !
index, an analysis parametric in these variables is per- j formed.
5.3.4 Kev Parameters and Ansivsis Assumotions (
\
I Table 5.3.41 contains a list of the key parameters assumed l in the full length CEA drop analysis, Assumptions used in the analysis include:
l OPPD NA 8303 NP. Rev. 02 i Page 29 of 129 i
5.0 TRANSIU;T A';D ACCIDENT ANALYSIS METHODS (Continued) 5.3 control Element Assemb1v Drop Event (Continued) l 5.3.4 Kev Parameters and Analysis Assuretions (Continued)
A. The rod block system is assumed to prevent any other rod motion during the transient.
B. The turbine ad. mission valves are maintained at a c.o-stant position during the transient. This is because the turbine admission valve position is set manually at Fort Calhoun Station and, therefore, the turbine admission valves will not automatically open in re-sponse to a reduced electrical generation output.
5.3.5 Analysis Method The analysis methods utilized by the District to analy:e the CEA drop event are discussed in Section 8 of Reference 5 3.
5.3.6 Analysis Results snd 10 CFR 50. 5o Criteria Typical analysis results are contained in Section 8 of Reference $.3 and in the 1987 update of the Fort Calhoun l Station Unit No. 1 USAR. The criteria of 10 CTR 50.59 are met if the required overpower margin calculated for this incident is less than the overpower margin being maintained by the current Technical Specifications.
OPFD NA 8303 NP, Rev. 02 Page 30 of 129
l Table 5.3.4 1 KEY PARAMETERS ASSUMED IN T11E WLL LE! 3' O.'
- ANALYC!S 1
l Parameter l' nit s Value Initial Core Power MWe 1500t l Initial Core Inlet 'T Maximum allowedt Temperature -
by Tech. Specs.
l Initial RCS Pressure psia Minimum allowedt by Tech. Specs.
Initial Core Mass Flow Rate x1061bm/hr Minimum allowedt l l by Tech. Specs.
l Moderator Temperature x10as/'F Most negative (
Coefficient allowed by Tech.
Specs.
CEA Insertion n Insertion Maximum allowed by Tech. Specs.
Radial Peaking Distortion '
Maximum value predicted Factor during core life l
l For DNER calculations, the effects of uncertainties on these parameters are combined statistically.
l l
l l
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OPPD NA 9303 NP, Rev 02 Page 31 of 129
P 5.0 TRANS!c:T AND ACCIDC;T ANA1.YSIS METHODS (Continued) 5.3 control Elecent Assenb1v Dron Event (Continued) {
P 5.3.7 Conservstism of Results The following area of conservatism is included in the anal-ysis:
A. The moderator temperature coefficient assumed in the analysis is the most negative value allowed by the ,
Technical Specifications. The actual end of Ilfe value, including measurement uncertainty, is less negative.
5.4 Four Pu?S Loss of Flev 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 failure c! 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 pressure drop.
, This signal is compared to a setpoint which is a function of the number of reactor coolant pumps in operation (which current Technical Specifications require to be four). A reactor trip would be initiated when the flow rate drops to 93% of full flow (956 minus 24 uncertainty).
0FFD NA 8303 NP. Rev. 02 lage 32 of 129
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l 5.0 TRANSIENT AND ACCIDENT ANALYSIS METH0vS (Continued) i 5.4 Four P's-o less of Flov Event I, Continue d) 5.4.2 Analysis Criteria The four pump loss of flow event is classified as an AOO for which the transient minimum DNBR must be greater than the 95/ 95 percent confidence interval limit using the CE 1 cor-relation.
5.4.3 Obteetives of the Analtill The objective of the analysis is to determine the required overpower margin that must be built into the DNB LCOs such that in conjunction with the lov flow trip the DNBR SAFDL is not' exceeded. Since the required overpower mar 81 n is de.
pendent upon both axial shape index and the CEA rod config-uration, an analysis parametric in these parameters is per-fo rme d.
5.4.4 Fev Para-eters and Analysis Assuretiona The closest approach to the DNBR SAFDL occurs for a loss of flov 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 eethod used by the District to analy:e the four-pump loss of coolant flow is discussed in Section 7 of Refer-ense 5 3. The District utili:es the CESEC TORC nethod to analy:e axial power distributions characteri:ed by both nega-tive and positive shape indices. The STRIKIN TORC method is not utilized by the Did?.rict because of the high rotational 2
energy of the pumps (N - 1185 rpm, I - 71,000 lb f t /puap) .
OPPD NA 8303 NP Rev. 02 Page 33 of 129
Table 5.4.4 1 P.EY PARAMETERS ASSUMED IN THE LOSS OF C00LRIT TLOV ANA1.YSIS Parn-eter h h Initial Core Power MVt 1500t l Moderator Temperature x10 ap/'T Ma':imurs allowed l Coefficient by Tech. Specs Tuel Tersperature x10 ap/'T Least negative l Coefficient predicted during Lore Itfe. ,
Low Flow Trip Delay Time see Maximua l' i
CEA Drop Time see Maximum allowed l j, by Tech. Specs. ;
Scram Reactivity Vorth as Minimum predicted I during core lifetime [
d I Scra:n Reactivity Consistent with axial !
Curve shape of interest L
i i
i t Tor DNBR calculations, effects of uncertainties on these parameters were }
, combined statistically, l l
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a l
! {
1 [
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OPPD NA 8303 NP. Rev. 02 [
- Page 34 of 129 L f
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5.0 TRMISIENT AND ACCIDENT ANA1.YSIS METit0DS (Continued) 5.4 Four Pume toss of Flev Evene (Continued) 5.4.6 Analysis Results and in cm $0.50 criteria f
. Expected analysis results are presented in Section 7.1 of f
Reference 5 3. The main difference between these results l L and the results for Tort Calhoun Station is that the ROPM vill be significant1/ reduced for Tort Calhoun Station.
This is because of the higher rotational energy of the {
Fort Calhoun reactor coolant pumps. ~
I The criteria of 10 CTR 50.59 are met if the required over-f power rargin calculated for the four pump loss of coolant flow event is less than the overpower margin being - ,- !
tained by the current Technical Specifications. !
5.4.7 knservatism of Results f The conservative nature of the DNBR R0PM values calculat. L ed for the four pump loss of flow event is demonstrated [
by the following conservative assumptions: j A. Field measurements of the CEA magnetic clutch decay [ [
is more rapid than assumed in the safety analysis. l 4
) ;
- 5. The availible scram worth is higher than assumed in l i the safety analysis. !
[
t C. The MTC at full power is more negative than the l value assumed in the safety analysis. -
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OPPD NA 8303 NP. Rev. 02 Page 35 of 129
5.0 TP,A.';SIENT MiD ACCIDENT M:ALYSIS MET 110DS (Continued) 5.4 Faur Puro less of Flev Event (Continued) 5.4.7 conservatism of Pesults (Continued)
D. The actual CEA drop time to 906 inserted is faster l than that assumed in the safety analysis.
E. The e nservatism of the CETOP calculations is dis. l cussed in Section 7 of Reference 5 3.
5.5 Asvenettie steam cenerator Esent 5.5.1 Definition of the Evtne The asymmetric transients arising from a secondary system malfunction in one steam generator result in changes in ,
core power distribution which are noe inherently covered ,
by the TM/LP or AFD LS$5. Consequently, these events must be analysed to determine the initial steady state thermal margin which is built into and'asincained by the Technical Specification LCO such that assurance is pro.
vided that the DNBR and peak linear heat rate SAFDL: are l not exceeded for these transients. The four events which effect the steam generator are: ;
A. Loss of load to one steam generator. !(r B. '
Loss of feedvater to one steam generator.
i C. Excess feedvater to one steam generator. ;
l 1
D. Excess load to one steam generator.
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5.0' TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.5 Asymmetric Steam Generator Eveng (Continue O 5.5.1 Definition of the Event (Continued)
The possible RPS trips which can occur to' mitigate the j consequences of these events include the lov ateam gen 7 -
l erator level, TM/LS, low steam generator pressure, and the asymmetric steam generator transient protection trip function (ASGTPTF). The particular trip which intervenes is dependent upon the event initiator and the initial i e
operating conditions. [
The ASGTPTF trip was installed in the Fart Calhoun Sea- l tion RPS prior to operation of Cycle 9 to reduce the mar-sin requirements associated with these asymmetric events and to insure that these events do not become a limiting i
- A00 for establishing initial margin which must be main-tained by the LCO. A system description of the ASGTPTF is presented in Appendix B of Reference 5 3. !
j i 5.5.2- Analysis Criteria ,
i- !
{. The asymmetric steam generator events are classified as
} A00s for which the following criteria must be met:
l A. The transient minimum DNBR must be greater than or equal to the 95/95 confidence interval limit using l the CE.1 correlation, and 1
l B. The peak linear heat must be less than or equal to 22 kw/ft.
)
N.
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Page 37 of 129 L _ ___ _
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.5 Aizm, metric Steam Generator Everi (Continued) 5.5.3 Obiectives of the Analysis The objectives of the analysis are to determine the re-quired overpower margin that must be built into the LCOs such that in conjunction with the ASGTPTF the DNBR and PLHGR SAFDLs is not exceeded.
5.5.4 Kev Parameters and Analysis Assumotiong Section 7 of Reference 5 3 demonstrates that the loss of l load to one steam generator (LL/1SG) is the limiting asym-metric steam generator transient for establishing initial steady state thermal margin which must be maintained by the Technical Specification LCO. Therefore, information is only provided for this e. symmetric steam generator event. The key parametera used in the analysis of the LL/1SG event are given in Table 5.5.4 1. The charging pumps and proport'ional heater systems are assumed to be inoperable during the transient. This maximizes the pres-sure drop during the event. The turbine admission valves are assumed to maintain a constant' position throughout h
the event since the turbine control system at Fort Cal-houn utilizes manual setting of the turbine admission valves.
5.5.5 Analysis Method l The method utilized by the District to analyze the LL/1SC is discussed in Section 7 of Reference 5 3.
OPPD NA 8303 NP, Rev. 02 Page 38 of 129
Table 5.5.4-1 KEY PARAMETERS ASSUMED IN THE LL/1SG EVENT Pa rmra te r Units Value Initial Core Power MV e 1530t l Initial Core Inlet 'F Maximum allowedt Temperature by Tech. Specs.
Initial Reactor Coolant psia Minimum allowedt System Pressore by Tech, Specs.
Core Average H gap BTU /hr ft2*F Maximwn value predict-ed during core life.
Initial Core Mass 6 x10 1bm/hr Best estimate flowt
- Flow Rate Scram Reactivity Worth top Minimum predicted during core life, t For DNBR calculations, effects of uncertainties on these parameters were coabined statistically.
OPPD tiA 8303 tiP, Rev. 02 Page 39 of 129
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5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5 '. 5 Asymmetric Steam Generator Event (Continued) 5.5.6 Analysis Results and 10 CFR 50.59 Criteria The results of the analysis for the LL/1SG event are dis.
cussed in Section 7 of Reference 5~3.
I The results for Fort Calheun Station are expected to be similar. The criteria of 10 CFR 50.59 are satisfied if the required overpowwr margin' calculated for the LL/1SG event.is less than the overpower margin being maintained by the current Technical Specifications.
5.6 Exc?ss Lead' Event l
5.6.1 Definition of Event An exc ss load transient is defihed as any rapi increase in the steam generator steam flow other than a steam line v break.
Such a rapid increase in steam flow results in a power mismatch between the reactor core and the steam gen-erator load demand. In addition, there is a decrease in
- the reactor coolant temperature and pressure. Under these conditions the negative moderator temperature coefficient reactivity causes an increase in core power.
The rapid opening of the turbine admission valves or the steam dump bypass to the condenser causes an excess load event. Turbine valvos are not si:ed to accommodate steam I
flow for powers much in excess of 1500 MVt. The steam
- dump valves and steam bypass valves to the condenser are l ,
OPPD.NA.8303.NP, Rev. 02 Page 40 of 129
r
, 5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.6 Excess Lead Event (Continued) k 5 6.1 Definition of Event (Continued) t sized to accommodate 334 and St. respectively.. )f the steam flow at 1500 MW. Therefore, the following load increase events are examined:
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 con-trol function is used to maintain lead, so unless valve failure occurs, the control valves will remain where positioned.
B. Opening of all dump and bypass valves at power due to steam dump control interlock failure: The cir.
cuit between the steam dunp controller and the dump valves is open when the turbine generator is on line. Accidental closing of the steam dump control p interlock under full load conditions, according to the temperature program of the controller, causes full opening of the dump and bypass valves. Since
, the reactor coolant temperature decreases during the l
event, these valves will be closed again after the l average reactor coolant temperature decreases to 535'F.
t l l C. Opening of the dump and bypass valves at hot standby l t
] conditions due to, low reference temperature setting in the steam dump controller: When the plant is in hot standby conditions the dump valve controller is I ,
b OPPD NA 8303 NP, Rev. 02 (:
Page 41 of 129
CTc n 5.0 TRANSIE!!T AND ACCIDENT ANALYSIS METHODS (Continued) 5.6 Excess Load Event (Continued) 5.6.1 Definition of Event (Continued)
C. (Continued) 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 cool. j ant temperature is 532'F, which is 8'F below the min-imum temperature required to open the dump and by.
pass valves (540'F). The maximum error that can be introduced in the referenced temperature setting is limited to 17'F since a narrow range instrument is used for this purpose. Reducing the dump valve con- 7 troller reference settin5 from 532' to 515' would re-sult in a partial opening of the valves but as soon H as the reactor coolant temperature dropped to 518'T the valves would again be completely closed.
i D. Opening the dump and bypass valves at hot standby due ta steam dump controller malfunction: The most' severe event 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 case C above but represents the most limiting event under hot standby condi-tions.
The possible RPS trips that might be encountered dur-ing this event are:
, , 1. Variable high power trip (VHPT).
OPPD.NA 8303.NP, Rev 04 Page 42 of 129
,1 8
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.6 Excess load Event (Continued) .
5.6.1 Definition of Event (Continued)
D. (Continued)
- 2. TM/LP trip.
- 3. Low steam generator water level trip.
- 4. Low steam generatot pressure trip.
The RPS trip initiated to mitigate the consequences ~of the event will depend upon the initial conditions and the rate of reactivity insertion due co' moderator feed-back 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 internal limit using the CE 1 correlation.
- 2. The peak linear heat rate (PLHR) east be less than or i equal to 22 kw/fc.
5.6.3 Obiectives of the Analysis The objectives of the analysis are to calculate a term which, l when incorporated in the TM/LP equation will ensure that the SAFDLs are not exceeded for those excess load events which require a TM/LP trip for protection and to ensure that the DNBR and LHR SAFDLs are not exceeded for excess load events
(
for which the TM/LP does not provide protection.
OPPD NA 8303 NP, Rev. 02 Page 43 of 129
d, '
?. 5.0' TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued)
- n. -
5.6 Excess Load Event (Continued) 5.6.4 Kev Parameters and Analysis Assumotions
. As discussed in Section 5 of Reference 5 3, sensitivity studies performed by CE have demonstrated that the maximum calculated term for the excess load event occurs at hot full power conditions. District sensitivity studies show similar results. Therefore, only the hot full power case is anal-yzed. The key parameters used in the analysis of the excess load event are given in Table 5.6.4 1. The remaining assump-tions are the came as those discussed in Reference 5 3.
5.6.5 Analysis Method l
The steps used for determining the value and calculatin5 the largest term for all excess load events which rely on the TM/LP trip for DNBR protection are given in Section 5 of Reference 5 3. The minimum transient DNBR value for excess load events protected by the variable High Power Trip is calculated us'ng the procedure discussed in the same Section.
l The PLHR is calculated by obtaining the core average linear heat rate at time of peak core power and multiplying it by the appropriate peaking factors and associated uncertain-ties.
I i
I I
l t
)
4 l
OPPD NA 8303 NP, Rev. 02 Page 44 of 129
Table 5.6.4-1 1
KEY PARAMETERS ASSUMED IN THE EXCESS LOAD EVENT ANALYSIS
, Parameter ED111 Ealut Initial Core Power MVt 1530t Initial Core Inlet *F At Power Maximun allowedt i Temperature by Tech. Specs.
Initial Reactor Coolant psia Minimum allowedt System Pressure by Tech. Specs.
Initial Core Mass 6 x10 1bm/hr . Minimum allowedt Flow Rate by Tech. Specs.
CEA Drop Time sec Maximum allowed by Tech. Specs.
Scram Reactivity. top Minimum predicted
- Worth during core life.
4 Moderator Temperature x10*'ap/'F Negative values up to Coefficient the most negative
- ^ value allowed by Tech.
Spees.
l t For DNBR calculations, effects of uncertainties on these parameters were combined statistically.
l
\
i i
(
1 l
l l
l l
l i
OPPD NA 8303 NP. Rev. 02 Page 45 of 129
o ', l l
l 5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS. (Continued) j
' 5. 6 Excess Load Event (Continued) 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 Reference 5 3. The cri- l teria of 10 CFR 50.59 are met if the term is less than or equal to the value used in the current TM/LP trip equa.
tion.
5.6.7 Conservatism of Results The following points demonstrace the conservatism of the overall results for the excess load event:
A. Field measurements demonstrate that the CEA magne. ,
tic clutch decay time is less than that assumed in
, the analysis.
B. The actual scram worths are higher than those in the analysis.
Whore the most negative MTC is used, the value is C.
more negative than that measured during plant oper-ation.
D. The actual Doppler reactivity is more negative than assumed in the analysis.
OPPD NA.8303 NP, Rev. 02 Page 46 of 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.6 Excess Lond Evant (Continued) 5.6.7 Conservatism of Pesults (Continued)
E. Field data damonstrates that the actual CEA drop .
time is less than that assumed in the analysis.
P. The conservatism of the term is discussed in Section 5 of Reference 5-3.
5.7 RCS Deeressurization Event 5.7.1 Definition of Event 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 single primary safety valve operating at rated thermal 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 compensate for the decreasing pressure the water in the pressuricer flashes to steam and the proportional heaters increase the heat added to the water in the pressuri:er in an attempt to maintain pressure. During this time the pres-surizer level also begins to decrease causing the letdown control valves to close and additional charging pumps to start su as to maintain level. As pressure continues to drop, the backup heaters energi:e to further assist in maintaining primary pressuse. A reactor trip is initiat-ed by the TM/LP trip to prevent exceeding the DNBR SAFDL.
OPPD.NA.8303.NP, Rev. 02 Page 47 of 129 i
l 5.0 TRM;SIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.7 RCS Deeressurication Event (Continued) 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 uains the CE 1 correlation. [
5.7.3 Obioetives of the Analysis l
r i
[
This event is classified as an A00 for which there must r 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 term for incorpor- [
- ation into the TM/LP equation. i
- , 5.7.4 Kev Parameters and Analysis Assumotions 1
I t The key parameters for the RCS depressurization event i
analysis are given in Table 5.7.4 1. Additional assump-
[
l tions are discussed in Section 5 of Reference 5 3.
- l. i I. l 5.7.5 Analysis Method '
I The methods used by cho District to analyze the RCS de+
pressurization event are contained in Section 5 of Refer- t ence 5 3.
I
- i l
! I l
L
+
f OPPD NA 8303 NP, Rev. 02 !
Page 48 of 129 l
. t
Table 5.7.4 1 KEY PAIULMETERS ASSUMED IN THE RCS DEPRESSURIZATION EVENT ANALYSIS P,rneeter Units Value Initial Core Power K'J e 1530t Initial Core Inlet 'F Maximum allowed by' nerature Tech. Specs.
Initial Reactor Coolant psia Upper limit of normc System Pressure operating range Moderator Temperature x10ap/'F Most negative allow.
Coefficient by Tech. Specs.
Fuel Temperature x10ap/*F Most negative predit Coefficient durin6 core life.
Core Average H gap BTU /hr ft2 'F Minimum predicted during core life.
Total Trip Delay see 1.4 Time t For DNBR calculations, effects of uncertainties on these parameters were combined statistically.
OPPD.NA.8303.NP, Rev. 02 Page 49 of 129
4 L' ,
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.7 RCS-Deeressurication Event (Continued) l 5.7.6 Analysis Results and 10 CFR 50.59 Results of the RCS depressuri:ation transient are discussed inReference5-3andinthe1984updateoftheFortCalhounl' Station Unit No. 1 USAR. The criteria of 10 CFR 50.59 are satisfied if the term is less than or equal to the value used in the current TM/LP trip equation.
4 5.7.7 Conservatism of Results The conservatism of the calculated pressure bias term is obtained by using the combination of the following conserva-I tive key parameters:
A.
Conservative scram reactivity characteristics are usedl in the analysis. .
B. Conservatively slow RPS response times are used. l C. Conservatively high primary relief or safety valve l areas are uscd.
D. The RCS pressure is initially assumed to be in its l upper limit as opposed to the normal operating pres-sure.
5.8 dpin 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 in-creased rate of heat extraction from the primary system.
OPPD.NA.8303 NP. Rtt. 02 Page 50 of 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS METi!0DS (Continued) 5.8 M11n Steam Line Break Accident (Continued) 5.8.1 Definition of the Event (Continued)
The resultant cooldown of the reactor coolant, in the pre-sence of a negative moderator temperature coefficient of reactivity, will cause an increase in nuclear pewer and trip the reactor. A severe decrease in main steam pressure will also initiate reactor trip and cause the main steam isola-tion valves to close. If the steam line rupture occurs be-tween the isolation valve and the steam generator outlet no:-
sle, blowdown of the affected steam generator will continue.
(However, closure of the check valve in the ruptured steam line, as well as closure of the isolation valves in both steam lines, will terminate 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.
Both full power and no load (hot standby) initial condition cases were considered for two !< .i operation (i.e., four reactor coolant pumps).
Since the steam generators are designed to withstand reactor coolant system operating pressure on the tube side with at.
mospheric pressure on the shell side, the continued inte-grity of the reactor coolant system barrier is assured.
The most probable trip signals resulting from an MSLB in-clude 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).
OPPD NA 8303 NP, Rev. 02 Page 51 of 129
l 5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.8 Main Steam Line Break Accident (Continued) 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 crit-ical heat flux is not exceeded.
4 5.8.3 Obiectives 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 ori-
! ginal FSAR. This is accomplished by demonstrating that i
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 An dvsis Assumotions The MSLB accident is assumed to start from steady state 4
conditions with the initial power being 1530 MVt (102%)
for the full power case and 1 MVt for the no load case.
l The reactor coolant system cooldown causes the greatest i 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 assurned in the analysis. Since the reactivity change associated with rnoderator feedback varies significantly OPPD NA 8303 NP, Rev. 02 Page 52 of 129
5.0 TPANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.8 Main Steam Line Break Accident (Continued) 5.8.4 Kev Parameter and Analysis Assumotions (Continued) '
over the tempera:ure 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 th6 basis that upon reactor trip the most reac-tive CEA is stuck in the fully withdrawn position thus yielding the most adverse combination of scram worth and ,
reactivity insertion. Although no single value of MTC is assumed in the analysis, the moderator cooldown reactiv-icy function is calculated assuming an initial MTC equal to the most negative Technical Specification limit.
Reactivity feedback effects from the variation of fuel temperature (i.e., Doppler) are included in the anal.
ysis. The most negative Doppler defect function, when used in conjunction with the decreasing fuel temperature causes the greatest positive reactivity insertion during l the MSI.B event. In addition to assuming the most nega-tive Doppler defect function, an additional 154,uncer-tainty is assumed, i.e., a 1.15 multiplier. This multi-l plier conservatively increases the suberitical multiplica- i tion and results in a larger return to power.
The delayed neutron precursor fraction, d, assumed is .
the maximum absolute value including uncertainties for j end of cycle conditions. This is conservative since it also maximizes suberitical multiplication and thus, en- I hances the potential for a return to power. ,
l OPPD NA 8303 NP, Rev. 02 '
Page $3 of 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.8 Main Steam Line Break Accident (Continued) 5.8.4 Kev Parameter and Analysis Assumotions (Continued)
The steam generator low pressure trip, which occurs at 478 l psia (including a 22 psia uncertainty below the nominal trip setsing 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 tino for the full power case. For the cases analyzed, it is assumed that the most reactive CEA l
is stuck in the fully withdrawn position. If all CEAs in.
sert (no stuck CEAs), there is no return.co critical and no power transient following trip.
t The cold edge temperatures are used to calculate moderator l reactivity insertion during the cooldown, thus maximi:ing l
the return.co critical and re' turn.co power potentials, i
l The Emergency Operating Procedures incorporate the Trip 2/ Leave 2 RCP operating strategy as indicated in Reference 5 8. For a steam line break, the Trip 2/ Leave 2 strategy will result in tripping two RCPs (at 1350 psia). If the l event was misdiagnosed as a LOCA, all four RCPs would be tripped. As discussed below, for a main steam line break the consequences of Trip 2/ Leave 2 is bounded by the loss of offsite power and the loss of offsite power case is bounded by tripping no RCPs. Consequently, the limiting main steam line break accident occurs with all RCPs operating.
The MSLB case with the RCPs tripped is similar to the MSLB case with a loss of offsite power since the RCPs coastdown in both events. As discussed in Reference 5 4 the loss of offsite power delays safety injection due to the time delay OPPD.NA.8303.NP. Rev. 02 Page 54 of 129
5.0 TPRJSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.8 Main Steam Line Break Accident (Continued) 5.8.4 Kev Parameter and Analv91s Assumotions (Continued) for the emergency diesel generators to restore power to the safety injection pumps and causes a coastdown of the RCPs.
The coastdown affects the degree of overcooling and in-creases the time for safety injection borated water to reach the core midplane. Because manual tripping of the RCPs results in a later coastdown of the RCPs and because safety injection is not delayed since offsite power is available (i.e., the diesel generator startup and pump loading delays are not present), the injected boron will arrive at the core midplane sooner for a MSLB with the RCPs tripped than for a
!!SLB with a loss of offsite power. Therefore, the reactiv-ity effects of a MSL3 with the RCPs tripped are less severe than for the MSLB with a loss of offsite power.
Referente !.4 states that *.he MSLB case with a loss of off-site power results in the injected boron being dominant over the RCS cooldown and concludes that the reactivity effects of a MSLB accident would be reduced in severity with a con-current loss of offsite power when compared to the same event with offsite power availabio and the RCPs operating.
Because the reactivity effects of a MSLB with the RCPs tripped after SIAS are less severe than a MSLB with a con-current loss of offsite power. it is concluded that the reactivity effects for the MSLB case with the RCPs tripped utilizing Trip 2/ Leave 2 at 1350 psia are less severe than for a MSLB with offsite power available and RCPs operating.
OPPD.NA 8303 NP. Rev. 02 Page 55 of 129
d d'
5.0 TRANSIE';T AND ACCIDENT ANALYSIS METHOD (Continued) 5.8 Main Steam Lfne Break Accident (Continued) 5.8.4 Kev Parameter and Analysis Assumotions (Continued)
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 adverse re-suits.
j A lower flow rate increases the initial fuel and average primary coolant temperatures and consequently results'in a higher steam generator pressure and a greater steam gen-erator mass inventory. These effects cause a longer blow-down, a greater blowdown rate and a greater decrease in average primary coolant temperature. After MSIV closu.'e l the lower flow rate decreases the rate of reverse heat i
transfer' from the intact steam generator, thereby increas-l . in5 the heat extracted from the primary steam by the rup-j tured steam generator. The overall effect is that the i potential for a return to power'is maxitnized.
1 i'
Maximum values for the heat transfer coefficient across l the steam generator are used for the no load initial con-i dition case, while nominal values are used for the full-load initial condition. These heat transfer coefficients result in the most severe conditions during the incident because of the shape of the reactivity versus moderator temperature function and the difference in average moder-ator temperature for the maximum and minimum values of the steam generator heat transfer coefficients.
OPPD NA 8303 NP Rev. 02 Page 56 ef 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.S Main Stamm Line Break Accident (Contintted) 5.8.4 Kev Parameter and Analysis Assumptions (Continued)
The fast cooldown following a MSLB results in a rapid shrink-ing of the reactor coolant. After the pressuri:er is em-ptied, the reactor coolant pressure is assumed to be equal to the saturation pressure corresponding to the highest temperature in the system.
Safety injection actuation occurs at 1578 psia (i.e., 1600 psia minus the 22 psia uncertainty) after the pressuri:er empties. Additional time is required for pump acceleration, valve opening, and flushing of the unborated part of the safety injection piping along with the requirement that the RCS pressure decrease below the shutsff head of the safety injection pumps (1376 psia for high pressure safety injec- ,
cion (HPSI) pumps and 201 psia for low pressure safety injec-tion 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 reac-tor coolant at the points of injection into the cold legs.
Slug flow is assumed for movemene, of the mixture through the piping, plena, and core. After the boron reaches the core midplane, the concentration within the core is assumed to increase as a step function after each loop transit inter-val.
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 conservatively chosen to be large to minimi:e the negative reactivity in.
sortion from safety injection.
OPPD NA 8303 NP, Rev. 02 Page 57 of 129
5.0 TPWISIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.8 Main Steam Line Break Accident (Continued) 5.8.4 Kev Parameter and Analysis Assunctions (Continued)
Since the rate of temperature reduction in the reactor coolant system increases with rupture size and with steam pressure at the point of rupture, it is assumed that a circumferential rupture of a 26. inch (inside diameter) s.eam line occurs at the steam generator main steam line nozzle, with unrestricted blowdown. Critical flow is assumed at the point of rupture, and all of the mass leaving the bruak is assumed to be in the steam phase.
This assumption results in the maximum heat removal fr'om
~
'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 generator will have steam flow throu6h 4
the unaffected steam line and back through and out the
! ruptured line. Based on sensitivity analyses perforced
- by the District, this is the most severe single failure for the steam line break event. The analysis credits a 1 choke which is installed in each steam line immediately
. above the steam generator and assumes the steam flow from j the intact steam generator is through a 50% area redue-tion choke installed in a 24 inch steam line. This flow will be terminated upon MSIV closure.
The feedvater flow at the start of the MSL3 corresponds to the initial steady state operation. For the full load initial condition, it is automatically reduced in accord-ance with the program used in the valve controller. For the no load initial condition, feedwater flow is assumed OPPD NA.8303 NP. Rev. 02 Page 58 of 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued)
- 5.8 Main Stenm Line Break Accident (Continued) 5.8.4 Kev Parameter and Analysis Assutotions (Continued) to match energy input by the reactor coolant pumps and the 1 MWt core power. Feedwater isolation upon the re-ceipt 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.
5.8.5 Analysis Method The analysis of the main steam line break accident is per-formed using CESEC which models neutron kinetics with fuel and moderator temperature feedback, the reactor pro-tective' system, the reactor coolant system, the steam gen-erators and the main steam and feedwater systems.
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
.983 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 current Technical Specification limit on shutdown margin and moderator tem-perature coefficient.
5.8.7 Conservatism of Results Conservatism is added to the analysis by inclusion of un-certainties in moderator and fuel temperature coeffi- I cients of reactivity, by taking no credit for void reac-OPPD NA 8303 NP, Rev. 02 Page 59 of 129 l-
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.8 Main Steam Line Break Accident (Continued) 5.8.7 Conservatism of Results (Continued) tivity feedback, by taking credit for only 1 HPSI pump, by assuming all RCPs operate instead of manually tripping two pumps and by taking no credit for the stuck CEA worth.
5.9 Seired Rotor Accident 5.9.1 Definition of Event The seized rotor accident is assumed to be caused by a mechanical failure of a single reactor coolant pump. It is assumed that the rotor shears instantaneously, leaving 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 reduction 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 seizure is classified as a postulated accident for which the dose rates must be within 10 CFR 100 guidelines.
5.9.3 Obiectiva of the Analysis i
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 in of the pins fail during the event.
OPPD NA.8303.NP, Rev 02 Page 60 of 129
5.0 TPA'iSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.9 Seired Rotor Accident (Continued) 5.9.4 Kev 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 con-servatively assumed to result in a 0.1 second rampdown of the core flow from its initial value to the 3 pump value.
5.4.5 Analysis Method Two methods of analyzing the seized rotor event are dis-
'l 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 transient anal-ysis input.
l l 5.9.5.1 Analysis Method Without Transient Analysis F
{ Resoonse Ineut i
. This method calculates the number of pin fail-
- j. ures assuming that the core flov instantan-l eously decreases to the 3 pump flow rate. This
- method utilizes the TCRC analysis with a 3 pump inlet flow distribution. The initial RCS pres.
sure and core inlet temperature are used as l input to TORC and the core average heat flux is 1
j conservatively assumed to remain at its initial l
t-4 i
l '
3 l
1
)
I l
OPPD NA.8303.NP. Rev 02 Page 61 of 129 1
Tabl. 5.9.4 1 KEY PARAMETERS ASSUMED IN THE SEIZED ROTOR ANALYSIS Paraceter Units Value Initial Core Power NVt 1530t l Initial Core Inlet *F Maximum allowedt Temperature by Tech. Specs.
Initial Reactor Coolant psia Minimum allowedt System Pressure by Tech. Specs.
Initial Core Mass 6 x10 1bm/hr Minimum allowedt l Flow Rate by Tech. Specs.
Moderator Temperature x10Ap/*F Most positive l Coefficient allowed by by Tech. Specs.
Fuel Temperature x10Ap/*F Least negative pre- l Coefficient dicted during core life.
Core Average H gap BTU /hr.ft2,.F Minimum predteced l ,
during core life.
CEA Drop Time see Maximum allowed by l Tech. Specs.
Scram Reactivity gap Minimum predicted
'Jo r th during core life.
t Uncertainties on these parameters are combined statistically.
J l
l OPPD.NA 8303.NP. Rev. 02 Page 62 of 129
+
5.0 TRANSIENT AND ACCIDENT ANA1.YSIS METHOD (Continued) 5.9 Seized Rotor Accident 'antinued) 5.9.5 Analysis Method (Continued) 5.9.5.1 Analysis Method k'ithout Transient Analysis Reseense ineut (Continued) value. The maximum value of FR is com.
bined with a conservatively flat power distri-bution. The TORC calculation determines the number of pins that have failed.
5.9.5.2 Ansivsis Methods ik.tne Transient Analysis L
This method utili:es the CESEC code to calcu.
late the transient response for the seized rotor event. The CETOP code is then used to I
~
determine the time of minimum DNBR. The TORC I
code utili:es the 3 pump inlet flow distribu-tion. 3. pump core flow rate, and the RCS pres-sure, core inlet temperature and core heat i
flux calculated at the time of minimum DNBR by '
CESEC.
I 5.9.6. Analysis Results and 10 CFR 50.59 Criteria l
l The results of the sei:ed rotor analysis are contained in f 5
Section 14.6.2 of the Fort Calhoun Station Unit No. 1 l (
USAR. The criteria of 10 CFR 50.59 are met, if the number I L
of pin failures is less than one percent.
r OPPD NA.8303.NP, Rev. 02 Page 63 of 129
1 5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.0 Seired Rotor Accident (Continued) 5.9.7 Conservatism of Re.sults Conservatism in the calculate number of fuel pins pre-dicted to experience DNBR is added through the use of the following assumptions:
A. The most positive MTC is assumed in the analysis. !
The actual MTC is more negative and would limit core power and heat flux rise.
B. A relatively flat pin census is assumed in the anal- l ysis. A core peaked pin census distribution would lower the number of pins predicted to experience '
DNB.
C. Fo : the case without transient analysis, no credit l is taken for the pressure increase during the trans- i isnt and calculating the minimum transient DNBR. I 5.10 CEA Eiection Accident 5.10.1 Defirittien of Ever$
i
. A CEA ejection accident is defined as a mechanical failure of a control rod raschanic61 pressure housing such that the
, coolant system pressare would sject the CEA and the drive shaft to a fully withdrawn position. The consequences of !
this mechanical failure is a rapid reactivity insertion i which when combined with ati adverse core Oover distribu. ,
i tion potentially leads to locali:ed fuel damage. The CEA CPPD.NA.8303.NP. Rev. 02 Page 64 of 129 l
. 5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) '
- r ,
I 5.10 riection Accident (Continued) 5.10.1 pefinition of Event (Continued)
(
ejection accident is the most rapid reactivity insertion i that can be reasonably postulated. The resultant core and [
thermal. power excursion is limited primarily by the f Doppler reactivity effect of the increased fuel l temperatures and is terminated by reactor trip of the (
remaining CEAs activated by the high power trip or f r
variable high power trip. j i
5.10.2 Analysis Criteria 2
The CEA ejection event is classified as a postulated acci.
dent. The design and limiting criteria are. ,
l A. Fuel claddin's and enthalpy thresholds (Reference. {
5 5) are: i j
4 i i
Clad Damage Threshold i l
Total Average Enthalpy - 200 cal / gram l l
Centerline Melting Threshold Total Centerline Enthalpy - 250 cal / gram t
i Fully Molten Centerline Threshold Total Centerline l t
4 Enthalpy - 310 cal / gram >
{
l i
j j
OPPD NA 8303 NP. Rev. 02 Page 65 of 129
5.0 TPANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.10 eEA E_iaccion Accident (Continued) 5.10.2 Analysis Criteria (Continued)
B. The peak reactor pressure during a portion of the l 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. This objection is achieved if the peak RCS pressure does not exceed 2750 psia.
C. Fuel melting vill be limited to keep the offsite dose l consequences well within the guidelines of 10 CFR 100.
l 5.10.3 Obiectives of the Ansivsis The objective of the analysis is to demonstrate that fuel failures are less than those reported in Section 14.3.4.1 of ths Fort Calhoun Station Unit No. 1 USAR of that site boun-dary doses are within the 10 CTR 100 limits.
5.10.4 Analysig Method The District utili:es the CEA Ejection Accident Ana).ysis of our current fuel vendor, Combustion Engineering. *his anal- .
ysis methodology is documented in Reference 5 5 and is per- l formed by Combustion Engineering. This methodology utili:es physics parameters, calculated by the District in accordance t ith the methods outlined in Reference 5 6, OPPD NA 8303 NP Rev. 02 Page 66 of 129
5.0 TPANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.10 CEA Eiection Accident (Continued) 5.10.5 Analvgis Results and 10 CFR 50.5o Criteria The results of the CEA Ejection Analynis 6.e reported in See-tion 14.13 of the Fort Calhoun Station Unit No. 1 USAR. Cri-teria of 10 CFR 50.59 are satisfied if fuel failures are less than those assumed for input to the Radiological Consequences portion of the analysis.
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 vorth and the ejected radial peak are calculated without any credit for Doppler or Xenon feedback. In addi-tion, the hot full power ejected worth and ejected peak are calculated assuming the no load temperature of 532'T. The
, lower temperature is more adverse since this causes a power J
roll to the core perip..ery which also happens to be the loca- l tion of the ejected CEA. Also, the ejected worth is calcu-lated assuming the CEAs are fully inserted for hot full power case regardless of PDIL. Thus, the ejected worth is conser-vative.
5.11 Loss of Coolant Accident The District does not perform the Loss of Coolant Accident Analysis.
The large and small break loss of coolant anclyses were performed by Combustion Engineering (CE). The large break topical is mentioned in Reference 5 7. The small break analysis shows the closest approach l to the 10 CFR 50.46 criteria for ECCS analysis. The actual differ-OPPD NA 8303 NP Rev, 02 Page 67 of 129
i 5.0 TRANSIENT AND ACCIDENT ANALYSIS Mb JD (Continued)~
5.11 Loss of Coolant Accident (Continued) r i ence is less than 10'F in PCT. The District verifies that the !
physics input assumptions and the maximum rod burnup are within the I bounds assumed in the CE large break analysis. ;
o 5.12 L234 of Load to Both Steam Generators Event f l
i 5.12.1 Definition of Event I
- A total loss of load to both steam genera'ers usually results from a turbine trip due to a loss of exteg..a1 electrical load -[
] or to abnormal variations in electrical network frequencies. f other possible causes include the simaltaneous closure of all f turbine stop valves or main steam isolation valves. All ini- j
! tiating mechanisms result in a corresponding reduction in i t
heat removal from the reactor coolant system due to the loss i of secondary steam flow. Although a Reactor Protective Sys.. !
i tem trip signal would normally result from a turbine trip, no l credit is taken in the analysis of this event for the turbine I
- trip signal. [
I I I i 5.12.2 Analysis Criteria' j
f '
t k
The loss of load to both steam generators event is classified
- . as an Anticipated Operational Occurrence (A00) for which the !
following criteria must be met: i l
1 l
1 i
i j
OPPD NA 8303 NP Rev. 02 Page 68 of 129
5.0 TRA!iSIENT AliD ACCIDENT ANALYSIS METHOD (Continued) 5.12 Loss nf Load to Both Steam Cenerators Event (Concinued) 5.12.2 Analysis Criteria (Continued)
A. The peak RCS pressure does not exceed 2750 psia l (110% of design pressure).
B. The transient minimum DNBR is greater than the 95/95 l confidence interval limit for the CE.1 correction limit.
C. The Peak Linear Heat Ceneration Rate (PLHCR) does j no't exceed 22 kw/ft.
Criteria B. and C. are not of major concern since DNBR in- l creases during the event and the PLHCR margin required is much less limiting than other A00s. Therefore, criterion A. is the main concern in analy:ing this event. The loss l o' load to both steam generators event is the limiting ADO event with respect to peak RCS pressure.
5.12.3 Obicetives of the Analysis The objective of the analysis is to demonstrate. for modi-fications to the plant which potentially degrade RCS heat removal capability (including steam generator plugging) that the peak RCS pressure stays within 1106 of the design pressure in accordance with Section III of the ASME Pres-sure Vessel Code. This objective is achieved if the pean RCS pressure does not exceed 2750 psia.
OPPD.NA.9303.NP, Rev. 02 Page 69 of 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.12 Loss of Load to Fpth Steam Generators Event (Continued) 5.12.4 Kev Parameters and Ana'.vsis Assumotions The key parameters used in the loss of load (to both steam generators) event are given in Table 5.12.4 1. Assumptions used in the analysis to maximize heat up of the RCS and con-sequently peak RCS pressure include:
A. The event is initiated by a sudden closure of tne tur- l bine stop valves without a simultaneous reactor trip.
B. No credit is taken for operation of the PORVs, pressur. l 1:er sprays, and the turbine steaa dump and bypass sys-tem, i.e., the pressuri:er pressure control system is assurned to be in MANUA1..
C. The rod block system is assumed to prevent rod motion l (other than scram) during the transient.
D. Maximum charging flow and :ero letdown flow are as- l sumed.
E. Termination of the event occurs as a result of a high (
pressurizer pressure trip.
5.12.5 Analysis Method The analysis methods utilized by the Distris.: to analyze the loss of load to both steam generators event consists of simu-lacing the event using the CESEC computer code, utili:ing the analysis assumptions listed in Section 5.12.4 (above) as in-put, and extracting the peak RCS pressure for comparison with the 2750 psia upper- limit. !
OPPD NA 8303 NP. Rev. 02 Page 70 of 129
Table 5.12.4 1 KEY PARAMETERS ASSUMED IN THE LOSS OF LOAD TO BOTH STEAM CENERATORS ANALYSIS farneater h Ea.l.ug Initial Core Power MVt 1530 (
Initial Core Inlet *F Maximum allowed Temperature by Tech Specs.
Initial RCS Preasure psia Minimum allowed by Tech. Specs Initial Steam Generator psia Minimum value cor.
Pressure responding to core inlet temperature operating range.
Initial Core Mass x106 lbm/hr Minimum allowed l Flow Race by Tech. Spe:s.
Moderator Temperature x10*'ap/*F Most positive l Coefficient allowed by Tech.
Specs.
Fuel Temperature x10*'ap/'F Least negative l Coefficient predicted during core life.
Fuel Temperature Coefficient Multiplier 0.85 CEA Drop Time see Maximum allowed l by Tech. Specs.
. Scram Reactivity Vorth nas Minimum predicted durin5 core lifetime Scram Reactivity Consistent with most Curve positive axial shape (bottom peaked) allowed by Tech.
Specs.
Core Average H gap BTU /hr ft2 'F Maximum predicted l during core lifetime.
Kinetics Parsmeters EOC parameters (minimum absolute 4).
RPS Response Time see 1.4 l OPPD NA 8303 NP. Rev. 02 Page 71 of 129
l 5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued)
I i
5.12 Loss af toad to Both steam cenerators Event (continued)
?
5.12.6 Analysis Results and 10 CFR 50.5o Criteria t
The results of the loss of load to both steam generators are !
t contained in the Fort Calhoun Station Unit No.1 USAR. The !
criteria of 10 CFR 50.59 are met, if the peak RCS pressure [
is less than 110% of design pressure in accordance with See- [
tion III of the ASME Pressure Vessel Code. This objective ;
is achieved if the RCS pressure does not exceed 2750 psia. :
1 ,
- i 5.12.7 Conservatism of Results i
The following areas of conservatism are included in the anal- f
- ysis to obtain a conservatively high peak RCS pressure
- [
- l
! F A. Field measurements demonstrate that the CEA magnetic ;
l i clutch decay time is less than that assumed in the l s
i analysis. L 1 1 e
} B. The actual scram worths are greatet than those assumed l
in the analysis. [
j C.
The actual MTC is more negative during power operation l f
] than assumed in the analysis. !
\
1 D. The steam dump and bypass systen and the pressurizer l ,
i pressure control system (PokVa and sprays) are oper- !
l I
ated in the AUTO mode rather than MANUAL as assumed in j the analysis.
I i .
- i d
i 1
a 4
I OPPD.NA 8303 NP Rev. 02 i Page 72 of 129 [
t !
i 5 l
5.0 TRR;SIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.12 Loss of Load to Both Steam Generators Event (Continued) 5.12.7 conservatism of Results (continued)
E. Actual secondary pressure is higher which results in l earlier secondary safety valve opening and earlier alleviation of the primary system temperature and pressure rises.
F. The maximum pressuri:er safety valve capacities are l assumed to be 90% of the ASME rated values.
C. A one percent pressure uncertainty is applied to the primary and secondary safetf valve setpoints, i.e., n 1.01 multiplier.
5.13 Loss of Feedvater Flow Event 5.13.1 Definition of Event A total loss of main feedwater flow event is defined as a loss of feedwater flow when operating at power without a corresponding reduction in steam flow from the steam gen-erators. The most likely causes for this event are the loss of all feedvater or condensate pumps or the inadver-tent closure of either the main feedwater regulating valves or the feedwater isolation valves due to a feedwater con-troller malfunction or manual positioning by the operator.
The result of this mismatch in which turbine demand remains at 1004, is a reduction of the steam generator liquid inven.
tories and a degrading RCS heat removal capability. As the heat removal capability is lost. through decreasing steam generator inventories (i.e., levels) the RCS temperatures OPPD NA 8303 NP, Rev. 02 Page 73 of 129
5.0 TPW;SIENT AND ACCIDENT ANAL.YSIS METHOD (Continued) 5.13 Loss of Feedvater Event (Continued) 5.13.1 Definition of Event (Continued) and pressure increase. Normally the event would be ter-minated by a reactor trip on low steam generator level.
Since no credit is taken in the analysis for the steam generator low level trip, a high pressuri:er trip even-cually results.
Automatic actuation of the auxiliary feedwater (afb') sys-tem will also eventually occur (after reactor trip) if either main feedwater is not restored or manual actuation of the AP=' system is not performed by the operator. The APJ system actuation ensures the maintenance of a second-ary heat sink.
5.13,.2 Analysis Criteria The loss of feedvater flow event is classified as an Anti-cipated Operational Occurrence (ADO) for which the follow-ing criteria must be met:
A. The peak RCS pressure does not exceed 2750 psia l (110% of design pressure).
B. The transient minimum DNER is greater than the l 95/95 confidence interval limit for the CE 1 cor-relation limit.
C. The Peak Linear Heat Generation Rate (PLHCR) does {
not exceed 22 kv/ft.
OPPD NA 8303 NP. Rev. 02 Page 74 of 129
i i
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) l a
5.13 Loss of Feedvater Flev Event (Continued) !
l i 5.13.2 Analysis Criteria (Continued) a t
Criteria B. and C. are not of major concern because DNBR l '
does not decrease below the initial steady stati value I and the F1)(CR margin required is much less limiting than [
other A00s. Therefore, only criterion A. requires re. l l evaluation should plant modifications (such as steam gen. ,
erator tube plugging) be made which result in degraded I
secondary heat transfer capability beyond that of this i l- event. For Fort Calhoun Station, this event is bounded
- by the loss of load incident, t o
t I
t 5.13.3 Obiectives of the Analysis I The objective of this analysis is to demonstrate, for I plant modifications which potentially degrade RCS heat removal capability (including steam generator tube plug.
l ging), that the peak RCS pressure stays within 110% of l 2 the design pressure in accordance with Section III of the i
ASME Pressure Vessel Code. This objective is achieved if l
[ the peak RCS pressure does not exceed 2750 psia. I d
f 1
! 5.13.4 Kev Parameters and Analysis Assumerions l
} [
I
- The key parameters used in the loss of feedwater flow j
{ event are given in Table 5.13.4 1. Assumptions in the
I
{
A. The event is initiated by an instantaneous loss of j main feedwater. No credit is taken for the low steam generator level trip.
I OPPD.NA.8303.NP, Rev. 02 i Page 75 of 129
5.0 TFWiSIE'iT AfiD ACCIDENT ANALYSIS METHOD (Continued) 5.13 Loss of Feedwiter Flow Event (Continued) 5.13.4 MeV Parameters and Analysis Assu?otions (Continued)
- 8. The steam dump and bypass system is assumed to be l in MANUAL (i.e., inoperative).
C. The pressurizer pressure control system is in l MANUAL (i.e.. PORVs and sprays are inoperable).
D. The pressurizer level control system is in MANUAL l vith maximum charging and zero letdown flows.
E. The rod block system is assumed to prevent rod l cotton (other than scram) during the transient.
5.13.5 Antlysis Method The analysis methods used by the District to analyze a loss of main feedvater flow event consists of using the CESEC computer code to simulate the event, utilizing the analysts assumption li:ted in Section 5.13.4 (above) as input, and extracting the peak RCS pressure for compari-son v* s the 2700 psia upper limit.
5.13.6 Analysis Results and 10 CFR 50.50 Criteria The results of the loss of feedwater flow are contained in the Fort Calhoun Station Unit No. 1 USAR. The cri-teria of 10 CFR 50.59 are met, if the peak RCS pressure is less than 110% of the design pressure in accordance with Section III of the ASME Pressure Vessel Code. This ,
objective is achieved if the peak RCS pressure does not exceed 2750 psia.
OPPD NA 8303 NP. Rev. 02 Page 76 of 129
r I
}
Table 5.13.4 1 I I
KEY PARAMETERS ASSUMED IN THE LOSS OF FEEDVATER FICW ANAL.YSIS
, Parameter h h
' i Initial Core Power MWe 1530 l Initial Core Inlet *F Maximum allowed Temperature by Tech. Specs.
} [
Initial RCS Pressure psia Minimum allowed I by Tech. Specs L t
Initial Steam Cenerator psia Minimum value cor. I Pressure responding to core inlet temperature [
operating range. f t
Initial Core Mass x106 tha/hr Minimum allowed j [
Flow Rate by Tech. Specs. ;
Moderator Temperature x10ap/'T Most positive l [
Coefficient allowed by Tech. :
Specs.
{
L l Fuel Temperature x10*'ap/*F Least negative i Coefficient predicted during l [
core life, i Fuel Temperature !
j Coefficient Multiplier 0.85 -
f
- CEA Drop Time see Maximum allowed I by Tech. Specs.
- i
!! Scram Reactivity Vorth 4Ap Minimum predicted
] during core lifetime I
- Scram Reactivity Consistent with most
] Curve positive axial shape
.j (bottom peaked) allowed by Tech.
Specs.
Core Average Hg ,p BTU /hr ft2 *F Maxim a predicted during core lifetime l )
t
- Kinetics Parameters EOC parameters ,
(afnimum absolute l J A). (
i !
] RPS Response Time see 1.4 l [
OPPD NA 8303 NP Rev. 02 l 1
Page 77 of 129 t
j k
_,_ ._ .~.- - _ __.,- _ _ _ _ . _ .__ _ ,__ 5 _ _ _ _ _ . - _ . . .
4 5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.13 Loss of reedvater Flov Event (Continued) 5.13.7 Conservatisms of Results A, Field measurements demonstrate that the GEA magnetic <
clutch decay time is less than that assumed in the ;
analysis.
- 5. The actual scram worths are greater than those as- l sumed in the analysis.
I
, C. The actual MTC is more negative during power oper- l ation than assumed in the analysis.
D. The steam dump and bypass system and the pressuri:er l
- pressure control system (FORVs and sprays) are oper-ated in the AUTO mode rather than the MANUAL mode as i assumed in the analysis.
E. Actual secondary pressure is higher which results in i.
earlier secondary safety valve opening and earlier alleviation of the primary system temperature and f
pressure rises.
F. No credit is taken for a steam generator low level I trip.
r L
i r
OPPD NA-8303 NP Rev. 02 -
Page 78 of 129
r-e f
t
( 6.0 TRANSIENT ANALYSIS CODE VERIFICATION !
6.1 Introduction fI s
The District utili:es the CESEC III computer code to calculate the transient response of the NSSS during events discussed in this doc-ument. Combustion Engineering has provided overall verification of h the CESEC III code in References 6 1 and 6 2. The purpose of the l t
work reported here is to demonstrate the District's ability to cor- j rectly utili:e the CESEC III code.
In order to demonstrate Omaha Public Power District's ability to cor- ,
rectly use the CESEC III computer code, verification work has been [
performed by benchmarking both actual plant transient data and in-dependent safety analyses previously accepted by the NRC. The plant !
I transients which were benchmarked were the Turbine Reactor trip and !
t Four Pump Loss of Coolant Flow events. The independent safety anal. [
yses which were benchmarked were the Dropped CEA, Main Steamline {
Break, and RCS Depressuri:ation events. Each of the comparisons !
I will be addressed below. (
i 6.2 Comoarison to Plant Data l
A prerequisite for beginning performance of transient analyses 1: f verification that ths code will stabili:e with the correct system parameters when simulating steady state operation. This step was performed following setup of the CESECeIII code and correct results l were obtained, f
I For plant transient benchmarking, the type of transients that have ;
occurred and both the quality and quantity of data existing for each is very limited. In nearly all cases, operators take actions which reduce the consequences of the event, introducing complicated pertur- f bacions in system response which cannot be easily modeled. because i the actions taken and the time at which they are performed are not i i h
. i OPPD NA 8303 NP. Rev. 02 !
Page 79 of 129 [
- l 1 ,
6.0 TPASSIENT ANALYSIS CODE VERIFICATION (Continued) 6.2 Co-carison to Plant Data (Continued) recorded. Strip chart recordings on an extremely compressed time scale are generally the only form of data available. This com-pressed time scale (with graduations typically of 10 minutes) does not permit adequate 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 354 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 permit accurate comparisons to the actual measured plant responses for both of the above cases. A summary of each of these comparisons follows.
6.2.1 Terbine Resetor Trie For the turbine-reactor trip case, the plant comparison data were obtained from the Cycle 1 startup testing per-formed May 10, 1974 The event was initiated fron 97% of full power, all rods out, and equilibrium xenon. The plant response data used in the CESEC III comparisons were obtained from vendor test recorders. No operator action was taken following the manual generator turbine trip (which provided the RPS "loss of load" trip) Prior to the trip the main feedvater, the pressuri:er pressure, and pressuri:er level control systems were all in the automatic mode, and the letdown backpressure control valve was in the manual mode. k'ith the exception of adjusting the letdown backpressure control valve at 20 seconds, no operator action was taken for 60 seconds following the trip.
OPPD NA 8303 NP, Rev. 02 Page 80 of 129
~
6.0 TPW:SIENT ANALYSIS CODE VERIFICATION (Continued) 6.2 cnenarisen en Plant Data (Continued) 6.2.1 Turbine Reactor Trio (Continued)
Figures 1 1 through 1 7 show plots of the comparisons be-tween the measured plant responses and the CESEC III pre-dicted responses. It should be noted that this test was performed based on a rated power level of 1420 L't rather r than the current limit of 1500 .MVt (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 I the same power decay rate, however, the endpoint residual I power is slightly higher, i.e., conservative. It should be noted that trip delays included in the CESEC III model-ing prevent the immediate power drop observed in the plant data; again this is conservative. The pressurizer pressure response predicted by CESEC III and shown in Figure 1 2 shows very good agreement with the plant re-sponse. The CESEC III case was initiated 10 psia above i che plant data and remained slightly above the plant re-sponse for the duration of the transient. The difference
, between the predicted and measured pressurizer pressures increased slightly due to the higher residual power after f trip as shown in Figure 1 1. This difference between the [
predicted and measured pressurizer pressures at 60 so-conds is only 19 psia, a value which is less than the ;
pressure measurement uncertainty. Figure 1 3 shows the [
J pressurizer leve' response. The comparison between the measured and predicted values shows excellent agreement.
i Figure 1 4 shows the RCS cold leg and hot leg temperature ,
i responses for each steam generator loop for the plant i
OPFD NA 8303 NP, Rev. 02 t Page 81 of 129
9 6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued) 6.2 Cow arison to Plant Data (Continued) 6.2.1 Turbine Reactor Trio (Continued)
) data and the CESEC III predicted average cold leg and hot.
les temperatures. The differences in the transient re-sponse of the two steam generator loops for the plant data is attributable to the differences in the main feed.
water flow rate rampdown after trip (see Figure 1 6),
T The CESEC III responses lead the loop seasurements be-
, cause of the measurement delays associated with the re.
I sponse time of the RTDs (resistance temperature devices)
- providing the temperature signals. Figure 1 5 shows the measured and predicted steam generator pressure re-
! sponses. Thes'e two plots show very good agreement 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 Fig.
I ure 1 1, a quicker turbine stop valve closure, and quick-er 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 agreement if
- the CESEC III input vare modified, however, the overall l
differences are small enough not to warrant the reanaly-sis.
In conclusion, the CESEC III predicted parameters for the turbine reactor trip show very good agreement with those measured in the Cycle 1 startup testing performed at nom-inal full power conditions.
1 OPPD NA 8303 NP, Rev. 02 Page 82 of 129
6.0 TRANSIE'iT ANA1.YSIS CODE VERIFICATION (Continued) 6.2 Comearison to Plant Dsta (Continued) 6.2.2 Four Pure Loss of Coolant Flow For the four pump loss of coolant flow case, the plant com-parison data were obtained from the Cycle 1 startup test performed March 6, 1974 This event was initiated from 356 power by manually and simultaneously tripping all four reac-tor coolant pumps. At the time of trip the pressurizer pressure, pressurizer level, main feedwater, and steam dump and bypass controllers were in the automatic mode. At ap.
proximately 20 seconds after the trip, the operators took manual control of feedwater in order to preclude overfeed-ing of the steam generators and too rapid of a cooldovn for i
the following natural circulation test.
The behavior of the various RCS and secondary parameters that were measured and the CESEC III predictions for the first 30 seconds following the RCP trips are shown in Fig-ures 2 1 through 2 8. These comparisons show excellent agreement. The minor differences that exist are discussed below.
i Figure 2 1 shows a plot of the measured total RCS flow ver-sus time and that predicted by the CESEC III code which in-corporates explicit modeling of the reactor coolant pumps.
These data show excellent agreement with the predicted flow being slightly cor.servative. Figures 2 2 and 2 3 show the pressuri:er pressure and level response comparisons which also show excellent agreement. Figure 2 4 shows plots of core nuclear power versus time. As *n the turbine reactor trip case, CESEC III shows a slightly higher residual power after trip. The predicted and Jeasured steam generator OPPD NA 8303 NP. Rev. 02 Page 81 of 129
6.0 TFx;SIENT ANAI.YSIS CODE VERIFICATION (Continued) 6.2 Catearison to Plant Data (Continued) 6.2.2 Four Pume loss of Coolant Flev (Continued) 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 Figure 2 6, is con-sistent with the data obtained from the turbine reactor trip case. Again the delay associated with the RTD re-sponse causes the predicted temperatures to lead those that were measured. Figure 2 7 shows that the main feed.
vater input function used in CESEC III was acceptable in terms of the actual feedvater system response, It should be noted that the operator action of assuming manual con.
trol of the main feedvater system at approximately 20 seconds had little effect on any of the other system para-meters examined, and that following a several second re-duction in flow the previous flow rate was reestablished.
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 something one would not expect and raises the question of the validity of the measure-ment or its uncertainty for this steam generator steam rate flow, because the two corresponding feedvater flow rates (in Figure 2 7) are consistent, In conclusion, the CESEC III predicted parameters for the 354 power total loss of coolant flow show very good agree-ment with those measured during Cycle 1 startup testing.
OPPD NA 8303-NP, Rev. 02 Page 84 of 129 i
6.0 TPANSIENT ANALYSIS CODE VERITICATION (Continued) 6.3 Co?earisons Petween OPPD Analysas and Inderendent Analyses Preriousl" Performed by the Fuel Venders of the transients analyzed by OPPD for reload core licensing (using CE methodology) no plant data existed, so comparison of the limiting events to previous independent analyses performed by either Advanced Nuclear fuels, formerly Exxon Nuclear Company (ENC), or Combustion Engineering (CE) was done. Since Exxon Nuclear Company performed some analyses in this section (used for comparison) prior to becoming ANT, all references to this company will be to ENC. For the compar.
ison cases, the assumptions used in the analyses were similar to those used by the District, i.e., the core physics parameters did not vary significantly between fuel cycles. The events chosen for compar-ison were:
(1) The Dropped CEA event is dependent upon the initial avail-able overpower margin to prevent exceeding the SAFDLs. The goal of the analysis is to determine the DNBR required over-power margin (ROPM).
(2) The Hot Zero Power (HZP) Main Steamline Break which deter-nines the minimum required shutdown margin.
(3) The Hot Full Power (HTP) Main Steamline Break which dete-raines the most negative moderator temperature coefficient of reactivity allowed.
(4) The RCS Depressurization event which is used in the deter-mination of the term. The term accounts for DNBR margin degradation in the thermal margin / low pressure (TM/LP) trip.
OPPD NA 8303 NP. Rev. 02 Page 85 of 129
6.0 TRASSIENT ANALYSIS CODE VERITICATION (Continued) 6.3 Corearisons Petvean OPPD Analyses and Inderendent Analyses Previous!"
Perforced bv the Fuel Venders (Continued) 6.3.1 Droered CEA The Cycle 8 Dropped CEA analysis performed by OPPD was com-pared to the previous analysis, contained in the Updated Safety Analysis Report (USAR). The USAR analysis was per.
formed by ENC for Cycle 6. Table 6.3 1 sumnarizes the para- l meters and their values for Cycles 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 vary similar prompt drop, to 69% versus 70%. respectively, and both cases show a return to a nominal 1006 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 Cal.
houn Station). The core heat flux plots are contained in Figure 3 2. Both are very similar, as was the case in the core power cases. Figure 3 3 contains plots of the coolant average temperature versus time. Both figures are in good agreement showing a drop in average coolant temperature to 567'T. Plots of the inlet and outlet temperatures for Cycle 8 are also included. Tigure 3 4 shows plots of the pres-suri:er pressure versus time. The minimum pressures pre.
dicted at 160 seconds are 1957 psia and if45 psia for Cycle 8 and Cycle 6. respectively. This difference is small enough to be less than the pressure measurement uncertainty.
In sumnary. 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.
OPPD NA 8303 SP. Rev. 02 Page 86 of 129
6.0 TRNiSIENT ANALYSIS CODE VERITICATION (Continued) 6.3 forearisons Batveen OPPD Ansivges and Inderandent Analyses Previnusiv Perfor:_ red by the Fuel Venders (Continued) 6.3.2 Hot Zero Power Main Stesmline Break The hot zero power (HZP) Main Steaaline Break, which is the basis for determination of the required shutdown margin, was analyzed by OPPD for Cycle 8. The results of this analysis have been compared to those of ENC in their Cycle 6 analysis and to those obtained by CE in their Cycle 6 control grade auxiliary feedwater (AFW) system analysis. Table 6.3 2 shows comparisons 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 max.
imum return to power is less for cycle 8 than for Cycle 6 and occurs later due to the use of a higher shutdown mar.
gin. The Cycle 6 CE ATW 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 aver.ge heat flux for OPPD ENC and CE, respectively. Both the CFFD and CE analyses, which were performed using CESEC.III and CESEC.I.
respectively show a slight heat flux increase at approxi.
mately 12 seconds. This is due to suberitical multiplica.
tion. Otherwise, the heat flux curves within the specific analyses are essentially the same as the core power curves with a slight decay. Tid ure 4 3 shows the total reactivity versus time for each of the aralyses. Vith very similar moderator cooldown curves. the peak reactivities occur chron-ologically with increasing shutdown margin as expected; i.e., for increased shutdown margin (CEAs) it takes Innger to be offset by the positive moderator cooldown reactivity insertion.
OPPD NA.8303.NP, Rev. 02 Page 87 of 129
(
I l 6.0' TRANSIENT ANA!.YSIS CODE VERIFICATION (Continued) !
l
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6.3 conearisons Eceveen OPPD Analyses and indecendent Ansivses Previnusiv Perforced by the Fuel Venders (Continued) !
6.3.2 Hot Zero Power Main Steamline Break (Continued)
Tigure 4 4 shows plots of RCS pressure versus time for Cycle 8 (OPPD) and Cycle 6 A N (CE). Also included in Figure 4 4 is the Cycle 1 (CE) results. All three of these curves show l excellent agreement. The Cycle 6 A N (CE) analysis shows a t lower and point 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, f I
Figure 4 5 shows plots of the steam generator pressv.res for {
Cye' 8 (OPPD) and Cycle 6 A W (CE), respectively. These f plo t show reasonable agreement between pressures and times.
The .1 crease in the intact steam generator's pressure is due to MSIV closure; i.e., failure of the reverse flow check f
valve on the intact steam generator was chosen as the most j adverse single failure. Following dryout of the ruptured !
steam generator, the pressure drops to atmospheric. The times of dryout are slightly different due to the increased l r
normal water level value used in the Cycle 8 analysis. [
6 f
In summary, the HZP Main Steamline Break analysis for Cycle I 8 shows trends similar to those in Cycle 6 as analyzed by both CE and ENC.
6.3.3 Hot Full Paver Main steam 11re Break i The hot full power (HTP) Main Steaaline Break provides an acceptance criteria for the most negative moderator tempera-i i
CPPD NA 8303 NP. Rev. 02 f Pate 88 of 129 I I
6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued) 6.3 Co-earisons Between OPPD Analyses and IndeDendent Analyses Previousiv Performed by the Fuel Vendors (Continued) 6.3.3 Mot Full Power Main Steamline Break (Continued) ture coefficient (MTC) of reactivity. If a return to crit.
ical 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 analy-sis. The Cycle 8 HFP analysis of this event was compared to the previous analyses performed by ENC in Cycle 6 and by CE in their Cycle 6 control grade ATV system analysis. Table 6.3 3 shows a comparison of the important input parameters for each of the analyses.
Tigures 51, 5 2, and 5 3 show plots of core power, core average heat flux, and total reactivity for Cycle 8 (OPPD).
Cycle 6 (ENC), and Cycle 6 ATV (CE). Vithin each cycle's analysis, the core average heat flux slightly lags the core power which peaks at a time several seconds aftat the peak reactivity is reached (for the return to critical cases).
The return to power peaks occur at different times due to the different scran worths used, as explained for the shut-down margin in the HZP Steamline Break analysis section.
Figure 5 4 shows plots of the RCS pressure versus time for the Cycle 8, Cycle 6 ATV, and Cycle 1 analyses. These plots are very similar and show excellent agreement. Figures 5 A and 5 5 show plots of the RCS temperatures for Cycle 8 and Cycle 6 ATV. Again good agreement exists to approximately 180 seconds. At this time, the Cycle 6 AFV analysis assumed runout flow from both ATV pumps to the ruptured steam gen.
OPPD NA 8303 NP. Rev. 02 Page 89 of 129
h 6.0 TPxiSIENT ANALYSIS CODE VF.RITICATION (Continued) 6.3 Cc-enrisons Between OPPD Ansiv5es and Indeoendent Ansivsis Previously Perforred by the Fuel Venders (Continued) l l 6.3.3 Het Full Pever Main Steamline Break (Continued) erator which resumed the RCS cooldown. This additional cool.
down caused by the AFV system is prevented from occurring in Cycle 8 by the logic of the never safety grade AFV system.
I Tigure 5 6 shows plots of steam generator pressures versus time for Cycle 8 and Cycle 6 AFV (CE). These results are very similar except that the intact steam generator pres-sure, in the CE analysis, begins to drop after 180 seconds due to the AFV induced RCS cooldovn.
6.3.4 RCS Beeressurization l The RCS Depressurization analysis is performed to calculate l a term for the TM/LP trip which accounts for the DNBR margin 1
degradation.
Because no figures from previous cycle analyses exist, com-parison was made between the transient analysis training manual sample analysis and tt.e figures generated by OPPD for Cycle 8. Pertinent input parameters are summarized in Table 6,3 4 l
l OPPD NA 8303 NP Rev. 02 Page 90 of 129
i 1
6.0 TRRISIENT ANA1.YSIS CODE VERITICATION (Continued) i I
6.3 Co-earisons Between OPPD Ansivses and Indet'endent Ansivses Previou91v I i Perfor-ed by the Fuel Vendors (Continued) I 6.3.4 RCS 3eeressurination (Continued) f A manual trip is next simulated at the time of maximum mar.
l gin degradation, i.e., at the time the maximum RCS Depres.
l eurization rate occurs. The uaximum RCS Depressurstation l rate occurs in approximately the first 20 seconds and is j constant. Therefore, the time at which a aan.isi trip should j occur is arbitrary but must be in the first % seconds. A l 1 trip time corresponding to a 100 psia drop is adequate to f
- perform the analysis, f
In the CE example, the initial pressure was 2300 psia a j value which corresponds to the maximum pressure before which ,
l the pressuriser sprays will be activated in a 2700 L'(th) (
l class planc (whose nornal RCS pressure is 2250 psia). In .
5 the Cycle 8 analysis, a value of 2172 psia was used for the !
j initial RCS pressure, since the normal operating RCS pres. )
sure at Fort Calhoun is 2100 psia. The Fort Calhoun pres. f suriser sprays are fully closed at 2175 psia and fully open !
at 2225 pata, i
i i
The comparison of the figures show good agreement in ene [
trends for tbd core power. core average heat flux, and RCS i pressure.
b
. I 1
OPPD.NA 8303.NP. Rev. 02 I 3
Page 91 of 129 t 1
1
6.0 TPWiS!ENT ANA1.YSIS CODE VERITICATICN (Continued) t 6.4 Su-rarv '
l The CESEC III computer cc,de var developed for the analy. sis of TSAR ;
transient and accident events for the two by.four loop Combustion l
Engineering plants. OPPD's engineering staff was trained for the !
proper uso of CESEC !!! and a close consultive relationship has i
been maintsitied with the CE development taan. Extensive man. hours i were involved in secting up the plant model, testing the input, f validating and verifying the output and quality assuring the base.
deck to ensure the applicability of the use of CZ$EC.III for steady state plant operation. Benchmarking aga bst Cycle 1 plant data for the Turbine '.leactor Trip and the Tour l' ump Los. of Cool-ant Flow was performed and excellent agreement between the predict-ed and observed re.ponses was obtainod, I Tor transients in which plant datt. were not available, comparisons vere performed between the OPPD Cycle 8 analyses of the limiting !
transients and the Cycle 6 analyses of the fuel tindors (CE and [
Dic) and, in one case, the transient analysis training manual. Sxam- i i
ple. In all cases, these benchmarking comparisons showed very j good agreement. !
t e District continues to maintain a quality assured model and pro- (
vides an update basedeck for each cycle, contained in an Opera.
l tions kupport Analysis Report (OSAR).
f i
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OPPD NA.8303 NP, Rev 02 [
Page 92 of 129 i t
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TABLE 6.3 1
. COMPARISON OF PARAMETERS INCLUDING UNCERTAINTIES USED IN THE CEA DROP ANALYSES FOR CYCLES 6 AND 8 Parameter h Cvele 6 Cvele 8 f
Initial Core Power Level MWe 102% of 1500 102% of 1500 i Core Inlet Temperature 'F 547 547 Pressuri::er Pressure psia 2053 2053 RCS Flow Rate gpm 190,000 197,000 1 ;
Moderator Temperature Coeff, 10 ap/*F 2.3 2.7 l Doppler Coeff Multiplier 1.20 1.15 CEA Insertion at Full Power n Insertion 0.0 25.0 Dropped CEA Worth top 0.34 0.28 I i
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, Page 93 of 129 i
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. TABLE 6.3 2 COMPARISON OF PARAMETERS INCLUDINC UNCERTAINTIES USED IN THE HZP MAIN STEAMLINE BREAK ANALYSIS FOR CYCLES 6 AND 8 Cycle 6 Parameter Unila Cvele 6 AFW Cvele 8 Initial Core Power Level MWe 0.0 1.0 1.0
?' Core Inlet Temperature 'F 532 '532 532 i Pressurizer Pressure psia 2053 2175 2171 i
RCS. Flow Rate gym 190,000- 190,000 197,000 Effective Moderator Temperature 10Ap/*F . 2. 3 2.3 2.5 Coefficient Doppler Coeff Multiplier 0.8 1.15 1.15 Minimum CEA Scram Worth 4 Ap 3.0 4.2 4.0
Initial Steam Generator Pressure psia N/A 900 895 ,
. Initial Steam Generator Mass t Narrow 63 63 70 l Inventory (Level) Range Scale '
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l OPPD NA 8303-NP, Rev. 02 Page 94 of 129 l
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TABLE 6.3 3 COMPARISON OF PARAMETERS INCLUDING UNCER1*AINTIES USED IN Tile HZP MAIN STEAMLINE BREAK ANALYSIS FOR CYCLES 6 AND 8 Cycle 6 Paramater Units Cvele 6 AW Cvela R Initial Core Power Level KWe 102% of 1500 102% of 1500 102% of 1500 Core Inlet Temperature 'F 547 547 547 Pressuri::er Pressure psia 2078 2175 2172 RCS Flow Rate gpm 190,000 190,000 197,000 Moderator Temperature 10ap/'F 2.3 2.3 2.5 Coefficient Doppler Coeff. Multiplier 0.8 1.15 1.15 Minimum CEA Scram Worth 4 op ?.81 5.81 6.68*
Initial Steam Generator psia N/A 880.5 890 Pressure Initial Steam Ger. ator 4 Narrov 63 63 70 Mass Inventory (' ' <a' ' Range Sca' a *
- Reduced to 6.57 to account for axial shape.
OPPD NA 8303.NP, Rev. 02 Page 95 of 129
4 I
TABLE 6.3 4 COMPARISON OF PARAMETERS INCLUDING UNCERTAINTIES USED IN THE RCS DEPRESSURIZATION ANALYSES FOR CYCLES 6 A?iD 8 EarnTeter Units Example Case
- Cvele 8 Initial Core Power Level MWt 102% of 1500 102% of 1500 Core Inlet Temperature *F 547 547 Pressuri::er Pressure psia 2300 2172 RCS Flow Rate gpm N/A 209.796 doderator Temperature 10ap/*F 2.5 -2.7 Coefficient Doppler coeff. Multiplier 1.15 1.15
- Example case input data consistent with 2700 MVt plant operating characteristics.
OPPD NA 8303 NP, Rev. 02 Page 96 of 129 s
7.0 REFERENCES
Si ction 4 References 4-1 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.
4-2 CEN 234(C) P. Louisiana Power and Light company, Waterford Unit 3 Docket 50 382, Response to Questions on CESEC, December, 1982.
43 Letter from A. E.Scherer(CE)toF.l.Miraglia(NRC), "Applica-bility of CESEC III to the Fort Calhoun Station," February 27, 1987.
44 CENPD 161 P, TORC Code, A Computer Code for Determining the Ther-mal Margin of a Reactor Core," July, 1975.
45 CEN-191(B) P, "CETOP D Code Structure and Modeling Methods for Calvert Cliffs 1 and 2 " December, 1981.
46 CENPD 162 P A, "CE Critical Heat Flux, Critical Heat Flux Correla-tion for CE Fuel Assemblies with Standard Spacer Grids Part 1 Uni-form Axial Power Discributions," September, 1976.
i 4-7 CENPD 207-P, "CE Critical Heat Flux, Critical Heat Flux Correla-tion for CE Fuel Assemblies with Standard Spacer Crids Part 2 Non-
, uniform Axial Power Distributions," June, 1978.
48 Letter'from E. C. Tourigny (NRC) to W, C. Jones (OPPD) dated March 15, 1983.
J 49 OPPD NA 8301, Rev. 03 "Reload Core Analysis Overview", April, 1988 4 10 CEN 257(0).P. "Statistical Combination of Uncertainties",
November, 1983.
Section 5 References
! 51 CEN 347(0).P. Rev. 01, "Omaha Batch M Reload Fuel Design Report",
{ January, 1987.
i 52 CEN 121(B) P, "CEAW, Method of Analyzing Sequential Control Ele-l ment Assembly. Group Withdrawal Event for Analog Protected Sys-t tems". November, 1979.
1 1 53 CENPD 199 P, Revision 1 P, "CE Setpoint Methodology", April, 1982.
l t
l i
i OPPD NA 8303 NP, Rev. 02 l Page 97 of 129 n
7.0 REFERENCES
(Continued)
Section 5 References (Continued) 5-4 Fort Calhoun SEh on Automatic Initiation of Auxiliary Feedwater, contained in the letter to W. C. Jones from Robert A. Clark, dated February 20, 1981.
55 CENPD 190 A "CE Method for Control E'.ement Assembly Ejection Anal-ysis", July 1976 56 OPPD NA 8302, Rev. 02, "Nuclear Design Methods and Verifications",
April, 1988.
57 Letter from D. M. Crutchfield (NRC) to A. E. Scherer (CE). "Safety Evaluation of Combustion Engineering ECCS Large Break Evaluation Model and Acceptance for Referencing of Related Licensing Topics Reports". July 31, 1986. l 58 Fort Calhoun SER on Ceneric Letter 86-06 (TMI Action Item II.K.3.5. '
"Automatic Trip Reactor Coolant Pumps During Loss of Coolant Acci-dent"), contained in the letter to R. L. Andrews from Anthony ~
Bornia, dated March 25, 1988.
Section 6 Rererences .
61 "CESEC Digital Simulation of a CE NSSS", Enclosure 1 P to i
LD 82 001, January 6, 1982.
62 Letter from A. E. Scherer to F. J. Miraglia, LD 87 013 dated February 27, 1981.
i i
t l
OPPD NA 8303 NP. Rev. 02 Page 98 of 129
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. IllITI AL POWER = 35.
PLAtlT DATA: TEST PERFORMED MARCH 6, 1974
.:-Pu:sLossOfFlow 0:anaPublicPowerDistrict figure usinFeeowaterflewvsiite FcrtCalhounStation-Unit!!o.1 2-7 OPPD.NA.8303.NP. Rev. 02 Page 112 of 129
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PLANT DATA: TEST PERFORMED MARCH 6, 1974
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_CorePowervsTi:e OPPD.NA 8303.NP, Rev. 02 Page 114 of 129
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Page 123 of 129
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i OmanaPublicPowerDistrict Ficure fullPowerStes:LineGreakIncicent 5-3 TotalReactivityvsTi:e FortCalhounStation-UnitNo.i OPPD NA-8303.NP. Rev. 02 Page 125 of 129
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React:r C oiant Te::eratures vs Ti:e I Fcrt Calhoun Station-Units-sA i
l D. i OPPD NA 8303 NP Rev. 02 Page 127 of 129
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Full Power Steam Line Break 1 0:anaPublicPowerDistrict Figure Inciden3 s-se i P.eetter Ccolant Te:ceratures vs Tite i Fort Calhoun Statica-Unit t!o. i OPPD NA 8303.NP, Rev. 02 l Page 128 of 129
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_ - . _ _ , _ . _ _ - _ _ _ . . _ _ _ _