ML20195F139
| ML20195F139 | |
| Person / Time | |
|---|---|
| Site: | Prairie Island  |
| Issue date: | 10/08/1998 |
| From: | Nelson H NORTHERN STATES POWER CO. |
| To: | |
| Shared Package | |
| ML20138L702 | List: |
| References | |
| NSPNAD-97002, NSPNAD-97002-R01, NSPNAD-97002-R1, NUDOCS 9811190200 | |
| Download: ML20195F139 (68) | |
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- Northern States Power Company's
- Steam Line Break o
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(Methodology-t i
- NSPNAD-97002 mO
- Re... -vision 1 0
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LOctober,1998 p
7 A[
LNorthern States Power Company NuclearAnalysis and Design NSP Non Proprietaryinformation t
9811190200 981016 PDR ADOCK 05000282 P
O Northero States rower's Steam Line Break Methodology NSPNAD-97002 Revision 1 October,1998 O
Principal Contributors Northern States Power:
H. O. Nelson Nuclear Engineering Technology Corp:
Dr. R. C. Kern
/8!9 Prepared By:
n Date:
///h f P Reviewed By: 8 e
Date:
m Approved By:
[
Date: /0 77[
t j
page 1 of 67
1 Abstract O
This report describes the Northem States Power Company reload safety analysis methodology for the core and containment responses during a Main Steam Line Break (MSLB) accident.
This report is based largely on report NSPNAD-95004, Rev.1, " Steam Line Break Core and Contairunent Integrity Analysis and Evaluation of the Removal of the Boric Acid Storage Tanks from the Safety Injection System for the Prairie Island Units". NSPNAD-95004 is a report that provides the safety analysis for a License Amendment Request (LAR) to justify eliminating the requirement that the Boric Acid Storage Tanks (BAST) be available for Safety Injection and allow the system to be realigned to use the Refueling Water Storage Tank (RWST) as the primary water source. The significant differences between the two reports is that NSPNAD-97002:
- 1. does not specify the source of borated water for the safety injection pumps,
- 2. states that the initial Steam Generator levels are higher then nominal,
- 3. Appendix A compares the new methodology to ANSI 56.4 instead of NUREG 0800,
- 4. explicitly states the acceptance criteria for a MSLB analysis, and
- 5. specifies the cycle specific parameters that must be reviewed each reload.
O Amendment No.133 to Facility operating License No. DPR-42 and Amendment No.125 to Facility Operating License No. DPR-60 for the Prairie Island Nuclear Generating Plant, Units Nos. I and 2, respective'y, revised the technical specifications for the implementation of the Voltage-Based Steam Generator Tube Repair Criteria. The NRC's safety evaluation for these amendments was based in part on an acceptable primary to secondary leakage calculations. These calculations were based in part on a differential pressure across the Steam Generator tubes of 2560 psi. Therefore Revision 1 of this report was published adding a fifth acceptance criteria related to the differential pressure across the Steam Generator tubes.
O page 2 of 67
l Legal Notice r10 l
This report was prepared by or on behalf of the Northern States Power Company (NSP). Neither NSP, nor any person acting on behalf ofNSP:
a) Makes any warranty or representation, express or implied, with respect to the accuracy, completeness, usefulness, or use of any information, apparatus, method, or process disclosed or contained in this report, or that the use of any such information, apparatus, method, or process may not infringe privately owned rights; or 1
b) Assumes any liabilities with respect to the use of, or for damages resulting from the use of, any information, apparatus, method, or process disclosed in the report.
Proprietary Data Clause AV This report is the property of the Northem States Power Company (NSP) and contains information that is proprietary to NSP. Any reproduction or copying of such information requires the express written consent of NSP. Disclosure of such information to parties outside of NSP i
requires the express written consent of NSP.
l l
l l
l i
4 O
l page 3 of 67 i
Table of Contents g
Cover Sheet 1
Abstract 2
Legal Notice and Proprietary Data Clause 3
Table of Contents 4
1.0 Introduction and Summary 6
2.0 Licensing Background 7
3.0 Code Description 8
3.1 DYNODE 8
3.2 CONTEMPT 9
3.3 VIPRE 9
0 4.0 MSLB Containment Response Methodology 9
4.1 Break Spectrum and Initial Conditions 9
4.2 Primary and Secondary Systems Modeling 11 4.3 Containment Modeling 13 4.4 Engineered Safeguards 14 5.0 MSLB Core Response Methodology 17 6.0 SSLB Core Response Methodology 18 7.0 Results of Analyses 19 7.1 Containment Response Cases 19 7.1.1 Benchmark Case 19 7.1.2 Case Spectrum 24 7.1.2.1 Offsite Power Availability 25 7.1.2.2 Break Location 25 7.1.2.3 Power Level, Break Type, and Entrainment 25 g
page 4 of 67
. -. ~ - -....
7.1.2.4 Single Failure Mode 27 t
7.2 Core Response Cases 46 7.2.1 MSLB Cases 46 i
j 7.2.2 SSLB Case 47 8.0 Reload Evaluation 55 8.1
- Cycle specific Physics Calculations 55 8.2 Reload Safety Evaluation 56 l
l-9.0 Acceptance criteria for Main Steam line Break 56
?
I l
10.0 References 58 j
Appendix A - " Comparison of the new MSLB Containment Response methodology with ANSI /ANS-56.4-1983" 60 l
l C'3'
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i l
l i
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- O page 5 of 67
1.0 Introduction andSummary This report describes the Northern States Power Company (NSP) reload safety evaluation (RSE) analysis methodology for the core and containment responses during a Steam Line Break (SLB) accident. It is NSP's intention to use this methodology for evaluation of future core reloads and plant modifications.
The results of the analyses show that the limiting MSLB containment pressure condition occurs at
[
]with a failure of[
] and [
] The peak pressure in this limiting case was less than the designed maximum containment pressure. The limiting containment temperature condition occurs at [
]with a failure of[
] and [
]. The peak temperature and temperature profile in this limiting case are used to ensure that containment equipment needed to mitigate the MSLB accident is environmentally qualified.
The analyses show that all containment design limits are met.
The results of the analyses show that the limiting MSLB minimum departure from nucleate boiling (MDNBR) for the core response case occurs at [
]with a failure of[
]and [
] The MDNBR for this case was higher than the fuel damage limit. Thus, no fuel damage was predicted and all other acceptance criteria for the MSLB core response are met.
The results for the limiting Small SLB (SSLB) case at [
]show that the core remains suberitical throughout the event, so that the acceptance criterion of no return to power for that case was also met.
It should be noted that the methods described in this report have been subjected to significant peer review by both members of the Nuclear Analysis and Department staff of Northern States Power and extemal consultants during the course of development.
Sensitivity studies were utilized to identify the most conservative value for certain parameters that impact different parts of the analysis in opposing directions. An example being the initial core flow, where a high value used in the system portion of the analysis results in a faster cooldown i
(and hence higher return to power) while a low value used in the core thermal-hydraulic model results in a lower thermal margin relative to the fuel damage limit.
2 O
page 6 of 67
i
' 2.0 Licensing Background The original licensing basis for the MSLB and SSLB events for the Prairie Island (PI) units j
consists of the analyses presented in the following sources:
PI Final Safety Analysis Report (FSAR, Reference 2) Section 14.2.5 NSP's response to IE Bulletin 80-04 and the corresponding Nuclear Regulatory Commission (NRC) Technical Evaluation Report (References 3 and 4), referred to as the May 1980 Submittal NSP's RSE Topical Report (Reference 5) e' The FSAR Section 14.2.5 contains the original analysis performed by Westinghouse for the MSLB and SSLB events. Those analyses utilized conservative and bounding assumptions for a large double ended guillotine break at the exit nozzle of the steam generator (SG) at HZP and EOC conditions in order to ensure limiting core and containment pressure responses for the MSLB case. A conservatively large break size was analyzed for the SSLB case to show that the core remained suberitical and that there was no return to power.
The NRC IE Bulletin 80-04 directed licensees to re-evaluate their MSLB analyses after taking into consideration new assumptions regarding additional water sources to the SGs. NSP responded to IE Bulletin 80-04 with the May 1980 Submittal which used codes and methods developed by the NSP Nuclear Analysis' and Department section (NAD). These methods were benchmarked to the FSAR analysis to demonstrate the capability of the methods to accurately model the event. The methods were then used to address the issues raised in IE Bulletin 80-04; specifically, the effect of high reactor coolant system (RCS) flow and auxiliary feedwater (Aux FW, or AFW) flow at run-out conditions. The NRC completed a review of the May 1980 Submittal and issued a Technical Evaluation Report concluding that the " analysis is acceptable and no further action is required of you [NSP] regarding this subject".
The RSE Topical Report (Reference 5) describes the codes and methods used by NAD in performing in-house reload safety analyses for the PI Units. For the MSLB event, the RSE Topical Report uses the same codes and methods as those in the May 1980 Submittal, and the methods and assumptions were consistent with the analysis in the FSAR Section 14.2.5. The j
NRC approved the RSE Topical Report and issued a Safety Evaluation Report (SER) (Reference i
- 6) enabling NSP to perform in-house RSEs. The NRC approved the core response methods for the MSLB event 'out did not complete the review of the containment response.
page 7 of 67
This report presents revised MSLB methodologies for both the core and containment responses.
3.0 Code Descriptions NSP uses three codes for the analysis of the MSLB core and containment responses, DYNODE-P, CONTEMPT-LT/028, and VIPRE-01. The DYNODE code is used to model the thermal hydraulic response of primary and secondary systems under transient conditions. For the MSLB, pipe break simulations are performed with DYNODE to detennine the mass and energy releases through the break as a function of time. The CONTEMPT code is used to model the containment systems and structures. For the MSLB, CONTEMPT uses the mass and energy releases from DYNODE as inputs to determine the containment pressure and temperature response. The VIPRE code is used to calculate the MDNBR during the MSLB using the DYNODE calculated core response boundary conditions that consist of the core average heat flux, the core average pressure and the core inlet flow and temperature. In addition, the VIPRE model utilizes the core power distribution based on the metrods described in Reference 5.
3.1 DYNODE O
A description of the DYNODE code may be found in the RSE Topical Report (Reference 5), the corresponding NRC SER (References 6), and in the code manual (Reference 7). Some specific code features used in DYNODE in order to evaluate the containment response for the MSLB event are:
[
]
[
]
[
]
[
]
3.2 CONTEMPT A description of the CONTEMPT code may be found in the RSE Topical Report (Reference 5),
page 8 of 67
and in the code manual (Reference 8).
- O 3.3 VIPRE A description of the VIPRE code may be found in the RSE Topical Report (Reference 5), and the code manual (Reference 13). The methods for utilizing VIPRE as part of the MSLB code package are contained in Reference 5. It should be noted that the W-3 Critical Heat Flux (CHF) correlation was used with a fuel damage limit that is pressure dependent as defined in Reference
- 5. No unusual features or options of VIPRE were used for the current analyses.
4.0 MSLB Containment Response Methodology The MSLB event results in a depressurization and cooldown of the RCS and main steam systems, and a corresponding pressurization and heatup of contamment. The mass and energy releases and containment response are dependent upon the following :
the break spectrum and initial conditions, e
primary and secondary systems, e
containment structures and systems, e
engineered safeguards systems.
e A spectrum of MSLB cases were run considering various power levels, break sizes and locations, single failures, and offsite power availability to determine the limiting conditions. -The spectrum of cases were selected to demonstrate the sensitivity on the containment pressure and temperature responses. Details of the models and assumptions used in the analyses are discussed below, i
4.1 Break Spectrum and Initial Conditions Reactor power levels of 0,30,70, and 102% ofrated were examined, ne 0% power level is referred to as HZP, and the 102% power level is referred to as HFP.
A spectrum of break locations and entrainment assumptions were examined. Different break areas were modeled, as a means of varying the discharge coefficient, to account for the uncertainties in the Moody critical flow model and to consider breaks in smaller piping connected c
to the main steam system. The following types of breaks were considered:
page 9 of 67
double ended (DE) guillotine with a break area the size of the SG exit nozzle and with liquid entrainment, double ended guillotine with a break area the size of the flow restrictor and with liquid entrainment, double ended guillotine with a break areajust large enough to allow liquid entrainment, double ended guillotine with a break areajust small enough to prevent liquid entrainment, and split break with a maximum area such that the main steam isolation valve (MSIV) e isolation results from a contamment signal and does not allow liquid entrainment.
The break flow rate was calculated using the Moody critical flow model under saturated conditions based on the enthalpy and pressure conditions at the break location.
Three different single failures of accident mitigating systems that are important for the MSLB were considered:
[
].
[
]..
[
].
The single failures were considered both with and without the availability of offsite power.
The models conservatively account for
] of the AFW pumps.
In some cases, liquid entrainment in the steam blowdown from the broken SG was modeled. The intact SG was [
]. Entrainment for the MSLB event is described in detail in the Westinghouse report WCAP-8822 (Reference 10). The WCAP conservatively calculates the break entrainment for non-preheat Model 51 Steam Generators. This analysis used the WCAP entrainment data with the uncertainties recommended by Westinghouse included. The break sizes and power levels used in the analysis were chosen to maximize the releases when entrainment was included.
Reactivity feedback coefficients for moderator temperature and density (including voiding),
Doppler, and boron worth were conservatively calculated at the limiting time in life for the MSLB page 10 of 67
i event, which was at either [
]as described in the RSE Topical Report (References 5 and 6). The minimum allowable shutdown margin (SDM) as calculated per reference 5 was assumed; however [
] in the SDM with [
] core temperatures were included in the model. The most reactive rod control cluster assembly was assumed to be stuck out of the core as either an initial condition (HZP) or during a scram for at power cases.
4.2 Primary and Secondary Systems Modeling The reactor coolant pumps (RCPs) were [
], since sensitivity studies have shown that any operator initiated RCP trips during the event were [
]. The RCPs were [
] for cases in which Loss Of Offsite Power (LOOP) was assumed. The deposition of RCP heat into the primary side coolant was included.
i Heat transfer from the primary to secondary side in the broken loop SG was modeled using the
[
] correlation at the SG tube surface to shell side water interface.
Reverse SG heat transfer from the intact loop was [
] the broken loop during the cooldown.
The Pressurizer pressure and level were [
].
The initial SG secondary side liquid inventory was [
].
For HZP cases, the initial decay heat level was calculated based on the [
] with uncertainties included and conservative assumptions of the timing and level of the decay heat following shutdown from full power. This same decay heat model was also used for at power cases.
All of the major primary and secondary side metal heat structures were [
].
Reverse steam flow from the intact SG was [
]-
page 11 of 67
The unisolated portion of the steam line volume [
].
The unisolated portions of the Main FW line were [
].
The steam generator tube height was [
1 1
The Main FW pumps were [
1 When the AFW pumps get the signal to start, it was [
).The AFW water was [
]. AFW flow to the affected SG was [
1 The Condensate pumps were [
].
The Main FW regulator valve on the broken loop was [
]. At the FW Isolation signal,
[
]. The Main FW bypass valves were [
The boron concentration of the SI water was [
] and [
]. The analysis accounted for the delay in boron transport between the cold leg injection and the reactor core.
The level of SG tube plugging was [
).
O page 12 of 67
l l
h.
[ ~ -] credit was taken for charging flow.
I' The turbine was tripped at the start of the event for HZP cases and at [
]-
l The reactor was tripped at [
L L
1 The effects of containment backpressure on the break flow [
].
The MSIVs were [
] after the appropriate setpoints and time delays were reached.
I The non-return check valves (NRCVs) were [
] in that steam line. Note that the NRCVs are passive safety related components.
which are not subject to the same single failure criterion as active components; however their failure has been considered as part of the sensitivity studies in conjunction with the assumed L
failure of the MSIVs to close.
I l.
i i
4.3 ContainmentModeling lO The containment volume, annulus, and atmospheric conditions were modeled. Concrete and metal heat structures were modeled. There was [
] between the containment and the annulus, or between the annulus and the outside atmosphere.
i The CONTEMPT default temperature flash model was used, which is appropriate for MSLBs when maximizing the containment pressure and temperature response.
l The initial containment pressure was assumed to be at the maximum allowed per Tech. Specs (Reference 9).
The containment spray and FCU systems and their actuations were modeled with conservative delay times. A conservatively [
] time was used for the time required to fill the piping of the Spray System. A conservatively [
] value was used for the capacity of the Spray Pumps. The temperature of the RWST water which was used by the Sprays was assumed to be conservatively
[
]. Conservatively [
] heat removal capability was assumed for the FCUs.
The heat structure heat transfer coefficients were calculated using the Uchida correlation as page 13 of 67 I
~
recommended in NUREG-0588 Appendix B (Reference 12).
O It was assumed that [ ] of the heat sink condensate revaporizes into the containment atmosphere.
This is consistent with the allowed value in NUREG-0588 (Reference 12).
4.4 EngineeredSafeguards The Engineered Safeguards (ES) system provides a number of signals that serve to mitigate the effects of a loss of coolant accident or a MSLB. The ES signals important for this analysis are:
SI Signal, designated 'S',
Steam Line Isolation Signal, e
Main FW Isolation Signal, designated 'FWI',
e Containment Pressure Signal, designated 'P', and e
Aux FW Pump Start Signal.
Plant process signals are generated when a safegaards setpoint is exceeded. These signals are the basic inputs that generate ES signals. They also serve as inputs to the Reactor Protection System (RPS). The process signals important to this analysis are:
Pressurizer Pressure:
- Exceeding the Low setpoint generates a 'S' signal.
Steam Generator Pressure:
- Exceeding the Low setpoint in either Main Steam Line generates a 'S' signal.
Average RCS Temperature (Tave):
- Exceeding the Low-Low setpoint in conjunction with High Steam Flow and
'S' isolates that loop's steam line.
Containment Pressure:
- Exceeding the High-1 setpoint generates a 'S' signal.
- Exceeding the High-2 setpoint isolates both steam lines.
- Exceeding the High-3 setpoint generates a 'P' signal.
- Steam Flow:
/
- Exceeding the High setpoint in conjunction with Low-Low Tave and
'S' isolates that loop's steam line.
- Exceeding the High-High setpoint in conjunction with 'S' isolates that loop's steam line.
page 14 of 67
-,_-~-.-..~.- -..
l
~
For all process signals and ES signals, conservative time delays were applied. Furthermore, delays associated with load rejection and restoration following 'S' or loss of offsite power (LOOP) were conservatively applied. For cases in which offsite power was lost, the timing of the ES signals in relation to the loading of the diesel generators was considered.
t The systems, components, and actuations relied upon for accident mitigation are listed below.
l i
Turbine Trip:
Generated by a 'S' signal. For this analysis no delays were assumed for stop valve closure. This [
] the SG inventory available for blowdown which is conservative.
Reactor Trip:
Generated by a 'S' signal. Delays were included to account for the opening time of the reactor trip breakers and control rod insertion.
]
MSIV and NRCV closure:
MSIV closure is generated by a steam line isolation signal. Delay times were included l
to account for closure times after an isolation signal is received. The NRCV [
O
] in the steam line ceases.
O SI Pump Start:
e Generated by a 'S' signal. Delays were incorporated to account for the load restoration and pump windup.
Turbine-driven AFW Pump Start-e Generated by a 'S' signal. The delay between the signal generation and actuation was included. No delay was assumed for the valve opening or line fill. This [
]
l the SG inveury available for blowdown which is conservative.
Motor-driven AFW Pump Start:
Generated by a 'S' signal. Delays were incorporated to account for the load restoration and pump windup. No delay was assumed for the valve opening or line fill. This [
] the SG inventory available for blowdown which is conservative.
Main FW Regulator Valve Closure:
Generated by a 'FWI' signal, which in tum is generated by the 'S' signal. [
i
] were explicitly modeled.
Main FW Bypass Valve Closure:
Generated by a 'FWl' signal, which in tum is generated by the 'S' signal. [
$. A
] were explicitly modeled.
4 U
page 15 of 67
Main FW Pump Trip.
Generated by a 'S' signal. A delay time was included and the pump coastdown was modeled.
Main FW / Containment Isolation Valve Closure:
Generated by a 'S' signal. Delays were incorporated to account for the load restoration, restoration of the 480V bus voltage, and valve closure time.
Containment FCUs Start:
Generated by a 'S' signal. Delays were incorporated to account for the load restoration and motor windup.
CS Pumps Start:
Generated by a 'P' signal. Delays were incorporated to account for the load restoration, pump windup, discharge valve opening, and line fill.
'lle purpose of rejecting and sequencing components following an 'S' signal or a LOOP is to ensure that an excessive number of coraponents do not start simultaneously such that an undervoltage occurs on a safeguards bus. The sequence opens permissive windows and, depending upon the presence of actuation signals, components may or may not start during that window of opportunity. The order in which loads are restored by the sequencer is:
O Step Window Component 1
'S' + 0 see
-SI pumps
-480V Motor Operated Valves 2
'S' + 5 sec
-CS Pumps (1st) 3
'S' + 10 sec None applicable to MSLB 4
'S' + 15 see
-FCUs
-CS Pumps (2nd) 5
'S' + 20 see
-Motor-driven Aux FW Prairie Island utilizes a sequencer that starts on either an 'S' signal or on a transfer ofload to a diesel generator (in the case of a LOOP). If offsite power is available, an 'S' signal causes loads on the safeguards buses to be rejected and the restoration sequence to begin. If offsite power is lost, the diesels will start on sensed undervoltage or an 'S' signal, whichever comes first. When the diesels are supplying power to the bus the restoration sequence begins. If an 'S' signal is generated aRer diesel startup, then all loads on the safeguards buses are rejected and the restoration sequence begins again.
O page 16 of 67
l 5.0 MSLB Core Response Methodology In order to maintain a more consistent methodology between the containment and core response l_
analysis for the MSLB event, the core response analysis as described in the RSE Topical Report l
(Reference 5) has been revised as follows.
The features and options that are pan of the containment response methodology described in Section 4.0 were retained for the core response method with the following exceptions and additions which have been made to maximize the severity of the thermal margin response.
The effects of core inlet and exit temperature asymmetries were [
]. The method conservatively [
].
[
] credit was taken for reverse hea*. transfer from the intact SG to provide the [
].
Sensitivity studies were performed to determine a consistent core inlet flow that was used in both DYNODE and VIPRE.
Sensitivity studies were performed to determine the limiting initial conditions and break spectrum and single failure assumption a!ong with the assumption concerning the availability of Offsite Power.
)
The timings of Safety Systems were adjusted in the most conservative direction to maximize the return to power aspects of the event.
The initial Containment pressure and humidity were set conservatively [
] and a sensitivity study was used to establish a conservative value for the initial temperature that results in the
[
] time for actuation of the Safety Systems.
\\
[
] condensate revaporization,[
] atmosphere / pool interaction and pressure flashing were modeled to [
] containment pressurization and [
] the actuation of the Safety Systems.
The VIPRE model corresponds to a full core representation which is necessary due to the stuck page 17 of 67
rod assumption. The hottest quarter assembly was represented on a subchannel basis similar to Figure C-2 of Reference 5.
Explicit conservative modeling of the effects of[
] on the power distribution within the VIPRE model was included.
The [
] CHF correlation of Reference 14 was used to calculate the MDNBR for LOOP cases in which the core inlet mass flux was outside the applicable range that has been established for the W-3 correlation.
Since the DYNODE boron concentration model in the current version assumes perfect mixing within each of the RCS volumes at each time point, a penalty of[ ] was applied to the calculated MDNBR to account for boron transport effects. This increment will be removed in the future should DYNODE be modified to explicitly incorporate boron transport.
Analyses were performed for breaks outside of Containment, since breaks of this nature must rely on the Nuclear Steam Supply System (NSSS) signal to actuate the Safety Systems.
The water temperature of the Safety injection flow was assumed conservatively [
], since tids results in a [
] return to power.
l 6.0 SSLB Core Responsei,[ethodology l
The SSLB methodology is described in the RSE topical report (Reference 5). None of the issues raised during the development of the new Steam Line Break methodology impacted that method.
Thus, the method which has been approved by the NRC in reference 5 was used to demonstrate that there is no return to power fut the limiting SSLB case that occurs at [
] when the boron concentration of the SI water was assumed to be [
).
The following items reinforce rather than modify the Reference 5 method.
A conservatively [
] RCS flow was used to [
] the cooldown rate.
.A conservatively [
] initial RCS temperature was used to [
] the Shut Down Margin.
O page 18 of 67
A conservatively [
] initial Pressurizer pressure was used to [
] the SI time e
(
delay and [
] the SI flow due to the [
] RCS back pressure.
[
] credit was taken for SG tube plugging.
7.0 Results ofAnalyses This section presents the results of the analyses using the methods described in Sections 4.0,5.0, and 6.0. These analyses utilized conservative core physics parameters that have been established i
using the parameters for the most recent PI cores calculated per reference 5.
7.1 Containment Response Cases Using the methods for evaluating the containment response described in Section 4.0, the containment response of the MSLB was evaluated for the PI Units. First, a benchmark case was run in order to verify that the new methods being used produced consistent results with the FSAR benchmark MSLB methods that are documented in the RSE Topical Report (Reference 5). Then, a spectrum of break cases was run in order to identify the most limiting conditions for fD containment pressure and temperature.
G 7.1.1 Benchmark Case The FSAR benchmark MSLB was run at HZP EOC conditions with a double ended guillotine break at the exit nozzle of the SG, with one safeg~uards train assumed to fail, and with offsite power available. This case was originally run and documented in Reference 5. The revised DYNODE benchmark model developed using the methods described in Section 4.0 differed from the original FSAR benchmark MSLB model in the following ways:
[
1
[
].
[
].
[
1
[
].
i e
[
].
page 19 of 67
The revised CONTEMPT benchmark model developed using the methods described in Section 4.0 differed from the original FSAR benchmark MSLB model in the following ways:
[
].
[
].
The results of the new benchmark case are shown in Figures 7.1-1 and 7.1-2 where they are plotted along with the results from the FSAR benchmark results from the RSE Topical Report (Reference 5). The peak containment pressure in the new benchmark case was [
] psia which is higher than the peak value of 55 8 psia from the FSAR Benchmark case. The [
] pressure response in the new benchmark was expected due to the modeling differences as described above.
Specifically, the following factors most significantly contribute to the [
] severe response in the new benchmark case:
[
].
[
1
[
].
[
1 Energy balances highlight the differences between the two cases and show how the modeling changes [
] the pressure response in the new benchmark case. An energy balance was done for both cases over the first [ ] seconds of the MSLB event, which is near the time of the peak pressure for both cases. The energy addition for the FSAR benchmark case is:
AUr,,, = {
] Btu The value of AU,,,, includes an entrainment multiplier of 0.85 which was the manner in which entrainment was accounted for in the original FSAR analysis. The total energy addition over the same [
] second time frame for the new benchmark case is:
AU_ = [
] Btu The energy addition difference can be explained by the four factors described above. [
]
g page 20 of 67
~.
available energy contributions in the new benchmark case are:
- ' C
[
]
AU. = [
] Btu
[
]
AUn, = [
]h
[
]
AU,,,,=[
] Btu
[
]
AU,,=[
] Btu total x 0.85 AU, = [
] Btu
\\
i lt can be seen that the difference of[
] Btu between the FSAR benchmark and the new benchmark is largely explained by these four factors. Note that the balances are not exact due to the fact that not all of the [
] available energy will have blown down by the time that the broken loop SG dries out.
An additional confirmation of the consistency of the two benchmark cases can be obtained from taking the ratio of the change in contamment pressures from the initial and the [
] second time points:
R, = change in pressure from new benchmark / change in pressure from FSAR benchmark O
=([
]- 15 )/( 55.8 - 15 ) = [
]
V This can be compared to the ratio of the total blowdown energy at the 140 second time point from the two cases:
R, = energy from new benchmark / energy from FSAR benchmark
=[
]/[
]= [
]
This was expected since Figure 14.3-24 from the FSAR shows that the change in containment pressure is linear with the total steam and air internal energy.
These results demonstrate that the methods described in Section 5.0 are conservative in comparison to the FSAR benchmark results, and will result in an analysis of the MSLB event that will conservatively calculate containment pressure and temperature response.
i page 21 of 67 4
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i 7.1.2 Case Spectrum O
As detailed in Section 4.0, a spectrum of MSLB cases were run. The following parameters were varied in order to generate this spectrum of cases:
offsite power availability, break location, e
power level, e
break type and entrainment, e
single failure mode, e
initial RCS temperature, and e
initial RCS flow.
e The analyses consider all of the combinations of the parameters listed above since the severity of the event is dependent upon a combination of all of the parameters taken together. It is possible to narrow the spectrum for determining the limiting containment pressure and temperature case by evaluating these parameters separately, as discussed below.
The results of the analyses show that the limiting containment pressure condition occurs at [
h
] with a failure of[
], and [
]. A [
] initial RCS temperature and [ ] initial RCS flow was found to be the most limiting set. The fact that the [
] temperature was more limiting was a consequence of the conservative physics parameters that tend to prodiice more severe results when the blowdown is extended over a longer period of time resulting from the [
] initial secondary side inventory.
The initial temperature assumption will need to be revisited if a different set of physics parameters are used in the future. The peak pressure in this limiting case was [
] psia at [
]
seconds into the event. The results of the limiting pressure case are shown in Figures 7.1-3 to 7.1-11. The limiting containment temperature condition occurs at [
]
with a failure of[
] and corresponds to [
]. The peak temperature in this limiting case was [
] F at [ ] seconds into the event. The results of the limiting temperature case are shown in Figures 7.1-12 to 7.1-20. The analyses show that all containment design limits are met.
O page 24 of 67
4 7.1.2.1 Offsite Power Availability Cases were run both with and without the availability of offsite power. The results showed that the [
] of offsite power resulted in a more severe containment pressure and temperature i
response. When offsite power was lost the RCPs were tripped and coasted down, this [
] the l
primary side cooldown and [
] the energy (and energy rate) that was transferred to the SGs.
4 In addition, the LOOP tripped the Main FW pumps, and delayed the start of the Motor Driven l
AFW pumps due to the load rejection and restoration sequencing, which resulted in a [
]in the amount of mass deposited in the SGs that was available for blowdown. The safeguard system actuations were also delayed due to the load rejection and restoration sequencing, but this effect
[
] the containment pressure and temperature [
] then the [
] ofprimary to secondary heat transfer and [
] feedwater additions [
] the containment pressure and temperature.
7.1.2.2 Break Location Two different break locations were analyzed, one at the SG exit nozzle and another upstream of the MSIV. The break at the exit nozzle of the SG has a large area of[
] ft. The break 2
2 upstream of the MSIV has an area of[ ] ft, which corresponds to the cross sectional area of the flow restrictor. The results showed that the [
] was the most severe for containment pressure and temperature response. The [
]
results in a very [
]. This does not allow [
]. In addition, the [
].
7.1.2.3 Power Level, Break Type, and Entrainment The relationship between the power level, the break type, and liquid entrainment is summarized below. The break area and the presence ofliquid entrainment are given as a function of power level and break type.
i A break that has an area that is the size of the flow restrictor is designated as a 'DE' type break.
A break with an area that is just large enough to allow liquid entrainment is designated as a 'DS' type break. A break with an area that isjust small enough to prevent liquid entrainment from occurring is designated as a 'DX' type break. A split break is one in which the steam line on L
page 25 of 67
either side of the split remains in thermal hydraulic communication and is sufficiently small such that there is no liquid entrainment, and is designated as a 'SP' type break.
h 2
Break Area (ft )
l Power Level (%)
Break Type 0
30 70 102 DE 1.4 1.4 1.4 1.4 DS 0.2 0.5 0.6 0.7 DX 0.1
- 0.4
- 0.5
- 0.6
- SP 0.2
- 0.472
- 0.454
- 0.43 *
- denotes no liquid entrainment Total SG liquid inventory is at its highest at the HZP condition, the inventory decreases as power level increases, and is at its lowest at the HFP condition. The inventory also increases as the secondary side temperature decreases. The total energy in the system is also at its highest at HZP due to the higher SG liquid inventory.
As power level increases, the minimum break area at which entrainment occurs [
]. This is due to the decrease in SG liquid inventory and pressure as power level increases.
As power level increases, the amount of entrainment for a given break size [
] due to the decrease SG liquid inventory and pressure.
The [
] cases were less severe for both containment pressure and temperature than the most limiting [
]. This is because the [
] behaves like a [
] until the MSIV isolation and then it essentially becomes a [
]in which [
] has been isolated. A comparable [
] will produce [
] limiting results because the blowdown area is [
] prior to the isolation.
The limiting containment pressure condition was at [
). Studies show that the peak containment pressure for a given power level was the [
]. In addition, for a given break size, studies show that peak pressure [
] as power level [
] due to the [
] SG liquid inventory. The [
] SG liquid inventory results in [
] liquid entrainment, which is a competing effect that attempts to mitigate the severity of the event but is overshadowed by the
[
] liquid inventory.
O page 26 of 67
4 The most limiting containment temperature condition was at [
]. Studies show that the peak temperature occurs at the [
], which is the [ ]
break. This is because [
] the amount of superheat in the containment atmosphere. The break area for a [ ] type break [
] with power level due to a ['
]
in SG liquid inventory and consequently in liquid entrainment.
7.1.2.4 Single Failure Mode As discussed in Section 4.1, there were [
). The most severe single failure for the most limiting containment pressure case was the [
]. The most severe single failure for the most limiting containrnent temperature case *vas the [
], however it should be noted that the [
] was also nearly as limiting. The limiting nature of the [
] failure is a consequence of the fact that this failure has a larger impact on the [.
] resulting from the other failure modes.
O i
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7.2 Core Response Cases The methodology described in Sections 5.0 and 6.0 has been used to analyze the MSLB and SSLB core responses. The results of these analyses are described in the following sections.
7.2.1 MSLB Cases For the MSLB core response, a set of sensitivity studies similar to those described in Section 7.1 has been performed to determine the limiting set ofinitial conditions and assumptions. The results of those studies have established that the most severe core response corresponds to the similar set of conditions as those associated with the limiting containment pressure response.
These conditions are: [
]. The MDNBR for this case was [
] based on the W-3 correlation that occurs at [
] see and includes the rod bow penalty of[
], an F engineering factor of[
], and a [
] increment due to neglect of boron transport o
effects in the DYNODE model. The MDNBR was significantly higher than the fuel damage limit of 1.45 from Reference 5 since the RCS pressure is < 1000 psia. The time of occurrence was much [
] than that obtained using the Reference 5 methodology, because the inclusion of[
] within the DYNODE model has shifted the limiting break size to [
]. The
[
] VIPRE power distribution was based on a [
] which is conservative relative to the [
].
The major DYNODE results are shown in Figures 7.2-1 to 7.2-6 for the limiting case.
The following discussions summarize the results of the important sensitivity studies.
Cases at [
] initial power were [
] severe due the [
] SG inventories and [
] stored energy in the RCS water. The assumption of [
] resulted in a [
] severe core transient, because the [
] and a [
] peak heat flux.
A[
] initial core flow also resulted in a [
] cooldown relative to a [
] flow case, however the [
] mass flux resulted in a [
] MDNBR. A [
] initial core temperature had a
{
] cooldown; however this was compensated by the [
] SG inventory and [
] retum to power resulting from the [
] core physics parameters and the [
] Shut Down Margin.
The failure of[
] was more severe relative to the other single fallures due to g
page 46 of 67
the much [
). Breaks outside containment were [
]
severe, since the maximum break size for that type of break is [
] inside containment. The assumption that the operators will trip the RCPs during a MSLB resulted in a core response that was [
3 7.2.2 SSLB Case The limiting SSLB has been analyzed using the methodology described in Section 6.0 with a conservatively [
]. This case also used the conservative core physics parameters. The resulting time dependent core k,,is presented in Figure 7.2-7 which shows that the maximum value was [
] which occurs at about [ ] sec. This response differs from that shown in Figure 3.14-22 of reference 5 due to the much [
). Since the core [
] throughout tids event, the acceptance criterion in Reference 2 was met.
J
' (/
i page 47 of 67
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l page 49 of 67
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8.0 Reload Evaluation
)
This section describes the physics parameters that are calculated each reload for comparison to the bounding values used in the safety analysis.
8.1 Cycle Specific Physics Calculations The cycle specific physics parameter calculations are performed at the core conditions that will generate the most limiting contaimnent and core response. Sensitivity studies are conducted to determine these limiting conditions accounting for the effects of control rods, xenon, power level, temperature, etc. for each parameter.
a.) Moderator Temperature Coefficient, au Calculations of a are performed in accordance with the general procedures described in u
reference 5. Cycle specific calculations are performed at [
] in order to obtain the most negative coefficient. Additionally, calculations are made with all rods in except for the most reactive RCCA stuck out at 1000 psia as a function of core average temperature. Using this functional value of a (T),
u K,y versus temperature is calculated consistent with the calculated [
] for the initial temperature. Model biases are included in all calculations.
O b.) Doppler Temperature Coefficient, ao Calculations of n are performed in accordance with the general procedures described in o
reference 5.
c.) Boron Reactivity Coefficient, c.
Calculations of a are performed in accordance with the general procedures described in s
reference 5.
d.) Shutdown Margin,SDM Calculations of SDM are perfonned in accordance with the general procedures described in reference 5.
e.) Nuclear Enthalpy rise Hot Channel Factor, Fa, The nuclear enthalpy rise hot channel factor Fa is calculated consistent with the general procedures in reference 5 accounting for the assumed stuck rod, [
] effects. The F is conservatively adjusted for allowed 3
quadrant tilt (T).
1 l
page 55 of 67
I i
f.) Fuel Pin Census Calculation of the number of pins (pin census) versus Fui s performed in accordance with i
the general procedures described in reference 5. The calculations determine the number of fuel pins above the Foi value at which the MDNBR exceeds the fuel damage limit.
8.2 Reload Safety Evaluation Each parameter calculated above is conservatively adjusted to include the model reliability factors RF, and biases, B,. These results are then compared to the bounding values assumed in the safety i
analysis. For K,versus temperature during cooldown, the reliability factors are applied to the calculation of the moderator temperature coefficient prior to the determination of Km.
Uncertainties applied to the shutdown margin (SDM) include reliability factors for the rod worth, moderator temperature defect, and Doppler temperature defect as described in reference 5. The cycle specific parameters are acceptable if the following inequalities are met:
l CYCLE SPECIFIC PARAMETER SAFETY ANALYSIS PARAMETER l
a.
K,(Tu)
[]
K,(Tu) b.
a *(1-RFo)
[]
a (least negative bounding value) o o
l c.
a+Mda
[]
a (least negative bounding value) s a
s d.
[]
SDM (bounding) e.
(F,+RFa,Gan)N16)
{]
F s (bounding) 3 3
f.
- fuel pins above the Fui value at
[]
- of fuel pins assumed to be failed in which the MDNBR exceeds the fuel dose calculations damage limit 9.0 Acctplance criteriafor the main Steam line Break The acceptance criteria for the main steam line break are as follows:
1 ) The maximum reactor coolant and main steam system pressures must not exceed 110%
of the design values when RCS temperature is above the minimum pressurization temperature listed in reference 15.
O page 56 of 67
The minimum core inlet temperature for the limiting case in section 7 remain above [
]*F which is higher then the current minimum pressurization temperature of 310 F listed in the Updated Safety Analysis Report. herefore the maximum reactor coolant and main steam system pressures must not exceed 110%
of the design values or 2750 psia and 1210 psia respectfully. The maximum pressures from the above cases were [
] psia for the RCS and [
] psia for the main steam system.
2.) The containment vessel intemal pressure must not exceed the designed maximum listed in reference 15.
l Currently the value listed in the Updated Safety analysis report is 46 psig (or 60.7 psia). The maximum containment pressure from the analysis in section 7 was [ ]
j psia. Therefore the analysis showed that this acceptance criteria was satisfied.
3.) The number of fuel rods calculated to experience a MDNBR ofless than 1.3 (W-3, when RCS pressure is > 1000 psia) or 1.45 (W-3, when RCS pressure is 2 500 psia but 5; 1000 psia) shall not exceed the number of failed rods assumed in the radiological dose calculations. These dose calculations shall not exceed relevant requirements of 10 CRF
- Part 100.
i i
l The core response cases in section'7 showed that no fuel experience a MDNBR of i
less than 1.45. Thus there were no fuel failures which assures that the dose l
calculations remain bounded. Serefore the analysis showed that this acceptance criteria was satisfied.
j 4.) The reactor core must remain in place and intact with no loss of core cooling capability.
The core response cases in secticn 7 showed that no fuel experience a MDNBR of less than 1.45. Thus there were no fuel failures which assures that the core remains coolable. Therefore the analysis showed that this acceptance criteria was satisfled.
)
O page 57 of 67 i
5.) The differential pressure across the Steam Generator tubes must remain less than or equal to 2560 psi.
The maximum RCS pressure from the above cases was [
] psia. Therefore the differential pressure across the Steam Generator tubes remained less than 2560 psi.
10.0 References Reference 1 Not used.
Reference 2 Prairie Island Final Safety Analysis Report.
Reference 3
" Main Steam Line Break Safety Analysis", Northem States Power report, May 8, 1980.
Reference 4
" Main Steam Line Break with Continued Feedwater Addition", NRC Safety Evaluation Report for Prairie Island, October 25,1982.
"PWR Main Steam Line Break with Continued Feedwater Addition", Technical Evaluation Report by the Franklin Research Center, September 21,1982.
Reference 5
" Prairie Island Nuclear Power Plant Reload Safety Evaluation Methods for Application to PI Units", NSPNAD-8102-A.
Reference 6
" Safety Evaluation by the Office of Nuclear Reactor Regulation of the Reactor Physics and Reload Safety Evaluation Methods Technical Reports NSPNAD-8101P and -8102P", February 17,1983.
l l
Reference 7 NSP DYNODE-P Code and Manual, Version 94305, Revision 0, January 1995.
Reference 8 NSP CONTEMPT-LT/028 Code and Manual, Version 94271, September 1994.
i Reference 9 Prairie Island Nuclear Generating Plant Technical Specifications.
O page 58 of 67
l i
Reference 10 Westinghouse Report WCAP-8822," Mass and Energy Releases Following a Steam Line Rupture", September 1976.
Westinghouse Report WCAP-8822-SI-P-A," Supplement 1 - Calculations of
)
Steam Superheat in Mass / Energy Releases Following a Steam Line Rupture",
l September 1986.
l Westinghouse Report WCAP-8822-S2-P-A, " Supplement 2 - Impact of Steam Superheat in Mass / Energy Releases Following a Steam Line Rupture for Dry and Subutmospheric Containment Designs", September 1986.
Reference 11 American Nuclear Society Proposed Standard, ANS 5.1 " Decay Energy Release Rates Following Shutdown of Uranium-Fueled Thermal Reactors," October (1971), Revised October (1973).
i Reference 12 NUREG-0588, Appendix B, Revision 1,"Model for Environmental Qualification O) for Loss of Coolant Accident and Main Steam Line Break Inside PWR and BWR t
1 %.
Dry Type of Containment".
Reference 13 NSP VIPRE Code and Manual, Version 95325, Revision 0, Mmh 1996.
Reference 14 E. Janssen and S Levy, " Burnout Limit Curves for Boiling Water Reactors" APED-3892, General Electric Company (1962).
Reference 15 Prairie Island Updated Safety Analysis Report.
l j
I i..
- O page 59 of 67 i
4-
i Appendix A Comparison of the new MSLB Containment Response methodology with ANSI /ANS-56.4-1983 American National Standards Institute / American Nuclear Society standard ANSI /ANS-56.4-1983 " pressure and temperature transient analysis for light water reactor containments" provides criteria and guidance for the analysis for MSLB accidents. Although the Prairie Island Units are not required to conform to this standard, the following is a discussion of how the new methodology meets each applicable subsection of the standard. This discussion is included in an attempt to provide additional supporting documentation to facilitate the review and acceptance by the NRC.
I O
9 i
page 60 of 67
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i 1'
i I
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