ML20096C393
| ML20096C393 | |
| Person / Time | |
|---|---|
| Site: | Monticello  |
| Issue date: | 10/16/1995 |
| From: | Bonneau C, Dean D, Matis L NORTHERN STATES POWER CO. |
| To: | |
| Shared Package | |
| ML20096C390 | List: |
| References | |
| NSPNAD-8608-A, NSPNAD-8608-A-R04, NSPNAD-8608-A-R4, NUDOCS 9601170279 | |
| Download: ML20096C393 (222) | |
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OCT 9 1995 pnuru "j t UNITED STATES NUCLEAR REGULATORY COMMISSION (%
'! WASHINGTON. D.C. 205W0001 M
- i # September 29, 1995 Mr. Roger 0. Anderson, Director Licensing and Management Issues Northern States Power Company 4I4 Nicollet Hall Minneapolis, Minnesota 55401
SUBJECT:
MONTICELLO NUCLEAR GENERATING PLANT - REVIEW 0F TOPICAL REPORT NSPNAD-8608, REVISION 3, " RELOAD SAFETY EVALUATION METHODS FOR APPLICATION TO THE MONTICELLO NUCLEAR GENERATING PLANT" (TAC NO. M93388)
Dear Mr. Anderson:
By letter dated February 2, 1995, Northern States Power Company submitted, for the staff review and approval, a Revision 3 of the Topical Report NSPNAD-8608,
" Reload Safety Evaluation Methods for Application to the Monticello Nuclear Generating Plant." This submittal describes the use of new methods, based on CASMO-3/ SIMULATE-3, to generate input for application to reload transient and safety evaluations for the Monticello plant.
The staff has reviewed the submittal and concluded that the safety evaluation methods using a SIMULATE-3 based on DYN0DE-B model is acceptable for use in the Monticello boiling-water reactor reload analyses. Details of our review are provided in the enclosed safety evaluation. This action closes TAC No. y M93388. If you have any questions concerning this action please call me at (301) 415-1392. Sincerely,
~
7 / . Ad e Tae Kim, Project Manager Project Directorate III-I Division of Reactor Projects - III/IV Office of Nuclear Reactor Regulation Docket No. 50-263
)
Enclosure:
Safety Evaluation cc w/ encl: See next page
)
- ~ _ . _ _ . . . . _ . . - ._ _ . . _ . _ _ . __ _. _ _ _ . . - _ . _ _ _ _ . . _ _ _ . _ _ _ . . 'O!
Mr. Roger 0. Anderson, Director Monticello Nuclear Generating Plant Northern States Power Company O cc: J. E. Silberg, Esquire Adonis A. Neblett Shaw, Pittman, Potts and Trowbridge Assistant Attorney General 2300 N Street, N. W. Office of the Attorney General Washington DC 20037 445 Minnesota Street e Suite 900 8: U.S. Nuclear Regulatory Commission St. Paul, Minnesota 55101-2127 Resident Inspector's Office 2807 W. County Road 75 Monticello, Minnesota 55362 Plant Manager Monticello Nuclear Generating Plant g ATTN: Site Licensing ; Northern States Power Company 2807 West County Road 75 Monticello, Minnesota 55362-9637 Robert Nelson, President g Minnesota Environmental Control Citizens Association (MECCA) 1051 South McKnight Road St. Paul, Minnesota 55119 Commissioner Minnesota Pollution Control Agency 9r 520 Lafayette Road St. Paul, Minnesota 55119 Regional Administrator, Region III U.S. Nuclear Regulatory Commission 801 Warrenville Road Lisle, Illinois 60532-4351 OI Commissioner of Health Minnesota Department of Health 717 Delaware Street, S. E. l Minneapolis, Minnesota 55440 Darla Groshens, Auditor / Treasurer # Wright County Government Center 10 NW Second Street Buffalo, Minnesota 55313 l 1 Kris Sanda, Commissioner Department of Public Service , 121 Seventh Place East Suite 200 $1 St. Paul, Minnesota 55101-2145 w=,y ms O
Ota uq y kl UNITED STATES a E E NUCLEAR REGULATORY COMMISSION f WASHINGTON, D.C. 20555-0001 4,*****yI SAFETY EVALUATION BY THE OFFICE OF NUCLEAR REACTOR REGULATION RELATING TO REVISION 3 0F TOPICAL REPORT NSPNAD-8608 RELOAD SAFETY EVALUATION METHODS FOR APPLICATION TO MONTICELLO FOR NORTHERN STATES POWER COMPANY MONTICEll0 NUCLEAR GENERATING PLANT DOCKET NO. 50-263
1.0 INTRODUCTION
By letter dated February 2, 1995 (Ref. 1), the Northern States Power Company (NSP) submitted Revision 3 of the Topical Report NSPNAD-8608, " Reload Safety Evaluation Plant," (Ref. Methods for Application 2 for NRC review. to the Monticello Nuclear Generating NSPNAD-8608-A, Rev. 2 describes the currently approv)ed reload safety evaluation (RSE) methodology for the Monticello Nuclear Generating Plant. This revision documents the qualification of the currently approved DYN00E-B model, based on newly approved CASM0-3/ SIMULATE-3 methodology (NSPNAD-8609, Rev. 2), to boiling water reactor (BWR) core RSE activities for the Monticello unit. NSP intends to use SIMULATE-3 for generation of physics input for the DYN0DE-B program in licensing applications of RSE analyses. 2.0
SUMMARY
OF THE TOPICAL REPORT Topical Report NSPNAD-8608, Revision 3, describes the NSP qualification of new SIMULATE-3 BWR. based DYN0DE-B transient analyses for application to the Monticello The qualification is addressed by comparing results of representative and limiting transients between the current nuclear data handling (NDH)-based DYN0DE-B model and the SIMULATE-based model. Specifically, the report addresses Rev. 1). The the results from the Monticello Cycle IS RSE analyses (NSPNAD-91001, consistent agreement between the previous NDH and the new SIMULATE based models presented in this topical report validates the NSP application of this revised input model for DYN00E-B analysis of the Monticello BWR unit. Section 1, providing an overview of the scope of the report, is modified to add the qualification of the SIMULATE-based DYN0DE-B model, as presented in Appendix A. Appendix A is added to the topical report to describe and qualify use of the NSP-specific SIMULATE-3 inputs for the DYN0DE-B transient model. ENCLOSURE i
'O;
3.0 TECHNICAL EVALUATION
g: The nuclear models in SIMULATE and DYN0DE-B used for scram reactivity, void reactivity, and Doppler reactivity are cross compared by using representative and limiting DYN0DE-B transients selected from Cycle 15 for comparison cf NDH-based versus SIMULATE-based analyses. The transients analyzed are slow - turbine control valve closure, turbine trip without bypass, feedwater controller failure, loss-of-feedwater heater, and main steam isolation valve e closure without position scram. 4.0
SUMMARY
AND CONCLUSIONS NSP has performed comparisons and benchmarking using the SIMULATE-3 based DYN0DE-B methodology. This effort consisted of detailed comparisons of SIMULATE and DYN0DE-B neutronics and comparisons of the current NDH-based 8 model with the SIMULATE-based model for representative and limiting Monticello transients. Based on the analyses and results presented in the topical report, the staff concludes that the SIMULATE-3/DYN0DE-B methodology, as validated by NSP, can be applied to transient BWR calculations for RSE applications as discussed in the above technical evaluation. The accuracy of this methodology has been demonstrated to be sufficient for use in licensing applications for safety and transient analysis. #
5.0 REFERENCES
- 1. Letter from R. O. Anderson (NSP) to Document Control Desk (USNRC),
regarding " Request for Approval of Revision 3 of Topical Report NSPNAD-8608, ' Reload Safety Evaluation Methods for Application to the Monticello y
- Nuclear Generating Plant'," February 2, 1995.
- 2. NSPNAD-8608, Revision 3, " Reload Safety Evaluation Methods for Application to the Monticello Nuclear Generating Plant," Northern States Power Company, January 1995. (Enclosure to Ref.1) 4 Principal Contributor: E. Kendrick e<
Date: September 29, 1995 O 1 l l
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- UNITED STATES U
i . NUCLEAR REGULATORY COMMISSION k ..... ,/ WASHINGTON, D.C. 2055M001 I SAFETY EVALUATION REPORT RELATED TO A RE0 VEST FOR APPROVAL OF REVISION 2 0F NSPNAD 8608-A RELOAD SAFETY EVALUATION REPORT l NORTHERN STATES POWER COMPANY MONTICELLO NUCLEAR GENERATING PLANT DOCKET NO. 50-263
)
1.0 INTRODUCTION
In a letter dated April 12, 1994, Northern States Power Company (NSP) requested U.S. Nuclear Regulatory Commission (NRC) review of Revision 2 to the topical report NSPNAD 8608-A, " Reload Safety Evaluation Methods for Application to the Monticello Nuclear Generating Plant." The current approved
) reload safety evaluation methods for Monticello are described in Revision 1 to NSPNAD 8608-A, in which a constant axial power shape is used to generate the minimum critical power ratio (MCPR) for transients. In proposed Revision 2, a time varying axial power shape would be used in all transients that utilize the one-dimensional kinetics option. The basic computer code for performing these analyses is the NSP BWR Nuclear Steam Supply System (NSSS) Transient Simulator Program, DYN0DE-B, which has previously been approved for NSP use by )
the NRC. 2.0 EVALUATION During BWR transients, the axial power distribution shifts toward the top of the core due to the effects of void collapse and control rod insertion
) (scram). Therefore, the use of a constant axial power shape for these transients is questionable. NSP has proposed using.a more realistic time varying axial power shape. The use of a time varying axial power shape for BWR transients has recently been approved by the NRC for Pennsylvania Power and Light Company in Susquehanna Steam Electric Station analyses. Yankee Atomic Electric Company has also found that varying the axial power shape in the hot channel analysis does affect the MCPR for Vermont Yankee Nuclear Power ) Station pressurization transients.
In response to a request for additional information from the staff dated July 14,1994, the licensee addressed the staff's questions at a meeting, which is summarized in a letter documenting the meeting dated August 9, 1994. During the meeting the licensee confirmed that comparisons of the feedwater control failure, turbine trip without bypass, generator trip without bypass, loss of auxiliary power, and turbine control valve slow closure transients for
) Monticello Cycle 16 at rated conditions resulted in a higher (more conservative) CPR using a time varying axial power shape. Therefore, the
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j - staff concurs that the use of a more realistic time varying axial power shape 8 relative to a constant shape is acceptable. 3.0 CdNCLUSIONS , Based on comparisons provided by NSP, and on results presented by other BWR licensees, the staff finds the use cf a time varying axial power shape g, acceptable for all Monticello safety evaluations. The proposed Revision 2 to NSPNAD 8608-A is, therefore, acceptable.
^
Principal Contributor: L. Kopp - Date: August 16, 1994
; ei e
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( ' e-..'. b UNITED STATES NUCLEAR REGULATORY COMMISSION f % WASHINGTO N. O. C. 20555
*, j g January 12. 1988 Docket No. 50-263 / 4,
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T <- I-Mr. D. M. Musolf, Manager - Nuclear Support Services
; q'$.e'@9j f
Northern States Power Company ' AffJf 6 e-414 Nicollet Mall M u'7 Minneapolis, Minnesota 55401 ,
Dear Mr. Musolf:
SUBJECT:
HRC SAFETY EVALUATION OF RELOAD METHODOLOGY FOR APPLICATION TO THE MONTICELLO NUCLEAR GENERATING PLANT (TAC NO. 62763) Your letters dated August 25, 1988, and October 2,1986 submitted Topical Report NSPNAD-8608, " Reload Safety Evaluation Methods for Application to the Monticello Nuclear Generating Plant", Revision 1 (August 1988) and Topical Report NSPNAD-8609, " Qualification of Reactor Physics Methods for Application to Monticello", Revision 0 (September 1986) for use in licensing calculations for Monticello. With the assistance of cur consultant, Brookhaven National Laboratory, we haveOur ccmpleted our review of the methodology described in those topical reports. Safety Evaluations relative to Report NSPNAD-8608 and Report NSPNAD-8609 are contained in Enclosures 1 and 2 to this letter, respectively. In summary, we find that the methods described in Report NSPNAD-8608, Revision 1, are acceptable for performing Monticello reload licensing calculations with the following restrictions and exceptions:
- The DYNODE-B (DNB) input mus.t be determined with the DNB/NDH code and the options and models described in Table ~ 4.1-1 and in Chapters 4 and 7 of ' Topical Report NSPNAD-8608.
- The DNB one-dimensional neutronics parameters should be conservatively adjusted by an amount AF to account for uncertainties in the collapsing of the core neutrenics from three dimensions to one dimension for each
( reload application, or included (as a bias) in the delta-critical power i
- ratio (delta-CPR) determination.
- An uncertainty allowance of 0.05 should be applied to all DNB delta-CPR calculations. (As discussed with your staff in a conference call on October 5,1988, the potential for. reduction of the magnitude of this uncertainty allowance through additional analysis and/or other n
comparisons would be considered by the NRC staff). If the uncertainties j or sensitivities used in the DNB uncertainty analysis increase, the effect of the changes should be included in the delta-CPR uncertainty 4 allowance.
- DNB stability analyses will be restricted to events with flow oscillations having frequencies below SHz to avoid difficulties in
g predicting core inlet flow oscillations. Acceptance of the methodology described in Report NSPNAD-8608 includes the DN8 calculation of transient delta-CPR and vessel pressure but . does not include the calculation of the core decay ratio. With respect to Topical Report NSPNAD-8609, we find the methodology described :. therein, and your use of that methodology, to be acceptable and have concluded that the qualification process covers an acceptable range of comparisons to e demonstrate that the methodology is capa,ble of satisfactory analysis of relevant reactor configurations and steady-state operating conditions, and of providing acceptable safety-related core parameters. However, the quoted 4.3". calculational uncertainty and resulting power distribution reliability factor is not acceptable to us since probe (TIP) measurement you have not uncertainty adequatelyisdemonstrated one-half of thethat TIP the traversing asymmetry measurement incere and sine'e sufficient justification for the deletion of the largest calculational / 9 measurement TIP comparison has not been provided. Therefore, until the , elimination of some of the TIP readings is acceptably justified, all TIP comparisons should be included in the determination of the DNB/NDH code power distribution calculation uncertainties and resulting average planar linea'r heat generation rate, the linear heat generation rate,'and the core power ratio. Should you have any questions concerning the enclosed Safety Evaluations, the 9-restrictions and exceptions discussed therein and highlighted above, and desire "further discussions with the NRC staff, please let me know. This completes our action under TAC No. 62763. ' Sincerely,
. g .
John" . Stefano , Profe Manager P je t Direct ate I -1 ivis of ctor ojects - III, IV, V
& Special Projects
Enclosures:
O! As stated l . l . l l
\ : ~ .' . :
Mr. D. M. Musolf .. Northern States Power Company Monticello Nuclear Generating Plant cc: ! Gerald Charnoff, Esquire 1 Shaw, Pittman, Potts and Trowbridge 2300 N Street, NW Washington, D. C. 20037 U. S. Nuclear Regulatory Comission Resident Inspector's Office Box 1200 Monticello, Minnesota 55367. Plant Manager , Monticello Nuclear Generating Plant Northern States Power Company Monticello, Minnesota 55367 Russell J. Hatling Minnesota Environmental Control Citi:: ens Association (MECCA) Energy Task Force 144 Melbourne Avenue, S. E. Minneapolis, Minnesota 55113 Dr. John W. Ferman Minnesota Pollution Control Agency 520 1.afayette Road . St. Paul, Minnesota 55155-3898 RegionalAdkinistrator,RegionIII U. S. Nuclear Regulatory Comission 799 Roosevelt Road f Glen Ellyn, Illinois 60137 - Comissioner,of Health Minnesota Department of Health 717 Delaware 4treet, S. E. Minneapolis, Minnesota 55440
- 0. J. Arlien, Auditor Wright County Board of Comissioners 10 NW Second Street Buffalo, Minnesota 55313
U
# D% h UNITED STATES NUCLEAR REGULATORY COMMISSION wAsHINCTcN, D. C. 20655 -*
I .,5
- f..c ...
e ENCLOSURE 1 SAFETY EVALUATION BY THE OFFICE OF NUCLEAR REACTOR REGULATION RELATING TO NORTHERN STATES POWER COMPANY , TOPICAL REPORT NSPNAD-8608, REVISION 1
" RELOAD SAFETY EVALUATION NETH005 FOR APPLICATION TO THE N0tiTICELLO NUCLEAR GENERATING PLANT" ,
DOCKET NO.,50-263 .
1.0 INTRODUCTION
By letter dated October 2,1986 (Reference 1), Northern States Power (NSP) submitted the Topical Report NSPNAD-8608, " Reload Safety Evaluation Methods for Application to the Monticello Nuclear' Generating Plant". This report describes the methods developed by Northern States Po'wer (the Licensee) for performing reload evaluations and other licensing transient analyses for the Monticello Nuclear Generating plant. I The report provides: (1) a brief description of the DYNODE-B (DNB) transient simulator (Reference 2) together with comparisons of DYNODE-B with other transient codes and with measu, red data a' n d (2) the procedures used to account for DYNODE-B calcu14tional une'ertainties. The NSP methods used to determine the DYNODE-B. physics, input and associated reliability factors are described in the NSP topical report NSPNAD-8609, " Qualification of Reactor Physics Methods for Application to Monticello". Prior NSP experience' has b'een the application of DYNODE-P (the PWR version of DYN0DE) to reload evaluations for Prairie Island, Units 3 and 2 which has been approved in Reference 3. - ) The purpose of this review is to evaluate the NSP reload methods for performing - safety analyses (including certain FSAR/ Chapter 15 transients) to determine operating limits and demonstrate acceptable margin to established safety limits for the Monticello plant. The topical report describes the methods for determining the transient reduction in thermal margin A(CPR) and transient
i s increase in pressure AP. The meth'ods for performing Nonticello st, ability g analyses including the determination of the core decay ratio are also included in NSPNAD-8f08, but these methods were considered outside the scope of this review. The topical report and reload methods are sumarized in Section 2 and the - 8 technical evaluation is presented in Section 3. The restrictions on application are given in Section 4. 2.0
SUMMARY
OF THE TOPICAL. REPORT The NSP reload methods define specific procedures for analyzing certain BWR thermal-margin /CPR, overpressurization and stability transients, and determining the reduction in margin to safety limits. The basic tool for performing the::e analyses is the NSP BWR Nuclear Steam Supply System (NSSS) Transient Simulator 9-Program; DYNODE-B. The topical report provides a sumary description of the DYNODE-B code, a description of the NSP qualification of DYN00E-B for performing the required transients, and the procedures used in the application of DYNODE-B - including: (1) a listing of the event's to be analyzed, (2) the model
^
selections, input data and initial conditions and (3) the procedures used for incorporating uncertainty allowance. The NSP DYNODE-B code description, qualification and application:are sumarized in the following. 2.1 DYNODE-B Code Descriotion e; The DYN0DE-B simulator was initially devel'oped by Nyclear Associates International (NAI) and the early development paralleled the development of the - General Electric (GE) point kinetics NSSS simulator - REDY (Reference 4). Following the GE development of ODYN (Reference 5), improved steam line O^
. hydraulics and one-dimensional kinetics were added to the initial point kinetics version of DYN0DE-B to better represent the steam line pressure wave phenomena and time dependence of the axial core power distribution observed in the Peach Bottom, Unit 2 turbine trip tests. These DYNODE-B improvements are similar in nature to the ODYN improvements but differ in the details of implementation as 8 O'
i described in the following. The DYNODE-B program provides a representation of the major components of the BWR NSSS. These include the reactor core, pressure vessel and internals, recirculation system, main steam lines and turbine, and control and safety systems. Each major component is represented by a set of time-dependent differential equations which are solved to determine the transient system response. The major features of each of these systems are. sunanarized below.
- Reactor Core Mode 1 The reactor core model includes a transient simulation of the fuel and cladding temperatures, the reactor power, and the energy distribution in an average coolant channel. In the calculation of the fuel rod temperature, the radial heat conduction equation is solved using a discrete radial nodalization within the oxide and cladding regions. The reactor power is calculated in either a point kinetics or one-dimensional axial kinetics model. The neutronics input for both of these models is determined from the NSP core simulator code-NDH (Reference 6). The one-dimensional neutronics representation is in tenns of nodal multiplication (k.o ) and migration area (M ). The nodal source equations
) are coupled to-the coolant enthalpy equations and are solved iteratively. The, channel coolant enthalpy is determined by solving an axially dependent heat -
balance equation assuming an, input core flow and inlet enthalpy. The fuel heat flux profile is as'sumed to be uniform axially when the point kinetics model is used. The local voids are calculated using a profile fit non-equilibrium flow I quality model. The local limits are determined using a hot-channel calculation which employs the GE GEXL correlation to calculate the critical power ratio (CPR). ~. . ) Pressure Yessel Fluid Region Model The pressure vessel fluid model includes the lower downcomer, lower plenum, bypass, separator, steam dome and upper downcomer regions. The conservation of mass, volume and energy equations are solved for each constant volume fluid ) region. For transients such as the loss of feedwater heating, feedwater flow l )
4 increase, and idle loop startup events, where significant temperature gradients g exist, an enthalpy distribution is calculated to provide a more detailed representation of the downcomer and recirculation lines, and jet pump diffuser. The vessel flow may be input or calculated using the conservation of momentum equations. The flow rate between the lower downcomer and lower plenum is determined with a dynamic model which includes the recirculation pump and - 8 determines the flow as a response to pump speeds. A feedwater system model is used to simulate transient feedwater flow and enthalpy through a simulated control system. 9 Main Steam System Model The main steam line model determines the main steam line flow and pressure dynamics by solving the fluid conservation equations using the method of characteristics (MOC). The relief, safety, bypass, MSIY and turbine valves are O~ represented together with a turbine model. In the turbine model, the turbine acceleration is determined from the change in steam flow and external toroue imbalance. i . Safety and Con' trol Systems Both the reactor s,afety syste'ms and control systems are modelled in DYNODE-B. The safety system includes the high pressure coolant injection system (HPCIS), the reactor core isolation cooling system, and the reactor protection system g (RPS) trip functions. The HPCIS pump is actuated (with appropriate delay) on low steam dome water level, vessel pressure or at a user-specified time and infects water into the upper downcomer. The RCICS is actuated by HPCIS, MSIV - closure or at a user-specified time, and pumps water into the downcomer. The DYNODE-B RPS trips (with user-specified delay time) on high neutron power, high O reactor vessel pressure, low steam dome water level, high steam dome water level and MSIY closure fraction. A turbine stop valve closure will also result in a l reactor trip. { l . . . . l
5 The DYN0DE-B model includes the principle BWR control systems; power level controller, tur'bine control valve and bypass controllers, pressure regulator controller, flow controller and feedwater controller. The operation of each of , these controllers is modelled including the input process signals,' controller
, setpoints, error signals and the lead / lag compensation. The controller output '
5 signals operate the associated valves or adjust the scoop tube position in the case of the M/G controller. , 2.2 DYNODE-B Oualification The NSP qualification of DYNODE-B is based on comparisons to (1) the GE REDY l Monticello FSAR analysis (Reference 7), (2) the GE Cycle-11 ODYN analysis for l Monticello (Reference 8), (3) the Peach Bottom-2 turbine trip tests (Reference
~
- 9) and (4) the Cycle-11 Monticello startup test data (Reference 10). Comparisons of the important core and system variables are presented as a function of time for available calculational and measurement benchmarks. DYNODE-B/REDY I-comparisons have been carried out' for the Monticello FSAR analysis of the i turbine trip, generator trip, MSIY closure, feedwater malfunction, pressure' regulator failure, and recirculation pump and controller events. The available Y~ DYNODE-B input data required for analyzing these FSAR eventswer'enot complete
~
and typical va1ues were used when datawerfenot available. The point kinetics and axial neutronics options were used as appropriate in analyzing these events. ,: DYNODE-B/REDY comparisons are' presented-for the core power and heat flux, i feedwater and- core inlet flow, and steam dome pressure and water level. ) . DYNODE-B/0DYN Comparisons
~ ~
DYNODE-B/0DYN benchmarking comparisons have been performed for the Cycle 11 , ) Monticello reload core for the load rejection without bypass, feedwater
., controller malfunction and MSIY closure events. The analysis of these events .
. required the DYNODE-B one-dimensional axial neutronics model. The 1-D neutronics input used in the GE ODYN analysis was not available to NSP and the corresponding DYN0DE-B input data was estimated. Comparisons are presented for
) the feedwater, core inlet and steam flow, vessel pressure, and core power and
O
, 6 -
heat flux. The core power transient plots do not show the DYNODE-B/0DYN comparison at the peak transient power. - 8 D1 NODE-B/ Peach Bottom-2 Turbine Trip Test Comparisons As a benchmarking of DYNODE-B to measurement, comparisons of DYNODE-B to the O three Peach Bottom-2 (PB-2) turbine trip : tests (at e 50%, 60%, and '70% of rated power) have been made. In these analyses the one-dimensional axial neutronics model was employed. However, since a three-dimensional PB-2 HDH neutronics model was not available and the DYNODE-B neutronics input data could
. not be determined accurately, the DYNODE-B model was adjusted so that it 0-reproduced both the measured transient peak core power and enthalpy. Transient comparisons are presented for the core power, steam dome pressure and turbine throttle pressure over the initial five seconds of the transient. MSP has performed hot-channel calculations using the GEXL correlation (Reference'11) and comparisons are presented for the transient reduction in critical power ratic (CPR) for the three tests.
l DYNODE-B/Monticello Startup Test Comparisons j d i ~ As an addition'al benchmarking of DYN00E-B to measerement, comparisons have been made to ti;e following six Cycle 1 Monticello startup tests: (1) turbine trip with bypass, (2) MSIY closure, (3) recirculation pump trip, (4) automatic flow decrease, (5) pressure regulator setpoint step and (6) feedwater controller level setpoint' step. Only the turbine trip and MSIV closure transients required 4 the one-dimensional axial neutronics model. The remaining tests are thermal-hydraulic / system transients involving relatively' mild core power , ,
! changes. Transient comparisons are presented for the vessel pressure, flows, I core power, and reactor water level for the six startup tests. g
- l. 2.3 DYNODE-B Aeolication i
The application of DYNODE-B to Monticello Reload Safety Evaluations is outlined in Chapter 4 of the topical report. The events to which DYNODE-B will be 4;
- j. . ,. . . . ,
i
.. 7 applied are listed for the thermal limits, vessel overpressure and system stability evaluations. Various DYNODE-B initial condition, input and model selections are provided including the choice (for each event) of the one-dimensional axial neutronics or point kinetics models.
The application of uncertainty factors to the kinetics input is accomplished using an additive bias together with a statistical s factor. An additional uncertainty allowance is applied to the control rod worth and void reactivity to account for the collapsing of neutronics data for the one-dimensional model. These uncertainty factors are combined statistically in the DYNODE-B uncertainty
) analysis, and an overall DYNODE-B uncertainty allowance is determined. Based on comparisons of DYNODE-B with the three Peach Bottom Turbine Trip Tests, NSP
! concludes that the DYN0DE-B predictions are always conservative and no uncertainty allowance is required for the DYNODE-B code modelling assumptions and approxitnations. 3.0
SUMMARY
OF TECHNICAL EVALUATION .- The Reload Safety Evaluation Methods described in the NSPNAD-8608 topical report consist of (1) the description of the DYNODE-B HSSS transient simulator code, ,
~
(2) the qualification of DYNODE-B for performing transient analyses and (3) the The i NSP procedures for applying DYNODE-B to Monticello Reload Safety Analyses. review of NSPNAD-8608 focusedon these three items: the DYNODE-B code and
' models, the NSP qualification and the DYNODE-B application. While the DYNODE-B l models were reviewed in detail, it is recognized that mcdeling approximations and assumptions are generally acceptable when they are accompanied by the application of appropriate uncertainty allowances and, consequently, special ~ ~
emphasis was placed on the DYNODE-B qualification and uncertainty analysis.
)l Several important technical issues were raised during the review which required This additional information and clarification from Northern States Power.
information was requested in Reference 12 and provided in the NSP response i i included in Reference 13 and in Chapter 7 of the revised topical report (Refer-i
?
O. 8 i ence17). The important technical issues raised duririg the review of g NSPNAD-8608 are discussed in the following. 3.1 DYNODE-B Models 3.1.1 Neutronics Methodolocy , 1 The DYNODE-B neutronics model i.s based on the one-dimensional time-dependent equations for the neutron source -Vl[. In the specific representation employed in DYNODE-B,the time dependence of the factordhis neglected, , however, the effect of this simplification is expected to be small for reactivity transients in which the. flux change is large. The DYNODE-B calculation assumes the radial power distribution does not change ". dur'ing the transient and accounts only for the time dependence of the axial e power distribution. Since the CPR calculation requires the instantaneous radial power distribution, NSP determined the change in the radial distribution of the LPRM si.gnals during the Peach Bottom turbine trip tests (Response to Question 23 of Reference 12). The results of this' analysis indicated that the effect of the
^
change in the radial power distribution during the turbine trip transient is negligible (f.D05AcpR/IcPR). The collapsing of'the core neutronics description from three-dimensions to one-dimension is a major part of the DYNODE-B neutronics methodology. The g. collapsing is performed by averaging the core neutronics properties over a horizontal plane for each axial elevation. Since .there is significant variation in the local neutronics properties, the collapsing'is sensitive to the flux . . weighting used to perform the averaging and the DYNODE-B 1-D neutronics param-eters are conservatively adjusted by an amount AF to account for uncertainties O in this 3-D to 1-D collapsing. The factor AF is determined by comparing three-dimensional NDH and one-dimensional DYNODE-B reactivity calculations. In the case of an overpressurization transient, NSP has shown in response to Question 24 (Reference 12) that the radial collapsing results in conservative 8 4'
e values for the void and Doppler coefficients and scram reactivity, and the AF may be taken to be zero.
- The DYNODE-B neutronics parameters are represented as a function of void ,
fraction and fuel temperature. The void fraction dependence, for example, is ' determined by performing perturbation calculations with both DYNODE-B,and the three-dimensional simulator program NDH. 'In the NSP procedure (described in the response to Question 9 of Reference 13)., the change in neutronics parameters cbserved in NDH (as a result of void changes in NDH) is correlated as a function of the void fraction calculated by DYN0DE-8. This procedure differs frc:r-standard collapsing procedures in which the change in NDH neutronics parameters would be correlated versus the void fraction change calculated by NDH. Since there are substantial differences between the DYNODE-B and NDH system modeling and the calculation of voids, it is expected this procedure will introduce a significant uncertainty into the DYNODE-B neutronics model. -
. The fuel temperature dependence of the DYNODE-B neutronics parameters is ,
calculat,ed using a similar procedure and will introduce additional uncertainty. ~ i In re,sponse to-Question 25 of Reference 13, NSP has evaluated the effect of this procedure on the calculation of &CPR/ICPR for the turbine trip without bypass transient. The NSP procedure,.was found to result in a t 10% conservative overpredictionofICPR/ICPR. 1 3.1.2 Thermal-Hydraulics . In the calculation of the core thermal-hydraulics, DYNODE-B assumes the core .. . pressure is axially uniform. Other transient codes solve the axial momentum - L squations and determine a local axial pressure. The basis for this assumption in DYNODE-B is that most of the core pressure drop occurs across the lower core : plate, and the effect of the nodal pressure variation is small compared to the effect of variations in fluid enthalpy. This assumption is expected to result . in a conservative increase in the rate of reactivity insertion and a decrease in ) axial power peaking in the case of an overpressurization transient. )
9
- 10 ,
The DYNODE-B local void fraction is detennined using a profile-fit non- g equilibrium flow quality model. While profile-fit models are gene' rally ' recomended for steady-state conditions, mechanistic void models include a detailed description of the void generation process for tracking dynamic effects, and are recomended for transient analyses (Reference 14). The uncertainty due to the use of the approximate profile-fit flow quality mode.1 has 8 been estimated by NSP and is included in the DYNODE-B uncert,ainty analysis ~ (Question 27, Reference 12). The DYNODE-B void-quality relation employs a conservative value for the concentration parameter C o =1.0 (Reference 5), and explicitly ' accounts for uncertainty in the required drift velocity in the DYNODE-B uncertainty analysis. The GEXL CPR correlation has been obtained from i GE and is implemented in DYNODE-B. In the calculation of the anergy transferred to the moderator, DYNODE-B uses conservative values for the gap heat transfer coefficient and direct moderator e: heating fraction (DMH) which are consistent with the GE ODYN values (Reference 5). The value of DMH will be revised to bound future fuel desiens if necessary. l _ 3.1.3 DYNODE-B/0DYN/REDY Differences *
*l The DYN00E-B program, to a large extent, followed the development of REDY and 1 ODYN. However, there are important modeling differences between these codes including: (1) the DYNODE-B profile-fit model vs. the REDY second order sweep model, (2) the. DYNODE-B 1971 ANS decay heat correlation vs. the REDY Stehn- g Clancy (1965) correlation, and the ODYN exponential model, (3) the DYN0DE-B l fission source neutronics equations vs. the ODYN one-group flux equations, (4) the DYN0DE-B MCC steam-line representation vs. the ODYN single phase nadal .. .
solution and (5) the DYNODE-B axially uniform core pressure model vs. the ODYN I axially dependent core pressure calculation. The effect of these differences is 9l included in the qualification code-code comparisons, f 8 4 e e
11 . 3.2 DYNODE-B Oualification , , As a result of the strong sensitivity of the DYNODE-B predictions of transient ACPR and pressure to the various modeling approximations and assumptions - described in Sections 2.1 and 3.1 and in Reference 2, a detailed and complete qualification of DYN0DE-B for application ,to Monticello reload analyses is ' required. This qualification provides the required (95% probability /95 confidence level) uncertainty allowance to be included in the DYNODE-B transient 4CPR and pressure calculations. , The qualification and uncertainty analysis of BWR transient methods is generally carried out using both (1) an overall approach in which the calculated ACPR is compared to benchmark calculations and/or measurements and (2) a components of uncertainty approach in w'hich the uncertainty of the various components is estimated and combined with appropriate sensitivities to determine the resultant CPR uncertainty (Reference 5). The DYNODE-B qualification presented in Chapters 3, 4 and 7 of NSPNAD-8608 includes both overall benchmark (code-code and code- . measurement) comparisons and a componen,t of uncertainties analysis. 3.2.1 DYN0DE-B' Uncertainty via Benchmark Comoarisons The DYNODE-B/REDY and DYNODE-B/0DYN comparisons presented in Chapter 3 demonstrate that these codes produce " comparable results," but differences in the input paradeters in these calculations are significant and prevent a cuantitative determination of the DYNODE-B uncertainty. The comparisons to the Peach Bottom turbine trip tests have been normalized to insure that DYNODE-B reproduces the measured transient peak power (TPP) and transient integrated ~ - power (TIP). Since the ACPR is determined, in large part, by the TIP, this normalization insures good agreement between the calculated and measured ACPR and prevents a quantitative determination of the DYNODE B transient ACPR (or pressure) uncertainty. en.:
O:
. l 12 ,
l l As a further test of the DYN0DE-B methods, NSP has calculated the BWR licensing O basis transient - the turbine trip without bypass (response to Question 30). , This transient is a design Ifmiting transient which has been calculated previously by both BNL and GE (Reference 15). The NSP model is based on the PB-2 model and is generally consistent with both the GE and BNL models. The licensing basis transient (LBT) response is determined by the negative' scram g! j reactivity insertion, the positive void reactivity introduced by the pressure increase, and the negative Doppler feedback. DYNODE-B predicts a more monotonic and strongly increasing pressure than both GE and BNL, which is believed to be the result of the use of the core dome' pressure rather the local pressure, and # a more rapid and stronger power transient. The calculated DYNODE-B A CPR for the LBT is therefore larger than both the BNL and GE predictions. It should be noted, however, that in this comparison the NSP model has been normalized to the Peach Bottom-2 peak and integrated power measurements and is, therefore, not a full test of the NSP reload methodology. e NSP has also made benchmark comparisons ,with the GE Monticello cycle-13 l predictions of ACPR/ICPR (in response to Question 27). These comparisons indicate a conservative ACPR/ICPR overp'ediction r by DYNODE-B of 0.012 for the turbine tr.ip transient without bypass, 0.012 for the feedwater controller 8 malfunction, and 0.003 for the feedwater heater failure transient. A comparison. to the GE calculated Monticel:lo cycle-13 maximum vessel pressure for the MSIY closure event also indicates a conservative DYN00E-B overprediction (1240 vs. 1227 psia). . g 3.2.2 DYNODE-B Uncertainty via Comeonent Estimates Estimates of the individual component (95/95) uncertainties in the neutronic model, including the void coefficient and control rod worth, are provided in 4 NSPNAD-8609 (Reference 6). The Chapter 4 calculation / measurement comparisons for the Peach Bottom turbine trip tests indicate that the DYNODE-B predictions are conservative, and NSP concludes that no additional allowance is required for uncertainties in the thermal-hydraulic and system modeling and input data. In el I l ei
) these comparisons HSP has attempted to eliminate the neutronic uncertainty and isolate the effect of the thermal-hydraulic / system modeling by adjusting the . DYNODE-B input parameters to force agreement between the calculated and measured TPP and TIP. However, since the neutronics and thermal-hydraulics are - strongly coupled and the TPP and TIP are determined by both the ) thermal-hydraulic / system and neutronic models, this DYN00E-B TPP and TIP norm-alization eliminates (to a large extent) t'he effect of th'e ' thermal-hydraulic / system modeling on the6CPR calculation. It is concluded, therefore, that these comparisons do not provide a valid determination of the j DYNODE-B transientlCPR (or pressure) uncertainty. In order to estimate the DYNODE-B accuracy in performing transient analyses, NSP has performed a detailed uncertainty analysis .(response to Question 27). The analysis included estimates for the uncertainties in the nuclear model (void, 3 Doppler and scram reactivity), thermal-hydraulic model (drift velocity and profile-fit correlation), and recirculation system and steam line models. The l i uncertainties employed were 95% probability /95% confidence level estimates based , on comparisons with measurement and/or engineering judgment. The uncertaintfes introduced by the radial collapsing of the NDH kg and M were shown to result in a conservative prediction of these parameters and no uncertainty was included. The effect of uncertainty in the scram reactivity was evaluated twice; first, assuming a uniform 10% reduction in the total bank worth ) and, second, reducing the scram reactivity insertion by 10% at the time of the power peak. The first method has a negligible effect on the transient, while the second approach results in a substantial increase inA CPR/ICPR (0.014) and 1 is recomended. The uncertainty analysis indicates that the DYNODE-B - - calculations are insensitive to changes in the recirculation loop and steam line ) modeling. NSP attributes this insensitivity to the use of the core dome , pressure rather than the local pressure, and states that this simplification is I conservative. When the (95/95) uncertainty estimates are propagated through ; DYNODE-B they result in a .024 ACPR/ICPR uncertainty allowance for the turbine
- trip without bypass transient (assuming the larger scram reactivity
) " uncertainty), and a 5 psi pressure uncertainty allowance for the MSIV closure event. J
_ _ _ _ _ - . . _ _. _ . . . ~ - _ __ _ _ _ _. .._. _ _ _ _ . _ _ _ _ _ _ 14 .- 3.3 DYN0DE-B Uncertainty Allowance *
- The NSP determination of the required DYNODE-B uncertainty allowance via component estimates (described in Question 27) assumes that (1) the effects of non-random uncertainties are always small or conservative and (2) all sources of g uncertainty are known and estimated accurately. In order to support these*
assumptions and provide confidence in the AcPR/ICPR andA P uncertainties, comparisons with benchmark calculations and/or measurements are also required. DYNODE-B calculations have been compared to both benchmark calculations of the licensing basis transient and the PB-2 benchmark measurements. However, the # DYNODE-B model used in both of these comparisons has been normalized to the PB-2 measurements and, consequently, does not allow a complete and independent ! assessment of the models and assumptions that will be employed in the Monticello ' reload analyses. Without this required independent benchmarking, in order to protect the CPR safety limit with.(95/95) confidence a oCPR/ICPR uncertainty l allowance of .050 is required. Since the estimated DYNODE-B pressure uncertainty of4P=5 psi is much smaller than the 310 psi bias included in the I determination of the ASME vessel overpressure limit (Reference 16), this , uncertainty may be neglected. ,, O 344 DYNODE-B Aeolication DYNODE-B is intended for application to the standard set of BWR events indicated j in Table 4.1-f, excluding the fuel loading error and control rod withdrawal O error which are analyzed with HDH (as indicated in response td Questions 20 and 21 of Reference 13, respectively). For ev'ents in which axial flux redistribution effects are important, the 1.'-D option of DYNODE-B will be used.- - l l In order to avoid difficulties in predicting core inlet flow esci11ations g' DYNODE-B stability analyses will be restricted to events with flow asci11ations below5Hj. The specific input and initial conditions used in t.he Monticello Reload analyses will be determined on a cycle-dependent basis. The reouired DYNODE-B kinetics , parameters including bundle power, control rod worth, void and Doppler l e
.. ; . .. . l 8 l l
15 l l 8 reactivity, delayed neutron fraction, and neutron lifetime are ' determined by NDH. NDH (95/95) reliability factors for these parameters together with an - allowance AF (when required) for HDH/ DYNODE-B collapsing uncertainty.are. included in the calculation of the DYNODE-B uncertainty allowance. The co'nservative ! , direction for the DYNODE-B component uncertainties is given in the response to Question 8 (Reference 13). The control rod scram time and uncertainty is determined using the same technique as used in preparing the ODYN input.
. The operating initial critical power ratio ICPR should be determined in three e steps:
- 1. The hCPR is determined for the transients of Table 4.1-1 using NDH or the appropriate 1-0 or point kinetics version of DYNODE-B. -
- 2. The base ICPR is determined by adding the transient dCPR and the safety limit minimum critical power ratio, SLMCPR, (typically 1.03-1.07)
ICPR y = SLMCPR + A CPR. 3 3. The operating initial critical power ratio is determined by including a . 6CPR/ICPR = 0.05 allowance for uncertainties ICPR = ICPR g (1+.05). 4.0 APPLICATION OF METHODOLOGY 3 The NSP Reload Safety Evaluation Topical Report - NSPNAD-8608, Revision 1, including the DYNODE-B models, qualification data base comparisons, and ~ supporting material provided in the Chapter-7 revision and in Reference 13 have g been reviewed in detail. Based on this review,we have concluded that the methods described in NSPNAD-8608 are acceptable for performing Monticello reload . licensing calculations with the following restrictions. .
- 1. The DYNODE-B input must be determined with the NDH code and the options and
, models described in Table 4.1-1 and in Chapters 4 and 7 of NSPNAD-8608. O
9
-- 16
- 2. The 1-D collapsing adjustment factors AF should be shown to be conservative for each reload application or included (as a bias) in the A CPR e
determination.
- 3. A 0.05 uncertainty allowance should be applied to a'll DYN0DE-B 6 CPR calculations, as described in Section 3.5. e
- 4. If the uncertainties or sensitivities used in the DYNODE-B uncertainty.
analysis increase, the effect of these changes should be included in the
- ~
A CPR uncertainty allowance. -
- 5. The application of DYNODE-8, is restricted to core inlet flow esci11ations with frequencies below 5 Hz.
- 6. This acceptance includes the DYNODE-B calculation of transient 4CPR and g.
vessel pressure but does not include the calculation of the core decay , ratio. i 5.0 CbNCLUSIONS
. O
~
The staff, with the' assistance of consultants from Brookhaven National Laboratory (BNL), has reviewed the NSP topical report NSPNAD-8608 which' presents the code description for the' DYNODE-B computer program, the code qualification benchmark analyses, and a description of the methodology that will be used to perform licensing analyses. The review has included the material prov,ided both ' in the original topical' report (Ref.'1) and the ? response to questions (Ref. 13). The acceptance applies to topical report NSPNAD-8608, Rev. 1, provided in Reference 17. The' methodology approval includes the acceptance of Topical Report NSPNAD-8609, Rev.1, " Qualification of Reactor Physics Methods for g Application to Monticello" which has been approved by'the staff with a companion Safety Evaluation which applies' to the overall reload methodology (Reference 18). Application of the' methodology to the Monticello'.Nd61 ear" Generating Plant subject to the restrictions identified in Section 4.0 of this SE. ! O lI e
. _ _ _ _ . - _ _ _ _ _ _ - . ~ __ _ _ _ _ . _ ._ _ .. _ . _ _ .__ _ _
3
6.0 REFERENCES
.1. " Submittal of NSP Reload Methodology Topical Reports," letter, David M'usolf (HSP) to Director,,0ffice of Nuclear Reactor Regulation, dated October 2, .
1986. ' ~)
- 2. "DN886098 Computer Code Users Manual,* Northern States Power Co., June i 1986. .
- 3. " Safety Evaluation by the. Office of Nuclear Reactor Regulation of the J
Reactor Physics and Reload Safety Evaluation Methods Technical Reports NSPNAD-8101P and 8102P for the Northern States Power Company," attachment to letter,. R.A. Clark (NRC) to D.M. Musolf (NSP), Docket Nos. 50-282 and 50-306, dated February 17, 1983. i 4. R.B. Linford, " Analytical Methods of Plant Transient Evaluations for the l General Electric Boiling Water Reactor," HEDO-10802, February 1973, and ) i Amendments 10802-01 and 10802-02. l - J. H 5. " Qualification of the One-Dimensional Core Transient Model for Boiling Water Reactors," NEDO-24154, October 1978. i
; 6. "Monticello duelear Gene ating Plant, Qualification of Reactor Physics j Methods for Application to Monticello," NSPNAD-8609, Rev 0, September 1986.
- 7. "Monticello Fuclear Generating Plant Final Safety Analysis Report,"
l Northern States Power Co., Docket 50-263, November 1968. .. - 3 8. " Supplemental Reload Licensing Submittal for Monticello Nuclear Generating Plant Reload 10 (Cycle 11)," 23A1673, January 1984. -
- 9. " Transient and Stability Tests at Peach Bottom Atomic Power Station Unit 2 at End of Cycle 2," EPRI NP-564, June 1968.
S
. o.
2
- 10. "Monticello Unit No.1 Startup Test Results," General Electric Topical Report, NEDE-10563.
4
- 11. ~
" General Electric BWR Thermal Analygis Basis (GETAB): Data, Correlation and Design Application," General Electric Topical Report, NEDE-10958-PA, January 1977. -
- 12. " Request For Additional Information on NSPNAD-8608 and NSPNAD-8609," Letter '
from Dino C. Scaletti (NRC) to Mr. D. Musolf (NSP), dated August 5,1987.
- 13. " Additional Information to Support the Submittal of NSPNAD-8608 and
- NSPNAD-8609," Letter, David Musolf (NSP) to Director,0ffice of Nuclear Reactor Regulation, dated September 29, 1987.
- 14. Lahey, R.T. Jr. and Moody, F.J., "The Thermal-Hydraulics of a Boiling Water Nuclear Reactor," ANS Monograph Series on Nuclear Science and Technology, 1977.
- 15. Lu, M.S., Cheng, N.S. , Shier, W.G. , Diamond, D.J. , Levine, M.M. , " Analysis of Licensing -Basis Tran51ents for a BWR/4," BNL-NUREG-26684, September 1979.
9
- 16. " Response to NRC Request for Information on ODYN Computer Model," letter, R.H. Bucholz (GE) to P.S. Check (NRC), MFN-155-80, September 5,1980.
- 17. " Revision 1 to NSP Topical Report NSPNAD-8608-Reload Safety Evaluation g Methods for Application to Monticello," Letter, D. Musolf (NSP) to Document Control Desk (USNRC), August 25, 1988.
- 18. Safety Evalutaion by the Office of Nuclear Reactor Regulation R' elating to Northern States Power Company Topical Report MSPNAD-8609 " Qualification of g Peactor Physics Methods for Application to Mont'icello.". ~ '
9
D O RELOAD SAFETY EVALUATION METHODS FOR APPLICATION TO THE MONTICELLO NUCLEAR GENERATING PLANT O NSPNAD-8608-A October 1995 Revision 4 Q Princioal Contributors Revision O Revision 1 Revision 2 Revision 3 C. S. Gantner R. O. Anderson C. A. Bonneau A. D. Bockelman C) J. K. Kapitz C. A. Bonneau K. E. Higar C. A. Bonneau Dr. R. C. Kern C. S. Gantner Dr. R. C. Kern D. W. Dean C. F. Nierode J. K. Kapitz T. E. Fieno D. A. Rautmann Dr. R. C. Kern Dr. R. C. Kern P. J. Riedel S. D. Montgomery , C. F. Hierode I J. S. Olson D. A. Rautmann j () P. J. Riedel P. M. Shah l 0 Prepared By WQ David W. Dean Date !sh/95-' l , Reviewed By [ p --_- ' - - '
. Date / / f[
ln l'"fi 'rd A'. onneau / / Approved By
- Date /O d [
Louis P Matis / I l l s
- O O
NSPNAD-8608-A Rev. 4 Page 1 of 194
. _ - . - - . . _ . . - - . . . . . . - . . . - - . . - . . . - ~ . - ~ . - - . . - - . . _ , , . - . . . .
O 2 LEGAL NOTICE ID ' This report was prepared by or on behalf of Northern States Power Company (NSP). Neither NSP, nor any person acting on behalf of NSPs
- 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 II rights; or
- 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. 1 (D l l i GD (D O l GD ! 4B - ED : NSPNAD-8608-A Rev. 4 Page 2 of 194
e . O Table of Contents Paae 1 () I
1.0 INTRODUCTION
AND
SUMMARY
. . . . . . . . . . . . . . . . . . . . . . 10 2.0 DYNODE-B CODE DESCRIPTION . . . . . . . . . . . . . . . . . . . . . 10 2.1 GENERAL DESCRIPTION . . . . . . . . . . . . . . . . . . . . 10 2.2 SPECIFIC MODEL DESCRIPTIONS . . . . . . . . . . . . . . . . . 11 2.2.1 CORE MODEL . . . . . . . . . . . . . . . . . . . . . . 11 2.2.2 REACTOR VESSEL FLUID MODEL . . . . . . . . . . . . . . 11
() 2.2.3 MAIN STEAM SYSTEM MODEL . . . . . . . . . . . . . . . 2.2.4 SAFETY SYSTEMS . . . . . . . . . . . . . . . . . . . . 12 13 2.2.5 CONTROL SYSTEMS . . . . . . . . . . . . . . . . . . . 13 2.2.6 INTEGRATION SCHEME . . . . . . . . . . . . . . . . . . 13 2.3 COMPARISONS WITH OTHER APPROVED LICENSING CODES . . . . . . . 14 2.3.1 CORE NEUTRONICS . . . . . . . . . . . . . . . . . . . 14 2.3.2 STEAM LINES . . . . . . . . . . . . . . . . . . . . . 14 2.3.3 REACTOR VESSEL PRESSURE DISTRIBUTION . . . . . . . . . 14 l C) 3.0 DYNODE-B CODE QUALIFICATION . . . . . . . . . . . . . . . . . . . . 16 i 3.1 NUCLEAR MODEL COMPARISONS . . . . . . . . . . . . . . . . . . 16 3.1.1 SCRAM REACTIVITY . . . . . . . . . . . . . . . . . . . 16 3.1.2 VOID REACTIVITY . . . . . . . . . . . . . . . . . . . 16 3.1.3 DOPPLER REACTIVITY . . . . . . . . . . . . . . . . . . 16 3.2 THERMAL - HYDRAULIC COMPARISONS . . . . . . . . . . . . . . . 17 3.2.1 CODE-CODE COMPARISONS . . . . . . . . . . . . . . . . 17 3.2.1.1 GENERAL ELECTRIC REDY CODE . . . . . . . . . . 17 () 3.2.1.1.1 TURBINE TRIP WITHOUT BYPASS . . . . . 17 3.2.1.1.2 TURBINE TRIP WITH BYPASS . . . . . . . 18 3.2.1.1.3 GENERATOR TRIP . . . . . . . . . . . . 18 3.2.1.1.4 CLOSURE OF ALL MAIN STEAM ISOLATION l VALVES . . . . . . . . . . . . . . . . . . 19 l 3.2.1.1.5 FEEDWATER CONTROLLER MALFUNCTION, MAXIMUM DEMAND . . . . . . . . . . . . . . 19 3.2.1.1.6 LOSS OF FEEDWATER . . . . . . . . . . 20 () 3.2.1.1.7 LOSS OF FEEDWATER HEATING . . . . . . 21 3.2.1.1.8 PRESSURE REGULATOR FAILS OPEN . . . . 21 3.2.1.1.9 RECIRCULATION PUMP SEIZURE . . . . . . 22 3.2.1.1.10 TWO RECIRCULATION PUMP DRIVE MOTOR TRIP . . .. . . . . . . . . . . . . . . . . 22 3.2.1.1.11 RECIRCULATION FLOW CONTROLLER FAILURE, INCREASE DEMAND . . . . . . . . . 22 3.2.1.1.12 RECIRCULATION FLOW CONTROLLER () FAILURE, DECREASE DEMAND . . . . . . . . . 23 3.2.1.1.13 IMPROPER START OF AN INACTIVE RECIRCULATION LOOP . . . . . . . . . . . . 24 3.2.1.2 GENERAL ELECTRIC ODYN CODE . . . . . . . . . . 25 3.2.1.2.1 LOAD REJECTION WITHOUT BYPASS . . . . 25 3.2.1.2.2 FEEDWATER CONTROLLER FAILURE - MAXIMUM DEMAND . . . . . . . . . . . . . . 26 3.2.1.2.3 MSIV CLOSURE (FLUX SCRAM) . . . . . . 27 C) 3.2.2 CODE-DATA COMPARISONS . . . . . . . . . . . . . . . . 27 3.2.2.1 PEACH BOTTOM 2 EOC 2 TURBINE TRIP TESTS . . . 27 3.2.2.1.1 TEST
SUMMARY
. . . . . . . . . . . . . 28 3.2.2.1.2 MODEL INPUTS . . . . . . . . . . . . . 28 3.2.2.1.3 DATA COMPARISONS . . . . . . . . . . . 28 3.2.2.2 MONTICELLO START UP TESTS . . . . . . . . . . 30 3.2.2.2.1 TURBINE TRIP WITH BYPASS AT 100%
POWER (STP 16) . . . . . . . . . . . . . . 30 3.2.2.2.2 CLOSURE OF 4/4 MAIN STEAM ISOLATION () VALVES AT 75% POWER (STP 11) . . . . . . . 31 3.2.2.2.3 2/2 RECIRCULATION PUMP TRIP (STP
- 14) . . . . . . . . . . . . . . . . . . . . 31 3.2.2.2.4 AUTOMATIC FLOW DECREASE AT 100% POWER (STP 15) . . . . . . . . . . . . . . . . . 32 3.2.2.2.5 PRESSURE REGULATOR SETPOINT STEP AT O
NSPNAD-8608-A Rev. 4 Page 3 of 194
. . - . . .. - . - - ~ - . . - ~ . . - ~ - - - . . . . . -- . -. ._.-. - . . - . - . . - , . - - ... .
O' 100% POWER (STP 18) . . . . . . . . .. . 32 3.2.2.2.6 FEEDWATER CONTROLLER LEVEL SETPOINT STEP AT'100% POWER (STP 20) . . . . . . . 33 gg 4.0 RELOAD SAFETY EVALUATION METHODS . . . . . . . . . . . . . . . . . . 123 4.1 MODEL/ EVENT APPLICATION . . . . . . . . . . . . . . . . . . . 123 4.2 INPUT PARAMETERS . . . . . . . . . . . . . . . . . . . . . . . 124 4.2.1 KINETICS PARAMETERS . . . . . . . . . . . . . . . . . 124 4.2.1.1 BUNDLE POWER . . . . . . . . . . . . . . . . . 124 4.2.1.2 CONTROL ROD WORTHS . . . . . . . . . . . . 124 4.2.1.3 VOID REACTIVITY . . . . . . . . . . . . . . . 125 4.2.1.4 DOPPLER COEFFICIENT . . . . . . . . . . . . . 125 II , 4.2.1.5 DELAYED NEUTRONS . . . . . . . . . . . . . . . 126 ' 4.2.1.6 NEUTRON SOURCE LIFETIME . . . . . . . . . . . 126 4.2.2 CRD SCRAM TIME . . . . . . . . . . . . . . . . . . . . 126 4.2.3 CRITICAL POWER RATIO . . . . . . . . . . . . . . . . . 127 4.3 LIMITING ACCEPTANCE CRITERIA . . . . . . . . . . . . . . . . . 127 4.3.1 THERMAL LIMITS . . . . . . . . . . . . . . . . . . . . 128 4.3.2 ASME VESSEL OVERPRESSURIZATION . . . . . . . . . . . . 128 4.3.3 SYSTEM STABILITY . . . . . . . . . . . . . . . . . . . 129 ID 4.4 EVALUATION AND APPLICATION OF UNCERTAINTIES . . . . . . . . . 129 4.4.1 THERMAL LIMITS . . . . . . . . . . . . . . . . . . . . 129 4.4.2-ASME VESSEL OVERPRESSURE . . . . . . . . . . . . . . . 130
5.0 CONCLUSION
S . . . . . . . . . . . . . . . . . . . . . . . . . . . . 136 ,
6.0 REFERENCES
. . . . . . . . . . . . . . . . . . . . . . . . . . . . 137 GD ;
7.0 RESPONSE TO NRC QUESTIONS . . . . . . . . . . . . . . . . . . . . . 138 I APPENDIX A QUALIFICATION OF SIMULATE-3 BASED DYNODE-B . . . . . . . . . 162 A.1 INTRODUCTION AND
SUMMARY
. . . . . . . . . . . . . . . . . . 162 A.2 NUCLEAR MODEL COMPARISONS . . . . . . . . . . . . . . . . . . 162 ,
A.2.1 SCRAM REACTIVITY . . . . . . . . . . . . . . . . . . . 162 A.2.2 VOID REACTIVITY . . . . . . . . . . . . . . . . . . . 162 ' A.2.3 DOPPLER REACTIVITY . . . . . . . . . . . . . . . . . . 163 (D ;' A.3 NDH TO SIMULATE-3 BASED DYNODE-B COMPARISONS . . . . . . . . 163 A.3.1 SLOW TURBINE CONTROL VALVE CLOSURE . . . . . . . . . . 163 A.3.2 TURBINE TRIP WITHOUT BYPASS (TTNB) . . . . . . . . . . 164 A.3.3 FEEDWATER CONTROLLER FAILURE (FWCF) . . . . . . . . . . 164 A.3.4 LOSS OF FEEDWATER HEATING (LFWH) . . . . . . . . . . . 164 A.3.5 MAIN STEAM ISOLATION VALVE (MSIV) CLOSURE WITHOUT POSITION SCRAM . . . . . . . . . . . . . . . . . . . . 164 A.4 CONCLUSIONS . . . . . . . . . . . . . . . . . . . . . . . . . 165 (D , A.5 REFERENCES . . . . . . . . . . . . . . . . . . . . . . . . . 166 9 4D i St NSPNAD-8608-A Rev. 4 Page 4 of 194
List of Figures Pace ) 2.1-1 3.1-1 Schematic of DYNODE-B BWR NSSS Representation . . . . . . . . . . NDH-DNB Comparison - Relative Power . . . . . . . . . . . . . . . 15 36 3.1-2 NDH-DNB Comparison - Relative Power . . . . . . . . . . . . . . . 36 3.1-3 NDH-DNB Comparison - Delta Rho . . . . . . . . . . . . . . . . . . 37 3.1-4 NDH-DNB Comparison - Delta Rho Error . . . . . . . . . . . . . . . 37 3.2-1 Turbine Trip without Bypass - Steam Dome Pressure . . . . . . . . 38 3.2-2 Turbine Trip without Bypass - Relative Power . . . . . . . . . . . 38 3.2-3 Turbine Trip without Bypass - Core Average Heat Flux . . . . . . . 39 i ) 3.2-4 3.2-5 Turbine Trip without Bypass - Core Inlet Flow . . . . . . . . . . Turbine Trip without Bypass - Main Steam Line Flow . . . . . . . . 39 40 3.2-6 Turbine Trip without Bypass - Feedwater Flow . . . . . . . . . . . 40 3.2-7 Turbine Trip without Bypass - Sensed Level . . . . . . . . . . . . 41 3.2-8 Turbine Trip with Bypass - Steam Dome Pressure . . . . . . . . . . 42 3.2-9 Turbine Trip with Bypass - Relative Power . . . . . . . . . . . . 42 3.2-10 Turbine Trip with Bypass - Core Average Heat Flux . . . . . . . . 43 3.2-11 Turbine Trip with Bypass - Core Inlet Flow . . . . . . . . . . . . 43 ) 3.2-12 3.2-13 Turbine Trip with Bypass - Main Steam Line Flow . . . . . . . . . Turbine Trip with Bypass - Feedwater Flow . . . . . . . . . . . . 44 44 a 3.2-14 Turbine Trip with Bypass - Sensed Level . . . . . . . . . . . . . 45 3.2-15 Generator Trip - Sensed Level . . . . . . . . . . . . . . . . . . 46 3.2-16 Generator Trip - Steam Dome Pressure . . . . . . . . . . . . . . . 46 3.2-17 Generator Trip - Core Inlet Flow . . . . . . . . . . . . . . . . . 47 l 3.2-18 Generator Trip - Core Average Heat Flux . . . . . . . . . . . . . 47 3.2-19 Generator Trip - Relative Power . . . . . . . . . . . . . . . . . 48 ) 3.2-20 3.2-21 Generator Trip - Feedwater Flow . . . . . . . . . . . . . . . . . Generator Trip - Main Steam Line Flow . . . . . . . . . . . . . . 48 49 3.2-22 Closure of All Main Steam Isolation Valves - Steam Dome Pressure . 50 3.2-23 Closure of All Main Steam Isolation Valves - Relative Power . . . 50 3.2-24 Closure of All Main Steam Isolation Valves - Core Average Heat Flux . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 51 3.2-25 Closure of All Main Steam Isolation Valves - Core Inlet Flow . . . 51 3.2-26 Closure of All Main Steam Isolation Valves - Main Steam Line Flow . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 52 ) 3.2-27 Closure of All Main Steam Isolation Valves - Feedwater Flow . . . 52 3.2-28 Closure of All Main Steam Isolation Valves - Sensed Level . . . . 53 3.2-29 Feedwater Controller Malfunction, Maximum Demand - Sensed Level . 54 3.2-30 Feedwater Controller Malfunction, Maximum Demand - Steam Dome ) Pressure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 54 1 3.2-31 Feedwater Controller Malfunction,' Maximum Demand - Core Inlet Flow . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 55 3.2-32 Feedwater Controller Malfunction, Maximum Demand - Core Average ) Heat Flux . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2-33 Feedwater Controller Malfunction, Maximum Demand - Relative 55 j 3 Power . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 56 3.2-34 Feedwater Controller Malfunction, Maximum Demand - Feedwater Flow . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 56 1 3.2-35 Feedwater Controller Malfunction, Maximum Demand - Main Steam ! Line Flow . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 57 3.2-36 Loss of Feedwater - Sensed Level . . . . . . . . . . . . . . . . . 58 I ) 3.2-37 Loss of Feedwater - Steam Dome Pressure . . . . . . . . . . . . . 3.2-38 Loss of Feedwater - Core Inlet Flow . . . . . . . . . . . . . . . 58 59 I i 3.2-39 Loss of Feedwater - Core Average Heat Flux . . . . . . . . . . . . 59 3.2-40 Loss of Feedwater - Relative Power . . . . . . . . . . . . . . . . 60 ; 3.2-41 Loss of Feedwater - Feedwater Flow . . . . . . . . . . . . . . . . 60 ; 3.2-42 Loss of Feedwater - Main Steam Line Flow . . . . . . . . . . . . . 61 3.2-43 Loss of Feedwater Heating - Sensed Level . . . . . . . . . . . . . 62 3.2-44 Loss of Feedwater Heating - Steam Dome Pressure . . . . . . . . . 62 j ) 3.2-45 Loss of Feedwater Heating - Core Inlet Flow . . . . . . . . . . . 63 3.2-46 Loss of Feedwater Heating - Core Average Heat Flux . . . . . . . . 63 3.2-47 Loss of Feedwater Heating - Relative Power . . . . . . . . . . . . 64 3.2-48 Loss of Feedwater Heating - Feedwater Flow . . . . . . . . . . . . 64 3.2-49 Loss of Feedwater Heating - Main Steam Line Flow . . . . . . . . . 65 3.2-50 Pressure Regulator Fails Open - Steam Dome Pressure . . . . . . . 66 3.2-51 Pressure Regulator Fails Open - Relative Power . . . . . . . . . . 66 D NSPNAD-8608-A Rev. 4 Page 5 of 194
() 3.2-52 Pressure Regulator Fails Open - Core Average Heat Flux . . . . . . 67 l 3.2-53 Pressure hegulator Fails Open - Core Inlet Flow . . . . . . . . . 67 1 3.2-54 Pressure Regulator Fails Open - Main Steam Line Flow . . . . . . . 68 II l 3.2-55 Pressure Regulator Fails Open - Feedwater Flow . . . . . . . . . . 68 3.2-56 Pressure Regulator Fails Open - Sensed Level . . . . . . . . . . . 69 3.2-57 Recirculation Pump Seizure - Sensed Level . . . . . . . . . . . . 70 3.2-58 Recirculation Pump Seizure - Steam Dome Pressure . . . . . . . . . 70 3.2-59 Recirculation Pump Seizure - Core Inlet Flow . . . . . . . . . . . 71 3.2-60 Recirculation Pump Seizure - Core Average Heat Flux . . . . . . . 71 3.2-61 Recirculation Pump Seizure - Relative Power . . . . . . . . . . . 72 3.2-62 Recirculation Pump Seizure - Feedwater Flow . . . . . . . . . . . 72 3.2-63 Recirculation Pump Seizure - Main Steam Line Flow . . . . . . . . 73 40' 3.2-64 Two Recirculation Pump Drive Motor Trip - Sensed Level . . . . . . 74 1 3.2-65 Two Recirculation Pump Drive Motor Trip - Steam Dome Pressure . . 74 1 3.2-66 Two Recirculation Pump Drive Motor Trip - Core Inlet Flow . . . . 75 3.2-67 Two Recirculation Pump Drive Motor Trip - Core Average Heat Flux . 75 3.2-68 Two Recirculation Pump Drive Motor Trip - Relative Power . . . . . 76 3.2-69 Two Recirculation Pump Drive Motor Trip - Feedwater Flow . . . . . 76 3.2-70 Two Recirculation Pump Drive Motor Trip - Main Steam Line Flow . . 77 3.2-71 Recirculation Flow Controller Failure, Increase Demand - Sensed 4D Level . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 78 3.2-72 Recirculation Flow Controller Failure, Increase Demand - Steam Dome Pressure . . . . . . . . . . . . . . . . . . . . . . . . . . . 78 3.2-73 Recirculation Flow Controller Failure, Increase Demand - Core Inlet Flow . . . . . . . . . . . . . . . . . . . . . . . . . . . . 79 3.2-74 Recirculation Flow Controller Failure, Increase Demand - Core Average Heat Flux . . . . . . . . . . . . . . . . . . . . . . . . . 79 3.2-75 Recirculation Flow Controller Failure, Increase Demand - Relative (D Power . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 80 3.2-76 Recirculation Flow Controller Failure, Increase Demand - Feedwater Flow . . . . . . . . . . . . . . . . . . . . . . . . . . 80 3.2-77 Recirculation Flow Controller Failure, Increase Demand - Main Steam Line Flow . . . . . . . . . . . . . . . . . . . . . . . . . . 81 3.2-78 Recirculation Flow Controller Failure, Decrease Demand - Sensed Level . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 82 3.2-79 Recirculation Flow Controller Failure, Decrease Demand - Steam () Dome Pressure . . . . . . . . . . . . . . . . . . . . . . . . . . . 82 3.2-80 Recirculation Flow Controller Failure, Decrease Demand - Core Inlet Flow . . . . . . . . . . . . . . . . . . . . . . . . . . . . 83 3.2-81 Recirculation Flow Controller Failure, Decrease Demand - Core Average Heat Flux . . . . . . . . . . . . . . . . . . . . . . . . . 83 3.2-82 Recirculation Flow Controller Failure, Decrease Demand - Relative Power . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 84 3.2-83 Recirculation Flow Controller Failure, Decrease Demand - g Feedwater Flow . . . . . . . . . . . . . . . . . . . . . . . . . . 84 3.2-84 Recirculation Flow Controller Failure, Decrease Demand - Main Steam Line Flow . . . . . . . . . . . . . . . . . . . . . . . . . . 85 3.2-85 Improper Start of an Idle Recirculation Loop - Sensed Level . . . 86 3.2-86 Improper Start of an Idle Recirculation Loop - Steam Dome Pressure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 86 3.2-87 Improper Start of an Idle Recirculation loop - Core Inlet Flow . . 87 3.2-88 Improper Start of an Idle Recirculation Loop - Core Average Heat gp Flux . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 87 3.2-89 Improper Start of an Idle Recirculation Loop - Relative Power . . 88 3.2-90 Improper Start of an Idle Recirculation Loop - Feedwater Flow . . 88 3.2-91 Improper Start of an Idle Recirculation Loop - Main Steam Line Flow . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 89 3.2-92 Load Rejection w/o Bypass - Vessel Pressure Rise . . . . . . . . . 90 3.2-93 Load Rejection w/o Bypass - Relative Power . . . . . . . . . . . . 90 3.2-94 Load Rejection w/o Bypass - Core Average Heat Flux . . . . . . . . 91 3.2-95 Load Rejection w/o Bypass - Core Inlet Flow . . . . . . . . . . . 91 ($ 3.2-96 Load Rejection w/o Bypass - Main Steam Line Flow . . . . . . . . . 92 3.2-97 Load Rejection w/o Bypass - Feedwater Flow . . . . . . . . . . . . 92 3.2-98 Load Rejection w/o Bypass - Sensed Reactor Water Level . . . . . . 93 3.2-99 Feedwater Controller Failure - Vessel Pressure Rise . . . . . . . 94 3.2-100 Feedwater Controller Failure - Relative Power . . . . . . . . . . 94 3.2-101 Feedwater Controller Failure - Core Average Heat Flux . . . . . . 95 9 NSPNAD-8608-A Rev. 4 Page 6 of 194
O 3.2-102 Feedwater Controller Failure - Core Inlet Flow . . . . . . . . . 95 3.2-103 Feedwater controller Failure - Core Inlet Subcooling . . . . . . 96 3.2-104 Feedwater Controller Failure - Main Steam Line Flow . . . . . . . 96 () 3.2-105 Feedwater Controller Failure - Sensed Reactor Water Level . . . . 97 3.2-106 MSIV Closure - Vessel Pressure Rise . . . . . . . . . , , . . . . 98 3.2-107 MSIV Closure - Relative Power . . . . . . . . . . . . . . . . . . 98 3.2-108 MSIV Closure - Core Average Heat Flux . . . . . . . . . . . . . . 99 3.2-109 MSIV Closure - Core Inlet Flow . . . . . . . . . . . . . . . . . 99 3.2-110 MSIV Closure - Main Steam Line Flow . . . . . . . . . . . . . . . 100 3.2-111 MSIV Closure - Feedwater Flow . . . . . . . . . . . . . . . . . . 100 3.2-112 MSIV Closure - Sensed Reactor Water Level . . . . . . . . . . . . 101 () 3.2-113 Peach Bottom Test TT1 - Steam Dome Pressure . . . . . . . . . . . 102 3.2-114 Peach Bottom Test TT1 - Relative Power . . . . . . . . . . . . . 102 3.2-115 Peach Bottom Test TT1 - Turbine Throttle Pressure . . . . . . . . 103 3.2-116 Peach Bottom Test TT2 - Steam Dome Pressure . . . . . . . . . . . 104 3.2-117 Peach Bottom Test TT2 - Relative Power . . . . . . . . . . . . . 104 3.2-118 Peach Bottom Test TT2 - Turbine Thrcttle Pressure . . . . . . . . 105 3.2-119 Peach Bottom Test TT3 - Steam Dome Pressure . . . . . . . . . . . 106 3.2-120 Peach Bottom Test TT3 - Relative Power . . . . . . . . . . . . . 106 3.2-121 Peach Bottom Test TT3 - Turbine Throttle Pressure . . . . . . . . 107
) 3.2-122 Turbine Trip with Bypass - Steam Dome Pressure . . . . . . . . . 108 l
j 3.2-123 Turbine Trip with Bypass - Total Vessel Flow Rate . . . . . . . . 108 1 3.2-124 Turbine Trip with Bypass - Relative Power . . . . . . . . . . . . 109 1 3.2-125 Turbine Trip with Bypass - Main Steam Line Flow . . . . . . . . . 109 I 3.2-126 Turbine Trip with Bypass - Feedwater Flow . . . . . . . . . . . . 110 ) 3.2-127 Turbine Trip with Bypass - Sensed Water Level . . . . . . . . . . 110 < 3.2-128 4/4 MSIV Closure - Steam Dome Pressure . . . . . . . . . . . . . 111 j m LJ 3.2-129 4/4 MSIV Closure - Relative Power . . . . . . . . . . . . . . . . 111 3.2-130 4/4 MSIV Closure - Steam Dome Water Level . . . . . . . . . . . . 112 I 3.2-131 4/4 MSIV Closure - Main Steam Line Flow . . . . . . . . . . . . . 112 3.2-132 4/4 MSIV Closure - Feedwater Flow . . . . . . . . . . . . . . . . 113 3.2-133 2/2 Pump Trip - Total Vessel Flow Rate . . . . . . . . . . . . . 114 i 3.2-134 2/2 Pump Trip - Core Average Heat Flux . . . . . . . . . . . . . 114 3.2-135 Automatic Flow Decrease - Steam Dome Pressure . . . . . . . . . . 115 l 3.2-136 Automatic Flow Decrease - Relative Power . . . . . . . . . . . 115 l 3.2-137 Automatic Flow Decrease - Main Steam Line Flow . . . . . . . . . 116 () 3.2-138 Automatic Flow Decrease - Feedwater Flow . . . . . . . . . . . . 116 3.2-139 Automatic Flow Decrease - Sensed Water Level . . . . . . . . . . 117 3.2-140 Automatic Flow Decrease - Total Vessel Flow Rate . . . . . . . . 117 3.2-141 Pressure Regulator Setpoint Step - Steam Dome Pressure . . . . . 118 3.2-142 Pressure Regulator Setpoint Step - Relative Power . . . . . . . . 118 3.2-143 Pressure Regulator Setpoint Step - Main Steam Line Flow . . . . . 119 3.2-144 Pressure Regulator Setpoint Step - Feedwater Flow . . . . . . . . 119 J 3.2-145 Pressure Regulator Setpoint Step - Sensed Water Level . . . . . . 120 l C) 3.2-146 Feedwater Controller Level Setpoint Step - Steam Dome Pressure . 121 3.2-147 Feedwater Controller Level Setpoint Step - Relative Power . . . . 121 3.2-148 Feedwater Controller Level Setpoint Step - Feedwater Flow . . . . 122 3.2-149 Feedwater Controller Level Setpoint Step - Sensed Water Level . . 122 7.30-1 Licensing Basis Transient - Peach Bottom TT3 . . . . . . . . . . 159 A.2-1 SIM-DNB Comparison-Relative Power 1040 psia . . . . . . . . . . . 167 A.2-2 SIM-DNB Comparison-Relative Power 1200 psia . . . . . . . . . . . 167 A.2-3 SIM-DNB Comparison-Delta Rho Scram . . . . . . . . . . . . . . . 168 () A.2-4 SIM-DNB Comparison-Delta Rho Scram Error . . . . . . . . . . . . 168 A.3.1-1 Slow TCV Closure - Steam Dome Pressure . . . . . . . . . . . . . 169 A.3.1-2 Slow TCV Closure - Relative Power . . . . . . . . . . . . . . . . 169 A.3.1-3 Slow TCV Closure - Core Average Heat Flux . . . . . . . . . . . . 170 A.3.1-4 Slow TCV Closure - Core Inlet Flow . . . . . . . . . . . . . . . 170 A.3.1-5 Slow TCV Closure - Main Steam Line Flow . . . . . . . . . . . . . 171 A.3.1-6 Slow TCV Closure - Feedwater Flow . . . . . . . . . . . . . . . . 171 A.3.1-7 Slow TCV Closure - Core Inlet Subcooling . . . . . . . . . . . . 172 A.3.1-8 Slow TCV Closure - Limiting ACPR/ICPR , . . . . . . . . . . . . . 172 () A.3.1-9 Slow TCV Closure - Sensed Reactor Water Level . . . . . . . . . . 173 A.3.2-1 Turbine Trip without Bypass - Steam Dome Pressure . . . . . . . . 174 A.3.2-2 Turbine Trip without Bypass - Relative Power . . . . . . . . . . 174 A.3.2-3 Turbine Trip without Bypass - Core Average Heat Flux . . . . . . 175 A.3.2-4 Turbine Trip without Bypass - Core Inlet Flow . . . . . . . . . . 175 A.3.2-5 Turbine Trip without Bypass - Main Steam Line Flow . . . . . . . 176 O NSPNAD-8608-A Rev. 4 Page 7 of 194
O A.3.2-6 Turbine Trip without Bypass - Feedwater Flow . . . . . . . . . . 176 A.3.2-7 Turbine Trip without Bypass - Sensed Level . .. . . . . . . . . 177 A.3.2-8 Turbine Trip without Bypass - Limiting 6CPR/ICPR . . . . . . . . 177 A.3.3-1 Feedwater Controller Failure-Sensed Reactor Water Level . . . . . 178 II A.3.3-2 Feedwater Controller Failure - Steam Dome Pressure . . . . . . . 178 A.3.3-3 Feedwater Controller Failure - Core Average Heat Flux . . . . . . 179 A.3.3-4 Feedwater Controller Failure - Core Inlet Flow . . . . . . . . . 179 , A.3.3-5 Feedwater Controller Failure - Relative Power . . . . . . . . . . 180 A.3.3-6 Feedwater Controller Failure - Feedwater Flow . . . . . . . . . . 180 A.3.3-7 Feedwater Controller Failure - Core Inlet Subcooling . . . . . . 181 A.3.3-8 Feedwater Controller Failure - Limiting 6CPR/ICPR . . . . . . . . 181
.. 3.3-9 Feedwater Controller Failure - Main Steam Line Flow . . . . . . . 182 (D A.3.4-1 Loss of Feedwater Heating - Sensed Reactor Water Level . . . . . 183 A.3.4-2 Loss of Feedwater Heating - Steam Dome Pressure . . . . . . . . . 183 A.3.4-3 Loss of Feedwater Heating - Core Average Heat Flux . . . . . . . 184 A.3.4-4 Loss of Feedwater Heating - Core Inlet Flow . . . . . . . . . . . 184 A.3.4-5 Loss of Feedwater Heating - Relative Power . . . . . . . . . . . 185 A.3.4-6 Loss of Feedwater Heating - Feedwater Flow . . . . . . . . . . . 185 A.3.4-7 Loss of Feedwater Heating - Core Inlet Subcooling . . . . . . . . 186 A.3.4-8 Loss of Feedwater Heating - Limiting 6CPR/ICPR . . . . . . . . . 186 (D A.3.4-9 Loss of Feedwater Heating - Main Steam Line Flow . . . . . . . . 187 A.3.5-1 MSIV Closure without Position Scram - Steam Dome Pressure . . . . 188 A.3.5-2 MSIV Closure without Position Scram - Relative Power . . . . . . 188 A.3.5-3 MSIV Closure without Position Scram - Core Average Heat Flux . . 189 A.3.5-4 MSIV Closure without Position Scram - Core Inlet Flow . . . . . . 189 A.3.5-5 MSIV Closure without Position Scram - Main Steam Line Flow . . . 190 A.3.5-6 MSIV Closure without Position Scram - Feedwater Flow . . . . . . 190 A.3.5-7 MSIV Closure without Position Scram - Core Inlet Subcooling . . . 191 gg A.3.5-8 MSIV Closure without Position Scram - Limiting 6CPR/ICPR . . . . 191 A.3.5-9 MSIV Closure without Position Scram - Sensed Reactor Water Level . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 192 >
d> ! 9 9 O El ~ NSPNAD-8608-A Rev. 4 Page 8 of 194
k , J l l l 1 l List of Tables l Pace l T) 3.1-1 NDH - DYNODE-B VOID REACTIVITY COMPARISON . . . . . . . . . . . . 35 1 3.2-1 PEACH BOTTOM-2 TURBINE TRIP TRANSIENT TEST ACTUAL CONDITIONS . . . 35 3.2-2 PEACH BOTTOM-2 TURBINE TRIP TRANSIENT TEST PEAK MEASURED AND CALCULATED RESPONSES . . . . . . . . . . . . . . . . . . . . . . . 35 l 3.2-3 PEACH BOTTOM-2 TURBINE TRIP TRANSIENT TEST CRITICAL POWER RATIO l RESPONSE . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 35 i SPECTRUM OF EVENTS FOR THERMAL LIMITS ACCEPTANCE CRITERIA 4.1-1 EVALUATION . . . . . . . . . . . . . . . . . . . . . . . . . . . . 131 [) 4.1-2 SPECTRUM OF EVENTS FOR ASME VESSEL OVERPRESSURE ACCEPTANCE ; CRITERIA EVALUATION . . . . . . . . . . . . . . . . . . . . . . . . 132 l 4.1-3 SPECTRUM OF EVENTS FOR SYSTEM STABILITY ACCEPTANCE CRITERIA l EVALUATION . . . . . . . . . . . . . . . . . . . . . . . . . . . . 132 l 4.1-4 DYNODE-B OPTION HODEL SELECTION FOR LICENSING APPLICATIONS . . . . 132 l 4.1-5 SENSITIVITY OF CPR TO VARIOUS THERMAL-HYDRAULIC PARAMETERS . . . . 133 4.2-1 INITIAL CONDITIONS AND INPUT PARAMETERS FOR DYNODE-B COMPARISONS i TO ODYN AND REDY RELOAD SAFETY EVALUATION OF MONTICELLO . . . . . . 133 ) [] 4.2-2 4.2-3 AXIAL POWER FACTORS FOR THE HOT CHANNEL MODEL . . . . . . . . . . STEADY STATE CRITICAL POWER RATIO COMPARISONS . . . . . . . . . . 134 134 4.3-1 FUEL CLADDING INTEGRITY LIMIT MCPR FOR MONTICELLO . . . . . . . . 134 4.4-1 COMPARISON OF MEASURED VERSUS CALCULATED TRANSIENT ACPR/ICPR . . . 134 4.4-2 COMPARISON OF MEASURED VERSUS CALCULATED MAXIMUM TRANSIENT STEAM DOME PRESSURE . . . . . . . . . . . . . . . . . . . . . . . . . . . 135 4.4-3 COMPARISON OF MEASURED VERSUS CALCULATED TRANSIENT POWER DECAY RATIO . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 135 <3 7.24.1 REACTIVITY CHANGES FOR DYNODE-B AND NDH . . . . . . . . . . . . . 148 7.27-1 INPUT PARAMETER UNCERTAINTIES USED FOR SENSITIVITY STUDIES . . . 153 7.27-2 TURBINE TRIP WITHOUT BYPASS SENSITIVITY STUDY . . . . . . . . . . 154 1 7.27-3 MSIV CLOSURE WITH HIGH FLUX SCRAM SENSITIVITY STUDY . . . . . . . 154 l 7.27-4 AUTOMATIC FLOW CONTROLLER DECREASE SENSITIVITY STUDY . . . . . . 155 7.30.1 INITIAL CONDITIONS FOR LICENSING BASIS TRANSIENT . . . . . . . . 158 l A.2-1 REACTIVITY CHANGES FOR DYNODE-B AND SIMULATE-3 . . . . . . . . . . 163 l A.3-1 TICPR FOR SIMULATE-3 AND NDH BASED DYNODE-B . . . . . . . . . . . 165 : () I I l l O O O O NSPNAD-8608-A Rev. 4 Page 9 of 194
O
1.0 INTRODUCTION
AND SUHMARY This report addresses the methods developed by Northern States Power Co. Nuclear Analysis Department (NSPNAD) to perform reload safety evaluations and gg other licensing transient analyses for the Monticello Nuclear Generating Plant. Section 2 of this report describes the DYNODE-B computer program. DYNODE-B is a transient simulator of the nuclear steam supply system (NSSS) of a boiling water reactor (BWR). This program simulates all the important features of a BWR design which significantly influence the response of the NSSS to transient conditions. The NSPNAD version of DYNODE-B includes a hot channel model which II uses the General Electric GEXL correlation to calculate critical power ratio (CPR). Section 3 describes the code qualification benchmark analysis done with DYNODE-B. This includes comparisons to other approved licensing codes (i.e., GE REDY and ODYN codes) as well as comparisons to data (i.e., Monticello startup tests and the Peach Bottom turbine trip tests). In all cases, DYNODE-B provides acceptable results. O Section 4 describes the methodology that will be used to perform licensing analyses. This includes a description of the models and input parameters used, the spectrum of events to which the methodology applies, a description of the application of uncertainties and conservatisms, a descriptien of the applicable acceptance criteria, and an evaluation of margin. Appendix A qualifies the use of DYNODE-B with SIMULATE-3 based inputs. I 9' The methodology described in this document used in conjunction with the DYNODE-B computer code provides a conservative evaluation of margins with j respect to thermal limits (CPR), ASME overpressure limits, and system i stability limits. l l 2.0 DYNODE-B CODE DESCRIPTION i 2.1 GENERAL DESCRIPTION $ The DYNODE-B computer program (6) is a transient simulator of the nuclear steam supply system (NSSS) of a boiling water reactor (BWR). This program represents all the important features of current types of BWR design which significantly influence the response of the NSSS to transient conditions. The major components of a BWR which are simulated are shown in Figure 2.1-1. Each major component is represented by a set of time-dependent differential equations. A self-initialization procedure is carried out for each of the $ component models in DYNODE-B at the beginning of each initial case. This self-initialization procedure is consistent with specified initial conditions. The major technical features of DYNODE-B are as follows:
- Provision for simulating a wide variety of transient conditions.
- Provisions for a representation of all current types of BWR gp design.
Multinode radial fuel rod and multinode axial coolant channel representations in core.
- Point kinetics, one-dimensional (axial) space-time kinetics, or power forced options for core power transients.
- Solution to conservation equations of mass, energy, volume, and GD momentum for the reactor vessel fluid and main steam system regions.
- Explicit representation of the main steam system relief, isolation, bypass, and turbine valves.
9 NSPNAD-8608-A Rev. 4 Page 10 of 194
U l D l
- Representation of the turbine.
Representation of heat transfer with the structural metal 1 J components of NSSS. l
- Representation of the reactor protective and safety injection systems.
Representation of the major control systems. Complete self-initialization.
- Full range of water properties.
2.2 SPECIFIC MODEL DESCRIPTIONS l 2.2.1 CORE MODEL I The core model consists of the neutronic and thermal-hydraulic analysis of the fuel and coolant. The average fuel rod is represented radially by a set of
] equal volume nodes in the uranium dioxide (maximum of 8) and two nodes in the cladding. The axial representation consists of a set of equal volume nodes with a maximum of 25. The heat conduction model allows temperature-dependent conductivity and heat capacity for the uranium dioxide. The gap is represented by an effective heat transfer coefficient which in a function of the average fuel temperature. The power distribution across the uranium dioxide is also modeled. The core heat generation in the moderator is dependent on the coolant void fraction. Heating of the bypass water region is 3 also represented. The surface heat transfer coefficient is based on the Thom correlation.
The conservation of mass and energy equations are solved in the coolant channel for a set of equal volume axial nodes subject to the core flow, pressure, and inlet subcooling boundary conditions which are obtained from the reactor vessel model which is described later. The core pressure is assumed to be spatially uniform. Several void fraction models are available; the m preferred model is a profile-fit model to compute the flow quality, which is
# then used to calculate the void fraction from a modified Zuber-Findlay drift-flux relationship.
The Critical Power Ratio (CPR) for a number of limiting bundles is obtained using the GEXL correlation to compute the critical quality. The core power transient is optionally based on a point or a one-dimensional
'g (axial) space-time kinetics model. The power transient can also be specified by the user. The kinetics models account for the important reactivity components which are void (density), fuel temperature (Doppler), and rod motion (scram). In addition, the user may specify a reactivity forcing function. The delayed neutrons are represented by a maximum of six precursors, and the decay heat is also explicitly modeled. The core may be ,
initially subcritical. The one-dimensional kinetics model is based on the I total fission source and a nodal representation for the average fuel and
-3 coolant channel. The nuclear parameters (K. and M 2) are obtained from a comparable three-dimensional model (1) in which a collapsing procedure is used to obtain the radially averaged values. Individual and groups of control rods (maximum of 10) may be represented. The collapsing procedure is used to obtain the initial condition parameters as well as the feedback parameters for the transient solution. Spatial variation of the total delayed neutron fraction and prompt neutron lifetime are represented. The initial power distribution is based on the solution of the neutron source equations with all time-derivatives set to zero.
2.2.2 BEACTOR VESSEL FLUID MODEL The reactor vessel (RV) excluding the core is represented by six fluid regions: upper downcomer, lower downcomer, lower plenum, bypass, riser (outlet plenum and separators), and steam dome. The conservation equations of mass and energy are solved for each region based on the boundary flows and enthalples. Heat conduction with metal structural components can also be O NSPNAD-8608-A Rev. 4 Page 11 of 194
C) , considered. The pressure distribution is assumed to be spatially uniform. The RV pressure is obtained from a consideration of the mass and energy balance in the steam dome which accounts for non-equilibrium effects. A gg separate model is provided to calcula+;e the water level in the steam dome which accounts for steam carryunder An the recirculation water and area variations due to the steam separator geometry. Level sensing is also represented. The RV flow is either user-specified or calculated from the conservation of momentum equations. In this latter case, the dynamics of the recirculation pumps (RP's) are also taken into consideration, and a wide variety of pump transients can be represented. The pump heat is included in the model. The II hydraulic model for the RP's is based on homologous relationships. The hydraulic model represents the flow in the two individual recirculation loops and considers forward and reverse suction and drive line flows. Automatic RP trip on low RV level or'high RV pressure can be specified. The initial suction and drive flows are specified, and the suction flow path loss coefficinnt is computed to provide momentum balance. This loss coefficient is assumed constant during the transient. Two-phase pressure drop and fluid inertial effects in the core, outlet plenum and steam separators are modeled. II 1 The transient core bypass flow fraction is computed based on the conservation of momentum equations. The initial pump status is user-specified as either on or off. The pump motor electrical torque is obtained from the output of the motor / generator (M/G) flow control system, which is described later. It should be noted that, when the dynamic flow calculation is used, the temperature (enthalpy) distribution within the downcomer, recirculation lines, and jet pumps is represented in detail to provide an accurate model of the GD . changes in core inlet subcooling due to changes in feedwater, high-pressure coolant injection system (HPCIS) and reactor core isolation cooling system
'ACICS) flows and enthalples, r2 feedwater flow is assumed to enter the top of the downcomer. The reedwater flow can be specified by the user or controlled by a three-element control system which is described later. The feedwater enthalpy is user-specified. (>
The safety / relief valves (S/RV's) are represented as individual valves. Account is taken for accumulation and blowdown, valve opening / closing delays, and valve stroking. 2.2.3 MAIN STEAM SYSTEM MODEL The main steam system consists of the main steam lines, main steam line isolation valves (MSIV's), bypass, turbine control and stop valves, and the g turbine. The main steam lines can be represented by either a lumped parameter model or a detailed model. , i For the lumped parameter model, the steam line portion on the RV side of the MSIV's is included in the definition of the steam dome, and the remainder is represented by a single volume. The flow through the MSIV's is calculated gg from an orifice equation so that the steam inertial effects are neglected. The steam dome pressure is obtained from the RV model as described earlier, and the steam line pressure is based on conservation of mass assuming saturation conditions. In the detailed model, the main steam system representation begins at the RV exit and consists of seven pipe segments per steam line. These seven segments consist of four in the main steam line (RV exit to S/RV location, S/RV to MSIV, MSIV to bypass valve location and bypass valve to turbine valve gp i~ location), the relief valve line, the safety valve line, and the bypass valve line. Each pipe segment is subdivided into a finer mesh with a maximum of 11 mesh points per segment. The conservation equations for mass, energy and momentum are solved at each mesh point based on the Method of Characteristics ; (MOC). The MOC model provides for realistic modeling of the pressure waves within the main steam system resulting from rapid valve motion with minimal GD ,1 l NSPNAD-8608-A Rev. 4 Page 12 of 194
__ _- ~- _ _ _ _ . __ . _ _. __ .- . . __. _ -._ numerical dispersion. The solution of the MOC model is based on the appropriate boundary conditions which consist of the steam dome pressure and the S/R, bypass, and turbine valve flow rates. Closure of the MSIV's is automatically initiated by any of the following three signals: low RV level, high steam flow, or low RV pressure. Appropriate time , delays and valve closure rates are user specified. l 1 The bypass valve flow is based on the bypass valve position (and hence area), which can be specified by the user or controlled by the Pressure Regulator Control System which is described later. Similar treatments are used for the
) turbine control and stop valves. Automatic bypass valve opening can be actuated on turbine stop valve closure.
The turbine model provides a representation of the turbine speed based on the angular momentum equation solution. The driving torque is related to the turbine inlet steam flow. The turbine speed can be used to simulate frequency I changes for the M/G drive motor torque. l ) l 2.2.4 SAFETY SYSTEMS The high pressure coolant injection system (HPCIS) model is based on a flow versus back pressure curve with a user-specified enthalpy. Automatic actuation with an appropriate time delay is provided based on either low RV i l water level or pressure. The reactor core isolation cooling system (RCICS) is modeled in a similar i manner to the HPCIS. The reactor protective system represents five explicit automatic trip l functions: high neutron power, high RV pressure, low RV level, high RV level, and MSIV closure fraction. A scram can also be initiated at a user-specified time. The flow dependence of the high neutron power trip is represented based on the sensed recirculation drive line flow. Each trip function has a unique time delay. 2.2.5 CONTROL SYSTEMS ]) The main feedwater controller is based on a three-element system with the sensed reactor vessel water level, sensed main steamline flow and the feedwater flow as the three input signals. The control system adjusts the feedwater valve position to attempt to obtain a zero error signal. RV back pressure effects on flow rate can be represented. The M/G flow controller accepts the coupler scoop tube position and output signal from the turbine speed governor as input and adjusts the scoop tube position to maintain the appropriate setpoint. Control is provided for each t l loop independently. The M/G dynamic model is based on solving the l conservation of angular momentum for the motor and generator separately. Idle l loop recirculation pump startup can be modeled. Automatic drive motor trip on high RV pressure can be specified. The pressure regulator control system accepts the sensed RV pressure and the turbine speed governor output signal as input. This system controls the turbine control and bypass valve positions to maintain the RV pressure at the l appropriate setpoint. RV back pressure effects on steam flow can be represented. 2.2.6 INTEGRATION SCHEME The reactor core model equations are integrated by using a fifth-order Runge-Kutta-Merson method in which the time step is automatically selected to achieve a user-specified accuracy limit. This same method is used to integrate the dynamic flow equations for the RV flow rates. Note that the
]) core and RV time step sizes are usually different with the former being smaller. The Moc solution in the steam lines is carried out over a user-specified fraction of the RV time step using an explicit integration technique and linear variation of the boundary conditions over the RV time step.
D NSPNAD-8608-A Rev. 4 Page 13 of 194
O 2.3 COMPARISONS WITH OTHER APPROVED LICENSING CODES This section provides a comparison with two other approved licensing codes: REDY (8) and ODYN [9). In general, the development of DYNODE-B paralleled the evolution of these two codes from the standpoint of applications to licensing analyses. Early versions of DYNODE-B were patterned after REDY, based on the information provided in Reference 8, so that the models, assumptions, and approximations in these early versions are similar to those of REDY. The advanced versions incorporated more sophisticated models in the areas of the core kinetics (1-D axial) and steam line hydraulics, following the improvements of ODYN over II, REDY. The nature of the DYNODE-B enhancements are similar to those of ODYN, but slightly different in implementation, as discussed later. The remainder of this section presents the major technical differences between ODYN and DYNODE-B. 2.3.1 CORE NEUTRONICS ODYN is based on one-group diffusion theory, in which the cross sections are a function of coolant density, fuel temperature, and control state. DYNODE-B is II based on the total fission source nodal formulation, in which neutron migration is represented by coupling coefficients between adjacent nodes. The coupling coefficients are functions of the migration area, M. Local neutron multiplication is given in terms of the infinite multiplication factor, K.. The forms of the equations for these two models are similar in nature, and the DYNODE-B nodal formulation can be derived from the one-group equations. The treatment of the delayed neutrons is identical in the two codes. The DYNODE-B model treats decay heat precursors in conformance with the 1971 ANS Standard $5 while ODYN uses a simple exponential decay nest model. The radial collapsing procedures used to develop the one-dimensional parameter are identical in nature. 2.3.2 STEAM LINES ODYN is based on a single-phase, one-dimensional nodal representation of the steam line (8 nodes) in which the steam is assumed to behave isentropically. DYNODE-B is based on a Method of Characteristic (MOC) solution to the () one-dimensional conservation equations for mass, energy, momentum, and state. The HOC methodology is more rigorous and does not assume a priori that the steam is isentropic. The MOC method was used as a reference in establishing the validity of the ODYN model. 2.3.3 REACTOR VESSEL PRESSURE DISTRIBUTION ODYN explicitly calculates the pressure at the reactor inlet and the RV dome. DYNODE-B calculates the preesure in the RV dome. The reactor core pressure is gp obtained by simulating the appropriate transport delay between the dome and the core outlet based on the sonic velocity and the distance between these two points. GD l l l 45 l l GP 1 NSPNAD-8608-A Rev. 4 Page 14 of 194 l
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3.0 DYNODE-B CODE OUALIFICATION This section discusses the benchmark analyses performed to qualify the DYNODE-B computer code for BRR analysis. O The nuclear models (1-D kinetics parameters) are compared to results from the 3-D simulator [1] in Section 3.1. The thermal-hydraulic models are compared to other approved licensing codes, i.e., GE's REDY and ODYN codes, and to test data, i.e., Peach Bottom turbine trip tests, and Monticello startup tests. These results are shown in Section 3.2. O It is the intent of this section to determine the applicability of the DYNODE-B code models to BWR analysis only. The application of DYNODE-B to licensing analysis and an evaluation of margin is contained in Section 4. 3.1 NUCLEAR MODEL COMPARISONS The 1-D kinetics parameters which are input to DYNODE-B are derived from a radial collapsing of the 3-D results from the corresponding NDH parameters (Reference 1). Since the NDH model has been qualified against plant GD measurements, the accuracy of the 3-D parameters is well founded, and thus the only other uncertainty associated with the 1-D parameters is that which is introduced by the collapsing process and the calculation of the feedback parameters. This additional uncertainty can be assessed by direct comparisons of reactivity changes between NDH and DYNODE-B results for comparable changes in the statepoints. This assessment has been made for the void, scram and Doppler reactivity com;,aents, since the dominant reactivity changes occur due to these three components during events analyzed with the 1-D model dp (overpressurization events). It should be noted that any uncertainty associated with these procedures will be included in the 6CPR/ICPR uncertainty due to input parameter uncertainties as applied in section 4.4.1. The results of specific comparisons for Monticello are presented in the following sections to demonstrate the nature of this process. 3.1.1 SCRAM REACTIVITY g The change in scram reactivities, calculated with NDH and DYNODE-B, for the case of all-rods-out (ARO) to all-rods-in (ARI) with the highest worth rod stuck out are shown in Figure 3.1-3. The percent difference between these two curves is presented in Figure 3.1-4. The DYNODE-B error is conservative when it is greater than zero. The vertical line on these curves shows the amount of control rod insertion when the minimum critical power ratio (MCPR) occurs for the turbine trip without bypass transient. For this case, no uncertainty need be applied for the scram reactivity component, since the control rod gp worth up to the time of MCPR is conservative relative to NDH. This is expected since the ik, input into DYNODE-B for the scram reactivity is determined between a fully converged, static, ARO NDH case and a source converged, NDH case with all but the highest worth rod inserted. The rods in NDH case held the independent core parameters, such as core power, flow, and void distribution constant. This does not take into account the flux redistribution effects as the control rods are stepped in. The neglect of this effect results in an added conservatism. gg 3.1.2 VOID REACTIVITY The calculated values for K-eff between NDH and DYNODE-B for a change in subcooling are shown in Table 3.1-1. The corresponding power distributions are shown on Figures 3.1-1 and 3.1-2. These results show that the DYNODE-B , result for void reactivity feedback is slightly more conservative. For each l DYNODE-B case shown in Table 3.1-1, the corresponding NDH collapsed 1-D parameters were used directly to determine the DYNODE-D K, and M2 distributions. In section 7.24 of this report, the DYNODE-B cases used the dp . feedback tables to determine the K-eff changes. The results of that analysis ) also confirm that no additional uncertainty needs to be applied due to the 3-D to 1-D collapsing process. j 3.1.3 DOPPLER REACTIVITY The comparison of the Doppler reactivity using the collapsed 1-D parameters Gb l l NSPNAD-8608-A Rev. 4 Page 16 of 194
i . directly from NDH is shown in section 7.3 of this report. The comparison using the transient tables is shown in section 7.24 of this report. The results of section 7.3 show that NDH and DYNCDE-B reactivities are the same ) when the collapsed 1-D parameters are input directly into DYNODE-B, and section 7.24 shows that an additional uncertainty is needed when the transient tables are used. l 3.2 THERMAL - HYDRAULIC COMPARISONS This section is intended to benchmark the thermal-hydraulic models in the DYNODE-B code. This is accomplished by benchmarking to other approved licensing codes and to test data. The test data benchmarks provide the ) primary checkout. The code benchmarks provide a secondary check of the overall behavior; however, exact comparisons are not expected, due to modeling i and input differences. Essentially, these benchmarks provide a full system ; checkout of the models in DYNODE-B, including the main models described in l Section 2. l In order to provide an accurate comparison, an effort was made to duplicate I the comparative conditions as closely as possible, i.e. models used in the ) case of benchmark codes and the measured test conditions for data comparisons. l These benchmarks are meant only to test the modeling of the DYNODE-B code. l The methodology used for Reload Safety Evaluations is described in Section 4, and these results are referenced where applicable. 3.2.1 CODE-CODE COMPARISONS In this section the DYNODE-B code is benchmarked to the General Electric REDY code (8] results in the FSAR analysis (2) and the ODYN code (9) results in the Cycle 11 reload safety analysis (4). 3.2.1.1 GENERAL ELECTRIC REDY CODE The models described in Reference 8 were duplicated as closely as possible for these cases. The major discrepancies in the code modeling are: REDY uses a second-order sweep model to calculate the dynamic void effects, whereas DYNODE-B uses a ] profile-fit non-equilibrium flow quality void model; REDY calculates decay heat from the stehn-Clancey correlation (8], whereas DYNODE-B calculates decay heat as part of the kinetics equations; and REDY assumes a constant value of cladding surface heat transfer coefficient throughout the transient, whereas DYNODE-B assumes the heat transfer coefficient to behave according to the Thom correlation. g The input from Reference 13 .as used wherever possible. In cases where the REDY input value was unknown, typical or actual plant values were used in the DYNODE-B models. In particular, the majority of the controller (flow, pressure, and level) model inputs to REDY were not known. Fourteen transients from the FSAR (2) were benchmarked. The following sections describe each transient and document the benchmark analysis results. 3.2.1.1.1 TURBINE TRIP WITHOUT BYPASS Description Of The Accident This transient is a severe abnormal event which results directly in a primary system pressure increase. It represents the sequence of events that would follow an instantaneous loss of condenser vacuum with failure of the low condenser vacuum scram signal, so that scram occurs upon automatic closure of the turbine stop valves. For this event, the turbine bypass valves are assumed to be inoperable because of the loss of condenser vacuum. Reactor scram is initiated by position switches on the turbine stop valves. Summary of Accident Analysis The reactor is assumed initially to be at rated power (1670 MWt). The turbine stop valves are taken to close in a conservatively short duration (0.1 second), and the insertion of scram reactivity is limited to the rate of insertion allowed by the Technical Specifications with the most reactive rod stuck out. The fuel temperature reactivity is assumed to correspond to its 4 NSPNAD-8608-A Rev. 4 Page 17 of 194
'O least negative time in life, while the void reactivity is taken as the most negative to maximize the power spike and pressure increases which result from the stop valve closure. The transient is mitigated by the action of the gg safety / relief valves, which are taken to open at the maximum pressure permitted by the Technical Specifications. Initial system pressure is conservatively placed 25 psi above the nominal operating pressure.
Figures 3.2-1 through 3.2-7 show the DYNODE-B versus GE REDY results for the Turbine Trip without Bypass transient. The results show excellent comparison. DYNODE-B slightly overpredicts the peak power, and hence pressure, response (344% nominal versus 321% from REDY). II Both codes predict the same initial water level response, though REDY predicts a stronger recovery of level. This is due to the differences in the transient void models. DYNODE-B accurately reproduces the REDY results. The minor discrepancies that exist are due to code modeling differences. This case therefore provides an acceptable benchmark. ID,r i 3.2.1.1.2 TURBINE TRIP WITH BYPASS Description of the Accident The sequence of events for the turbine trip with bypass malfunction are similar to that for the turbine trip without bypass (Section 3.2.1.1.1), except that the condenser heat sink is presumed to be available and hence the turbine bypass valves are operable. Following stop valve closure, the pressure regulator controls react to open the bypass valves and relieve steam to the condenser. ID Summary of Accident Analysis The reactor is assumed initially to be at rated power (1670 MWt). The turbine stop valves are taken to close in a conservatively short duration (0.1 second), and the insertion of scram reactivity is limited to the rate of insertion allowed by the Technical Specifications with the most reactive rod stuck out. The fuel temperature reactivity is taken to be at its least negative time in life while the void reactivity is at its most negative value. (D A rapid power spike and pressure increase follow the valve closure. The transient is terminated by the reactor scram, opening of the turbine bypass valves, and by the safety / relief valves whose opening setpoints are presumed at the maximum value permitted by the Technical Specifications. The initial system pressure is conservatively assumed to be 25 psi above the nominal ' setpoint. Figures 3.2-8 through 3.2-14 show the DYNODE-B versus GE REDY results for the ($ ! Turbine Trip with Bypass transient. The results show excellent comparison. DYNODE-B slightly overpredicts the power, and hence pressure, response. Both codes predict the same initial water level response, though REDY predicts a stronger recovery of level. This
- is due to the differences in the transient void models. The increased level predicted by REDY in turn causes the feedwater controller to cut back on flow.
This effect is not seen in DYNODE-B, since the level does not recover to the gp same degree. DYNODE-B accurately reproduces the REDY results. The minor discrepancies that exist are due to code modeling differences. This case therefore provides an acceptable benchmark. 3.2.1.1.3 GENERATOR TRIP Description of the Accident The generator trip transient is a severe overpressurization transient which is (p : similar in nature to the turbine trip. It represents the sequence of events which would follow rapid closing of the turbine control valves, which could follow a complete loss of electrical load. Reactor scram is initiated automatically by relays which sense the fast turbine control valve closure. GD NSPNAD-8608-A Rev. 4 Page 18 of 194
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Summary of Accident Analysis The reactor is assumed initially to be at rated power (1670 MWt). The turbine , control valves are taken to close in a conservatively short duration (0.2 , J second), and the insertion of scram reactivity is limited to the rate of ' insertion allowed by the Technical Specifications with the most reactive rod stuck out. A rapid pressure increase follows the valve closure, the magnitude , of which depends principally on the scram reactivity insertion rate and the ! void reactivity. The fuel temperature reactivity is conservatively taken at its least negative time in life, while the void reactivity is at its most negative to maximize the power excursion. The transient is mitigated by
, opening of the turbine bypass valves and the safety / relief valves.
J Figures 3.2-15 through 3.2-21 show the DYNODE-B versus GE REDY results for the l Generator Trip with Bypass transient. The results show excellent comparison for all variables except water level and l feedwater flow. Both codes predict the same initial water level response, though REDY predicts a stronger recovery of level. This is due to the l differences in the transient void models. The increased level predicted by l ,J REDY in turn causes the feedwater controller to cut back on flow. This effect l 1s not seen in DYNODE-B, since the level does not recover to the same degree. l DYNODE-B accurately reproduces the REDY results. The minor discrepancies that exist are due to code modeling differences. This case thgrefore provides an , acceptable benchmark. l 3.2.1.1.4 CLOSURE OF ALL MAIN STEAM ISOLATION VALVES j l Description of the Accident l ~ Closure of all the main steam isolation valves (MSIV) while the reactor is at ! power can result in a significant overpressure transient in the reactor l vecsel. The MSIV's can be closed directly by operator action while at power. Normally, as the valves close in all four steamlines, a reactor scram is initiated by position switches which sense the closure. As the system isolates, pressure rises in the vessel until the safety / relief valves open to mitigate the accident. O Summary of Accident Analysis The reactor is assumed to be at rated power (1670 MWt). The MSIVs are taken to close in three seconds with a nonlinear valve flow characteristic. Fuel temperature reactivity is taken to be at its least negative time in life, while the void coefficient is at the most negative to maximize pressurization. Position switches on each MSIV will cause a reactor scram when valves in three of the four steamlines reach approximately 10% closed. This results in a gg scram before any significant steam flow interruption takes place. System pressure rises due to heat stored in the core. Insertion of scram reactivity is limited to the rate allowed by the Technical Specifications with the most reactive rod stuck out. The transient is mitigated by opening of the safety / relief valves, with lift setpoints assumed at the maximum Technical Specification limit. Initial system pressure is assumed 25 psi above the nominal value. gg Figures 3.2-22 through 3.2-28 show the DYNODE-B versus GE REDY results for the MSIV Closure transient. The DYNODE-B results show excellent comparison for all variables. The peak pressure predicted by REDY is approximately 10 psi greater than predicted by DYNODE-B. This difference can be caused by slight differences in the non-linear MSIV position versus area curve input. DYNODE-B accurately reproduces the REDY results. The minor discrepancies that gg exist are due to code modeling and input differences. This case therefore provides an acceptable benchmark. 3.2.1.1.5 FEEDWATER CONTROLLER MALFUNCTION, MAXIMUM DEMAND Description of the Accident Failure of the feedwater controller in the direction of increased feedwater flow results in a moderator temperature and void decrease and a reactor power
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NSPNAD-8608-A Rev. 4 Page 19 of 194
_ - .. . ~- .- -- - - - - - - . - - --- O increase through the effect of the negative reactivity void coefficient. Water level increases during the initial part of the transient until the high water level turbine trip setpoint is reached. The turbine and feedwater pump gg trips, the reactor scrams, and an overpower and overpressure transient occurs. i I Summary of Accident Analysis i The reactor is taken to be at 65% rated power (1085.5 MWt). The feedwater
- controller is assumed to fail in such a manner as to cause the feedwater flow
; to increase to its full run-out value. The water level and core inlet
- subcooling increase, causing reactor power to increase until the high water level turbine trip setpoint is reached, causing a turbine trip, feedwater pump II i trip, and a subsequent reactor scram due to turbine stop valve closure. Fuel 1 temperature reactivity is taken at its least negative time in life, while the
! void reactivity is at its most negative to maximize the power and pressure j transient. The insertion of scram reactivity is limited to the rate of insertion allowed by the Technical Specifications with the most reactive rod stuck out. The transient is mitigated by opening the turbine bypass valves and j the safety / relief valves. The safety / relief valves are assumed to open at the t maximum pressure permitted by the Technical Specifications. O Figures 3.2-29 through 3.2-35 show the DYNODE-B versus GE REDY results for the Feedwater Controller Failure - Maximum Demand transient. The results show excellent comparison for all variables. The predicted scram I time is slightly later in DYNODE-B. This is due to the fact that REDY assumes l an instantaneous increase in feedwater flow whereas DYNODE-B assumes the feedwater control valve opens at the maximum rate. The peak pressure predicted by REDY is approximately 25 psi greater than predicted by DYNODE-B. ID j This is due to the fact that the slight differences in the steam line models j; cause DYNODE-B to open the bypass valves, whereas REDY predicts they stay closed. DYNODE-B accurately reproduces the REDY results. The minor discrepancies that exist are due to code modeling differences. This case i therefore provides an acceptable benchmark. I 3.2.1.1.6 LOSS OF FEEDWATER l Descriotion of the Accident GD
- A loss of feedwater flow results in a situation where the mass of steam leaving the reactor vessel exceeds the mass of water entering the vessel, i
1 resulting in a net decrease in the coolant inventory available to cool the 1 core. Feedwater control system failures or feedwater pump trips can lead to . partial or complete loss of feedwater flow. Feedwater flow would decay over a j few seconds and the recirculation flow control system would ramp the recirculation pumps down to about 20% speed when the feedwater flow falls below 20% of rated. Water level declines rapidly and a reactor scram takes (p', place when the low level trip setpoint is reached. The system subsequently closes the main steam isolation valves (MSIV's), and actuation of the high 4 pressure coolant injection (HPCI) and reactor core isolation cooling (RCIC)
- systems on low level setpoints terminate the transient.
- Summary of Accident Analysis
! The reactor is taken to be at rated power (1670 MWt). The loss of feedwater ! is modeled as taking place over a three second period. When the feedwater gp l flow reaches 20% of normal, the recirculation pump speed demand is set to 20%. 3 4 The power level and system pressure decline in a fashion which depends principally on the void reactivity coefficient. Figures 3.2-36 through 3.2-42 show the DYNODE-B versus GE REDY results for the ! Loss of Feedwater transient. The DYNODE-B results show excellent comparison for all variables. The only minor discrepancy is that REDY predicts a reactor scram on low water level at dp approximately 13 seconds. DYNODE-B does not predict the scram throughout the 16 seconds simulated. This is due to the fact that the low water level scram input to REDY is 9 inches above that used in DYNODE-B. DYNODE-B accurately reproduces the REDY results for this transient. The only discrepancy is caused by the input difference mentioned above. This case G NSPNAD-8608-A Rev. 4 Page 20 of 194
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) therefore provides an acceptable benchmark. 3.2.1.1.7 LOSS OF FEEDWATER HEATING J Description of the Accident A loss of feedwater heating event can occur as the result of a loss of extraction steam to a feedwater heater. An alternative, but generally less severe, loss of heating can result from inadvertent actuation of high pressure ! coolant injection (HPCI), which delivers relatively cool water to the reactor through the feedwater sparger. Reduction in feedwater temperature follows, with a gradual rise in reactor power as the moderator temperature declines and reduces the core void fraction. If neutron power exceeds the reactor trip j ) setpoint, a scram occurs; otherwise the system settles to a steady-state high-power condition until the operator intervenes. j Summary of Accident Analysis The reactor is taken to be at rated power (1670 Mwt). The feedwater i temperature change is modeled as a 100 'F decline with a 30-second exponen'ial I time constant. This is more severe than any loss of feedwater heating which I can result from a single system malfunction. Fuel temperature reactivity is T) chosen at its least negative time in life, while void reactivity is at its most negative to maximize the power increase. Scram insertion, should it l occur, is limited to the insertion rate permitted by the Technical Specifications. Figures 3.2-43 through 3.2-49 show the DYNODE-B versus GE REDY results for the Loss of Feedwater Heating transient. 7 The DYNODE-B results show excellent comparison for all variables. DYNODE-B predicts a more conservative increase in power and average surface heat flux. This case therefore provides an acceptable benchmark of DYNODE-B's capabilities to reproduce the REDY code predictions. 3.2.1.1.8 PRESSURE REGULATOR FAILS OPEN Description of the Accident In the event that either the electrical or mechanical pressure regulator were m to fail such that the turbine control and/or bypass valves were opened, steam " flow from the reactor would increase. System pressure would drop, causing an increase in core voids and a consequent drop in reactor power. Depressurization would continue until the main steam isolation valve closure setpoint was reached, resulting in closure of the valves and a reactor scram. Decay heat then would cause the system to repressurize, limited by opening of the automatic safety relief valves until cooldown was initiated. m Summary of Accident Analysis The reactor is taken to be at rated power (1670 MWt). The pressure regulator is taken to fail in such a way that the turbine control valves and/or bypass valves are opened to 110% steam demand (the maximum permitted by the control system). The excess demand depressurizes the system until the main steam isolation valves close and the reactor scrams. Thereafter, the pressure rises due to decay heat, and the automatic safety / relief valves lift intermittently until cooldown is initiated. q '> Figures 3.2-50 through 3.2-56 show the DYNODE-B versus GE REDY results for the Pressure Regulator Fails Open transient. DYNODE-B follows the REDY predicted steam dome pressure very closely with the exception that DYNODE-B predicts the MSIV closure on low turbine throttle pressure to occur approximately 2 seconds after REDY. This is due to differences in the steam line model. The same initial depressurization causes a greater void increase in REDY than in DYNODE-B due to the differences in the void models. This in turn causes the REDY predicted core power and core inlet ') flow to drop faster than DYNODE-B. DYNODE-B accurately reproduces the REDY results for this transient. The discrepancies are caused by code modeling differences. This case therefore provides an acceptable benchmark. O NSPNAD-8608-A Rev. 4 Page 21 of 194
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(? l 3.2.1.1.9 RECIRCULATION PUMP SEIZURE l Description of the Accident The recirculation pump seizure is a nearly instantaneous stoppage of a g recirculation pump shaft and impeller. This stoppage results in a very rapid reduction in core flow and a subsequent decline in core power. Because the heat flux at the fuel pin surface declines more slowly than the core flow, there is a potential degradation of thermal margin. No reactor scram results, I and the system settles to reduced power. Summary of Accident Analysis , The reactor is taken to be at rated power (1670 MWt). The affected pump speed 1 II is instantaneously set to zero and the drive flow abruptly decays. Jet pump flow in the seized loop reverses in less than 1 second. As a result, core flow decreases, causing an increase in void fraction and a consequent reduction in reactor power. The degree of reduction in power depends principally on the void coefficient of reactivity, which is taken at its least negative time in life, while Doppler reactivity is at its most negative to maximize the heat flux to flow ratio during the transient. Heat flux from the fuel pins lags the core power decline, and the system relaxes to a reduced power steady state. GD Figures 3.2-57 through 3.2-63 show the DYNODE-B versus GE REDY results for the Recirculation Pump Seizure transient. The DYNODE-B results show excellent comparison for all variables. This case therefore provides an acceptable benchmark of DYNODE-B's capabilities to reproduce the REDY code predictions. O 3.2.1.1.10 TWO RECIRCULATION PUMP DRIVE MOTOR TRIP Description of the Accident In the event that the power supply to both recirculation pump motor / generator (M/G) sets were lost, the pumps would coast down and coolant flow to the core would decline. Core voids would then increase and power would decline. The system settles to a natural circulation condition where core flow is provided through the jet pump suction path by the weight of subcooled water in the downcomer. Heat flux decline lags power and core flow, so there is a (D potential degradation of thermal margin limits. Summary of Accident Analysis The reactor is taken to be at rated power (1670 MWt). The transient is initiated by setting the recirculation pump drive motor torques to zero. The inertia of the M/G sets is included in the analysis because there is no single event which would result in simultaneously opening the pump generator breakers to both pumps. The void reactivity is taken at its least negative time in () life, while fuel temperature reactivity is at its most negative to maximize the power to flow ratio during the event. The pumps, core flow, power, pressure, and steam flow all decline to steady-state, natural-circulation conditions. Figures 3.2-64 through 3.2-70 show the DYNODE-B versus GE REDY results for the Two Recirculation Pump Drive Motor Trip transient. DYNODE-B compares very well with the REDY-predicted core inlet flow during the initial flow coastdown. As the transient progresses, REDY predicts a slightly lower core inlet flow. This could be caused by differences in the pump model input. The lower core inlet flow predicted by REDY causes increased voiding and hence a greater power decrease and more level hol Np. The lower power causes lower heat flux and lower pressure. The higher level causes a greater feedwater decrease in an attempt by the feedwater controller to compensate. In general, DYNODE-B follows the same trends and reproduces the REDY result () for this transient. The discrepancies are caused by code modeling difference. This case therefore provides an acceptance benchmark. 3.2.1.1.11 RECTRCULATION FLOW CONTROLLER FAlLURE, INCREASE DEMAND Description of the Accident There are several possible failures which can result in an increase in core O NSPNAD-8608-A Rev. 4 Page 22 of 194
D l l l coolant flow. The most severe of these occurs when a motor / generator (M/G) l set fluid coupler for one recirculation pump attempts to achieve full speed at
, maximum acceleration. The result is a surge of additional coolant through the %) core and a consequent power increase. If the neutron flux increases to the l high power trip setpoint, the reactor scrams. The possibility of a large ]
power increase allows for potential degradation of thermal margin. j l Summary of Accident Analysis l The most severe initial condition for the increasing recirculation flow I transient is near the low end of the automatic recirculation flow control range, where reactor power is approximately 65% of rated power and core flow , .() is approximately 50% of rated flow. The pumps are operating at approximately l 45% speed, and the relative M-G set fluid coupler scoop tube position is i approximately 20%. The transient is modeled by moving the scoop tube position j at its maximum rate to the maximum coupling position. Void reactivity is l taken at its most negative time in life, while fuel temperature reactivity is i at its least negative to maximize the power increase. As the pump speed increases, core flow and power increase and, if the power increase is sufficient, a scram occurs. The system then settles to a steady state until -() the operator intervenes. Figures 3.2-71 through 3.2-77 show the DYNODE-B versus GE REDY results for the Recirculation Flow Controller Failure - Increased Demand transient. l The DYNODE-B results show excellent comparison for all variables. Minor discrepancies in the level response are caused by the transient void model differences. () DYNODE-B accurately reproduces the REDY results for this transient. The minor i discrepancies are caused by code modeling differences. This case therefore provides an acceptance benchmark. 3.2.1.1.12 RECIRCULATION FLOW CONTROLLER FAILURE, DECREASE DFMAND Description of the Accident , The failure of one recirculation pump motor / generator (M/G) set speed 77 controller could cause the scoop tube position to move at its maximum speed in the direction of zero pump speed and flow. As a result, core flow, power, steam flow, and pressure all decrease. Because the decline in heat flux lags that of core flow and power, there is a potential degradation in thermal margin. Summary of Accident Analysis The reactor is taken to be at rated power (1670 MWt). The transient is fm initiated by forcing the scoop tube position of the affected loop M-G set from its initial value to zero at the maximum rate. Core flow, power, steam generation, and pressure all decline, and the system settles to a steady state at reduced power with reverse flow through the inactive jet pumps. Void reactivity is taken at the least negative time in life while fuel temperature reactivity is at the most negative to maximize the power to flow ratio during the transient. Figures 3.2-78 through 3.2-84 show the DYNODE-B versus GE REDY results for the () Recirculation Flow Controller Failure - Decreased Demand transient. The DYNODE-B results show excellent comparison for all variables. The DYNODE-B predicted core inlet flow drops slightly lower than the REDY-predicted flow. This can be due to minor input differences in the coupler torque versus slip and coupling function. All of the other discrepancies are insignificant and attributable to the void model differences. O DYNODE-B accurately reproduces the REDY results for this transient. The slight discrepancies are caused by code modeling or input differences. This case therefore provides an acceptance benchmark. O NSPNAD-8608-A Rev. 4 Page 23 of 194
() ' 3.2.1.1.13 IMPROPER START OF AN INACTIVE RECIRCULATION LOOP pescriotion of the Accident Improper start of an inactive recirculation loop involves activating an II improperly warmed idle recirculation pump while the reactor is at power. Depending on the initial reactor condition, this incident can cause a significant power increase and reduction of thermal margin. The system settles out to an increased power steady state or, in the event the high neutron power trip setpoint is reached, the reactor scrams. Summary of Accident Analysis The initial conditions of the system substantially affect the results of the II , transient. One recirculation pump is presumed operating at full speed while - the second pump is stopped. The idle loop pump discharge valve is taken to be initially closed with the discharge bypass valve open. The inactive drive line is assumed to be filled with cold (100 *F) water. Reactor power and core flow are conservatively placed at midrange values with analyses performed to determine the most adverse conditions. The motor / generator (M/G) set fluid coupler for the idle pump is initially set for 50% speed demand. The transient sequence of events is as follows: ID l A. At t = 0, the idle M/G set drive motor breaker is closed. 1 B. The drive motor reaches near-synchronous speed quickly, while the l generator reaches approximately 80% speed in 5 seconds. C. At 5 seconds the generator field breaker is closed, loading the generator and applying starting torque to the pump motor. Generator (D speed decreases, the pump breaks into rotation and builds up speed. D. Generator speed. demand is programmed back to 20% starting at 8 seconds. E. The pump discharge valve / drive motor interlock is cleared, and the valve opens with a 60-second stroke time. The transient system behavior depends, to a great degree, on initial system (p power. At relatively high power, the pressure drop across the core becomes large and the starting pump does not develop sufficient head to reverse the I I backflow through the idle loop diffusers. Consequently, the water injected out the idle drive lines is swept back into the downcomer, where it is heated before eventually returning to the lower. plenum through the active loop. In contrast, at low power, the starting pump may cause the jet pump flow to become positive, sending the cold water directly into the immediate core flow path and resulting in a substantial core power increase and a possible reactor gp trip on high flux. Following the pump start, core flow abruptly increases, causing a power increase. If the reactor does not trip, the power peaks and then settles to a new level. As the pump discharge valve opens, power will increase as the valve permits flow to increase. If the reactor does trip, the scram terminates the power increase and causes the system to settle to zero power conditions. In either case, heat flux will increase to a maximum value at which time gg thermal margin will reach a minimum, and then the heat flux will decline. Figures 3.2-85 through 3.2-91 show the DYNODE-B versus GE REDY results for the Improper Start of an Inactive Recirculation Loop transient. The DYNODE-B results show the same general trends for all parameters as those predicted by GE REDY, although the magnitudes of the responses are different. This is due primarily to an apparent discrepancy between GE's written description of this transient and the plotted results (Reference 2). In the gp description, it is stated that the active pump initially produces 115% of normal rated flow in its associated jet pumps; in the figure, a flow of 150% is indicated. In either case, the core receives 54% of its normal rated flow, and all remaining flow from the active loop appears as reverse flow through the inactive loop. Therefore, the reverse flow through the inactive loop is much higher in the case of 150% active pump flow. The DYNODE-B analysis uses GD . - NSPNAD-8608-A Rev. 4 Page 24 of 194
e 4 0 115% active loop flow, per the written description. Because of the initially high reverse flow in the GE REDY case, the inactive (3 pump is unable to establish positive flow during the transient. As a result, the cold water in the loop is swept up into the downcomer, where it mixes with the bulk water before being pumped through the core by the active loop. This causes a relatively gradual reactivity insertion, so that the resulting power spike and rise in heat flux are mild. In contrast, the reverse flow in the DYNODE-B case is low enough so that the idle pump does establish positive flow upon starting up. The cold water is () therefore pumped directly into the lower plenum and through the core, causing a faster reactivity. insertion than in the GE REDY case. At the same time, establishing positive flow through the idle loop means an additional power increase because of the higher core flow. As a result, the power spike and the rise in heat flux are higher in the DYNODE-B case. The responses of other parameters are correspondingly altered. Despite this input difference which causes DYNODE-B to predict a more severe transient than REDY,.both codes predict the same trends and show the same () general results for this transient. This case therefore provides an acceptance benchmark. 3.2.1.2 GENERAL ELECTRIC ODYN CODE The models described in Reference 9 were duplicated as closely as possible for these cases. Major differences in the code models are described in Section 2.3. Wherever possible, the input from the Monticello Cycle 11 ODYN analysis (4) was used. The major input discrepancy in these cases is that the 1-D kinetics inputs used by the ODYN code were unknown and had to be estimated. Three transients from the Cycle 11 Supplemental Reload Analysis (4) were benchmarked. The following sections describe each transient and document the benchmark analysis results. b) 3.2.1.2.1 LOAD REJECTION WITHOUT BYPASS l Description of the Accident I Fast closure of the turbine control valves is initiated whenever electrical l grid disturbances occur which result in significant loss of load on the generator. The turbine control valves are required to close as rapidly as ; possible to prevent overspeed of the turbine generator rotor. The closing causes a sudden reduction of steam flow which results in a nuclear system pressure increase. The reactor is scrammed by the fast closure of the turbine () control valves. Summary of Accident Analysis The reactor and turbine / generator are initially operating at full power when the load rejection occurs. The power / load unbalance device steps the load . reference signal to zero and closes the turbine control valves at the earliest possible time. The turbine accelerates at a maximum rate until the valves start to close. The turbine control valves close in 0.25 sec.
)
Reactor scram is initiated upon sensing control valve fast closure. The insertion of scram reactivity is limited to the rate of insertion allowed by the Technical Specifications with the most reactive rod stuck out. A rapid pressure increase follows the valve closure, the magnitude of which principally depends on the scram reactivity insertion rate and the void reactivity. If the pressure rises to the pressure relief set point, some or all of the relief valves open, discharging steam to the suppression pool. If the pressure rises to 2 1150 psig, trip of the M/G set breaker occurs, causing () both recirculation pumps to coast down. Figures 3.2-92 through 3.2-98 show the DYNODE-B calculated results versus the General Electric ODYN results. DYNODE-B underpredicts the ODYN power and hence heat flux, increase. This is O NSPNAD-8608-A Rev. 4 Page 25 of 194
O due to differences in the void and scram reactivity functions. It is impossible to accurately reproduce the 1-D reactivity inputs used by General Electric based on the limited information available. A better response can be II achieved by performing sensitivity studies for the kinetics parameters. However, since the purpose of these benchmarks is only to perform a general check on the models used in DYNODE-B (the data comparisons in Section 3.2.2 perform the primary check), the benchmarks were left as is. It is sufficient to understand discrepancies due to input differences in this case. Most of the remaining differences are attributable to the difference in the heat flux response. The larger heat flux predicted by ODYN causes the pressure to hang up for a longer time and the water level to recover more 45 quickly due to a larger core resistance. Note that GE plots actual water level and DYNODE-B sensed water level which is the source of the difference in the initial values. Faster recovery of the water level causes the feedwater controller to ramp down the feedwater sooner. The differences in core inlet flow response are partially attributable to the heat flux differences and partially due to code modeling differences. In DYNODE-B, the pressure difference between the core outlet plenum and the steam 45 dome is not explicitly calculated. Thus, for the recirculation flow rate calculation, this pressure difference is computed from the momentum equation in which the steam separator flow acceleration term is obtained by assuming that the separator flow is replaced by the total core flow. This assumption is equivalent to assuming that the core fluid is incompressible. Thus, in cases of rapid void collapse in the core, this acceleration term does not play a significant role. The effect of this assumption is expected to be small, since the core void fraction is primarily responding to changes in pressure (p which are being taken into account properly. For this transient, this assumption results in DYNODE-B underpredicting the core inlet flow increase during the initial pressurization. The impact of this effect on ACPR is insignificant. This benchmark represents a positive check of DYNODE-B's capabilities to perform BWR transient calculations. The differences between the DYNODE-B and ODYN results are well understood and do not reflect deficiencies in the gp DYNODE-B code. 3.2.1.2.2 FEEDWATER CONTROLLER FAILURE - MAXIMUM DEHAND Description of the Accident This event is postulated on the basis of a single failure of a control device, specifically one which can directly cause an increase in coolant inventory by increasing the feedwater flow. The most severe applicable event is a feedwater controller failure resulting in maximum flow demand, which causes an gp increase of feedwater flow to the reactor vessel. This excess flow results in an increase in core subcooling, which results in a core power rise, and a rise in the reactor vessel water level. The rise in the reactor vessel water level eventually leads to high water level trip of the feedwater pumps and turbine, in turn causing a reactor scram. Summary of Accident Analysis The reactor is taken to be initially at 98% rated power (1634 MWt) and 100% flow. This point was found to be more conservative than the 100% power /100% flow point for the cycle 11 core (Ref. 4). The reactor is operating in a manual flow control mode which provides for the most severe transient. The feedwater controller is assumed to fail during the maximum flow demand. Maximum feedwater pump run out is assumed. The influx of excess feedwater flow results in an increase in core subcooling which gp reduces the void fraction and thus induces an increase in reactor power. The excess feedwater flow also results in a rise in the reactor vessel water level which eventually leads to high water level; main turbine and feedwater trip and turbine bypass valves are actuated. Reactor scram trip is actuated from main turbine stop valve position switches. Relief valves open as steamline pressures reach relief valve setpoints. If the pressure rises to 21150 psig, 9 NSPNAD-8608-A Rev. 4 Page 26 of 194
l J l l trip of the M/G set breakers occurs, causing the recirculation pumps to coast down.
- 2) Figures 3.2-99 through 3.2-105 show the DYNODE-B calculated results versus the General Electric ODYN results.
The DYNODE-B results compare very well to the ODYN results. The same input l and modeling differences exist as in the previous benchmark (Section 3.2.1.2.1 Load Rejection without Bypass). The Feedwater Controller Failure transient is not as sensitive to void reactivity as is the Load Rejection transient and j hence provides a much better code comparison. i m , J The DYNODE-B results are slightly time shifted (approximately 0.5 sec). This is due to the fact that General Electric assumes instantaneous feedwater runout flow, whereas DYNODE-B opens the feedwater control valves at the maximum rate to runout flow. Time to runout in DYNODE-B is 1.1 sec. This case provides an excellent check of DYNODE-B's capabilities to perform BWR transient analysis. The differences between the ODYN and DYNODE-B [j results, for this transient, are insignificant. 3.2.1.2.3 MSIV CLOSURE (FLUX SCRAM) Description of the Accident This event is performed to show compliance with the ASME Vessel Pressure Code. The MSIV's can be closed directly by operator action while at power. Closure of all main steam isolation valves (MSIV) while at power can result in a significant overpressure transient in the reactor vessel. Normally, as the c) MSIV's close, a reactor scram is initiated by position switches which sense closure. In addition, a secondary reactor scram will be initiated on high neutron flux. As the system isolates, pressure rises in the vessel until the safety / relief valves open to mitigate the accident. Summary of Accident Analysis The reactor is assumed initially to be at rated power (1670 MWt). The MSIV's are taken to close in three seconds with a non-linear valve flow r, characteristic. A reactor scram on MSIV position is conservatively ignored. '# Reactor scram is initiated on high neutron flux. The insertion of scram reactivity is limited to the rate of insertion allowed by the Technical Specifications with the most reactive rod stuck out. A rapid pressure increase follows closure of the MSIV's. If the pressure rises to the pressure relief set point, some or all of the relief valves open, discharging steam to the suppression pool. If the pressure rises to 2 1150 psig, trip of the M/G set breakers occurs. L> Figures 3.2-106 through 3.2-112 show the DYNODE-B calculated results versus the General Electric ODYN results. The DYNODE-B results compare favorably to the ODYN results. The same input and modeling differences exist as in the Load Rejection without Bypass benchmark (Section 3.2.1.2.1) since the two transients are very similar in response. The MSIV closure transient pressurizes more slowly and therefore is
,s less sensitive to void reactivity and hence the DYNODE-B and ODYN results (J compare more closely. The maximum increase in the reactor vessel pressures are within 10 psi. The differences between the DYNODE-B and ODYN responses to this transient are insignificant.
3.2.2 CODE-DATA COMPARISONS In this section the DYNODE-B code is benchmarked to three Peach Bottom turbine trip tests (16] and six Monticello start-up tests [14]. The purpose of these benchmarks is to qualify the models used in DYNODE-B and to gaantify the es conservatism in the DYNODE-B code. Section 4 discusses the quantification of
'> the code conservatisms.
3.2.2.1 PEACH BOTTOM 2 EOC 2 TURBINE TRIP TESTS Three instrumented turbine trips were carried out at the Peach Bottom-2 reactor during April 1977. These tests were conducted with the direct scram on stop valve position bypassed so that a trip on high flux was obtained. O NSPNAD-8608-A Rev. 4 Page 27 of 194
. . _ _ - ~ ~ _- -- _ _ _ -- . - - - , - - . _
() : This departure from the normal reactor condition was required to obtain a sufficiently large flux response to allow a more complete model-test comparison. A detailed description of the test conditions and measurement gg -. process can be found in Reference 16. - 3.2.2.1.1 TEST
SUMMARY
The initial power and flow conditions for each test are shown in Table 3.2-1. These test conditions were selected in order of increasing power along a line of. constant reactor flow. Prior to the second turbine trip test, it was necessary to reduce core flow to obtain the power to within 1% of planned test power level due to the xenon level in the core at the time of the test. In each of the three tests, the turbine stop valve position scram was disabled II : and the flux scram setpoint was reduced. The scram setpoints are also listed in Table 3.2-1. A total of 153 signals were recorded by a digital data acquisition system. The comparisons presented here will concentrate on those parameters which f affect the transient ACPR. 3.2.2.1.2 MODEL INPUTS (D : The DYNODE-B program has been used to model the three Peach Bottom 2 End of Cycle 2 (PB2EOC2) turbine trip tests (TT1, TT2, and TT3) for the purpose of benchmarking against overpressure transients which result in a rapid power increase. This benchmark effort began with an early version of DYNODE-B which did not have a one-dimensional kinetics or a detailed steam line model, so that these results were based on point kinetics and the lumped steam line models. This same model was used in pre-test predictions which validated the corresponding REDY results. Subsequently, the latest version of DYNODE-B was (D : used to incorporate spatial kinetics and steam line momentum effects. The development process is described below. The initial modeling of PB2 EOC2 was accomplished by utilizing design data and operational characteristics published in Reference 11 for the Reactor Coolant System (RCS). The point kinetics parameters were generated with a full 3-D nodal model of the core, similar to the models for Monticello described in Reference 1, using the actual initial test conditions. This work was (D [ performed by UAI (formerly NAI) and documented in Reference 10. This Best Estimate model utilized the MOC solution for the steam line momentum effects i as well as actual APRM trip setpoints, actual turbine and bypass valve positions, scram velocities, and recorded initial test conditions (RV pressure and flow, core power, steam flow, and core inlet subcooling) from Reference 16. The results of these comparisons provided satisfactory agreement with the (p ; measured core power and pressure transient data. Later on, after the one-dimensional kinetics model had been implemented in DYNODE-B, the benchmarks were repeated. However, for these analyses, the one-dimensional kinetics could not be obtained directly from the 3-D model, since the model is no longer available. Thus, an approximate approach was , taken in which the reactivity dependencies on void, fuel temperature, and i control state were established to give results which were comparable with the (g point kinetics data. The void dependency was then adjusted until the peak power matched the test data, and the scram worth was adjusted until the integrated power matched the test data. This procedure effectively eliminates ; the uncertainty due to the kinetics parameters. The test data comparisons thus represent differences due to the DYNODE-B computer code uncertainties , only. Therefore, these tests therefore represent a way to quantify these i uncertainties (See section 4.4), 3.2.2.1.3 DATA COMPARISONS $ This section describes the calculated to measured comparisons for the most important transient parameters; neutron flux, steam dome pressure, turbine , throttle pressure, and critical power ratio. Table 3.2-1 summarizes the i comparisons. A detailed description of each of the above four parameters follows. 9 NSPNAD-8608-A Rev. 4 Page 28 of 194 l
. - . . - - - - - . . - . . ~ - _ - ~ . - . - . . - - - _ - . . ~ - _ ~ . -
O Neutron Flux Comoarisons The neutron flux transient is initiated by the main steam line pressure rise due to turbine stop valve (TSV) closure. Normally a reactor scram on TSV , () position would occur at 10% closure of three out of four valves. This scram signal was bypassed for these tests. A pressure wave, due to TCV closure, travels down the steamline and into the core, causing void collapse and a flux increase. The largest flux rise occurs near the top of the core, which has the largest void fraction and the largest void coefficient. The flux increase causes a reactor scram on high neutron flux. The power peaks and turns around I due to the insertion of scram reactivity as well as a decrease in the void reactivity and an increase in the negative Doppler reactivity. For the Peach () Bottom Test conditions, the scram reactivity is the dominant contributor to the flux transient turn-around. This is due to the fact that many control rods are inserted in the core, initially giving rise to a strong scram reactivity. Figures 3.2-114, 3.2-117 and 3.2-120 show the calculated versus measured responses of the relative neutron flux (APRM Channel A from Reference 16) for tests TT1, TT2, and TT3, respectively. The uncertainties in void reactivity () and scram reactivity have been factored out as discussed previously. Therefore, as would be expected, the DYNODE-B results show excellent comparison in the peak flux, flux slopes, and widths of the flux peak. Transient Pressure Comoarisons Dynamic pressure measurements were recorded at the turbine inlet, in the steamline 90 ft downstream from the vessel, the vessel dome, and near the core exit plenum. In all of the pressure comparisons listed in this section, the () data shown are the unfiltered data as recorded by the pressure sensors. The sensors are connected to the appropriate measurement locations by water-filled sensor lines. These sensor lines have their own second-order response which can often give rise to oscillations in the recorded data. Further discussion of the sensor line effects is contained in Reference 16. I Figures 3.2-115, 3.2-118 and 3.2-121 show the calculated versus measured response of the turbine throttle pressure for TT1, TT2, and TT3, respectively. DYNODE-B accurately predicts the initial pressure oscillation in both timing () and magnitude, indicating that the initial time effects; i.e. delays, rise times, and frequencies; are well modeled. As the transient progresses, the calculated wave frequencies are accurately predicted, though the wave amplitudes are greater. The increased amplitude does not appreciably affect the transient results with respect to CPR. The overall magnitude of the turbine throttle pressure is conservatively overpredicted for the latter part of the transients, b) Figures 3.2-113, 3.2-116 and 3.2-119 show the calculated and measured steam dome pressures follow the same trends as the turbine throttle pressures; the initial pressure rise and wave frequency are well predicted, the wave amplitudes are slightly overpredicted, and the overall magnitude is conservatively overpredicted. Critical Power Ratio Critical Power Ratio is defined as the ratio of the bundle power which would () produce onset of transition boiling to the actual bundle power. A good measure of the relative severity of a particular reactor transient is the maximum change of CPR, divided by the initial or steady-state CPR (ICPR). The
" measured" CPR is taken from Reference 9 and is determined as follows:
"For the Peach Bottom turbine trips, the CPR comparisons have been made by driving a hot channel transient thermal-hydraulic calculation with experimentally determined inlet flow, pressure, and fuel heat generation rate.
The pressure input was taken from the core pressure signal, which was filtered () with a 5 Hz low pass filter. The transient fuel heat generation rate was taken to be proportional to the total APRM response. Core flow was obtained from pressure drop measurements taken across four of the jet pumps throughout the three turbine trips. Changes in core flow can be detected by assuming the jet pump pressure drop to be proportional to the square of the flow. In practice, however, this is not an accurate measure of core flow because of the O NSPNAD-S608-A Rev. 4 Page 29 of 194
D. large amount of noise in the jet pump pressure drop signal. In this case, a5 Hz filter was applied to the four jet pump signals to reduce the noise component and then averaged to obtain a pressure drop. The steady-state flow gg was normalized to the recorded flow at the beginning of each transient."
"For the transient CPR calculations driven by the experimental data, uncertainties in the input quantities will contribute to an uncertainty in the ratio CPR/ICPR. (Reference 16) quotes a 22 psi uncertainty in core pressure.
This pressure uncertainty, coupled with a 3% uncertainty in flow, results in a 0.01 uncertainty in the ratio 6CPR/ICPR. This CPR uncertainty is obtained from sensitivity calculations carried out on pressurization type transients." O The calculated CPR is determined from a hot-channel model in DYNODE-B using the GEXL correlation. The hot-channel dimensions are taken from Reference 11. The initial hot-channel bundle power was forced to give the correct ICPR. In both cases, the initial conditions, channel properties, and the CPR correlation are identical. Only the transient forcing functions, i.e., power, pressure, flow, and inlet enthalpy are different, so that a good measure of the CPR uncertainty due to code model uncertainty is obtained. II The calculated versus measured CPR results are shown in Table 3.2-3. For each transient, the calculated ACPR/ICPR is approximately 10% greater than the measured value. This indicates that the DYNODE-B code model uncertainties provide a conservative 10% bias on transient ACPR/ICPR. 3.2.2.2 MONTICELLO START UP TESTS The DYNODE-B code has been used to model six Monticello Start-Up Tests. These GP tests are described and documented in Reference 14. The modeling of these tests was done using best-estimate input parameters. A 1-D kinetics model was used for the MSIV closure and Turbine Trip transients. Point kinetics were used for the remaining four cases. This is in accordance with the guidelines in Section 4.1. The results for each transient are discussed in the following sections. (D 3.2.2.2.1 TURBINE TRIP WITH BYPASS AT 100% POWER (STP 16) DescriDtion of the Test The purpose of this test was to determine the response of the reactor system to a turbine trip. The turbine was tripped with the Turbine Emergency Trip Switch at 1656 MWt. Reactor pressure peaked at 1115 psig, an increase of 105 psi. The M/G set (p breakers were tripped on turbine trip causing a flow coastdown. All four relief valves opened to terminate the pressure transient. A power increase was not observed on the APRMs. Summary of the Test Analysis Figures 3.2-122 through 3.2-127 show the calculated versus measured results i for the Turbine Trip Start Up Test. (g - DYNODE-B overpredicts the core power response with a peak relative power of approximately 300 percent. The data does not show a power increase during the initial pressurization. This is probably due to a faster / stronger scram than was assumed in the analysis. Since DYNODE-B overpredicts the integrated power, it also overpredicts the vessel pressure response. Both the calculated and measured results show that all four relief valves open, but DYNODE-B predicts a peak vessel pressure of 1154 psia compared to the measured value of 1130 psia. gp : DYNODE-B conservatively predicts the vessel flow coastdown and does a good job of tracking level. A " measured" Critical Power Ratio was calculated by forcing the DYNODE-B hot channel model with the measured data. This resulted in a " measured" ACPR/ICPR 4D - NSPNAD-8608-A Rev. 4 Page 30 of 194
I , of 0.003. The DYNODE-B calculated results give a 6CPR/ICPR of 0.156. This is due mainly to the difference in the power response. ) I This transient provides a good benchmark of DYNODE-B's capability to conservatively predict reactor vessel pressure and transient 6CPR. It also provides an excellent benchmark of DYNODE-B's capability to model transient l vessel flow response. 3.2.2.2.2 CLOSURE OF 4/4 MAIN STEAM ISOLATION VALVES AT 75% POWER (STP 11) Description of the Test The purpose of this test was to functionally check the main steam line [) isolation valves for proper operation, demonstrate the capability to perform isolation valve test closures without threatening reactor safety or causing a I reactor scram, determine reactor transient behavior following simultaneous full closure of all MSIV's, and determine isolation valve closure times. The full isolation test was done at 75% power by tripping the relays in the RCICS circuit with a special test switch to give a full isolation and subsequent scram. Following the full isolation at 75% power, reactor pressure ,[) increased 69 psi to 1069 psig two seconds after the MSIV's had closed. ) Summary of the Test Analysis Figures 3.2-128 through 3.2-132 show the calculated versus measured results i for the MSIV Closure Start Up Test. l The measured feedwater flow did not behave as would be expected from automatic controller action. Therefore it was presumed that the feedwater was [] controlled manually during the test and the measured feedwater flow was forced onto the DYNODE-B solution. The measured steam flow shows unexplainable , behavior and was assumed to be bad data. The two most important input I parameters, MSIV closure time and scram time, are unknown and were assumed to l match the nominal values in the DYNODE-B calculation. l Both the measured and calculated results show a rapid increase in pressure due to the MSIV closure. DYNODE-B predicts a faster initial rise than the data {y (50 psi /sec versus 30 psi /sec). The data shows that the pressure peaks at about 1080 psia (15 psi below the relief valve setpoint) and then slowly j decays. DYNODE-B predicts that the pressure rises to the relief valve i setpoint, cycling the relief valves to control pressure. The differences in I pressure response could be attributable to several different factors; the test may have a slower MSIV closure than assumed, a faster scram than assumed, the MSIV valves may not have closed completely, or there may be a steam I condensation effect due to uncovery of the feedwater sparger. In any case, there is insufficient data available to determine a cause and effect.
)
This test does not provide a very good benchmark due to the poor quality of the data. It does show that DYNODE-B tracks the water level very well during the initial pressurization and that DYNODE-B conservatively overpredicts the peak transient pressure. 3.2.2.2.3 2/2 RECIRCULATION PUMP TRIP (STP 14) Description of the Test I) The purposes of this test were to evaluate the recirculation flow and core power transients following trips of both of the recirculation pumps, calibrate the reactor core flow measurement system, and measure the reactor core flow by performing mass and energy balances on the reactor downcomer. Both individual and dual pump trip transients were recorded. For the purposes l of this analysis, only the two-pump trip case was examined, since this l represents a more severe transient than the single pump trip. Prior to tripping the pumps, core performance data were taken to enable the peak heat
) flux and MCHFR to be evaluated. A recording was taken which included a trace of the core flow and the simulated heat flux.
I Summary of the Test Analysis Figures 3.2-133 and 3.2-134 show the calculated versus measured results for 2/2 Pump Trip Start Up Test. O NSPNAD-8608-A Rev. 4 Page 31 of 194
- - . - .- - - - - .. -. .= .- - . - - . -~ -- -. . -. . () - DYNODE-B conservatively overpredicts the vet:.el flow coastdown, i.e., DYNODE-B predicts lower flow than the data. Also, DYNODE-B conservatively overpredicts the core average heat flux. Since these are the only variables available for gg - ' comparison, it is concluded that DYNODE-B provides a conservative prediction of Critical Power Ratio for this transient. 3.2.2.2.4 AUTOHATIC FLOW DECREASE AT 100% POWER (STP 15) Description of the Test The purpose of this test was to determine the plant response to changes in the recirculation flow and to demonstrate the plant load-following capability. To determine the plant response to changes in the recirculation flow, the IIi master flow controller setpoint was stepped a nominal 2 10% of full scale. At each test condition, the test was repeated with several controller settings to aid in optimizing the response of the system. Initial individual loop control settings were at 500% proportional band and 15 repeats per minute reset. The optimized controller settings were arrived at during testing at 50% power, where the proportional band of both loops were set to 450%, loop A reset was set at 40 repeats per minute, and loop B reset was set at 20 repeats per minute. The initial master controller settings were 400% proportional band ID . and 8 repeats per minute based on results obtained during 50% power testing. Instabilities which occurred during flow ramp testing between 75 and 100% power on the 100% power-flow line were corrected by reducing the resets to 2 repeats per minute. The upper and lower speed demand limits were set to 93% and 58% speed, respectively. This placed the limits of automatic and master manual flow control to a range from 75% to 100% power in the 100% power rod pattern. d> : The automatic flow control flow ramp tests were performed with the Electrical Pressure Regulator (EPR) setpoint adjuster gain (POT 150P) at 3.0 psi /t and a time constant (POT 162P) of 20 seconds. To demonstrate plant load-following capability, the fast flow changes were made with the final control system settings described above. The load changes were made first in the Master Manual mode and then in the Automatic Flow QD : Control mode. In the Automatic mode, the load changes were caused by ramping the turbine speed / load changer. Turbine load could be dropped very rapidly by the automatic opening of the bypass valves. l For the purposes of this analysis only the Pump Flow Decrease in Automatic j Manual from 100% to 75% power was examined since this represents the most 1 severe transient in the series. 4 As the flow controller responds to the setpoint step, the vessel flew ramps down and core power decreases due to the increased void feedback. The entire system decays to a new steady state. An APRM decay ratio of 0.25 was calculated based on the measured data. Summary of the Test Analysis Figures 3.2-135 through 3.2-140 show the calculated versus measured results for the Auto Flow Decrease Start Up Test. () The DYNODE-B predicted results show excellent comparison during the first 20 see of the transient, tracking all variables very closely. Beyond this point, the DYNODE-B results deviate slightly. The data settles out to a new steady state condition very quickly, with a decay ratio of 0.25 calculated from the i APRM response. DYNODE-B conservatively predicts a decay ratio of 0.89 and a l longer frequency (approximately 30 see versus 10 see from the data). This transient provides an excellent benchmark of DYNODE-B's capability to dB , conservatively predict decay ratio. 3.2.2.2.5 PRESSURE REGULATOR SETPOINT STEP AT 100% POWER (STP 18) , Description of the Test l The purpose of this test was to determine the reactor and pressure control system responses to pressure regulator setpoint changes, to demonstrate the 9 NSPNAD- 0608-A Rev. 4 Page 32 of 194
D stability of the reactivity void feedback loop to pressure perturbations, and to optimize the pressure regulator setpoints. Pressure setpoint changes were made'with both the Electrical Pressure Regulator (EPR) and the Mechanical Pressure Regulator (MPR) to determine reactor and turbine system responses and to demonstrate the stability of the reactivity void feedback loop to pressure perturbations. The pressure disturbances were ebtained by changing the regulator setpoint downward and then upward as fast as possible to produce a nominal 10 psi change in reactor pressure. This was done with the load limits out of the way ) (load limiter set well above the reactor power level) and repeated with the load limits incipient (load limiter set at the reactor power level). The changes in the EPR setpoint were made from a special test circuit located in the cable spreading room, which initiated a step change in the setpoint. For the purposes of this analysis, only the -10 psi step of the EPR setpoint at 100% power was examined, since this represents the most severe transient in ) the series. The pressure controller responds to -10 psi step by opening the turbine control valves to drop the turbine throttle pressure by 10 psi. As the pressure drops the core power drops due to increased voiding. The entire system decays to a new steady state with a period of approximately 8 seconds. The stability of the reactivity void feedback loop was clearly demonstrated at ) this test condition. Summary of the Test Analysis Figures 3.2-141 through 3.2-145 show the calculated versus measured results for the Pressure Setpoint Step Start Up Test. The DYNODE-B predicted results show excellent comparison to the test results. The predicted results exhibit approximately the same period with a slightly
- ) greater amplitude. The reported APRM decay ratio is zero, based on the measured data. DYNODE-B calculates a decay ratio of 0.21 for relative power.
This transient provides an excellent benchmark of DYNODE-B's capability to conservatively predict decay ratio. ; 1 3.2.2.2.6 FEEDWATER CONTROLLER LEVEL SETPOINT STEP AT 100% POWER (STP 20) I Description of the Test ) } The purposes of this test were to determine the effect of changes in ! l subcooling on reactor power and steam pressure and to demonstrate that reactor responses to changes in subcooling are stable at all power levels. The changes in subcooling were introduced by varying the vessel water level setpoint (3 and 6-inch changes), and the resulting transients were recorded. Testing at all power levels, in three-element and one-element level control and automatic and manual recirculation flow control, yielded stable plant () responses to changes in subcooling. Decay ration of primary variables were less than 0.25 for all of these tests. For the purposes of this analysis, only the 6 inch level step at 100% power in three-element control and automatic flow control was examined, since this represents the most severe transient in this series. The feedwater controller responds to the 6-inch setpoint drop by cutting back Q the feedwater flow to attempt to balance the level error. As the feedwater flow drops, the core power and vessel pressure drop slightly due to the decreased inlet subcooling. As the level drops, the controller error goes to zero and the feedwater flow and all other core parameters return to their original level. The stability of the reactor in response to subcooling changes was clearly NSPNAD-8608-A Rev. 4 Page 33 of 194
. . .. -. _ . . . .. ..--. =--... .~.... - .~.-. -.,.--.-~ ~.. .. .....~ .~ -
() demonstrated at this test condition. Summary of the Test Analysis Figures 3.2-146 through 3.2-149 show the calculated versus measured results I. for the Feedwater Controller Level Setpoint Step Start Up Test. The DYNODE-B predicted results show excellent comparison to the test results. The predicted level response and feedwater flow track very closely. Both DYNODE-B and the data show a decay ratio of zero for the core power. There are no significant deviations between the predicted and measured results. This transient provides an excellent benchmark of DYNODE-B's capability to II predict decay ratio and to track water level. (D . q> i i 4D i l l I 4 l G J l l 9 i i NSPNAD-8608-A Rev. 4 Page 34 of 194 ! _.. . . . _ . . . . _ _ l
. . l
') l I i i Table 3.1-1 NDH - DYNODE-B VOID REACTIVITY COMPARISON , l f3 Dome Pressure Subcooling K,, t/ (osia) (Btu /lbm) NDH DYNODE-B 1038 24.42 0.99833 0.99811 1238 53.20 1.01991 1.02003 Reactivity Change = 0.0212 0.0215 AP = (K i -K) /K i K 2 2 l l C) Table 3.2-1 PEACH BOTTOM-2 TURBINE TRIP TRANSIENT TEST ACTUAL CONDITIONS Test Reactor Power Core Flow Rate Core Pressure , Number (MWt) (% Rated) ( 10'1b / h ) (% Rated) (osia) l TT1 1562 47.4 101.3 98.8 1005 l TT2 2030 61.6 82.9 80.9 995 l C.) TT3 2275 69.1 101.9 99.4 1005 Test Core Inlet Enthalpy APRM Trip Set Point l Number (Btu /lb) !% Rated) TT1 528.4 85 TT2 519.8 95 TT3 523.6 77 O Table 3.2-2 PEACH BOTTOM-2 TURBINE TRIP TRANSIENT TEST PEAK MEASURED AND CALCULATED RESPONSES Variable TT1 TT2 TT3 Date DYNODE-B Data DYNODE-B Data DYNODE-B e) Average Neutron Flux (% Rated) 239 239 280 281 339 342 Core Exit Pressure (psia) 1036 1053 1034 1066 1072 1088 Reactor Vessel Pressure (psia) 1031 1047 1038 1062 1061 1082 Data from Reference 16 o# PEACH BOTTOM-2 TURBINE TRIP TRANSIENT TEST CRITICAL POWER RATIO Table 3.2-3
RESPONSE
Variable TT1 TT2 TT3 Data DYNODE-B Data DYNODE-B Data DYNODE-B ICPR 2.536 2.536 2.115 2.115 2.048 2.048 6CPR 0.431 0.474 0.288 0.315 0.270 0.305
- gg ACPR/ICPR 0.170 0.187 0.136 0.149 0.132 0.149 Data from Reference 9 I
l C) lO NSPNAD-8608-A Rev. 4 Page 35 of 194
, . 4 I
O 1 4 1
. Monticello Cycle 10 -
NDH - DNB Comparison
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h\onticello Cycle 13
. NDH - DNB Comparison i
Rgure 3.1-3 DYNODE-B i Delta Rho NDH
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I n U j NSPNAD-8608-A Rev. 4 Page 37 of 194 i l
O Monticello FSAR Benchmark . ! Turbine Trip w/o Bypass e Figure 3.2-1 DN8063/86 Stoom Dome Pressure FsA.n.. .. soo
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o Monticello FSAR Benchmark o Turbine Trip w/o Bypass Figure 3.2-3 DNao63/s6 Cor AveroSe Heat Flux _FS_A_R_ _ _ _ . 12 0 m)
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O Monticello FSAR Benchmark - Turbine Trip w/o Bypass e DNB063/86 Figure 3.2-5 McIn Stoom Line Flow _FS_A_R_ _ _ _ . 150 l
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- 4-NSPNAD-8608-A Rev. 4 Page 40 of 194
Monticello FSAR Benchmark - n Turbine Trip w/o Bypass Figure 3.2-7 DN8063/86 Sensed Level ps_A_n_ 29 i : : : :- : 3 . O s'ss . : : : : J 'N ' j : ! ! :, ,- -
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E. NSPNAD-8608-A Rev. 4 Page 41 of 194
C-Monticello FSAR Benchmark ~-
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Figure 3.2-8 DN8070/86 Steam Dome Pressure _FS.A.R_ _ _ _ . en i i i : :
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1 Monticello FSAR Benchmark l J Turbine Trip w/ Bypass ) l D N B 070/86 l Figure 3.2-10 l Core Averas. Heat Flux F5_AJ,___, 15 0-l l 1 1 i,0 g..' ..... .. . s .. . . . ..s.. . . . .s... ..
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o; i Monticello FSAR Benchmark - Turbine Trip w/ Bypass e DNB070/86 Figure 3.2-12 Main Stea:n Line Flow -.AR-----. FS Iso g i l : : ; .
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r . . o Monticello FSAR Benchmark - a Turbine Trip w/ Bypass DNBO N f6 Figure 3.2-14 4 Sensed Level _FS_A_R_ . . . . ] 20
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0 ' a NSPNAD-8608-A Rev. 4 Page 45 of 194
o Monticello FSAR Benchmark Generator Trip w/ Bypass e DN8049/86 Rgure 3.2-15 Sensed Level F.sAR 100
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e NSPNAD-8608-A Rev. 4 Page 46 of 194 1 _____________________1
(. Monticello FSAR Benchmark ) ' Generator Trip w/ Bypass DNB069/86 i Rgure 3.2-17 ] Cor. Inlat Flow F5_AJ_ _ _ _ . I$o - ) : . : :
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3 )- .! 4 NSPNAD-8608-A Rev. 4 Page 47 of 194
Monticello FSAR Benchmark - Generator Trip w/ Bypass
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e : . O r. b,--- -----.- - ..___ 0 0 2 4 6 8 10 l'2 14 le Tim. (s.c.adi) O, D N B 069/86 Figure 3.2-20 Feedwater Flow FS A.R. - - - . IS O SQQ-dA- * . *
. s , s * .~.%.* .
g . s s: .
. %. s s
s .iG x g N s 94 *% s s-
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, Monticello FSAR Benchmcrk -
Generator Trip w/ Bypass DNB069/86 Figure 3.2-21 Vessel Stoom Flow psAn..... i 150 * '
,) . . .
- : i i 0 . . . . . . . .. . . : .i. . . . .:: . . . . .: .i 4
i \
)M l
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.--2a**t**+++****#.***********t*********t********
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) : :
;---- s.-----+-----
o , , , . 4 6 8 10 12 14 le ; 0 2 ' Time (Seconds) 3 4 O O O O NSPNAD-8608-A Rev. 4 Page 49 of 194
o Monticello FSAR Benchmark 100% MSIV Closure
- DNB064/86 Rgure 3.2-22 Steam Dome Pressure FS,A,R, _ _ _ ,
300
- : i l : ,
i e 100- **- ++4+- 24 * = -4
- 4 * * * * * * * * - '4-*-******* *4* * '*
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e . 'h , m . I
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' s! ! O E : : P= y,_,--.
o : . .-- . c . . U : . 0 . . i- i i c:
-too . . ,
0 2 4 6 8 10 12 14 le Time (Seconds) e. DNB064/86 Figure 3.2-23 Reactive Power FS.A.E. _ _ _ . iso 100 . . - . .
} ' '
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en
= o 2 4 a a lo #2 w la ; Time (Seconas) .c e.
n.s g-NSPNAD-8608-A Rev. 4 Page 50 of 194
l Monticello FSAR Benchmark I D 100% MSIV Cosure ! l Figure 3.2-24
\
DNeo6us6 l Core Average Heat Flux p gg i iso _s______. D -
- +
son- ~ . . :. . . . . ... . .. . . . :. . . . . ;. . k s M : 's s * ) : l
% =
so- .. . h. ....i. 1 ., ,... .i.. , . .;.. ..4.. 4__ _4. , __ 6 - o 8 2 4 e a 30 is (4 se Time (Seconds) D Figure 3.2-25 DN806u86 Core inlet Flow ps4, eso -------. 9 l . ___ ico 'V'
- 1 H
~
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- 5. ui- .!
t I
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o Monticello FSAR Benchmark *
- 100% MSIV Closure
- Figure 3.2-26 DNB06&86 McIn Steam Line Flow .F.S.A.R. . .
150 i . . O i . 100- * -=- = = = *
- 4.=***
,-4* ,
. - = 4. -
*
- 4- ***
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u ( .ai'[ :
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=/ *
.-i . : a W
g l
- : ---r-- :
o . . . o 2 4 e a 30 12 14 le Time (Seconds)
*l ,
i Figure 3.2-27 DNB064/86 Feedwater Flow F.s A.R. . . . . I iso l e\ I so0 a.:. , ., , .. . .j . .
..l
- g '
g M N ce i O
- i I
E a so-,
. . . . s I C i-
= .i
> l w
i 2 i u .e sa a e S s l i w 0' I
- o 2 4 6 a in i2 i4 ia
= Time (Seconds) l c
w 2 s Q1 NSPNAD-8608-A Rev. 4 Page 52 of 194 1 l
L. o Monticello FSAR Benchmark 100% MSIV Closure m J Rgure 3.2-28 DNB06#86 Sensed Level A J, , , _ , 20 FS,,
- i 1
i g
- 0 -
1
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. -
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E9fP 0 2 4 6 8 10 l'1 14 le Time (Seconds) m J O k O < E e G O z *
.c d .c w
C_ O 5
.f G E v
NSPNAD-8608-A Rev. 4 Page 53 of 194
o Monticello FSAR Benchmark
' Feedwater Control Molfunction (Max Demand)
- DNB065/86 Figure 3.2-29 Sensed Level Es_^J. _ _ _ .
100 l l .
. . . 1
- : 0 i
go. . . . . . . . . . . . . ...i..... . ..;.. .. .... . .i......... . .
. . . : i e j
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o is r'o 3o 4o Time (Seconds) O DNB065/86 Figure 3.2 e.0 . , Steam Dome Pressure Es_Ay, _ _ _ , l 200 <
. i 9'
]
l . r l 10 0. , . . , I s 1 v' t \
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.a l 'e l e ? O # NSPNAD-8608-A Rev. 4 Page 54 of 194
[. . n 1 l l Monticello FSAR Benchmark i e Feedwater Control Failure (Max Demand) \ I l l D NB 065/86 ; Figure 3.2-31 Core inlet Flow ps42 l 100 l
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73 . . ....). . .
..g. . 3 . . .
l
. 1
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. ~ . . ..
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... .i. .
1 l 3 i l l o . . . O 10 20 30 40 1 Time (Seconds) ' l l 1 Figure 3.2-32 DNB065/86 Average Surf ace Heat Flux F,5,A,2, , , , , 100 3 J 15- s .N l l s , , M 30-
\
L le h e I (' .
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l
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, m, .
o 20 30 40
.a Time (Secon=s) # m ?
NSPNAD-8608-A Rev. 4 Page 55 of 194
~
O Monticello FSAR Benchmark - e Feedwater Control Failure (Max Demand) DNB065/86 Fi ure 3.2-33 Re$ative Power FS_A_R, , _ _ , ISO
- ,i e
,00 . . . . . . . . . . . . . . . . . . , , . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
,' I i
m L-s : :
. i $
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- I : :
i i :
- i . :
s . .
\ I -
e s 0 0 l'0 2'O 3'O 40 Time (Seconds)
@<1 I
i DNB065/86 Figure 3.2-34 j Feedwater Flow FS_A_n _ _ _ _ . l$ 0 e
.00 . . . .;. .
M .
. 4 4 f T e g $ 1 N .
vi 5 0.l. . . 3 4
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e o - e k
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- c. O, O O 10 20 20 de s Tim e (Seconos) l
- l I
".'S '
s I el C l NSPNAD-8608-A Rev. 4 Page 56 of 194 '
e Monticello FSAR Benchmark -
)
n Feedwater Control Feilure (Max Demand) l Rgure 3.2-35 DNB065/86 l Vessel Steam Flow p34,
-------- l iso I ')
- I l
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l l 4 . .} . p a _..--.__J.., l I g , o o l'0 2'0 30 40 l Time (Seconds) ' O b) _) e s e o E O i k
=
2
) I C NSPNAD-8608-A Rev. 4 Page 57 of 194
o. 1
.Monticello FSAR Benchmark
- Loss of Feedwater DNSO46/86 Figure 3.2-36 Sensed Level FS_A.R.
so
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op. .. ..<... . j , 20- * -l '
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-so . . . .
i4 se o 2 4 6 to it Time (Seconds) O l i DNB066/86 Figure 3.2-37 Steam Dome Pressure F.S A R--- .. I so 1 1 e: l m Cl. ! e - 1.. e '. ' b . l O.
. m . .,, . e -
s s a , - e e 3 s o v D -s c I -
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% Tim e (Seco n= s) 4 l
3 e 4 NSPNAD-8608-A Rev. 4 Page 58 of 194 l
1 I Monticello FSAR Benchmark - 3 Loss of Feedwater DH8066/86 Figure 3.2-38 Core inlet Flow _FS.A.R IS O l 1 i : ; : i
- : . . 1 a00 *** *=***
' ****.*********i'**** *.,******i** *
) x
's' 50- *********<*******t** * ** ******'*}**
**** *<***********N:
. - . . 1. -
) : . o . , , . . 0 2 4 4 8 to 42 14 16 Time (Seconds) ) DNB066/86 Figure 3.2-39 Average Surface Heat Flux F S,A,R, , , , , is 0 J . 10 0 , --- - ~ . . x - I ] * '
'* 4..;%~
E o
> 30.1. . .; . . %g s:
4 3 2 ! , e C l 4 ) i
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!C@
- n. O 4 14 16 O O 2 4 6 to 12
; Time (Seconce) .a 3 L 4 NSPNAD-8608-A Rev. 4 Page 59 of 194
. . 1 o
Monticello FSAR Benchmark Loss of Feedwater Figure 3.2-40 DHa o66/s6 Relative Power F.S.A.R. . . . . ISO
! ! . i : i e
l : :
- l . .
100- 4- * * ***+4***** *a *4 - **********4- * * *
,-*****,~ ,
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= . : . ! #
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.s
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. . '. ?
o . . o 2 4 6 a to it 14 36 Time (Seconds) O Figure 3.2-41 DNeo66/s6 Feedwater Flow r.s.A.n. . . . . G Doom - r * * ** * , = .s- .
,I .
e M
- 1 .
O O N
@ b i \ .
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m o} [ nip 0 2 d 6 9 to 11 14 16
; Tim. (s.conas) o l =
t O , NSPNAD-8608-A Rev. 4 Page 60 of 194
o Monticello FSAR Benchmark - a Loss of Feedwater Figure 3.2-42 DNao66/a6 Vessel Steam Flow FSAR tso O ! : i i
.i .
10 0 - +
,-----.." *"*."*~"a "" " " " "4" * ".
4". y g i
,'5,*.,'s
()
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go. . . . . . . . < . . . . . . . .p..
.. ..<..........t........<...........g........<........
O : - i
- : . . . ggip D .
0 2 4 6 8 to 82 14 le Time (Seconds) O O e O . N 4 o s 2 e e
.^
O M A e
=
m w i O " NSPNAD-8608-A Rev. 4 Page 61 of 194 a
a Monticello FSAR Benchmark
' Loss of Feedwater Heating
- DNB052/86 Figure 3.2-43 Sensed Level _FS_A - R. _ _ _ .
100
- : i :
i : i i ! ! : 8 00 . . . . ..<........3... ... ..:.. .. .....>.. . . .<...........;. ... .. .<.. .... .. i i : : i E . : O S 0 _
=> : : .
Jv :' ~ 00 ..<.......>..
- g--
-100 . . . .
0 20 40 40 00 100 12 0 140 140 Tim e (Seconds) O DNB052/86 Rgure 3.2-44 Stecm Dome Pressure FSAR.___. n00 e 004. ., , i. l t/5 . .
- a. * ,
e e 0 g e e ci c IV s o en .C a v s, t i i
- , 004. . * .>.. ..;.. .s.. .
1 e I s 1 3 i G
! nip
- m. i00 C 0 70 80 60 to 10 0 l'0 '40 140
', 'im e (S co ne s) .a I.
e 1 4 NSPNAD-8608-A Rev. 4 Page 62 of 194
l p 1 1 4 Monticello FSAR Benchmark l J Loss of Feedwater Heating l 1 Figure 3.2-45 DN B 052/86 Core inlet Flow psAn O i i 1 l
. \
1 i . i l 100 ---,----- ----------- --- .---------- --- --
. l l
. j .
J .
- : . . : l
.0 . . . . . 4. . . . . .; .. , ..;. . . ;. . .;. . .;. 'l 0
0 2'O d'0 $0 B'O 1de 12 0 140 150
- h. (s.co nds) m a
Figure 3.2-46 DNB052/86 Average Surface Hect Flux FSAR,,,,, ise m . J
. ,j 100- ' . s . ,
I n n . Q . 6 l A 1
~
s v 50- 1 1. s 4
..(
.: 1
-l: :
y < s E j q .* v 3 g 0 g 0 0 20 40 60 to 100 120 64 0 160 1 a Tim. (S condo
.O v .&
C NSPNAD-8608-A Rev. 4 Page 63 of 194
o Monticello FSAR Benchmark Loss of Feedwater Heating
- DNB052/86 Fig u r, 3.2-47 Relative Power _FS_A_R_ _ _ _ .
tso
- I
. i. .
. i. G 125- - ++ t... .++e*t--**+-- *++l- + + - .-
................e----.......
,V.'~,,,,,.~.. ,,
,e. .
100- ~- - '. - - * *= 4-
- < * * - oc -* - - *4****-*
g pg. . . . . . . . . . . . . . . . ..... . . ..,.... . . .. . . ...... . . . . . . . . . .. . . .- . . . . . . . 50- -*<
* - * * * - + + * -
-- +#.- *-+.t--***<**=t*
. .t- . . . .
2s- -
----t---*-t- -
9-0 . . . . . . . 0 20 40 60 a0 100 120 i40 se0 Time (Seconds) O DNB052/86 Figure 3.2-48 Feedwater Flow F_S_A_R_ _ _ _ . 200
- e.
150-
- i 1-
- k
- e a ,00-l,. - . ....:---
e. e N n I . g N v I
$01- - -
r ,
, t
- a. A
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- c. O C 0 to 40 60 80 10 0 120 84 0 1e0
~,. Tim. (3.co nsi) .a e = NSPNAD-8608-A Rev. 4 Page 64 of 194
. Monticello FSAR Benchmark -
> Loss of Feedwater Heating Rgure 3.2-49 DNeos2/s6 Vessel Steam Row psre. .
,3 ,
h , 8 . . ... - -----: . - 100- M n:nMS_. .*.e.. ....v. . i. -...i.
- i : : :
, x 00 . . . . . . . . .
0
) . .
. . ggs 50 .
0 20 40 40 80 100 120 14 0 14 0 Time (Seconds) e f4
-e*
~.
e C I i 6 l M 1
=
m
.3 s.
) [ a ggpNAD-8608.A Rev. 4 Page 65 of 194
Ol l 1 l 1
.Monticello FSAR Benchmark -
Pressure Regulator Fails Open *j DNB071/86 Rgure 3.2-50 Steam Dome Pressure Fs_A_R_ , , , 200
- j g iOO. . . . . ...... .. . . . . . . . . ...... ..>.............................
m ,~~ s . a : , r,, .. e : '
- E 0 ,
8 Oi e : ' ;
.c : ' ; :
U % . g . .
- ;f *
- s. .
.i00 .. .. ... . . . r. . ... .
.3..............i...
i O
-200 0 l'0 20 3'O 40 Time (Seco nds)
O DN'O ** Fi ure 3.2-51 Re1ative Power F_S_A_R_ _ _ _ 150 1 g l l 00 1 . .; . . . . s
's .,
' l j
e s . R H ' \ h
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- s O
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' . . , 's Ggip !
2 0 to 20 30 so
; %. (Seconeo o
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i e
. Monticello FSAR Benchmark -
- O Pressure Regulator Fails Open I
a j Figure 3.2-52 DNBO N86 j Core Average Heat Flux ps.A_g. __ 1
- : i j0 -
a iOO. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ...... . . .... j ' -
- : l
, 1 g
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'i -
a
- s :
t .s . .
- s .
1, . s . .
.0 ...............,.....s...
s, .......y.............................. s - s, o o, 0 l'0 20 3'O 40 Time (Seconds) O Figure 3.2-53 D N B071/86 Core inlet Flow p54.g.... O : 100 _ , . b.... 3 s .
. , , ,, M e . .:s . s Q {
c. M
-h
'% h
. ~..
y 50 - . i.. .. *?.......,,
>, . I Z *
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us s.
- u. .
O <
?
. 0 5
0 to to so 40
$ Time (Seconos) .
o E O b E NSPNAD-8608-A Rev. 4 Page 67 of 194
o Monticello FSAR Benchmark Pressure Regulator Fails Open
- DNB071/86 Figure 3.2-54 Main Steam t.ine Flow FS,AJ,,,,
IS O j : g
. : i 100- - * . - - .- -- - * * * - -
C . g x ; : : g a : . 1 . lt i ' 50- - y- -- - -
-[t
- - t- * * -
lt : :
- s :
- s : .
s . {%t ,/~~~'f9, .
$i
-[
5, , s
--l j [ l----{
- o. .
o to 20 30 4o Time (Seconds) i l l O i i i D N B 071/86 Figure 3.2-55 Feedwoner Flow F.S.A.R. . . . 150 1 Q
^
l 100 . , - - 'a - -i - . g-
/
~
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/ ,
i i* g! f h 'g s 3 i
^
,; t \J ', l i - 50 e a s J z I < 1 = e l UEP ., 0 = 0 10 20 30 40
; Time (Secones) s C
E NSPNAD-8608-A Rev. 4 Page 68 of 194
Monticello FSAR Benchmark
)
Pressure Regulator Fails Open Rgure 3.2-56 DNB071/86 Sensed Level FSAR
.-so . ...
,-4 : . i
- s
*s :. : .
- t .
) o s . . t i : : I
- a. t a . s, .
. I e .
s j * * * * * * * * * * * * * * * * * ** ***g** ********* ****************************d*
- c : g j j j c . \ :
- 5 : :.
e s # m : % . ,# c
- s "
g *40- ******************************\********!***,"*,'*-www=*--*>*** * * * * * * * * * * * * * * *
.c . s' ;'
W . ": 6o. ...................,..................e...................s.................
- : i
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- '~
o to to 30 4e . Tim e (Seconds) i
) ) ) )
1
)
NSPf3AD-8608-A Rev. 4 Page 69 of 194 I
-ol l l
I l
. Monticello FSAR Benchmark Recirculation Pump Seizure *)
l DNB067/86 Figure 3.2-57 1 Sensed Level FsAR 40
- i i O
- l
**i*+*a ***i+-***.-
10- a. - a. * +-
- i . - **** *
.- - t. - - - -
- - .: . ~
,r, -
- a. ,
J ' 5 o'
, O
= :
e C. : e : : : u . . 2o. - ....i. . -- .i..-i....- -.- g
-40 ,
o 2 4 0 T:m. (secondo O D N B 067/86 Figure 3.2-58 Steam Dome Pressure psAn 600
- 9 50-* * . . .
a ..4 Jo . c e 3 . 0 O s ci i ' ko o c 5 '- , 3 V f .' N , * . ,, ,
- l
* [
T .s o j. . . . . ..:. s, N ,~'
- a z ~s < l. - ,e e d
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-100 I '
!MiP
= 0 2 e a e
; Tim. (S.cocc0 oa
*; S:
E NSPNAD-8608-A Rev. 4 Page 70 of 194
. Monticello FSAR Benchmark i
) Recirculation Pump Seizure DNB067/86 Rgure 3.2-59 Core inlet Flow Fs,Ay,,,,, 15 0 l ) l io.. ... .. .. . . . . . . . . . . . . . . . . . . . . .... ..
...3.... . ...
) x ! !
- I s . .
s : : : _________4_____...__.w----------- 3t. . . . .N....... g ..
. ......p
. . l
) . o . . . 0 2 4 6 e Time (Seconds) l l ) D N B 067/86 Rgure 3.2-60 Average Surface He=t Flux F,5,Ay, , , , , ISO 9 oo. . . . . . . ) e x --- i
- N 1 5 '- - .....,:
s ..l e so. . 5* h w l *
* }
i ) $ e
- a. 0
- O 2 4 a a
= Tim e (Se con =s)
.a e
s
- NSPNAD-8608-A Rev. 4 Page 71 of 194 n.
o Monticello FSAR Benchmark - Recirculation Pump Seizure DNB067/86 Fi ure 3.2-61 Rebative Power Ps.A.R. . . . . . 150 e i i i 100- - * * * * - * * + * * - - -- --- . - - - - -
- i e
t x ', : : i : . s : : 30 . . . . . . . ... ..>. . ... ...... ... ..).. . ..........g..
. . ... I s \
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s
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s s , .
, n.
. : 9; 1
o . . . 0 2 4 6 4 Time (Seconds) O D N B 067/86 Figure 3.2-62 Feedwater Flow F.S.A.R . . . . . ISO
- . 9 I . .
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4 N e
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/ M e.
g o 2 4 6 : x Time (Seconds) .a 1 . O 6 t NSPNAD-8608-A Rev. 4 Page 72 of 194
)
Monticello FSAR Benchmark Recirculation Pump Seizure DNB067/86 Figure 3.2-63 Yessel Steam Flow FS
- A.R..--.
iso j : .
) -
s : :
^~~- ~- ~- - ~* . ~ - ~ '
100-r---^.
' ~,,_'~~; .
i N - ) M
) .
g o . 2 4 6 8 o Time (Second0 ,D 3 J e. h a e _=, . E
) 2-
_a e e a. m w J
) $
a. Page 73 of 194
- MSPNAD-8608-A Rev. 4
. O Monticello FSAR Benchmark ,
2/2 Recirculation Pump Trip DNB073/86 Rgure 3.2-64 Sensed Level FS_AR 100 e
.. ...........i.,...........i............i... -
m . . . m . . . x . U . z ------ ------- j . - .----------.. g e i z . x v : : i
.,C. ................... . .. .. . . . . . . . . . . . . . . . . . . . . . . . . . .. ....
- i :
i - e
-10 0 < . , .
O s to IS 20 TIME (SECON DS) e^ D N B 073/86 Rgure 3.2-65 Steam Dome Pressure FS,A),,,,, 50
- g 0 --. ,
\ .
a ', . l m i , . 2 1 % g,- . . 30 1. . . . . .; % m .
- 8 e c ,
l
,a v . , 4
.".. *. t = '
- i 5 .iOO. .,,.....-.-.,i 2 a f-
> a C ! g. m g -85 0
. 0 $ 10 15 to - Tim . (S. cone s) a o
E 2 e: U E NSPNAD-8608-A Rev. 4 Page 74 of 194
D a Monticello FSAR Benchmaric ~- 2/2 Recirculation Pump Trip
\
Rgure 3.2-66 D NB 073/86 Cor. Inlet Flow ' l FsA.R. ) ! i j -
?30 . .... . .. ... ..... . .. .....
i M ! .
. ~
50- .. "" .." +"t. "."." '
" * ~ v;;' & ~ . ' '
."""l"".".""""
- . . . :- - - . . . . . . I 1
] l D 0 s (0 is no Time (Seconds) D Rgure 3.2-67 D N B 073/86 Average Surface Heat Rux eso rs,Ay,,_,, ] i l
.0 .. .. . .. . . .
) *
" l ?.
e .
.\ '
s00* . m 1'
- t tv....' '
< ]
z I. . 3 - 4 S 3 m
' 1, a
z c O S 5 30 IS to C,* Tim e (seconos) oa M G " E NSPNAD-8608-A Rev. 4 Page 75 of 194
o; i Monticello FSAR Benchmark . 2/2 Recirculation Pump Trip DNeo73/86 R ure 3.2-68 Re$ctive Power Es_Af,__.. ISO
. i 9 i : :
100- + * * - u : : : i . : . i : : : s : .
.0 . . . .g. . ... ..,...... ........ .... .>..
.. s s : :
's ' . :
., , , -l O
0 , . 20 O s no es Time (Seconds) O DNs 73/s6 Rgure 3.2-69 Feedwater Flow gS A), , , , , ese .
- : 9 a
io . .. ', . . , . . . . ..I s
's N
- O
\ \\.
i w s s 1
, N:# N l C
3 0.l . .. . . si
. .s . . . . .... .
.. .. .....I
-.: 4
-'s .
-1t 2- s . I e s s
. .." g e , *,- g
.. - l q
}
g
- a. Of O 5 10 15 20
*; Tim. (S. con = s) ; 9 a.
- NSPNAD-8608-A Rev. 4 Page 76 of 194 m.
e a ' Monticello FSAR Benchmark - 2/2 Recirculation Pump Trip DNB0W86 Figure 3.2-70 Vessel Steam Flow FSAf,,,,, 15e 7, ios _ .
~~. . ) ? ' 'b '
x ,
's, .
,o. . .. .
e e %
' g~~. .,_
3 : . o . . 5 o s to is 20 Time (Seconds)
)
T r's
)
0 1 E. N e :
=
I O ! I NSPNAD-8608-A Rev. 4 Page 77 of 194 2
O Monticello FSAR Benchmark - e Recirculation Controller Fails (Increase) Rgure 3.2-71 DN 8072/86 Sensed Level _FS_A_R_ _ _ _ . loo O co- -- - - - + - - u.*l-- -
.c u
- c
~,~ -
- y. - ()
e m ~~..:: ' : g - s, : , x ,. : v : 50- -- *- ***b*****--*t- - * ** * - *** I- *- * * *
- 5 5
-too .
o to 20 to 40 Tim. (s.c.nds) O Rgure 3.2-72 D N eo72/s 6 Stvom Dome Pressure F.S_A.R. . . _ . too S so- - - H- - j- --
]- - -
i G a. 5 O1 e e m o, :^ g R - e E .N s 1 q v \s- __ 3 i _; ) ..]
.S o 4- - '-t- -
- i- -
- - 4.
. I = l 1-1 i' S
0 s -ico ' S O o lo to 3o so ' ~, 'im e (Seco nes) ' a
- 9 9
$ NSPNAD-8608-A Rev. 4 Page 78 of 194 h
o i l o Monticello FSAR Benchmark i Recirculation Controller Fails (Increase) i 1 DNB 072/86 Figure 3.2-73 { Core inlet Flow _FS A.R. _ _ _ _ . 100 i
- : : (
C) :. i i re. ........,..<.g....,.,n-.a,i;................;.;..................,.,
~
~~~. - l 4 . .
1 1
- : j 0 M SO=
* * =* *=*= **
=== **** = = = = * = =
*-*=*=:***==***=***** l
- .i :
- 2. . . .
O : : .
. . l 0 . .
1 I O 10 20 30 40 I Time (Seco nds) O D N B 072/86 Figure 3.2-74 Average Surf ace Heat Flux FS,A,R, , , , , 100 O . rs.. ,. .p . .i. .;. . i
.-.,.w s
O x so. .
\
.l l
. \s r
- s: .
s e \.v.'g s 9 i l m* s g 23 I. . . .: . . . ... . ......
! \ %., '* a I 4 j ,
- i 0 -
1
'. w'- ..
--....----.. y
- a. O l
T 0 80 20 30 40 m Tim e (Seconas)
.a
, l O i I NSPNAD-8608-A Rev. 4 Page 79 of 194 2
o Monticello FSAR Benchmark e Recirculation Controller Fails (Increase) Fi ure 3.2-75 DNBO72/86 Re1otive Power F_S_A.R. _ _ _ _ . 300 g ISO- - ** * * * * * 'l** ** - * ** -
- I- ** -- * **3* * * * * * -
- I . .
i . . . i P y
- q L
x 00 .. . l I ; , ; f t l i . . ' s0- -v- - .*s* - . >- .. - -*l- - *- -
' i i* O!
s, -- 0 , O :D 20 30 40 Time (Seconds) ei! l l i Figure 3.2-76 D N B 072/86 l Feedwater Flow F_S_A_R_ _ _ _ . Ist h' 125- + t*** * --)*- -* ** - 100-
*n
.) . e . o e g- . .
I i s i O x rs. . 3 l $.. . s e
\ ?
5 so- .s.
~
t
. . .u' . .
I
\'3 ..i -
_,= . E C
\
< :s _. h.
e j g 2 j -s
- a. 0, 8
', \ @
0 0 10 to 30 40 s
.a Tim (Seconed
. O a
2 NSPNAD-8608-A Rev. 4 Page 80 of 194
o 4 0 Monticello FSAR Benchmark - l Recirculation Controller Fails (Increase) DN B072/86 Rgure 3.2-77 Vessel Stoorn Flow ' FS_A_R loo O .
. : i l
. . i 75- -- -l- -- l- l- - - - - i 5 $ $
g : : .
% . i
() M So- - ~
-~ ~~
i ~~"
; ~ "- -
g : . .
\ ! I s
s: : s: . . v . : ss- - ~- 4.s~ ~
-~ - - l~ ~ - ~~~ ~i ~- - -
O : , . .
- 's . :
o
. ______. __ h. E o io 2's so 40 Time (Seconds)
O O O e 2 R e b
=
E O i A wm 2 m O i. - m NSPNAD-8608-A Rev. 4 Page 81 of 194
o Monticello FSAR Benchmark e Recirculation Controller Fails (Decrease) Rgure 3.2-78 No FSAR Data Avollable Sensed Level D N B 074/86 to
- e e
e . : e . . J . : i ! O 3 . . . .' . . ........ .... ...;. . .......... .......;.. ..... ... .. .....j. ..... . .. ... .. m : : ; e : o . d :. :. : U -
...................................3.....................
-40 . .
o 10 20 so ao Time (Seconds) e-Roure 3.2-79 D N B 074/86 Steam Dome Pressure F.S.A_R. . . . . 100 e so. ..>.
. . i, . .
G. en . . . S e e m ~%' e e % . . N e c 'k
.c .*
N s e y -.*. .* ,
;\ %'
. *~.............. ..........-
=S0 * * * * * * * . * .r. ,
~e
= 1
< a l O-J a. 0
-100-1 0 10 20 l@
l 30 so .a Ti,n. (Seco n ce) s a. e:
- a. NSPNAD-8608-A Rev. 4 Page 82 of 194
3 3 Monticello FSAR Benchmark Recirculation Controller Fails (Decrease) Figure 3.2-80 D N80N86 Core inlet Flow ps_A n. . . . . l . ! ' 4
- j 100-
==s-++**++,a. * * - * * * * * - * * * * *
- * * * * . . * - * . . . . . . . 1 a s +
g
- t
% j
% l .
M '% l
~i ' !. !
, : i l
S0- +++....-+-).<*
- ~~..".i"*"*-----.1....._.....
-*-+.t**. -
*- . . . . - t-**+-
i .
- : i 3 i i
0 , l 0 10 i 20 30 40 ; Time (Seconds) ! O ) Figure 3.2-81 DN80W86 Average Surface Heat Flux
,,, F.S A_R. . . . .
b . l00a .w.g. .;.. . ..;. . ..
- s. -
'I s o :
n : ~. l. l .K ,
) 4 y 1
,' w , .-.....___.....
.....----(
<a 50J., . . . .i. .
.3
.* 8
*ll l
= t e -
c I
< 0
-a b n
v - i g A 0 I-0 { le 20 30 .0 m Tim e (Secongs) 3 O i
= NSPNAD-8608-A Rev. 4 Page 83 of 194
o Monticello FSAR Benc'n mark - e Recirculation Controller Fails (Decrecce) Figure 3.2-82 D N B 07&86 Relative Power FS_A_R_ _ _ _ . g
. j : j 100- - - - * * * * . * * * * * - . * * - . . - . . . . - . . . . . .
's s . :
e N - M \ l f I s . . .
\i !
30 . . . . . . . . . (.
..g...........
j
~.*,,-?"',,-.
*- l g
0 , 0 10 20 30 40 Time (Seconds) O' l Figure 3.2-83 D N B 07#86 ; Feedwcter Flow FSAR l 200
- I Gl i,0 .
.i . .;. I O
x iOQ. .
. I.
- g, s' s : , - -- , '
i l \'i- ,
,i j
_J l -- ..
- i . ,
$ 0 -g . .). ' :: .' , ,
_e I s l e J. e a [ Nip L 0 g 0 10 20 30 40 .a
- *:m. (5.co ne s) 5 9 8.
- NSPNAD-8608-A Rev. 4 Page 84 of 194 h
O Monticello FSAR Benchmark O Recirculation Controller Fails (Decrease) Rgure 3.2-84 DNBO7#86 Yessel Steam Row ise _FS_A_R_ _ _ _ . O ! : : [ i 100 * * - - - - - -- - - -- --
': :s O = :
s i
- s, : : i
' : i 50-
.---4-t-- -
--t---- '
i
- 1 O :
0 52P 0 l'c 2'O 3'O 40 Time (Seconds) O i i lO ' O
\
eo N e b~ e E l k O i e t M l G- , e z w
.S lO i.
l a-NSPNAD-8608-A Rev. 4 Page 85 of 194 i
O 1 l i Monticello FSAR Benchmark e. Idle Loop Startup ; 4 Figure 3.2-85 DNB 068/86 Sensed Level _FSAR____. l
~
e!
- : 1 toi ~ " ' ' - "5-; * " '"5' *- - '5- - - .
4 . u 1 e ,.
-~~ -
- e, e ,q,,.
l'e . '
.s; v
.i.. . . ..>.. .
....y. ... . . . .. ... . . ..
- 152. = " -- + -
- ~"
"""h
""t e
-to' 0 10 20 30 40 Time (Seconds) l e'
Figure 3.2-86 D N B 068/86 Steam Dome Pressure r$An____. e' s o I. . . m . c. E ,~ O
, , _~
4 + 1 e 8 s o N .C 5 l e f -J l
- 4
-501- +
o 1 8
- c. I.
t. ~
! j e S
- a. -10 0 S
O o to go ao .o .a Tim. (Seconos) I NSPNAD-8608-A Rev. 4 Page 86 of 194 C
O 4 4 I - Monticello FSAR Benchmark
- O Idle Loop Startup DNs06s/s6 Figure 3.2-87 Core inlet Flow FSAR
= 150
!O i i i l
i00 ... ... ..... ..... . . .. ...... . ....:.. ... .. . .. ....... . . . . . . l 4 jO x ! . i 4 : _M.. ---_-_-__ .
; . -_ - I---- -
................>...................g...........
gg. . .....t........ .. .. ! . : : i
- o
~
t 0 . . . 3
, 0 to 20 30 40 s Tim. (S.conds) i b
i 1 t 40 I , l l 1
- DNB068/86
- Fiaure 3.2-88 j A7erage Surface Heat Flux FSA,R,,,,,
ISO i in : : iV 4 i
+
t .* 10 01 + + .
- I M I a / .
=
} - ------------------- -- -----------
U *
+ $0 4 l * '
- t
= 1 4 e 1 1 < c *1 J .O - 4 l
1 S
- l 6 0' C 0 10 20 30 40 1
.a Time (Seconss)
O Page 87 of 194
$ NSPNAD-8608-A Rev. 4 a
OI. 1 l l Monticello FSAR Benchmark .I Idle Loop Startup l FI ure 3.2-89 DNB068/86 j Relative Power FSAR ISO 9 l .
,00 .. . . . .. : .. ....
I ig
'\
l $ H . . g . . s: l -F%s --------- n-
$0- *- -) - .t- -------l*-* -
l l l
- e-0 0 l'0 2'O 3'O 40 Time (Seconds) e.
Figure 3.2-90 D N B 068/f3 Feedwater Flow rs.A_R-9- 100- - i . . m M
% .~ .
3 _, .. . _-. -+ - ----- __{
- 50 - . . .!
. 4 I . 4 a
-w l 1 - c ) l I l $' l l a 0 *I ! O G 10 20 30 40
' ," Tim. (5.co ne s) a e $ NSPNAD-8608-A Rev. 4 Page 88 of 194 m.
. - . . - - -- . - . . . _ . -. . _ . . . -- -. - - = _ - - . ._
'o s 4 0 Monticello FSAR Benchmaric - 1 4 Idle Loop Startup t Figure 3.2-91 D H8068/86 Vessel Steam Flow FSAR j
, eso
'O 1
- 1 e
f . e I - 1 100- = = - * * * ** - * *- - *- *= * * - * ** ==**. * =*** * = = *
- 4 e : :
q .
- e .
- O x i -
1 :e :
~
.s.o - - - ~ . .
~
e g$a ......e ee ee ee e e ee . ... e e e s ee } e e eeeeeeeese eeeeee ee e e e ee e en e e ee e se l j : : i
. e : e i
i i i : I
;O -
- 1 l
O M 0 10 20 30 40 Time (Seconds) O O l
'O i e,
N D s s
. l 2 \
e C O e
- i e I we 4
en h eh 3 0 e 7 NSPNAD-8608-A Reve 4 Page 89 of 194 4
O' Monticello Cycle 11 l
. Load Rejection without Bypass e
Figure 3.2-92 DWODE-B Vessel Pressure R1:e _G_E _A_n I _al s_f s. ! 200 1
- _-:--:_s . :
e 180- ~"~*~"i.~~"y"4."~~~~-"~}""*-- ~~~a"
~*4.~*~~~.-.
/ . . .
/
/ .
./
sn 100- ~**"~~";y"~~~--"~~*~~;""~*- ~;- "~+"--~~~~""~
- a. . : : : : m
- T j, .
, 1
/ : : : : :
60- ~~~~4.~*- .
~.~4.~- ~~~1""~~"""4.~"~~~.~>~~~
l o 0 1 2 3 4 s s s Time (Seconds) . I
- 1 i
l l e Figure 3.2-93 DYNODE-B Relative Power G_E_A_n_el
- L sis.
1,s
.
- JL I g
/I I 3 i
i,s. 1. . . # 1. 1. . . ..:: . . . . . -. . . . . :: . . . . . . . .. . . :'. :.......... ............
.....I....
- vv
. l :. : : : =-
W' t .
\ .
- . .... l. . .. .:.. . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. . ... . . . . . . . . . . . . . . . . . .
j . l . 100- a' - - 4.. - - -l . *l-- --
. l
. .4.-***..:*.....- .
M : . : : : g. 60- *~
"~~~k."~~** .
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0 ~-- - ' h 0 4 2 a k s s Time (Seconds) e' NSPNAD-8608-A Rev. 4 Page 90 of 194
e Monticello Cycle 11 l
. Load Rejection without Bypass e
i Figure 3.2-94 DYNODE-B Core Avera9e Heat Flux G_E _A_n_alI s_ls. ISO _ 125- +-< -- - - - - - -- t-n *t- **
. - I V - :
- l
/e:s - ( -
l 1 .
. %s .
/ N l .
100- --- --k- - i
. *.k-- -
. j g 73 .. . ... . . .:. . . . . . . . . . . .. .. .% ,, 5. y .*. . . . .
. .... i
. . . N
%s s~
- l. .
l
-*'.;*--**---*'t-S0- . .
**f-*,***
l l . 25- *-.-
* * - - k- **= - l * - * * * * **
. .- .k
- . : . )
) . :
0
)
0 i : 4 s e 1 Time (Seconds) J Figure 3.2-95 DYNODE-B Core inlet Flow IS O _o.s. .A -1n i i.ts. ns. . . . . . . . . , .. . . . . . .. . . . .
.) .
;%t s .
f.P N N f *? . . 100 s.
\.-e. *
* - .- = =e I : l ,
. . l g . .
. \
N g yg. ...- -.. .. .
........\....' N . + : N: ) :
- s* s -*- ?. % *w
$0- -
--.P. -
i.
- ' ' ,'5 - _
29- . . - . . . . . . . . . 0 0 0 1 2 3 4 5 e Time (Seconds) n J NSPNAD-8608-A Rev. 4 Page 91 of 194
err. . 91 Monticello Cycle 11
' load Rejection without Bypass .-
DYNODE-B Figure 3.2-96 Moln Steam Line Flow GE Anal1 sis. ise A-w; 125- .* *=*=*i**==*-*=***** = * * =>= = = = =*1==* *= =>******=**
- :n :
100- - + - * . # I -- 7 .* . : ' -i *F * * * - *
- r*
- k * -* ** * *
- (
o\ j: i I \ : _c #g e /:\ t ' 3
. - rk c
,+ .... lv.. 1 ....... -
- 1:. : :
1s. .- 7...........3 O-
- I . : . :
80- I
- i. -- - g - *.---.)- ----i.--->-----
. l . . .
- g. j . . . .
s: : : : : 25- -- - I
\' -
l 1 :
- a f I : w 3 . . . .
0
< I: i 0 1 2 3 S S 6 Tim. (Seconds)
DYNODE-B Figure 3.2-97 Feedwater Flow GE Anoi
---- 1 sis.
IS O p'. L.,
*=-- ~ N: s I
125- -
,< 7 -*+.--*!*---t.**\-+.
N
/ .
. %. N
/ * *
; ; N 100- **-s * * - * * * * - *-
- - . = = = == * , * * * *-
. *.e*=*\==**=
l l l l N
; \
l l \
- : : N y ys. . . . . .. .. . . . . . . . . . . . . . . . . ..... .. ... . . . .... ..........
* . h
. . . T
. . . . . . < . . . . . . . . . . . . .y..
s0 . . .
...> . . . .......q...... .. ,.... . . . . . .
25- . . . .. - ........ . L.-
*. .4 :. .....--J-----
-*L-...-- .
i
- . : : l
=l
. . . . . pl 0 , . . . :
o 2 3 4 s a Tim e (Seco nds) e NSPNAD-8608-A Rev. 4 Page 92 of 194
4-7 T1
- O I Monticello Cycle 11 ya Load Rejection without Bypass Figure 3.2-98 DYNODE-B Sensed Reactor Water Level .n o.i .
.o. .s. .A..1.
100 90- ** ***
**t.**a*- **?.'***a****t**--
-*****-.?**-
yr5 00 4" " " " 4 702 * "-" "*a "4."-"" "" a.i"
""""i.~a"""- .
~4.""~~
=. = =**=* - * . .- ---
- 60- * *==* *=
-=*~=.**
.-*'***=*.*** * *=*
j .
=.
e 80-
- - - - - -- -* * * * * * + + + * - - *** - -- ** e f': r -
j w : . : : ,, ; - d O = g 40
......................y........................
..p,'.....7....
. . . s .
m : : :
=
- 30- w
- 4+ a -" ,
* -" ".:J#"s-"
a 20- *"-~~""D.-CU>"""*"*".I""*""*4."""""4.+ i, 10- * + - - - * * * - - - - -
- a 0 . . . .
0 2 3 4 S 6 i Time (Second ) 1, 4 4 . 10 I i 1 i l l I O O . O O Page 93 of 194 NSPNAD-8608-A Rev. 4
1 o Monticello Cycle 11 - Feedwater Controller Failure e DYNODE-B Figure 3.2-99 Vessel Pressure Rise .G.E.A.n.al1s.is. ISO
. . . l
- :^
. . 1 s is. ... . . .. . ..<..................>.. ..... ... . .. ...............4...g . .y " . ,
e,
- : s
, . k %
. : ; s l
- : : N l l::
i.e.
...............'...............'.................................I
..............\.s
- : l
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.- rs- - .. .
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- : 1 :
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l . I . 20- " *"""
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- as "j"".~="'""" '
- . :. / :
. -,,,,-_ - J .
o
- i
. e r
o s.s so.s 54 tr.s Tim (Seconds) ei Figure 3.2-100 DYNODE-B Relative Power .G.E.A.n.e.l1 sls. 175
. . i ,1
. . g i
- : : I g -:
- : : I : L 150- """"'*~='"""="""="="*"*"*"""="l" g
! I N O1 l :l is , . . g .. ...........
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~
e, - J :( 100- 9 ,,, .""""""4""""+ :""""""" "A" " * " -
- : : :I
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+ . + i.
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- : : : I 25- " " "- ".""""".""..."..t""."""...;".1"..""
\s
. . . N
~~- e 0 3.5 F 10.5 14 17.5 Time (Seconds) a:
NSPNAD-8608-A Rev. 4 Page 94 of 194
.,V . k.- I
. . 1 e
i Monticello Cycle 11
.Feedwater Controller Failure O
Figure 3.2-101 DWODE-B Core Average . Heat Flux G
- _E _A.n.a.lyd_ s.
ISO 125- ++++ <-***..+ * .t+ *-*
**t.*-
* + - . . . . .
O .
*>/.-'
/ ' \s w :
100- - -
"~"""*"""a"""""4.-""s
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' Feedwater Controller Failure o DYNODE-B Rgure 3.2-103 Core inlet Subcoolin9 -o.s. .A.n.a.l .i.1.
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l l i 8l NSPNAD-8608-A Rev. 4 Page 96 of 194
e Monticello Cycle 11 Feedwater Controller Failure 3 DYNODE-B Rgure 3.2-105 Sensed Reactor Water Level os An.1.. ..
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O i O nsenAo-esos-A nev. 4 Pase 97 of 194 i
(( ' O Monticello Cycle 11
'100% MSIV Closure o DYNODE-B Rgure 3.2-106 Vessel Pressure Rise _G_E _A_N_A_LY_S_I S.
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_. . m_ _ __ i .r Monticello Cycle 11 j .100% MSIV Closure
- o
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r Q l Monticello Cycle 11 100% MSIV Cosure *l l Figure 3.2-110 DYNODE-B McIn Steam Line Flow GE A. _N_A_LYS.IS. l ISO e 125- <* = = = * * - -
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4 , .
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l J 1 I Monticello Cycle 11
- o
'100% MSIV Closure DYNODE-B Figure 3.2-112 Sensed Reactor Water Level G. .E. A.N.---
. ALYSIS.
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o Peach Bottom Turbine Trip - e Test TT1 Figure 3.2-113 DNB 001/86 Steam Dome Pressure
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' Peach Bottom Turbine Trip -
Test TTl l DNB001/86 Fi ure 3.2-115 Tu'rbine Throttle Pressure T_e_s __ t oo_t_a_ . 1100 1 b : l I 10$0- - b- s *-i2 '>' - * *
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2 1 e y NSPNAD-8608-A Rev. 4 Page 103 of 194 4
o
' Peach Bottom Turbine Trip -
e Test TT2 Figure 3.2-116 DNB002/86 Steam Dome Pressure T..... e .o. o., .. 1100
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i, NSPNAD-8608-A Rev. 4 Page 104 of 194
Peach Bottom Turbine Trip Test TT2 Figure 3.2-H8 DNaco2/s6 Turbine Throttle Pressure T._. _ _o_o_, ._ . k F 1960- **
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NSPNAD-8608-A Rev. 4 Page 105 of 194 4
a: 1 1 Peach Bottom Turbine Trip . Test TT3 i Figure 3.2-119 DNeoo3/s6 Steam Dome Pressure T._. _ _ _D_ ._ , ._ . 1300 g> i i [ i s-y_----. ioso. . . . . . . . . . . . . . . . . . . . . . . . . . . .. ,... . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
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4 4 Id Peach Bottom Turbine Trip Test TT3 Figure 3.2-121 DN B 003/86 Turbine Throttle Pressure T.,et p a r a_ , neo m ' u)
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o Monticello Cycle l h Turbine Trip Startup Test D N B038/86 Agure 3.2-122 Steam Dome Pressure _T._ _, _c a_ t_._ . 1200
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NSPNAD-8608-A Rev. 4 Page 108 of 194
i i a l 1 i i l Monticello Cycle l D ' Turbine Trip Startup Test l 1 I DNB038/86 Figure 3.2-124 Relative Power Ie_ e. D a i_a_ , soo , D $
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.? NSPNAD-8608-A Rev. 4 Page 109 of 194 t__________
O Sonticello Cycle 1 e Turbine Trip Startup Test DN8038/86 Figure 3.2-126 Fe.dwater Flow T i_e
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tiSPNAD-8608-A Rev. 4 Page 110 of 194
i 1
' I
) Monticello Cycle l ' l 4/4 MSIV Closure Startup Test ! Figure 3.2-128 D N a o35,*s 6 Steem Dome Pressure T. _ ._, c.. o_ i_._ . 182 0 ) : I 1lco. ........ .. .....g. . . . . .......p.... ....
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o Monticello Cycle I
- 4/4 MSIV Closure Startup Test Figure 3.2-130 DNB035/86 Steam Dem. Water Level _T._ _,_ D ._ _._ .
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S. e e. NSPNAD-8608-A Rev. 4 Page 112 of 194
e , l e Monticello Cycle l
- 4/4 MSlV Closure Startup Test l
l Rgure 3.2-132 DNB035/86 l Feedweter Row T._. _,_ 0._1_._ . 12 0 9 .
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NSPNAD-8608-A Rev. 4 Page 113 of 194 S.
O Monticello Cycle l 2/2 Pump Trip 5tartup Test Figure 3.2-133 DNe o36/a6 Total Vessel Flow Rote _T._ _,___
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g 50 . 0 2 4 4 8 10 Time (Seconds) e Fiaure 3.2-134 DNB036/86 Co~re Avere9e Heat Flux T_.___ oo_,._. 100 %
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Mo, licello Cyde l 3 - Auto Flow Decrease Startup Test DNB 037/86 Figure 3.2-135 Steam Dome Pressure .T.. s.i _Dat. .. le30
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. NSPNAD-8608-A Rev. 4 -
3. e.
D-9 Monticello Cycle l ' Auto Flow Decrease Startup Test Figure 3.2-137 D N8 0 37'8' Main Steam 1.ine Flow 12 0 T._i_t _D a_t_o_ .
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43 g 0 , O le 20 30 40 50 Tim. (s.conds) e' i 1 Figure 3.2-138 osso37/s6 Feedwaier Flow 120 _T._._ _o o., .. ! l i : :
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e 3 AAonticello Cycle 1 Auto Flow Decrease Startup Test Agure 3.2-139 D N B 037/86 Sensed Reactor Water Level Te_._, o o_,o_ . ) . I i. 30 ... . . . . . . . . . . . , . . . . . .. ..............;.............. s : - J ~ .
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O Monticello Cycle 1 '- e Pressure Step Startup Test DNB039/86 Figure 3.2-141 Steam Dem. Pressure T._s_e _o.,o_ to29 . i i i *
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- Page 118 of 194 4
a NSPNAD-8608-A Rev. 4 e
a r Monticello Cycle 1
- o Pressure Step Startup Test ,
DNB039/86 Figure 3.2-143 T. on,. I Main Steam Line Flow -------. i 12 0 1
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c :o ao ae 4o Tim e (S.conds) 0 D N B 039/86 Figure 3.2-144 Feedwater Flow T.. D.... . . . 12e O . too "" _-,__. :.;w w' ewever%q:: icy ;;;;;; z I o,. . . . . . . . . . . . O . . . . .. . .. . . . ..I M .o.
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A s O - i NSPNAD-8608-A Rev. 4 Page 119 of 194 A
. O Monticello Cvele 1 - Pressure Step startup Test Figure 3.2-145 DNE DW86 Sensed Reactor Water I.evel _T._ _, .o._.. ._, . e i i e 4 .....................)......................! p.....................p..................
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0 k NSPNAD-8608-A Rev. 4 Page 120 of 194
p , Monticello Cycle 1 - FW Level Setpoint Step Startup Test Rgure 3.2-146 DNB040/86 Steam Dome Pressure T. D ,. 1030 D :.
.........q.........<.................).........p.......<......... ........y......
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1000 0 S 10 15 20 25 30 35 40 d5 Tim. (s.c.nds) V Figure 3.2-147 DNB040/86 Relative Power T.
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) NSPNAD-8608-A Rev. 4 Page 121 of 194
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n u Monticello Cyde 1 - e FW Level Setpoint Step Startup Test DNB040/86 Rgure 3.2-148 Feedwater Flow T_e_s_t __Dat_o_ .
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