ML20153E171
| ML20153E171 | |
| Person / Time | |
|---|---|
| Site: | Fort Calhoun  |
| Issue date: | 04/30/1988 |
| From: | OMAHA PUBLIC POWER DISTRICT |
| To: | |
| Shared Package | |
| ML20153E168 | List: |
| References | |
| OPPD-NA-8301-NP-R03, OPPD-NA-8301-NP-R3, NUDOCS 8809060230 | |
| Download: ML20153E171 (27) | |
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OMAliA PUBLIC PO'a'ER DISTRICT NUCLEAR ANALYSIS RELOAD CORE ANALYSIS METHODOLOGY OVERVIE
l OPPD NA 8301 NP Rev. 03 I
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e APRIL 1988 '
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8009060230 000830 i i PDR ADOCK 05000205
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--- - - - - - -- - -- '-I
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l l ABSTRACT
, This document is a Topical Report describing Omaha Public Power District's re- '
f load core analysis methodology for application to the Fort Calhoun Station Unit !
No. 1. '
]
The report povides an overview of the District's reload core methodology. An- I
- alyses performed by the District and its contractors are described. Details of -
l the thermal hydraulic vtethodology which were previously submitted to the NRC l are provided.
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TABLE OF CONTENTS
!9
) SECTION ffGE I
1.0 INTRODUCTION
1 1
- 2.0 IVEL SYSTEM DESIGN 1 3.0 NUCLEAR DESIGN 2 I
j 4.0 THERMAL HYDRAULIC DESIGN 3
)
l j 5.0 POSTUIATED ACCIDENTS AND TRANSIENTS 8 i i 6.0 SETPOINT CENERATION 8 i
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7.0 REFERENCES
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4 LIST OF TABLES TAe t.E _ TITLE E6EE I 41 PARAMETER RANCES OF TliE SOURCE DATA FOR TllE CE 1 14 CliF CORRELATION AND THE PANCE OF ANF AND CE 14 X 14 FOR FORT CAU10UN VALUES i
l l 42 COMPARISONS BETVEEN TORC AND CETOP D 15 ,
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1 OPPD NA 8301 NP, Rev. 03 ;
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l LIST OF FIGURES FIOl'RE TITLE PAGE 41 CE FUEL SPACER GRID 16
. 42 ANF FUEL SPACER GRID 17 1
43 AXIAL LOCATION OF FUEL ASSEMBLY SPACER GRIDS 18 44 STAGE 2 TORC CHANNEL CEOMETRY FOR FORT Caul 0UN .
UNIT NO. 1 19 45 AXIAL POWER DISTRIBUTIONS FORT Caul 0UN CYCLE 8 20 1
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OMAllA PUBLIC PO'n'ER DISTRICT j RELOAD CORE HETil0DOLOGY OVERVIEW l REVISION DATE 1
00 septertber 1983 01 June 1'35 02 Noveinber 1986
! 03 April 1988 l I
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l OPPD NA 8301 NP, Rev. 03 j v 1
O OMAHA PUBLIC P0WER DISTRICT RELOAD CORE HETHODOLOCY OVERVIEW l.0 ISTRODUCTION Analyses done to license reload cores for Fort Calhoun Station consists of the analysis performed by the Omaha Public Power District and the analysis performed by the nuclear fuel vendor. The current nuclear fuel vendor is Combustion Engineering (CE); however, future reload fuel may be supplied by any of the fear PVR nuclear fuel vendors: Advanced Nuclear Fuels Corp. (ANF), combustion Engineering (CE), Westinghouse, or Babcock and Wilcox. The following sections discuss the reload analyses and consolidate information about the District's methodology previously submitted.
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2.0 FUEL SYSTEM DESIGN The fuel assembly mechanical design and analysis are performed by the nuclear fuel vendor. The fuel mechanical design and design methods
] utilized for Fort Calhoun Station by Combustion Engineering are de.
, scribed in Reference 2 1, Advanced Nuclear Fuels Corp, fuel mechanical design and design methods are discussed in References 2 2, 2 3 and 2 4 In an effort to further reduce the flux to the reactor vessel welds, a
l part length poison rods may be incorporated into the fuel loading pat-tern. The poison rods are composed of B 4C over the middle 50% of the ,
1ength. Inert material comprises the balance of the rod. When used they will reside in the guide tubes cf the selected assemblies. For example, in Cycle 10 the poison rods were in the four peripheral guide l tubes of the a* smblies in quarter core locations )., 2, 3, and 8.
OPFD NA 8301 NP, Rev. 03 ;
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3.0 SL' CLEAR DESIGN Q The District's nuclear design methodology is discussed in Reference 31. 'l 3.1 Fuel Manarement The reload core fuel management is perf ormed by the District. Cur-rent fuel management schemes are selected to reduce flux to the reactor pressure vessel welds.
3.2 fever Distribution Measurement 1
The District utili:es the CE methodology (Reference 3 2) to mea- l sure the power distributions. This methodology is discussed in the Cycles 5 and 6 reload submittals and approved in the SER's for these fuel cycles (References 3 3 and 3 4), k 3.3 Uncertainties and Allowances The power distribution uncertainties which are included in the overall analysis of reload cores are:
Parameter Uncertainty 3D Peak. Fq3 D 6.2%
1 Integrated Radial Peak F 6.0%
R !
Planar Radial Peak. Fxy 5.3%
These values are approved for use in CENPD 153 P (Reference 3 5). l A more detailed discussion of the treatment of uncertainties and allovances can be found in Reference 3 6. !
4 OPPD NA 8301 NP, Rev. 03 Page 2 of 20
3 3.0 NUCLEAR DESIGN (Continued) i 3.4 Physics Safety Related Data The physics safety related data are produced using the methodology discussed in Reference 3 1. !
4 4,0 TilERMAL HYDPJULIC DESIGN
. 4.1 Steadv State DNBR Analysis l
The steady state DNBR analysis is performed by the District using the TORC /CETOP/CE 1 methodology (References 4 1, 4 2, 4 3, 4 4 and ,
i 4 5). This methodology was approved for use by the District in l Reference 4 6,
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4.1.1 Crid Spacer Loss Coefficients l
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The analysis utilizes a D TORC model with explicit repre-
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sentation of the loss coef71cients associated with ANF and l CE fuel assemblies, i i j
j Th6 values were obtained by ANF using single phase pres-sure drop testing of an ANF test assembly and a typical CE 14 x 14 assembly, i
1 Because of the sensitivity of DNBR calculations to the [
3 difference in spacer grid loss coefficients, the District utilizes the Reynold's number expression for loss coeffi- I j I eients. This provides s'.e most accurate representation of L the pressure drop across each spacer grid in the assembly, i Thus, the cross flows between adjacent assemblies in the
] region of the spacer grid are accurately modeled, a
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4.0 TliEPJtAL liYDRAULIC DESIGN , Continued)
O 4.1 Steadv Stare DNBR Analysis (Continued) 4.1.1 crid Spacer Loss coefficients (Continued)
The spacer grid geometries for the CE and Mir spacer grids are shown in the attached Figures 41 and 4 2. The spacer 3
grid envelope far both the CE and MIF grids is 8.115 inches by 8.115 is.ches. The axial location of the CE and ANF spacer guds is shown in Figure 4 3. t In D TORC calculations, the spacer grid loss coefficient j for a channel corresponds to the assembly type whenever a
, channel represents a single assembly or a portion of an assembly. The choice of loss coefficient for lumped chan-nels in D TORC is made such that the minimum flow is pro-vided to the limiting fuel assembly. The CETOP model em-O ploys th9 spacer grid loss coefficient for the limiting assembly calculated in D TORC. The inlet flow fraction of the CETOP model is tuned such that the CETOP model pro.
duces conservative results with respect to the D TORC model, which models all fuel assemblies. ,
! 4.1.2 CE 1 Correlation l
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The District utilizes the CE 1 correlation for DNBR calcu- l lations. The ranga of data in the data base for the CE 1 correlation is contained in References 4 1 and 4 2. The
- range of parameters for the CE 1 correlation and corre.
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sponding ranges for the CE and MiF assemblies are shown in 8 Table 4 1. Because the data for the ANF fuel assembly are k
) within those specified in CE 1 data bsse, the use of CE 1 i
j correlation is appropriate for the ANT fuel. l 1
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l OPPD NA-8301 NP, Rev. 03 Page 4 of 20 i
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4.0 TilEPX\L !!YDPAULIC DESIGN (Continued)
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\s- 4.1 Stendy State DNBR Analysis (Continued) 4.1.3 D TORC and CETOP Hedels The District utilizes the D TORC code (Reference 4 3) and '
the CETOP code (Reference 4 4) to perform thermal hydraul- '
ic analysis for the Fort Calhoun reload core. The frac-tion of inlet flow to the hot assembly in the CETOP model ;
is adjusted such that the model yleids appropriate MDNBR results when compared with results of D TORC analysis for a given range of operating conditions. The fraction of in-let flow is determined for each reload core. The use of this methodolo6y was approved for use by the District in Reference 4 6.
],
The followin5 paragraphs discuss the application of the CETOP code to the Fort Calhoun reactor. Examples are for a the Cycle 8 core, n
Thermal margin analysis utilizing the CETOP model is sup-ported by comparing its predictions for Fort Calhoun Sta-tion with those obtained from a detailed TORC analysis.
J Several operating conditions were arbitrarily selected for [
] this demonstration; they are representative, but not the complete set, of conditions which would be considered for a normal DNB analysis. ;
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4.0 TilERMAL llYDRAULIC DESIGN (Continued)
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4.1 Steadv State DNBR Analysis (Continued) 4.1.3 D TORC and CETOP iiodels (Continued)
A thermal margin model for 1500 5'T for Fort Calhoun Unit No. 1 was developed for the following operatin5 ranges.
- Inlet Temperature 450 to 600*F ,
System Pressure 1750 to 2400 psia Primary System 4 Punp Flow Rate, (LCO - 197,000 gpm) 804 to 1204 Axial Power Distribution 0.517 to +0.526 ASI
.O The detailed thermal margin analyses were performed for the sample core using the TORC Channel 1 geometry shown in
) Figure 4 4. The appropriate spacer grid loss coefficient l vas applied to each "assembly" channel or partial assembly channel in each stage. In stage 1, lumped channel 28 util-1:ed the CE spacer grid loss coefficient because the chan-i nel was predominantly composed of CE fuel. Lunped chan-nels 26 and 27 utilized the ANT spacer grid loss coeffi-
' l cient because either the channel was composed of entirely ANF fuel or contained a singla CE assembly not on a bound- l ary between channels. The axial power distributions are given in Figure 4 5. These distributions were the most l limiting ones generated for the length of the cycle and i
for the various power dependent insertion limits examined.
The results of the detailed TORC analyses are given in Table 4 2.
O OPPD NA 8301 NP, Rev. 03 Page 6 of 20
4.0 THERMAL HYDRAULIC DESIGN (Continued) 5
( 4.1 Steadv State sDNBR Analysis (Continued) 4.1.3 D TORC and CETOP Models (Continued)
The CETOP design model was a total of four thermal hydraul -
ic channels to model the open core fluid phenomena. Chan-nel 2 is a quadrant of the hottest assembly which repre-sents the average coolant conditions for the remaining por-tion of the core. The boundary between channels 1 and 2 is open for crossflow; the remaining outer boundaries of channel 2 are assumed to be impermeable and adiabatic. .
Channel 2 includes channels 3 and 4 Channel 3 lumps the subchannels adjacent to the MDNBR hot channel 4. The "hot" assembly determined from D TORC analysis was an ANF
- assembly. Since CETOP modela a quadrant of the "not"
- assembly, the ANF apacer grid loss coefficient was used in l
l the analysis.
The CETOP model described above was applied to the same cases as the detailed TORC analyses. The results from the j CETOP model analyses are compared with those from the de-tailed analyses in Table 4 2. The uncertainties associ. l
) ated with the thermal hydraulic analysis are ecmbined statistically (Reference 4 7). In this method, the impact l I
of component uncertainties on DNBR is assessed and the '
i SAFDL is increased to include the effects of the uncer-taintics.
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- 5.0 POSTULATED ACCIDENTS AND TRANSIENTS
'O v The postulated accidents and transients are analyzed using the methodol-ogy discussed in Reference 5 1.
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6.0 SETPOINT GENER$ TION
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Ihe District utilizes the methodology discussed in CENPD 199 P (Refer-ence 6 1) to generate setpoints for Fort Calhoun Station- The Dis- !
trict's reactor physics methodology is discussed in Reference 6 2.
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The scrarn reactivity curves are produced using the QUIX code, The power-
- to fuel design limit on centerline melt is derived using the QUIX code with the appropriate combinations of planar radial peaking factor, T
Txy . and axial power distribution.
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l The thermal margin analysis is done using the CETOP code with the appro- ;
priate combinations of the integrated radial peaking factor, TR'
- axial power distribution, RCS inlet temperature, and RC3 pressure.
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j The Fort Calhoun RPS utilizes the "standard" local power density trip and TM/LP trip. The sequential CEA withdrawal is analyzed using the 4
methods described in Reference 6 1 and not included in the TM/LP trip ;
considerations. The RCS depressuri:ation and excess load events provide -
i the transient analysis input into the TM/LP trip. !
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_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - - - _ _ _ _ - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - -- - - - - - - - - - - - - - - ~ - -
6 6.0 SETPOINT GENERATION (Continued)
- V O 6.1 Incora Lc0 Monitoring The Better Axial Shape Selection System (BASSS) monitors the 1.im- [
iting conditions for Operat. ion on peak linear hes.t rate and de-J parture from nucleate boiling usin;; as input the data available l from the MINI CECOR code and the p3 ant ecmputer. This arrangement is similar to the one uced by Balt.imore cas and Electric at their
] Calvert Cliffs Unscs, and described in the Combustion Engineering
{ Setpoint Topical (Ref. 6 1),
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MINI CECOR is a mini computer version of Combustion Engineering,,'s ;
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CECOR code. It is used in this application to symthesiae the following parameters from readings of the fixed in core detectors:
! 1. The three dimensional power peaking factor (Fq )
, 2. The core average axial shape index (I) 4 i
3.
The total planar radial peaking factor (FxyT) l ,
t 4.
The total integrated radial peaking factor (FR) ,
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"I These inputs to BASSS r;e descriptive of the existing core power 1
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j The inputs to BASSS obtained from the plant computer are the fol- !
loving:
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- 1. Measured core power level '
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SETPOINT CENERATION (Continued) 6.1 Incore tco Monitorine (Continued) i i
j BASSS consists of two algorithms: one for peak linear heat rate monitoring and another for DNB monitoring. The peak linear heat rate algorithm uses the 3 D power peaking factor and the measured l
- core power level to calculate the core peak linear heat rate. The I j algorithm applies appropriate uncertainties and allowances (per i
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the Technical Specifications) to the 3 D peaking factor. The mea.
j sured peak linear heat rate is compared to the monitoring limit. [
which is based on both LOCA and A00 transient analysis considera. ;
- tions, and an alarm is activated when the monitoring limit is ;
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exceeded. The power operating limit on linear heat rate is also calculated and displayed as an indication of the available oper. l j ating margin. The DNB algorithm is an improvement over the excore !
1 i ASI nonitoring system in that it uses in. core axial shape index, l CEA group position and the radial peaking factors to establish the !
l plant's power operating limit. An alarm is activated when the
]O power operating limit is exceeded. A gain in operating margin re-suits from the following:
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j 1. A reduction in ASI uncertainty due to the use of in core l r
j ASI versus ex. core ASI. I 1
- 2. Knowledge of the actual CEA group position versus the
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- 3. Knowledge of the actual radial peaking factors versus the
] ex core system's assweptions that radial peaks are at the l j Technical Specification limits. I i
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Page 10 of 20
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6.0 SETPOINT CENERATION (Continued) 6.1 Incore LCO Monitorine (Continued)
BASSS is also provided with the capability to monitor the Limiting Conditions for Operation on F xyT and FR.I If the Tech-nical Specification for Operation on F xy T or FR are j exceeded during normal plant operation, BASSS will activate an !
alarm and calculate the proper trade off with maximum allowed
- power that ensures chat the Axial Power Distribution and Thermal t
- Margin /Lew Pressure Trips remain conservative. An alarm is !
] activated if the measured power level is higher than the allowed power level.
i 6.2 Uneerrainties lj l l :
The uncertainties are treated statistically in the District's set-point analysis. :
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7.0 REFERENCES
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/ lts.t.ien 2 References 21 "Omaha Batch M Reload Fuel Design Report". CEN 347(0) P, Rev. 01, January 1987, 22 "Ceneric Mechanical Design Report for Exxon Nuclear Fort Calhoun 14 x 14 Reload Fuel Assembly", XN NF 79 70 P September 1979. !
23 "Ft. Calhoun Design Report Extended Burnup Analysis", XT hF 82 61, 3 October 1982.
4
'4 "Extended Burnup Report for Fort Calhoun Reloads XN 4 and XN 5 I
(Batches K and L)," ANF 87 139(P), October 1987.
Section 3 References 1
l 31 "Reload Core Analysis Methodology, Neutronics Design Methods and Verification", OPPD NA 8302, Rev. 02, April 1988.
4
- 32 "INCA, Method of Analyzing Incore Detector Data in Pressurized Water !
1 Reactors" CENPD 145 P April 1,1975.
1 33 Letter from R.W. Reid (NRC) to T.E.Short (OPPD), December 5, 1978.
34 Letter from R.W. Reid (NRC) to W.C. Jones (OPPD). April 1, 1980.
! 35 "INCA /CCCOR Power Peak Uncertainty", CENPD 153 P, Revision 1 P A, ,
i May 1980. .
l l 35 "Statistical Combination of Uncertainties", CEN 257(0) P, Parts 1 and 3, November 1983. ,
)
j Section 4 References i
! 41 "CE Critical Heat Flux", CENPD 162 P A, Part 1. Combustion Engineer-1 ing, September 1976. l l '
i 42 "CE Cri; teal Heat Flux", Part 2. CENPD 207 P A, Combustion Enginesr.
1 ing, June 1976. ;
43 "A Computer Code for Determining the Thermal Margin of a Reactor Core," CENPD 161 P, TORC Code, Combustion Engineering, July 1975. ;
1 j 44 "CETOP D Code Structure and Modeling Methods for Calvert C1.?ffs 1
Units 1 & 2," CEN 191(B).P. Combustion Engineering December 1981. '
)
j 45 Letter from Cecil Thomas (NRC) to Mr. A. E Scherer (CE), November i
- 2. 1984 )
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7.0 REFERENCES
(Continued)
Section 4 References (Continued)
- 46 Letter from R. A. Clark (tfRC) to W. C. Jones (OPPD), March 15, 1983. ;
1 j 47 "Statistical Combination of Urcertainties," CEN 257(0).P. Part 2, November 1983. I 1
Section 5 References !
i f 51
- Reload Core Analysis Methodology, Transient and Accident Analysis I
- Methods and Verification," OPPD.NA.8303. Pev. 02, April 1988. E
+ .
] Section 6 References [
61 "CE Setpoint Methodology,' CENPD 199 P. Revision 1.P, April 1982. l l
62 "Reload Core Analysis Methodology Neutronics Design Methods and l Verification." OPPD.NA.8302 Rev 02, April 1988. l t
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i TABLE 4 1 '
j PARAMETER RANGES OF THE SOURCE DATA FOR THE CE 1 Citr CORRE1ATION AND THE RANCE OF Mir AND CE 14 x 14 FOR FORT Call 10UN VALUES l
CORRELATION CE ANF { I l PARAMETER RANGE PANGE PriGE t
Pressure (psia) 1765 to 2415 N/A N/A Local coolant Quality . 16 to .20 N/A N/A I
1 LocalMassXelocity (1bm/hr ft*) 0.87x106 to 3.21x10 6 i
N/A N/A l I Subchannel Vetted Equiv. .3588 to .4043 to .4010 to !
Dia.'reter (in) 54L7 .5449 .5402 .
l Subchannel Heated Equiv. .471) to .5334 to .5270 to !
Diameter (ir) .7837 .7840 .7760 i
Heated Length (in) 84 to 150 128 128 f
Crio Spacing (in) 14.2 to 18.25 16.8 16.8 I l
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038%RISONS IE3M23e 10RC APO GIOP-D Axial Elev.
Operatim Parameters PDER O=lity at PDER of PD81R finL-Detailed MIOP-D Detailed QIOP-D 10RC Inlet *IORC Inlet Inlet Avg Mass Core Avg. Shape Relative Fiw Relative Fim Pressure T + ature velocity HeatFlg Index Flw in Factor Fl w in Factor Detailed (pala) 6 2 Incation 5 Incation 5
(*F) (10 11mVhe-ft ) (51U/hr-Ir) (ASI) 'ICRC IOP
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1750 450 1.7432 242409 .517 2100 450 1.7432 257008 .517 2250 450 1.7432 261195 .517 2400 450 1.7432 264118 .517 2100 545 2.1790 216494 .517 1750 600 1.7432 14928'l .206 2100 600 1.7432 168727 .206 2250 600 1.7432 176398 .206 2400 600 1.7432 184260 .206 2100 545 2.1790 257118 .206 2100 545 2.1790 282778 .001 2100 545 2.1790 298644 .203 1750 450 1.7432 295262 .521 1750 545 1.7432 227063 .527 1750 COO 1.7432 147319 .527 2100 545 2.1790 255014 .527 _
OPPD .%-8301-NP, Rev. 03 Page 15 of 20
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'XIAL LOCATION OF FUEL ASSEP.BLY SPACER GR105 OPFD NA.d301 NP, sev, 03
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i i
l Omaha Public Power District Nuclear Analysis Reload Core Analysis Methodology Neutronics Design Methods And Verification ;
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, OPPD NA-8302-NP i Rev, 02 i
April 1988 I
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Copy No.
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ABSTPACT
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This document is a Topical Report describing Omaha Public Power District's re-load core neutronics design methods for application to Fort Calhoun Station Unit No. 1.
The report addresses the District's neutronics design methodology and its appli-cation to the esiculation of specific physics parameters for reload cores. In addition, comparisons of results obtained using this methodology to results from experimental measurements and independent calculations are provided, D
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l OPPD NA 8302 NP, Rev. 02 i i
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Table nf contents s
v Section f.afa
1.0 INTRODUCTION
1 2.0 BASIC PHYSICS MODELS 1 2.1 Neutron Cross-Sections 2 2.2 Diffusion Theory Models 3
- 2.2.1 MC 3 l
2.2.2 ROCS 4 2.2.3 QUIX 5 3.0 FORT CAUiOUN PMYSICS MODELS 7 3.1 Neutron Cross Sections 7 3.2 Diffusion Theory Models 8 3.2.1 MC 8 3.2.2 ROCS 9 3.2.3 QUIX 9 4.0 APPLICATION OF PHYSICS METHODS 10 4.1 Radial Peaking Factors 10 4.2 Reactivity Coefficients 11 4.3 Neutron Kinetics Parameters 12 4.4 Dropped CEA Data 12 4.5 CEA Ejection Data 14 4.6 CEA Reactivity '.
4.7 CEA Withdrawal Data 16 4.8 Reactivity insertion for Stearn Line Break Cooldown 17 4.9 Asywnetric Steam Centrator Event Data 18 4.10 QUIX Calculations 19 5.0 VERIFICATION OF NEUTRONICS MODELS FOR FORT CAUiOUN STATION 19 f
5.1 ROCS MC Generated Albedos 20 I 5.2 Planar Radial Peaking Factors 21 5.3 Integrated Radial Peaking Factors 22 I j C
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OPPD NA 8302.NP Rev. 02 11 l
Table of Contents (Continued)
O Section Eaga 5.0 VERIFICATION OF NEUTRONICS MODELS FOR FORT Call 10UN STATION 5.4 The District's ongoing Benchmarking Program 22 5.5 Summary 23
6.0 REFERENCES
24 APPENDIX A CEPAX/DIT Verification Program APPENDIX B ongoing Benchmarking Program ilistorical Information O
O OPPD NA 8302 NP, Rev. 02 iii
LIST OF TABLES a TABLE TITLE E6EE 51 Unrodded HZP Critical Boron Concentrations Calculations 26 5-2 Low Power Physics Isothermal Temperature Coefficients 27 53 Comparison of Calculated and Measured Isothermal Temper. 28 aturo Co fficients i
54 Comparison of Calculated and Measured Power Coefficients 30 5-5 Cycle 11 CEA Worths 31 4
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l GPPD !!A 8302 !;P, Rev, 02 l l
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. . . . _ . _ - - - ~ . .- . .. -- . _ _ . _ . . . - - - .- . _ . - . ..
OMAHA PUBLIC PO'n'ER DISTRICT NEUTRONICS DESIGN METHODS AND VERIFICATION i
i REVISION DATE 00 September 1983 01 November 1986 t
02 April 1988 1
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l OPPD NA 8302 NP. Rev. 02 i V l
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Omaha Public Power Di:;rict Reload Core Analysis Methodology
(""T Neutronics Design Methods and Verification L]
1.0 INTRODUCTION
The District's neutronics design calculation methods ara described along with results obtained when these methods are compared to experimental measurements and independent calculations. The discussion of calcula-tional methods includes descriptions of the basic computer codes and pro-cedures for applying these codes. Comparison of the calculations to measurements and independent calculations are performed using the same codes and computational methods used in the Fort Calhoun reload core de-sign efforts. The basic physics models, supplied by Combustion Engineer-ing (CE), are described in Section 2.0. Section 3.0 describes the Dis-trict's application of these models to the Fort Calhoun f.eactor. Sec-tion 4.0 presents the application of these physics models to the reload core analysis. Section 5.0 discusses the District's latest verification program which includes the recent cycle by cycle comparisons of District calculated data to measured data and data from independent calculations.
() Section 6.0 contains the individual references. Appendix A discusses the CEPAK to DIT cross sections Verification Program, and Appendix B pre- i sents the historical information, collected from beginning of core life, from previous ongoing benchmarking programs mentioned in previous neu-tronics methodology submittals. ;
I I
2.0 BASIC PHYSICS MODELS The District's neutronics design analysis for the Fort Calhoun core is based on a combination of multi group neutron spectrum calculations, which provide cross sections appropriately averaged over a few broad energy groups and few group one . two and three- dimensional diffusion theory calculations, which result in integral and differential l reactivity effects and power distributions. Calculations are performed .
l with the aid of computer progrards embodying analytical procedures and l fundamental nuclear data consistent with the current State of the Art. l l
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l 2.0 BASIC PHYSICS MODELS (Continued)
(s .
2.1 Neutron Cross-Sections The data base for both fast and thermal neutron cross sections is de-rived from ENDF/B IV with changes recommended by the Cross-Section .
Evaluacion Working Group (Reference 2 1). These recommendations con-sist of changes to the shielded resonance of U238, and the Watt fission spectrums of U 235 and Pu239, and changes in A for U 235 and Pu239 Few group cross sections, for subregions of the core that are represented in spatial diffusion calculations, (e.g.,
l.
fuel pin eclls, moderator channels, structural member cells, etc.) l are calculated by the DIT 1attico program. These cross sections are generated as a function of fuel temperature and moderatcr tempera-ture to accommodate the temperature feedback routines within the diffusion theory models.
The DIT code performs all the functions of the traditional transport methods which attempt co represent the complexities of the PVR fuel assembly geometry, including neutron energy spectrum interactions in the fuel, control rods, control rod locations (water holes), burn-able absorber rods, and incore flux detectors. The essential fea-ture of DIT, which distinguishes it from the traditional methodology is that the spectrum spatial averaging procedures are based on calcu-lations in two dimensional geometry. Hence few approximations to ,
i the geometry representation are necessary. The use of nodal trans-port theory has made it feasible to retain discrete pin geometry in l
both the fine and broad energy group calculations. A more complete {
description of the DIT procedures for generating few group neutron
]
cross sections can be found in References 2-2 and 2 3.
i d
i Previously, the District utilized the CEPAX program to produce fev-
! group cross sections. These cross sections were also generated as j functions of fuel and moderator temperature. Cor:parisons of calcu.
l lated and measured data reported in Section 5.0 include calculations performed using the CEPAK program.
OPPD NA 8302 NP. Rev., 02
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2.0 BASIC PHYSICS MODELS (Continued) 1 2.1 Neutron Cross Sections (Continued)
The CEPAX program is the synthesis of a number of computer codes, many of which were developed at other laboratories, e.g., FORM, THERMOS and CINDER. These programs are interlinked in a consistent way with inputs from differential cross section daca from an exten-sive library. A description of the CEPAX procedures'used to gener- ,
ate few group neutron cross sections can be found in Reference 2-4 2.2 Diffusion Theorv Models The diffusion theory models package used to calculate core physics parameters for Fort Calhoun Station consist of the MC, ROCS, and QUIX computer codes. The MC (fine mesh) and ROCS (coarse mesh) codes can be executed in one, two or three dimensions to calculate ,
j static and depletion dependent parameters such as reacravities, flux, nuclide and power distributions and CEA worths. The QUIX code is executed in one dimension to calculate axial power distributions ,
4 and CEA worth.
Previously, the District utilized the PDQ X code to produce tradt-tional two dimensional fine meth core design information. The PDQ X j program was an extension of PDQ 7 and HARMONY programs mentioned in Reference 2 5. Comparisons of calculated data between PDQ X and MC are reported in Section 5.0.
l l 2.2.1 HG !
. I j The MC program is a fine mesh method used to solve the two-group neutron diffusion equation. MC uses the 3 D coarse mesh analysis (ROCS) to recover local information on power, burnup and flux by performing fine mesh, imbedded diffusion i
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l 2.0 BASIC PHYSICS MODEl.S (Continued) 2.2 D.iffusion Theorv Model (Continued) 2.2.1 HQ (Continued) theory calculations within the coarse mesh nodes. The capa-bilities of MC offer a more computationally efficient alter-native to conventional fine mesh diffusion theory computer codes (i.e., PDQ-X) which in practice are limited to 2 D core analyses (Reference 2 6). MC air.o eliminates the PDQ X proble;n of representing large gradients in the vicinity of CEA guide tubes (water holes) and burnable absorber pins.
The PDQ X diffusion theory formulation does not provide the correct flux levels for fuel adjacent to water holes or burn-able absorbers, in fact, when PDQ X makes adjustments to ab-sorption ar.d removal cross sections necessary to obtain cor-rect reaction rates for non fuc1 cells, results are made O
d worse. By using the imbedded nodal calculational technique, MC removes the need to use PDQ X type adjustments.
MC employs macroscopic (static) and microscopic (depletion) cross section data generated by methods described in Section 2.1.
2.2.2 EQQ.1 The ROCS program is a coarse mesh two group solution of the l neutron diffusion equation based upon a mesh centered higher order finite difference formulation. It incorporates closed channol thermal hydraulic modeling into its evaluation of the interaction of neutron flux effects and the macroscopic physical and thermal properties of distributed materials.
Because of MC dependency on ROCS for input information and l O
OPPD NA 8102 NP. Rev. 02 Page 4 of 31
2.0 BASIC PHYSICS H0DELS (Continued) ~'
2.2 Diffusion Theorv Model (Continued) 2.2.2 ILQCS. (Continued) ,
h the ROCS coarse mesh nodal structure, ROCS is more efficient i than HC for evaluating a core's static and depletion depen-(
dent properties. ROCS also employs macroscopie (static) and i
microscopic (depletion) cross sections generated by the f 1 methods described in Section 2.1. A more complete descrip-l tion of the ROCS program is found in References 2-3 and 2 7. l 1
, 2.2.3 Q1113 i
The QUIX program is a one-dimensional (axial) representation ;
- of the core used to determine static and time dependent reac- I tivities and power distributions at selected stages of-deple-l tion. This program solves the neutron flux and associated eigenvalue in problems containing up to 140 distinct regicas
]
or compositions with variable mesh intervals. The macro-scopic cross section distributions, fission product yields, '
and xenon and boron microscopic cross sections required as input to QUIX are obtained from a three dimensionsi ROCS calculation. Local power density (fuel temperature) feed-i i back is included by modifying the point vise macroscopic !
absorption and removal cross sections. The change in cross- ;
sections is represented by a function of the difference be- ,
i tween the local axial power density and the referenced power j density. Moderator density feedback is included by relatit",
changes in the macroscopic absorption and removal cross see-tions to the local hydrogen number density which is calculat-ed from enthalpy at each axial segment. These cross sec.
- tions are generated in such a way that the fuel and rnoder-ator temperature coefficients calculated by QUIX arc equal
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2.0 BASIC PHYSICS MODELS (Continued)
O 2.2 Diffusion Theory Model (Continued) 2.2.3 QEIX (Continued) ;
to or conservative with respect to the fuel and moderator temperature coefficients calculated by ROCS. The axial-l reflector cross sections input to QUIX are determined in such a way that the steady state axial power distribution generated by QUIX matches the axial power distribution gen- ,
erated by ROCS, Details of the above treatments are given ;
i in Reference 2 8, !
j In addition to the eigenvalue problem, QUIX will perform four types of searches to obtain a specific ei envalue, 8
viz., a uniform poison search, buckling search, CEA region boundary search, and a moderator density dependent poison search. The uniform poison search assumes an axially con-I O stant macroscopic absorption cross section whereas the mod-orator density dependent poison search assumec a distributed macroscopic absorption cross-section dependent upon the axial moderator density. The moderator density dependent search is used to simulate the reactivity effects of the soluble boron in the reactor coolant.
Through the use of rod shadowing factors, shape annealing factors and shape index biases, the QUIX program has the capability of simulating excore detector response expected during normal operation. The procedures used for those simulacions are described in Reference 2 9.
O OPPD NA 8302 NP Rev. 02 Page 6 of 31
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! 3.0 FORT CALHOUN PHYSICS MODELS l !
The District utilizes the basic CE physics models described in Section 2.0 to model the Fort Calhoun reactor core. The computer codes which embody '
these basic physics models are maintained on the CE camputer system at-Windsor, Connecticut. The District accesses these computer codes through ,
a time sharing system. CE maintains all documentation and quality assur-ance programs related to these computer codes. The following paragraphs discuss the specifics of the Fort Calhoun models. "
i 3.1 Neutron Cross Sections The two group neutron cross sections utilized in the ROCS and MC '
, models of the Fort Calhoun reactor core are generated using the DIT code. Cross sections have been generated for unshimmed ANF and CE fuel assemblies and shimmed ANF and CE fuel assemblies. The cross-sections have been generated for the District by CE and are ,
based on information supplied by the District.
O P
1 The cross sections utilized to model the Fort Calhoun reactor are in the form of universal table sets. The two group cross sections are generated as functions of enrichment, fuel temperature, moderator temperature, burnup and in the case of shimmed fuel assemblies, !
B4 C shim number density. The table sets are applicable over a !
fuel temperature range from room temperature to 1800'K and a mod-erator temperature range from room temperature to 600*K. The fine l mesh table sets include explicit treatment of the pin cells immed-iately around the CEA guide tube (water hole) to properly account !
for the peaking of thermal flux in these water holes. Therefore, no !
] corrections need be applied to the pin powers produced by the fine mesh model, i
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- 3.0 FORT CALHOUN PHYSICS MODELS-(Continued) 3.2 Diffusion Theorv Models o
The District utilizes the MC, ROCS and QUIX models described in Sec- ,.
t j tion 2.0. The District utilizes both two dimensional and three- !
dimensional ROCS /MC models. The QUIX model is a one dimensional ,
model. '
i 3.2.1 HQ [
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] The District's MC model is a two group fine mesh model "ch i
generates its solutions based upon a two dimensional dep L.
tion in the x y plane. MC uses imbedded fine mesh calcula. [
tions within coarse mesh ROCS nodes to produce solutions in ,
the axial (z) dimension. Each fuel pin cell and shim pin j
- cell is represented by a single mesh point. The model also includes explicit representation of the CEA guide tube l l (water holes). The macroscopic cross section to node assign- ;
I ment is located in a geometry file which also provides shim j loadings and uranium metal weight for depletion calcula.
f tions. The model is representative of the core between 15% i 4
and 85% of full core height.
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The MC model is used to simulate the expected mode of oper-scion in the cycle being analyzed. This calculation results i in natarial distributions and radial peaking factors which !
are used in the safety analysis and setpoint generation. ;
Un)1ke PDQ X, MC utilizes exposure and geometry information 4
to create a library of precalculated coefficients for the .
i incore monitoring system.
I The mode of operation at the Fort Calhoun reactor is base ;
j loaded operation. Base loaded operation consists of reactor '
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1 3.0 FORT CAU10UN PHYSICS MODELS (Continued)
J' 3.2 Diffusion Theorv Models i
3.2.1 lig (Continued) ,
I operation at or very near rated thermal power throughout the cycle. The lead CEA bank insertion is held to a minimum.
Historically the lead CEA bank at Fort Calhoun has been in- !
sorted less than 5% of the time whenever the reactor is at a l 1 steady power level. Reference 7 1 discusses the impact of l
1 operation with a time averaged lead bank insertion. Due to !
i ;
its dependency on ROCS coarse mesh information, the model J J must be depleted in the same number and magnitude of ROCS
- j time steps, which are typically depleted in time steps of i j 1,000 MWD /MTU.
l l In order to use MC beginning with the District's Cycle-12 i
Reload License Submittal, MC prediction calculations were !
a benchmarked against PDQ X prediction calculations for Cycles j 10 and 11. An overview of the benchmarking results are re-j ported in Section 5.0.
t j 3.2.2 EqQ1 f
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The District utilizes a three dimensional and a two dimen- l sional two group ROCS model. The boundary conditions are !
l derived in accordance with the methodology discussed in !
I Refererce 3 2. }
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j 3.2.3 QLfD; i
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The District utilizes a one hundred and twenty five axial
] node QUIX model. The data for the QUIX model is obtained l l from the three dimensional ROCS calculations. !
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4.0 APPLICATION OF PHYSICS METHODS m
IV) Previous sections have focused on the reactor physics models utilized by ,
the District to model the Fort Calhoun reactor. In this section, calcu.
lations of the various core parameters used in the safety analysis are dascribed. The main parameters considered are the radial peaking factors (FR and Fxy), the moderator temperature coefficient, the fuel temper-ature or Doppler coefficient, the neutron kinetics parameters, CEA drop data, CEA ejection data, CEA scram reactivity, reactivity insertion for the steamline break cooldown, radis.1 peaking data for the asymmetric stesm generator event, and axial power distributions.
4.1 Radial Peakinc Factors .
The radial peaking factors, FR rnd Fxy, are calculated using the MC and 3 D ROCS models. Values of FR and F xy f r both unrodded and rodded core configurations are obtained directly from the MC power distribution. Since the MC model utilizes a pin power correc-tion edit implicitly accounting for the peakin5 of the thermal flux in the CEA guide tubes (water holes) no correction is required to the peaking factors calculated by MC. Unlike PDQ X, the values of F
xy and FR for unrodded and rodded cores are reported as core peaking edits. MC reads in the 3 D ROCS power histories, calculates F
xy for the plane of depletion and then synthesizes the calcula-tion plane by plane to obtain the maxi.aum core F xy. MC also utilizes the 3 D ROCS power information to calculate F Rbased on the axial integration of the planar power distribution obtained from 1 ROCS. The uncertainties for the radial peaking factors are given in
- Reference 4 1.
)
l The physics models are used to calculate the expected values of PR and F xy.
The actual values of FR and Fxy used in the safety
, analysis are chosen to be conservatively high with respect to those anticipated during the core life, i
OPPD NA 8302 NP. Rev. 02 Page 10 of 31
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4.0 APPLICATION OF PHYSICS METHODS (Continued) 4.2 Reactivity Coefficients -
- The ROCS models are used to calculate the moderator temperature coefficient (MTC) and the fuel temperature coefficient (FTC). The
, HTC is defined as the change in resetivity per degree change in moderator temperature. Calculationally, the MTC at a temperature of I
- Teod is determined by running three calculations; one at Teod' !
one et Taod + 10'F and one at Teod .10*F. The MTC at a f temperature of Teod is the average of the two calculated values. ,
i The reactivity change is calculated with the ROCS model by varying
- the inlet temperature while holding all other parameters such as the ,
fuel temperature and nuclide concentrations constant. i' The FIC or Doppler coefficient is defined as a change in reactivity
! per degree change in the effective fuel temperature. The effect of I ,
j fuel temperature upon resonance neutron energy absorption is account- t j ed for in the ROCS /MC models by means of power feedback options. l r j
The representation of the variation in the few. group cross sections c with fuel temperature involves two main segments. The first is to i
- represent the variation in cross section with fuel temperature, the f l second is to relate fuel temperature to reactor power density. The !
first portion is included in the basic methods employed to generate the few group cross sections. The second portion requires establish-ment of correlations between fuel temperature (i.e., effective fuel i
i
- temperature to be used in generation of cross sections) and the reac. [
- tor power density. The relationship between fuel temperature and reactor power density employs direct fits to FATES (Reference 4 2) I fuel data. This method results in the fuel temperature correlation i l
for each fuel type which is both local power density and fuel expo. I i
r sure dependent. !
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1 1 4
)
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4.0 APPLICATION OF PHYSICS METHODS (Continued)
('
4.2 Reactivity Coefficients (Continued)
The reduction in reactivity resulting from an increase in effective st e} tenperature is determined by ROCS. Typically, a temperature interval of $0* / is used to determine this coefficient The ph"sics movels are used to calculate the expected values of the h'tC throughout the cycle. The actual values of the HTC used in the safety analysis are chosen to conservatively bound expected values '
<!* o MTC The ceasurements of the MTC made during the operation of the reactor include uncertainties to assure that the actual MTC does not exceed ths 'alues used in the safety analysis. A fifteen percent uncertainty is applied to the Doppler coefficient when it is used in the safety analysis calculations.
4.3 I O Neutron Kinetics Parameters i
The neutron kinetics parameters A, A and the neutron lifotime, 1*, are calculated using Combustion Engineering's ROCS computer code. The technique utilized.co calculate the kinetics para-meters and the neutron lifetime is based on first order perturbation i theory. Details of the perturbation approach are discussed in 1 References 4.3 and 4.4 4.4 Droered CEA Data The neutronics data unique to the dropped CEA analysis are the i
values of Fg and Fxy following the drop of a CEA and the reac-tivity worth of the dropped CEA. The values of FR and F in-
' xy crease due to a large azimuthal tilt caused by the drop of a CEA and i
o OPPD NA 8302 NP, Rev. 02 Page 12 of 31
4.0 APPLICATION OF PHYSICS HETHODS (Continued)
\ 4.4 Drooned CEA Data (Continued) occur on the side of the core opposite the dropped CEA. Because the maximum FR and Fxy occur far away from the dropped CEA, the intra assembly power distribution is not perturbed. Therefore, the "post drop" value of FR and Fxy can be calculated by multiplying the "pre drop" values of FR and Fxy by the ratio of the assembly power after and before the drop of the CEA. This ratio is the distortion factor. The distortion factor is defined as the ratio of the assembly RPD at a given power level and time in core life containing a dropped CEA to the same assembly RPD without a dropped CEA.
The distortion factor and dropped CEA reactivity worth can be cal-culated using the 2 D or 3 D ROCS model. The 2 D ROCS calcu'.ations yield the Fxy distortion factor as a function of CEA bank inser-tion (i.e., ARO, Bank 4 In, Banks 4+3 In) and power level. The 3 D FR distortion factor is calculated for a specific CEA insertion and power level. Sufficient margin exists at the lower power levels such that F Rdependent DNBR calculations do not adversely effect operating margin. The "post drop" value of FR using the 3 D FR distortion factor is calculated by multiplying the "pre drop" value of FRfor the particular CEA insertion and power level by the 3 D FR distortion factor.
l 1
The 2 D and 3 D ROCS "post drop" power distributions are calculated l l
l with fuel temperature and moderator temperature feedback. The cal-culations assume that the core average Axial Shape Index (ASI) is being controlled within the "constant ASI" limits in accordance with the Fort Calhoun Operating Manual.
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4.0 APPLICATION OF PHYSICS METHODS (Continued)
(m)
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4.4 p.rppoed CEA Data (Continued)
An uncertainty is applied to the reactivity worth of the dropped CEA based on the verification contained in Reference 4-5. !
4.5 CEA Election Data The neutronics data unique to the CEA ejection analysis are the j values pre ejected and post ejected radial peaking factors and the !
reactivity worth of the ejected CEA. The maximum post ejection l 4
radial peaking factor and maximum ejected CEA reactivity worr.hs are {i calculated for the maximum CEA insertion allowed by the PDIL at HFP r
j and HZP. The neutronics parameters are calculated using HFP and HZP
- l 2 D ROCE and MC models. The post ejection radial peaking factor is !
calculated by multiplying the 2 D ROCS post ejection assembly RPD by f the corresponding pin to box ratio from MC. The ejected CEA *eactiv-icy worth is obtained directly from ROCS calculations. ROCS post
- ejection power distributions are calculated without moderator or ;
fuel temperature feedback. !
4 l
The post ejection value of Fq which is obtained using MC, is cal-culated by multiplyin5 the post ejection value of Fxy by the max- ;
imum value of gF , the a:imuthal tilt allowance, the augmentation !
l factor, the engineering heat flux factor, the fuel densification +
I factor, and the F quncertainty documented in Reference 4 1. An [
l uncertainty is applied to the ejected CEA worth. l l
i j 4.6 CEA Reactivity l
l t j The CEA reactivity calculations done in a reload core safety analy-sis are the calculation of the total reactivity of CEA's inserted O !
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4.0 APPLICATION OF PHYSICS METHODS (Continued)
(3
\s- 4.6 CEA Reactivity (Continued) into the core during a reactor trip (CEA scram reactivity), the generation of the scram reactivity curves, and the calculation of required shutdown margin.
f The CEA scram reactivity worth at HZP is calculated by obtaining the net worth for all CEA's between the HZP PDIL CEA position and the fully inserted position and subtracting the worth of the highest worth stuck CEA. These calculations are done using the ROCS model.
An uncertainty is applied to the HZP CEA scram reactivity worth.
The HZP CEA screm reactivity for the CEA ejection transient is calculated in a similar fashion except that the vocch of the ejected and highest stuck worth CEA's are subtracted from the net worth.
The scram CEA worth at HFP is calculated by obtaining the HFP not worth for all CEA's between the HFP PDIL CEA position and the fully
(N inserted position, subtracting the worth of the highest worth stuck CEA and subtracting the moderator void collapse allowance. The ther-mal hydraulic axial gradient reduction allowance, and the loss of worth between HFP and HZP are also subtracted from the HFP net worth for the scram CEA worth to be used in all transients except the four ,
pump loss of flow event and the steam line break incident. These are not applied to the four pump loss of flow scram CEA vorth be-cause the closest approach to the SAFDL during the four pump loss of flow event occurs prior to significant CEA insertion. These allow-ances are not applied to the steam line break (SLB) incident HFP CEA 4
scram worth because the HFP SL3 reactivity insertion curves implicit-ly account for these effects.
O OPPD NA 8302 NP, Rev. 02 Page 15 of 31
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4.0 APPLICA110N OF PHYSICS METHOD 5 (Continued) 4 4.6 CEA Reactivitv (Continued) i The moderator void collapse allowance is 0.0% op at BOC and 0.1% !
As at EOC. The thermal hydraulic axial gradient reduction allow- !
ance is 0.2% ap at BOC and 0.4% op at EOC. An uncertainty is applied to the HFP CEA scram reactivity worth. The HFP scram reactivity for the CEA ejection transient is calculated in a similar ;
fashion except that the worth of the ejected and highest stuck worth ;
, CEA's are subtracted from the net worth. All CEA worth calculations f assume the ASI is being controlled within the "constant ASI" limits in accordance with the Fort Calhoun Operating Manual, t 1 i l
The generation of the scram reaccivity curves utilizes the methodo1 ,
! ogy discussed in Reference 4 6.
1
< l The calculation of the required shutdown margin is only performed at -
t l HZP since the shutdown margin at power is controlled by the PDIL. !
4 The available HZP shutdown margin is equivalent to the HZP CEA scram
) teactivity. .j i ,
i 4.7 CEA Vithdrawal Data 4
[
The reactor core physics data ur.ique to the CEA withdrawal analysis f j is the maximum differential CEA worth. This is the maximum amount I
., of reactivity at any time in core life that can be added to the core per inch of CEA motion. When the maximum differential CEA worth is J combined with the maximum CEA withdrawal rate of 46 inches / minute, a conservative withdrawal rate expressed in top /see is obtained ,
and ased as input to the CEA withdrawal analysis. !
l l I ,
OPPD.NA.8302.NP, Rev. 02
, Page 16 of 31 ,
g 9 I
't 4.0 APPLICATION OF PHYSICS METHODS (Continued)
! 4.7 CEA Withdrawal Data (Continued) i l
r The maximum differential CEA worth is obtained for the'sequantial.
I withdrawal of the CEA banks from the HZP PDIL to an all rods out con-dition. The 3 D ROCS model is utilized to calculate this parameter.
j The calculations are performed assuming that the reactor is being i a
controlled within the "constant ASI" limits in accordance with the Fort Calhoun Operating Manual, i 4.8 Reactivity Insertion for Steam Line Break Cooldown i i
i The reactor core physics data unique to the steam line break trans-j ient analysis is the reactivity insertion due to the cooldown of the !
moderator. There are two sources of this reactivity insertion. The ;
first is the positive reactivity insertion due to the increasing }
density of the moderator as the cooldown progresses. The second is '
the reactivity insertion due to the Doppler coefficient as the effec- L
) tive fuel temperature changes. (
I f
i j
Reactivity insertions due to the moderator density increase and the l l Doppler coefficient are both calculated using a full core ROCS f model. The axial leakage or buckling is adjusted such that the mod.
j
(
erator temperature coefficient calculated by the ROCS model corre-j sponds to the most negative Technical Specification limit. The reac- !
l' tivity insertion calculations are performed with all CEA's except !
i the most reactive CEA inserted in the core. l I
i r i i 4
1 The moderator density reactivity inser$1on curve for the hot :ero !
power steam line break case is calculated by successively lowering f
the inlet temperature of the ROCS model from 532*F and allowing only I moderator temperature feedback in the model. The calculations typi- j cally result in a curve of reactivity insertion vs. moderator temp-l erature from a hot zero power temperature of 532'T to 212'F.
1O 0FFD NA 8302 NP. Rev. 02 j Page 17 of 31 l
1 I
_~
t 4.0 APPLICATION OF PHYSICS METHODS (Continued) 4.8 Reactivity Insertion for Steam Line Break Cooldown (Continued) i The Doppler reactivity insertion curve for the hot zero power case is also calculated by steadily decreasing the inlet temperature of
- 1 the ROCS model. The fuel temperature feedback in the model allows the production of a curve of Doppler reactivity as s function of fuel temperature. All zero power calculations are performed assum-ing there is no decay heat and no credit is tanen for. local voiding in the region of the stuck CEA.
The moderator density reactivity insertion curve for the full power
] case is calculated by decreasing the power level and core average avers y .*oolant temperature from full power to the hot zero power inlet temperature and then successively lowering the inlet tempera-ture as in the hot zero power case. Only moderator temperature feed-i back is utilized in the ROCS model. The Doppler reactivity inser-tion curve is calculated by a similar procedure utilizing the fuel temperature feedback in the model.
6 I
Since the modtrator reactivity insertien curve corresponds to an MTC (
l which is at the Technical Specification limi.t. no additional uncer- f
{ tainty is added to this curve. A fifteen percent uncertainty is I applied to the Doppler reactivity insertion curve.
! 4.9 Asymetric Steam Generator Event Data l
I For the range of temperatures considered, the intra assembly peaking f
! does not vary as the inlet temperature is changed.
1 !
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) OPPD.NA.8302.NP. Rev. 02 l 1
j Page 18 of 31 l i :
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4.0 APPLICATION OF PHYSICS HETHODS (Continued) 4.10 OUIX calculations 6
The District utilizes the QUIX model to perform various axial shape analyses related to the generation of the reactor protective system setpoints. The QUIX calculations are carried out in accordance with ;
the methodology discussed in Reference 4 6.
5.0 VERIFICATION OF NEUTRONICS MODELS FOR FORT CALHOUN STATION l i
1 1
The District has performed extensive verification of the neutronics models
] used in the reload core analyses. The results of the previous District
! verification efforts were reported in References 5 1 and 5 2. These ;
i effvrts consisted of utilizing cross sections produced by CEPAK and DIT and confirming the District's ability to use the models with DIT cross- ;
sections. Extensive verification of the use of DIT cross sections was t also done by Combustion Engineering (CE) and reparted in Reference 5 3.
!o
- In order to present the most recent verification efforts in this section, I
a the previously mentioned verification programs have been organized into an ;
t appendix. The verification of CEPAK to DIT cross sections, consisting of 2 '
predictions of overall core reactivity, power distributions, reactivity
! coefficients, CEA worth and xenon reactivity, along with the appropriate j figures and tables, is reported in Appendix A. l I
i 1'
r
- The most current neutronics verification program performed by the District j involved the benchmarking of the fine mesh, imbedded nodal diffusion
] theory code, MC, with PDQ.X. The purpose of this program was to replace 1 l PDQ.X with the more cost efficient MC code and demonstrate the District's i ability to effectively use MC models, i
J
]
1 1
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j OPPD.NA 8302.NP. Rev 02 j Page 19 of 31
5.0 VERIFICATION OF NEUTRONICS MODELS FOR FORT CAlll0UN STATION (Continued)
This verification is in addition to the comprehensive verification of the methods done by CE and described in References 5 4 and 5 .'.. It is not the intention of the District to repeat CE's verification effort which in-cludes a statistical assessment of the adequacy of the uncertainties used by both CE and the District. Rather, it is the District's intent to demon-strate that the District can adequately model ths Fort Calhoun core and that the results of the District's verification effort are consistent with those reported in References 5 4 and 5 5.
The District's verification using MC incorporates predictions made for Cyclo 6 10 and 11. Benchmarking was performed for the predictions of ROCS /
MC generated albedo boundary conditions, planar radial peaking factors (Fxy) and integrated radial peaking factors (FR ). The results of the verification effort include data for both PDQ X and MC predictions.
5.1 ROCS MC-Cenerated Albedos O Prior to the Cycle 12 Reactor Physics Reload Analysis, ?.he Fort Calhoun Station core boundary conditions for ROCS model were gener-ated by CE using the CEFIACS computer code. The MC code has the capability of calculating coarse mesh (ROCS) boundary albedos. The MC generated albedos are fed into ROCS prior to perforcing fuel de.
pletion.
A comparison of the MC generated albedos against the CEF1ACS gener-ated albedos is necessary to confirm the validity of the HC gener-ated albedos. These radial power distribution differences between the MC generated albedos and CEFIACS generated albedos are within
- h. The differences at the periphery are consistently higher than the ones for the inner assemblies. This behavior is mainly due to the direct effect imposed on the peripheral assemblies by the virtue OPPD NA 8302 NP, Rev. 02 Page 20 of 31 i
_ _ . . . _ .._. __ ____ , -_ . ~ . _ _
.j 5.0 VERIFICATION OF NEUTRONICS MODELS FOR FORT CAIJ10UN SYATION (Continued)
O 5.1 ROCS MC Generated Albedos (Continued) i d
of their location. The low neutron radial leakage core loading patterns employed by the Districe result in a low neutron flux and power distribution at the periphery which 'n turn calcula ;e higher ;
peripheral power dis':ribution differences. The magnitude of these differences is consistent with those reported by CE in Reference
] 5 5.
a f
5.2 Planar Radial Peakina Factors MC is & 3 D code that performs 2 D depletion calculations. The third dimension is gener.6ted by synthesis with ROCS. Planar radial !
- 1,eaking factors (Fxy) are thus calculated for the plane of inter- t est and then synthesized axially in order to determine the maximum i l; planar radial peaking factor in the core. '
I l
The difference between the MC calculated and PDQ calculated Fxy's i j do not follow a unified pattern throughout the core. The planar l radial peaking factor distributions are influenced by several fac. l tors, some of which are working in opposite directions, combustion
) Engineering, in their independent review of the District's bench- l l marking analysis, stated that the results of the District's bench- ;
l l marking ef MC are consistent vith similar work done at CE on ether i j plants.
i The planar radial peaking factors predicted by MC for the plane of l a
depletion compare closely to PDQ predictions. MC predictions of corevide peaking factors, as expected. produce slightly more con-
, servative results than PDQ midplane predictions. !
, i l i J
iO !
1 OPPD NA.8302 NP, Rev. 02 l Tage 21 of 31 {
f
)
.-_______m. - . _ _ - _ _ _ , . . . . . . _ , . _ _ , _ _ . _ , _ _ _ _ _ . . _ _ , _
l !
l l 5.0 I VERIFICATION OF NEUTRONICS MODEl.S FOR FORT CAIJ10UN STATION (Continued) i 5.3 Intecrated Radial Peakinz Fact 2IA i
MC performs the calculations of the integrated radial peaking fac-tors (FR) for all assemblies. The MC calculated FR 's are com- !
pared to PDQ F R 's which are calculated by multiplying each assem. -
bly pin to box ratio by its correspond.ing relative power density [
from ROCS core edits. The calculation of the integrated radial peaking factors in MC is performed by multiplying the assembly axial integral of the planar radial power distribution.
The fluctuations between MC calculated and PDQ calculated FR 's :
cannot be represented by a uniform pattern throughout the core.
Similar to the planar radial peaking factors, the integrated radial :
peaking factor variations are influenced by a number of factors, f
However, the overall findings show acceptable results. Combustion '
Engineering, in their independent review of the District's MC benchmarking analysis stated that the results of the District's O benchmarking of MC are consistent with similar work done at CE on other plants.
The FR 's predicted by MC for the plane of depletion compare close-ly to PDQ predictions.
5.4 The District's Onreine Benchmarkine Program The data reported in this section consists of both historical and current information for the District's ongoing benchmarking pro-gram. The historical data, collected through Cycle 10 and reported in Reference 5 2, is located in Appendix B. The verification pro.
gram has been updated th augh mid Cycle 11. Tables within Section 5 have been updated to include only recent cycle by cycle information (Cycles 9, 10 and 11).
O OPPD NA 8302 NP, Rev 02 Page 22 of 31
5.0 VERIFICATION OF NEUTRONICS MODELS FOR FORT CAulOUN STATION (Continued)
O
\'
5.4 The District's onceine Benchmarkinc Procram (Continued)
Both verification of program segments include information consisting of startup physics testing predictions, reactor testing analysis and a core follow effort. This program vill continue to provide verifi-cation data in the future. i 5.5 Summary The District has an ongoing neutronics methodology verification pro-gram. The results of this verification program for previous cycles demonstrate the ability of the District to utilize the neutronics methods described in this document.
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i 1 OPPD NA 8302 NP, Rev. 02 j Page 23 of 31 1
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1
6.0 REFERENCES
Section 2.0 References 21 ENDF 313. "Benchmark Testing of ENDF/B Data for Thermal Reactors.
Archival Volume," July,1981, 22 A. Jonsson, J. R. Rec and l',. N. Singh, "Verification of a Fuel Assembly Spectrum Code Based on Integral Transport Theory,"
Trans. Am. Nucl. Soc., 28, 778 (1978),
23 CENPD 226 P, "The ROCS and DIT Computer Codes for Nuclear Design "
December, 1981. l l
24 System 80 PSAR, CESSAR, Vol. I, Chapter 4.3.3, Amendment No. 3, f June 3, 1974. ;
25 W. R. Cadwell, "PDQ 7 Reference Manual," WAPD TM 678, January, i
1968. .
] L 1 26 S. F. Crill, A. Jonsson and M. W. Crump, "Recent Developments in I
- the ROCS /MC Code for Retrieving 1.ocal Power Informacion in Coarre- '
Mesh Analysis," CNS/ANS International Conference on Numerical L Methods in Nuclear Engineering, Montreal, Canada September 6 9, ;
1983. !
i, 27 T. G. Ober, J. C. Stark, I. C. Richard and J. K. Gasper "Theory, [ ;
O Capabilities, and Use of the Three Dimensional Reactor Operation and Control Simulator (ROCS)," Nucl. Sci. Eng., 64, 605 (1977).
i 28 System 80 PSAR, CESSAR, Vol. 1. Appendix 4A, Amendment No. 3 June
); 3, 1974. l [
t 29 (
CENPD 199 P, Revision 1 P. "CE Setpoint Methodology," April 1982. l r
1 i
Section 3.0 References f i
j 31 CENPD 199 P, Revision 1 P, "CE Setpoine Methodology," April 1982. l 4
j 32 CENPD 226 P, "The ROCS and DIT Computer Codes for Nuclear Design," '
, December, 1981.
j
}
j Section 4.0 References !
41 CENPD 153. Revision 1 P A, "Evaluation of Uncertainties in the !
Nuclear Power Peaking Measured by the Self Powered Fixed In Core l
)
,a Detector System " May, 1980.
i OPPD NA 8302 NP. Rev. 02 !
Page 24 of 31 l
t
6.0 REFERENCES
(Vontinued)
Section 4.0 References (Continued) 42 "Development and Verification of a Fusi Temperature Correlation for Power Feedback and Reactivity Coefficient Application," P. H.
Cavin and P. C. Rohr, Trans Am. Nucl. Eng. 30, p. 765, 1978.
43 A. F. Henry, "Computation of Parameters Appearing in the Reactor Kinetic Equations," WAPD 142, December 1955.
44 R. W. Hardie W. W. Litke, Jr. , "PERT V, A Two Dimensional Portur.
bation Code for Fast Reactor Analysis," BNVL 1162.
45 CENPD 226 P, "The ROCS and DIT Computer Codes for Nuclear Design,"
December, 1981.
46 CENPD 199 P, Revision 1 P, "CE Setpoint Methodology," April 1982.
Section 5.0 References 51 CEN 242-(0).P. OPPD Responses to NRC Questions on Fort Calhoun Cycle 8, February 18, 1983.
52 "Reload Core Analysis Methodology, Neutronics Design Methods and Verification," OPPD NA 8302 P, Rev. 01.
( 53 CENPD 153 P, "INCA /CECOR Power Peaking Uncertainty," May,1980, 1
j 54 CENPD 226 P, "The ROCS and DIT Computer Codes for 1:uclear Design "
December, 1981. l 1
55 S. F. Crill, A. .'onsson and M. W. Crump, "Recent Developments in i the ROCS /MC Cot, for Retrieving Local Power Information in Coarse-Mesh Analysis," CNS/ANS Inte* national Conference on Numerical l Methods in Nuclear Engineering, Montreal, Canada, September 6 9, l 1983.
1 I
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OPPD NA 8302 NP, Rev. 07 Page 25 of 31
/
TABLE 5 1 Unrodded HZP Critical Boron Concentrations Calculations ,
3-D Heasured ROCS Ovele vem (DIT) -
{
J 9 1518 1
l 10 1474 11 1502 1
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OPPD NA 8302 NP, Rev. 02 Page 26 of 31
)
4
- - - _ _ _ . _ , _-_,_.._____,.,.,_______-..,.,_,.,__,-_..._,__--_,m- . , . , , . _,____-____...-.-.,_.m-,-._---.,
TABLE 5 2 Low Power Physics Isothermal Tempers ' ure Coefficients (f
Boron DIT t Concentration Measured ROCS CZtig (rrm) (Ae/'F) (Ar/'F)
^ ""
9 1457 0.30 x 10 .
10 1457 0.23 x 10 '
i 11 1496 0.20 x 10 -
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l I
l 1
l l
9 OPPL-dA 8302 NP, Rev 02 Page 27 of 31
_ _ _ - _ __ =_ - _ _ . - - . - . - . . . _ . .~ - - - - - . . -
1 TABLE 5 3 4
I
( Comparison of Calculated and Measured Isothermal Temperature Coefficients J
BOC a Calculated j Critical Boron Measured DIT. ROCS I
Percent of Concentration ITC ITC (x100 Ae/'F)
Cvele 0 Rated Power (rem) (x10 Ae/'F)
J 9 94 1036 0.39 I 10 95 1017 0.48 11 93 1073 0.43 ~
~
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l, OPPD NA 8302 NP, Rev. 02 Page 28 of 31
TABLE 5 3
() Comparison of Calculated and Measured Isothermal Temperature Coefficients (Continued)
EOC Calculated Critical Boron Measured DIT ROCS Percent of Concentration ITC ITC Cvele Rated Po.eer (orm) (x100 Ar/'F) (x100Ac/* F) 9 96 300 1.46 10 95 302 1.53 11 95 301 1,62 d
I
. O l
i ,
a l
1 NOTE: Full Rated Power - 1500 MWt !
1 l
l l
- O j OPPD NA 8302 NP, Rev, 02 j Page 29 of 31 l
I TABLE 5 4 !
Comparison of Calculated and Measured Power Coefficients a
DIT. ROCS ;
Percent Critical Measured Calculated ;
of Boron Power Power -
Burnup Rated Cone. Coeff. Coeff.
CZalg fik'D/MTU Power fermi (Ae/4 Power) (Ae/t Power) 1
~ ~~
9 420 94 1036 1.64 x 10 ;
1
, 9 9663 96 300 1.57 x 10 !
l
- 10 583 95 1017 1.24 x 10*4 3 ,
10 9261 95 302 1.40 x 10*'
11 433 93 1073 0.95 x 10
11 9765 95 301 1.52 x 10 -- .
1 i
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i 1
4 i
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i l
.I L I i 1
NOTE: Full Rated Power - 1500 Mk't l
I 1O J
CPPD.NA 8302.NP. Rev. 02 i Page 30 of 31 l
i 1
TABLE 5 5 Cycle 11 CEA Worths i Calculated DIT ROCS 3D EEsm2 4 60) Dae) i ,_, _
- 4 0,76 1
I i 3 0.46 ,
i 2 0.97 l l
1 0.58 l 1 i Total i
- (4+3+2+1) 2,77 l
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i j OPPD NA-8302 NP, Rev. 02 Page 31 of 31 !
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OMAHA PUBLIC PO'n'ER DISTRICT NUCLEAR ANALYSIS RELOAD CORE ANALYSIS METHODOLOGY
,s l CEPAX/DIT VERIFICATION PROGPRt i
l I OPPD NA 8302 NP I
APPENDIX A Rev, 02 O
I 1
l April 1988 !
l i
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1 4
1 0
~.-- - _ - - - - - _ . . . . . . - _ , - . - - - ,. - - - .
1 3
Table of Contents 3ECTION fAGE A.0 VERIFICATION OF NEUTRONICS MODELS FOR FORT l- l Caul 0UN STATION -
1 '
A.1 Core Reactivity 1 ,
A.2 Power Distributions 2 i
A.2.1 Radial Power Distributions 2 i i
A.2.2 Axial Power Distributions 3 r
A.3 Reactivity Coefficients 4 A.4 CEA Reactivity Worth 4 A5 Comparisons to Critical CEA Positions 4
, Tollowing a Reactor Trip i
A.6 Comparison to Independent Rad1 Power 5
] Distribution Calculations REFERENCES f.
1 i
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]
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, OPPD NA 8302 NP, App. A, Fev. 02 4
i a
j
LIST OF TABLES l
TABLE II.Ill }.'.625 i
A1 Unrodded HZP Critical Boron Concentrations Calculations 7 i
- A2 Summary of Comparisons of Measured and Calculated Integral 8 Assembly Relative Power Densities j A3 Low Power Physics Isothermal Temperature Coefficients 11 l l A4 Comparison of Calculated and Measured Isothermal 12
! Temperature Coefficients !
A5 Comparison of Calculated and Measured Power Coefficient 14 i s
A6 Cycle 1 CEA Worths 15 A7 Cycle 2 CEA Worth- 15 t
l l A.8 Cycle 3 CEA Vorths 16 J !
A9 Cycle 4 CEA Worths 16 l A.10 Cycle 5 CEA Vorths 17 5
3 A ll Cycle 6 CEA Worths 17 i j A 12 Cycle 7 CEA Vorths
> 18 ,
) A 13 Cycle 8 CEA Vortii. 18 l
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j OPPD NA 8302 NP App. A. Rev. 02 j 11 1
i A,0 VERIFICATION OF NEUTRONICS MODEl.S FOR FORT Call!OUN STATION O The District has performed extensive verification of the neutronics models ,
used in the reload core analyses. The results of the previous District verification efforts were reported in Reference A 1. This effort utill::ed l, cross sections produced by CEPAK, The methodology discussed in this re-port utili:es cross sections produced by DIT. In order to demonstrate the j l District's ability to utilize the models with the DIT cross sections, addi. ,
, tional verification was undertaken. I I I a
] This verification is in addition to the extensive verification of these I
methods done by Combustion Engineering (CE) and reported in Reference A 2.
It is not the District's intent to repeat CE's extensive verification ef- i fort which includes a statistical assessment of the adequacy of the uncer- l
- tainties used by both CE and the District. Rather, it is the District's i
- intent to demonstrate that the District can adequately model the Fort Cal. I houn core and that the results of the District's verification effort are ;
consistent with those reported in Reference A 2.
O The District's verification using DIT cross sections utilizes data record-l ed for Cycles 6, 7 and 8. Benchmarking was performed for the prediction !
of overall core reactivity, power distributions, reactivity coefficients, Cf.A worth and Xenon reactivity. The results of the verification efforts include data for both CEPAR and DIT cross sections. I
,I 1 The verification uses experimental data from the Fort Calhoun reactor and i
\
independent calculations performed by CE and Exxon Nuclear Company (ENC).
Experimental data is obtained from startup tests and core follow programs. ;
Calcu14tional data is obtained from startup predictions, special analysis of startup tests or design lifetime computations.
j A.1 cnre Reactivity 3
The analysis of predicted reactivity for Fort Calhaun Station util-1:es studies of startup tests and plant data obtained during opera-tion at power. The parameter used to measure reactivity is the i
critical boron concentration.
i j OFFD NA 8302 NP, App. A, Rev. 02 j Page 1 of 18 l
A.0 VERIFICATION OF NEUTRONICS MODELS FOR FORT CA1)!OUN STATION (Continued) l O A.1 Core Renetivity (Continued) t Comparisons between measured and calculated critical boron concentra- !
tions for the unrodded HZP core are presented in Table A.1. The re.
l sults using the DIT cross sections arc consistent with those reported '
in Reference A.2. '
Operating plant data has been analyzed and evaluations of core reac.
tivity predictions carried out. There is little difference between calculated curves utilizing either CEPAK or DIT cross sections for cycles 6, 7, and 8. Results for the operating plant data comparisons demonstrate the District's ability to calculate core reactivity.
A.2 Power Distributions Extensive comparisont of power distributions have boen performed for Fort Calhoun and other CE reactors. Tnese <:omparisons are contained in References A.2 and A.3. The data given for Fort Calhoun in Refer.
enee A 3 vere supplied by the District.
A.2.1 Radial Power Distributions The District has performed comprehensive core follow calcula.
tions since the start of Cycle 3 in 1976. Table A.2 sum:nar.
1:es the results of comparisons between the axiaMy integrat.
ed assembly power as calculated by ROCS and that measured by CECOR usin5 the self.povered rhodium detector for Cycles 3, 4, 5, 6, 7 and 8. These cortparisons are only performed for instrumented assemblies because CECOR calculates the power for non instrumented assemblies using coupling coefficients O
OPPD NA 8302.NP, App. A, Rev. 02 Page 2 of 18
1 A.0 ERIFICATION OF NEUTRONICS MODELS FOR FORT Caul 0UN STATION (Continued) l A.2 Power Distributions (Continued)
A.2.1 Radial Power Distributions (Continued) derived from the physics codes. The instrumented assembly powers are calculated by a method independent of the predic-tive code. Sample comparisons for Cycle 8 are included in thAs document.
The extensive comparison between the calculated and measured radial power distributions verifies the capability of the Dis-trict to calculate these power distributions.
A.2.2 Axial Power Distributi.qng f The District has benchraarked the ROCS code against CECOR mea.
sured axial power distributions. This document contains com-parisons of core average and selected assembly axial power distributions fer Cycles 5 through 8.
The District has benchmarked the QUIX code against measured data by comparing the QUIX calculated ASI and the CECOR mea-sured ASI during an axial oscillation test peric, rued during Cycle 8 power ascension testing. The lead bank CEA's re.
mained in the core during the entire test.
The comparisons demonstrate the District's capability to cal.
culace axial power distributions using both ROCS and QUIX.
O OPPD NA 8302 NP. App. A. Rev. 02 Page 3 of 18
a A.0 ?"RIPICATION OF NZUTRONICS MODELS POR PORT CAIJ10UN STATION (Continued) l A,3 Renetivity coeffici m l
i I
The capability of the District's ROCS model to ptedict the Isothermal Temperature Coefficient (ITC) and the Power Coefficient (PC) has been '
l j benchmarked against physics tests conducted at Port Calhoun for all j operating cycles. Table A 3 shows the comparison between calculated l l and measured ITC's for zero power startun testing at the beginning of 4
the cycle. Also included are calculations performed by ENC, using '
XTC. The comparison of measured and calculated ITC's for "at power" 2
conditions is shown in Table A 4. The comparison of measured and ,
I
, calculated Power coefficients is shown in Table A.S. In all cases,
) the ROCS code accurately predicts the behavior of the core and the
- results using the DIT cross sections are consistent with results 7
j reported in Reference A 2. l j A.4 CEA Reactivity Vorth l
'O The District has extensively benchmarked the ROCS code against mea. >
sured and independently calculated values of CEA reactivity worth. ,
1 Tables A 6 through A 13 show the results of this benchmarking effort. l CE performed the PDQ calculations for Cycles 1, 2 and 4. ENC per. i formed the XTC calculations for Cycles 6, 7 and 8. The District per.
formed all 2 D ar.d 3 D ROCS calculations and the cycle 5 PDQ calcula-
! tions. These results demonstrate the District's capability to calcu-late CEA vorths and the results using DIT cross sections are consis- ,
tens with the results reported in Reference A 2.
}
A.5 Co-carisons en f'ritical W Positions Fo11 ovine a Reseter Trio l
]
1 l Another measure of the ability of the 3 D ROCS model to accurately
] needict reactivity changes is its ability to predict the critical I
boron concentration and CEA position following a reactor trip. A l
- study of this type was done for criticalities during the recovery I
I
} l 1 1 1
OPPD SA 8302 NP. App A. Rev. 02 Page 4 of 18
,1
A.0 VERIFICATION OF NEUTRONICS MODE 1.S FOR FORT CAUIOUN STATION (Continued) l O A.5 Comnarisons to Critical CFA Positions Fo11 ovine a Reactor Trio l (Continued) from a reactor trip for Cycle 2. This study demonstrates the ability of the District's ROCS model to ac.urately model the power defect and xenon buildup and decay.
A6 Comnarison to indeoendent Radial Po.,'er Distribution Calculations l Comparisons between the District's ROCS model calculations and ENC XTC model calculations of the !!FP radial power distributions have been performed. The comparisons show good agreement between the independent models.
O O
OPPD NA 8302 NP, App. A, Rev. 02 Page 5 of 18
REFERENCES O A1 CEN 242 (0) P. OPPD Responses to NRC Questions on Fort Calhoun Cycle 8 February 18. 1983, A2 CEN 226 P, "The ROCS and DIT Cornputer Codes for Nuclear Design." l Decceber, 1981.
A3 CESPD 153 P. "INCA /CECOR Power Peaking Uncertainty." May, 1980, {
i
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O OPPD NA 8302 NP. App. A. Rev. 02 Page 6 of 18 l
Table A 1 Unrodded HZP Critical Boron Concentrations calculations 3 D* 2 D* 3D Measured ROCS ROCS ROCS PDQ CXI.l.t een (CEPA)) (CEPA)) (DIT) (CEPT #1 KIG ,
I 1 933 1
2 1240 3 1000 4 1027 5 1242 1
6 1230 7 1241 8 1240
- A 20 ppm bias has been applied to these calculations, i
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OPPD NA 8302 SP. App. A. Rev. 02 Page 7 of 1S L .-. _ _ - - .- - - -
Table A 2 Sum:nsry of Comparisons of Measured and Calculated Integral tussembly Relative Power Densities Cycle CEPAK DIT ?!ominal Burnup Power CYele (%'D/Mnn (6) m n of Full Power 3 80 45 3 177.5 60 3 510 95 3 800 100 3 1000 100 3 1428 100 3 2510 100 3 3100 100 >
3 3500 100 3 4000 100 3 4500 100 i
3 5200 100 3 5800 100 3 6400 100 3 7200 90 3 7715 80 4 200
] 100 4 1000
) 100 1
4 2000 100 1 4 3000 100 4 4000 100 4 5000 100 !
4 6000 100 i
i OPPD !;A 8302 ??P, App. A. Rev. 02 Page 8 of IS l
i
!L Table A 2 Surmary of Cornparisons of Measured and Calculated Integral Asserbly Relative Power Densities (Continued)
(
Cycle CEPAX DIT Nominal i Burnup Power Cif,,lt, f M'.JD/MTU) (t3 it), 4 of Ful1 Power
! 4 7000 100 4 8200 100 !
5 300 100 l t
5 1000 100
' l 5 2000 100 4
F q 5 3000 100 j 5 4000 100 1 4
5 5000 100 I j 5 6000 '
100 j 6 50 66 d
6 500 100 l 6 1000 100
- 6 2000 i
90 ,
6 i 3000 65 t 6 4000 75 i 6 5000 75 i
j 6 5800
\
75 6 6500 l 100 1 6 7500 50 !
)
l 6 8500 95 !
6 9500 95 I 6
(
10500 95 !
l 7 135 _ _ 70 -
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OPPD NA 8302 NP, App. A, Rev. 02 i j
Page 9 of 18 '
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Table A 2 9 Su:raary of Comparisons of Measured and Calculated Integral Assembly Relative Power Densities (Continued) l j
, Cycle CEPAK DIT Nomini.1 Burnup Power m
cycle (WD/MTt') (g) n of Full Power r
~
7 500 100
- 7 1000 100 7 2000 100 ,
7 3000 100 7 4000 100 l 7 5000 100 l l 7 6000 100 [
J 7 7000 100 l
\
7 8000 l 100 j
- 7 a
9725 100 '
l I 8 50 45 l 8 250 100 8 1000 '
4 100 l 8 2000 ,,,, 100 I l
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OPPD NA 4302 NP, App. A. Rev. 02 Page 10 of 18
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Table A .1, Low I'over Physics Isothermal Tertperaturo Coef ficiones P,o ro n CCPAK* DIT Concentration !!easured ROCS ROCS XTG cveln vern) (?re'F) (Ae/*F) (Ac/'Fi (Ar/*Pi 1 993 0 26 x 10
1 854 0.11 x 10
2 1240 0.41 x 10
2 1198 0.32 x 10
2 1164 0.09 x 10
l 3 1000 0.078 x 10 j i
4 1020 0.14 x 10*' !
l 5 1228 0.20 x 10*'
6 1213 0.23 x 10
7 1213 0.12 x 10*'
8 1240 0.16 x 10
i Calculated results were biased by 0.20
- 10 ap/'r l I i i I
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OPPD ??A 8302 NP. App. A. Rev. 02 Page 11 of 18
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- Table A 4
\
, Comparison of Calculated and Measured '
q Isothermal Temperature Coefficients a
BOC f
I !
- Calculated
- Calculated !
1 Critical Boron Measured CEPAX ROCS DIT. ROCS i Percent of Concentration jTC TC ITC Cycle Rated Power (ppm) (x10 op/'F) (x10{op/'F) (x10 op/'r) I 1
I ~
4 2 69(1) 927 0,28 j 3 46(I) 720 0.41 '
i 4 92(1) 690 0.42 5 93(1) 876 0.19 j 6 95(1) 848 0.46 l 7 96(2) 817 0.52 !
l 8 79(2) 817 0.84 -
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l i OPPD NA 8302 NP, App, A. Rev. O.
1
) Page 12 of 18 '
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l Table A 4 Comparison of Calculated and Measured
.l Ov Isothermal Temperature coefficients (Continued) 3 EOC l
2 4
Calculated
- Calculated i
' Critical Boron Measured CEPAX. ROCS DIT. ROCS ,
Percent of Concentration iTC iTC ITC Cycle Rated Power (ppm) (x10"ap/*F) ( x10'*ar/
- F) (x10'ap/'F) i
) 1 75(1) 239 0.98 3 2 46(1) 104 1.62 1
3 90(1) 62 1.65 I,
4 95(1) 44 1.41 1 5 94(1) 296 0.97 6 96(2) 307 1.51 7 95(2) 192 1.85 -
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I (1) Full Rated Power - 1420 M'Je '
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j (2) Tull Rated Power - 1500 M'Je
' t BOC calculated results were biased by 0.20 x 10 ar/'T and EOC l i
1 calculated results were biased by 0.40 x 10*' ar/'T.
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- OPFD.NA 8302.SP, App. A. Rev. 02 Page 13 of 18
l i Table A 5 -
Coreparison of Calculated and Measured Power Coefficients CEPAL ROCS DIT. ROCS Percent Measured Calculated Calculated j of Critical Power Power Power Burnup Rated Boron Coeff. Coeff. Coeff.
M %'D /MTU h Cone. (Ar/4 Power) (or/4 Power) (6e/4 Power)
J 2 10877 46(1) 104 1.95 x 10*'
j 3 157 46(1) 720 1.47 x 10*'
) 3 1513 90(I) 535 1,12 x 10*'
3 4183 90(1) 309 1.31 r. 10*'
3 7208 90(1) 62 1.48 x 10*'
i 4 267 92(1) 690 1.04 x 10*'
4 4690 94 ( '* ) 288 1.12 x 10*' !
l 4 8027 95(1) 44 1.10 x 10*'
5 426 93(1) 876 1.05 x 10*'
l 5 6815 94(1) 296 1.25 x 10*'
6 400 95(1) 848 1.11 x 10*'
6 6467 96(2) 307 1.45 x 10*' {
7 450 96(2) 817 0.98 x 10*'
, 7 6900 95(2) 283 1.30 x 10*' [
7 7800 95(2) 192 1.57 x 10*'
8 459 79(2) 817 1.18 x 10*' -
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(1) Tull Rated Power - 1420 K.'t '
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(2) Pull Rated Power - 1500 K.'t '
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i i l 3 OPPD NA.8 302.NP, App. A. Rev. 02 j Page 14 of 18 l
)
Table A 6
[%.j j Cycle i CEA Worths Calculated Calculated Calculated CEPAK ROCS CEPAK-ROCS CEPAK-PDQ Measured 3D 2D 2D Group (S60) (S60) J o) (too) 4 0.58 3 0.57 2 2.01 A 3.06 B 2.10 Total (4+3+7+A+B) 8.32 J
l Table A 7
^
\, Cycle 2 CEA Worths *
(Q Calculated Calculated Calculated CEPAK ROCS CEPAK ROCS CEPAK PDQ Measured 3D 2-D 2-D Q.Ipq92 (460) (too) (too) (too) 4 0.65 3 0.41 2 1.67 1 0.95 Total (4+3+2+1) 3.68 1
D (G
OPPD NA 8302 NP, App. A. Rev. 02 Page 15 of 18 L _ _ _ _ _ _ . _ _ _ _ _ _ - - - -
Table A 8
) Cycle 3 CEA Worths a
Calculated Calculated CEPAK-ROCS CEPAK-ROCS Measured 3-D 2D Group (460) (Soo) ,__f*or) 4 0.74 I 3 0.59 2 1.96 1 0.80 Total (4+3+2+1) 4.09 Table A 9 O Cycle 4 CEA Worths Calculated Calculated CEPAK ROCS CEPAK ROCS Measured 3D 2D Group (g (gao) (goo)
, ml 4 0.63 3 0.60 2 1.90 1 0.92 Total (4+3+2+1) 4.05 l l
(O d i 1
OPPD NA 83('2-NP. App. A Rev. 02 Page 16 of 18
, Table A 10 l
( Cycle 5 CEA Worths Calculated Calculated Calculated CEPAK ROCS CEPAK ROCS CEPAK PDQ Measured 3-D 2D 2D Group (560) (560) (S60) (SAo) 4 0.57 3 0.67 2 1.40 1 0.99 Total (4+3+2+1) 3.63 -
Table A ll Cycle 6 CEA Worths O Calculated Calculated Calculated Calculated XTC CEPAK ROCS CEt'AK PDQ CEPAK PDQ Measured 3D 2-D 3D Groun (Soo)_ (S60) (SAo) (S60) (SAo) 4 0.52 3 0.66 2 1.57 1 0.93 Total (4+3+2+1) 3.68 I
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OPPD !A 8302 !!P, App. A. Rev. 02 Page 17 of 18
~
Table A 12 Cycle 7 CEA Worths Calculated Calculated Calculated Calculated XTG CEPAK ROCS CEPAK-PDQ CEPAK-PDQ Measured 3D 2-D 3D Group (*60) (660) (S6o) (SAo) (%Ao) 4 0.49 3 0.47 2 1.65 1 0.66 Total (4+3+2+1) 3.27 _
Table A 13 Cycle 8 CEA Vorths
(^)S
\.
Calculated Calculated Calculated Calculated XTG CEPAK-ROCS CEPAK-PDQ CEPAK PDQ Measured 3D 2-D 3-D Croim (S6o) (Soo) (Soo) ($6o) (too)
~
4 0.58 3 0.63 2 0.99 1 1.00 Total (4+3+2+1) 3.20 ..
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OPPD NA 8302 NP, App. A. Rev. 02 Page 18 of 18
. O OMAHA PUBLIC P0k'ER DISTRICT i NUCLEAR ANALYSIS l
ONGOING BENCHMARKING PROGRAM llISTORICAT. INFORETION i.
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l OPPD NA 8302 NP APPENDIX B Rev. 02 O :
, t April 1988 f l
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- - . . ~~ ,-~~~,-a--~~. --
t LIST OF TABLES l Ci U TABLE TITLE EAEE B1 Unrodded HZP Critical Boron concentrations calculations 1 B2 Low Power Physics Isothermal Temperature coefficients 2 B3 Comparison of Calculated and Measured Isothermal 3 Temperature Coefficient t
B-4 Comparison of Calculated and Measurei Power Coefficient 5 B5 Cycle 1 CEA Worths 6 i B6 Cycle 2 CEA Worths 6 ,
B7 Cycl? 3 CEA Worths 7 B-8 Cycle 4 CEA Wotths 7 B9 Cycle 5 CEA Worths 8 B 10 Cycle 6 CEA Worths 8 B ll Cycle 7 CEA Worths 9
) B 12 Cycle 8 CEA Worths 9 B 13 Cycle 9 CEA Worths 10 l
B 14 Cycle 10 CEA Worths to l 1
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l OPPD NA 8302 NP App. B. Rev. 02 l 1
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Table B 1 Unredded ll2P Critical Boron Concentrations calculations l
l 3 D* 2 D* 3D .
!teasured ROCS ROCS ROCS PDQ I Cvele non (CEPAK) (CEPA}3 (DIT) 10EPAP) KIG 1 933 2 1240 1
3 1000 4 1027 5 1242 6 1230 7 1241 8 1240 9 1518 l 10 1474
~
- A 20 ppm bias has been applied to these calculations.
r i
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OPPD !!A 8302 tiP, App. B, Rev. 02 Page 1 of 10
I 1
Table B 2 1
/-
( }j Low Power Physics Isothermal Temperature Coefficients l
Baron CEPAK* DIT Concentration Measured ROCS ROCS XTC Cvele 'momi (6 c /' F) (Ac/'F) (6c/'F) (or/'F) 1 993 0.26 x 10
1 854 0.11 x 10
2 1240 0.41 x 10
2 1198 0.32 x 10*'
2 1164 0.09 x 10
3 1000 0.078 x 10
4 1020 0.14 x 10
5 1228 0.20 x 10""
6 1213 0.23 x 10
() 7 1213 0.12 x 10
8 1240 0.16 x 10
9 1457 0.30 x 10
10 1457 0.23 x 10
- Calculated results were biased by 0.20 x 10 op/*F i I
1 i /T l
OPPD NA 8302 NP App. B, Rev. 02 Page 2 of 10 l
Table B 3
(,-~') Comparison of Calculated and !deasured
Isothermal Temperature Coefficient BOC Calculated
- Calculated Critical Boron Measured CEPAK ROCS DIT ROCS Percent of Concentration .
TC ITC Cycle Rated Power (ppm) (x10{TC Ap/'F) (x10{Ap/*F) (x100 Ap/'F) 1 -
2 69(1) 927 0.23 3 46(1I 720 0.41 4 92(1) 690 0.42 5 93(1) 876 0.19 6 95(1) 848 -0.46 7 96(2) 817 -0.52 8 79(2) 817 0.84 9 94(2) 1036 0.39 10 95(2) 1017 0.48 _ ,,,
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G OPPD t;A 8302 NP, App. B, Rev. 02 Page 3 of 10 1
Table B 3 p)
- , Comparison of Calculated and Measured Isothermal Temperature coefficient (Continued)
EOC Calculated
- Calculated Critical Boron Heasured CEPAK-ROCS DIT ROCS Percent of Concentration TC ITC Cycle Rated Power (ppm) (x10{ap/*F) (x10{TC ap/'F) (x10 ap/*F) 1 75(1) 239 -0.98 2 46(1) 104 1.62 3 90(1) 62 1.65 4 95(1) 44 -1.41 5 94(1) 296 0.97 6 96(2) 307 -1.51 7 95(2) 192 -1.85 8 95(2) 292 1.86 m 9 96(2) 300 -1.46
, 10 95(2) 302 1.53 -
~
(1) Full Rated Power - 1420 FNt l (2) Full Rated Power - 1500 !C.'t
- BOCcalculatedresultswerebiasedby0.20x10'Iap/'Fand EOC calculated results were hiased by 0.40 x 10'" ap/'T.
O 1 U !
l OPPD NA 8302 NP, App. B, Rev. 02 Page 4 of 10 l
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Table B 4 n l Comparison of Calculated and Measured Power Coefficient (v)
CEPAK ROCS DIT ROCS Percent Critical Measured Calculated Calculated of Boron Power Power Power Burnup Rated Conc. Coeff. Coeff. Coeff.
Cvele MWD /MT0 Power (com) (60/6 Power) ( Ao /S Power) (60/* Power) 2 10877 46(1) 104 -1.95 x 10 4 3 157 46(1) 720 1.47 x 10
3 1513 90(1) 535 1.12 x 10
3 4183 90(1) 309 1.31 x 10*'
3 7208 90(1) 62 -1.48 x 10*'
4 267 92(1) 690 1.04 x 10
4 4690 94(1) 288 1.12 x 10
4 8027 95(1) 44 -1.10 x 10
5 426 93(1) 876 1.05 x 10
5 6815 94(1) 296 1.25 x 10
6 400 95(1) 848 -1.11 x 10
p 6 6467 96(2) 307 -1.45 x 10
O 7 450 96(2) 817 -0.98 x 10
7 6900 95(2) 283 -1,30 x 10
7 7800 95(2) 192 -1.57 x 10
8 459 79(2) 817 1.18 x 10
8 6150 95(2) 292 1.70 x 10
9 420 94(2) 1036 -1.64 x 10
9 9663 96(2) 300 1.57 x 10
10 583 95(2) 1017 1.24 x 10"4 10 9261 95(2) 302 1.40 x 10 - J t
(1) Full Rated Power - 1420 MVt (2) Full Rated Power - 1500 !Nt tO V
OPPD t;A 8302 tiP. Anp. B. Rev. 02 Page 5 of 10
Table B 5 f3 Cycle 1 CEA Worths
's /
calculated Calculated calculated CEPAK ROCS CEPAK ROCS CEPAK-PDQ Measured 3D 2D 2-D Croun (Soo) ($60) (*Ao) (SAo) 4 0.58 3 0.57 2 2.01 A 3.06 B 2.10 Total
( 4 + .' + 2 .-A + B ) 8.32 -
Table B-6 O Cycle 2 CEA Worths V
Calculated Calculated Calculated CEPAK ROCS CEPAK ROCS CEPAK PDQ Measured 3D 2-D 2-D Group (too) (sao) (Soo) (too)
~
4 0.65 3 0.41 2 1.67 1 0,95 Total (4+3+2+1) 3.68 _ ,
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i OPPD NA 8302 NP, App. B Rev 02 Page 6 of 10 1
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Table B-7 i Cycle 3 CEA Worths l Calculated Calculated CEPAK-ROCS CEPAK-ROCS Measured 3D 2D Group (sor) (tool (Soo) 4 0.74 3 0.59 2 1.96 1 0.80 Total (4+3+2+1) 4.09 '
Table B 8 Cycle 4 CEA Worths Calculated Calculated CEPAK ROCS CEPAK ROCS Measured 3-D 2-D Group (too) (soo) (soo) .
4 0.63 3 0.60 2 1.90 1 0.92 Total (4+3+2+1) 4.05 l
OPPD !,A 8302 t1P, App. B Rev. 02 Page 7 of 10
l l
Table B 9 4
V) Cycle 5 CEA Worths i 1
Calculated Calculated Calculated CEPAK ROCS CEPAK ROCS utPAK-PDQ Measured 3D 2D 2D Grnun (S60) (Soc) (Soc) (S60) 4 0.57 3 0.67 2 1.40 1 0.99 Total (4+3+2+1) 3.63
~
Table B 10 D
(V Cycle 6 CEA Worths Calcula.ed Calculated Calculated Calculated XTC CEPAK ROCS CEPAK PDQ CEPAK PDQ Measured 3-D 2-D 3-D Group (SAO) (S60) (S60) (460) (S60)
~
4 0.52 3 0.66 2 1.57 1 0.93 Total (4+3+2+1) 3.68 ~
l
(
U OPPD NA 8302 NP, App. B. Rev. 02 i Page 8 of 10 l
Table B 11 7-( ) Cycle 7 CEA Worths
\_)
Calculated Calculated Calculated Calculated XTC CEPAK ROCS CEPAK-PDQ CEPAK PDQ Measured 3-D 2D 3D C rot _m (S Ae( (S60) (S60) (Sao) (Sao)
. f 4 0.49 3 0.47 2 1.65 1 0.66 Total (4+3+2+1) 3.27 -
Table B 12 Cycle 8 CEA Worths
/~S b
Calculated Calculated Calculated Calculated XTG CEPAK ROCS CEPAK PDQ CEPAK PDQ Measured 3D 2D 3D Group (160) (160) (%Ao) (too) (SAo)
~
4 0.58 3 0.63 2 0.99 1 1.00 Total (4+3+2+1) 3.20 _ _
) (OQ OPPD NA 8302 NP, App. B, Rev. 02 Page 9 of 10
l TABLE B-13 ;
1 Cycle 9 CEA Worths Calculated DIT ROCS Measured 3D Grotm (160) (560)
- ~,
4 0.67 3 0.72 2 1,63 1 0.84 Total (4+3+2+1) 3.86 TABLE B 14 Cycle 10 CEA Worths Calculated DIT ROCS Measured 3D Groun (sor) (soo) 4 0.61 3 0.62 1
2 1.94 1 0.85 t
Total '
(4+3+2+1) 4.02 _
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OPPD tiA 8302 tiP, App. B Rev. 02 Page 10 of 10 i
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2 Omaha Public Power District l l Nuclear Analysis l l Reload Core Analysis Methodology ;
i
.a i i l i Transient and Accident Methods and Verification !
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i OPPD NA 8303 NP l 1 .
4 Rev, 02 i
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i April 1988 1
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f ABSTRACT DI This document is a Topical Report describing Omaha Public Power District's re-load core transient and accident methods for application to Fort Calhoun Station Unit No. 1.
The report addresses the District's transient and accident analysis n.ethodology and its application to the analysis of reload cores. In addition, comparisons of results using the NSSS simulation code to results from experimental measure- i ments and independent calculations are provided.
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OPPD NA 8303.NP, Rev. 02 i
Table of Contents V 1.0 INTRODUL.' ION AND
SUMMARY
l 2.0 CHAPTER 14 EVENTS CONSIDERED IN THE RELOAD CORE ANALYSES 1 2.1 Criteria 2 2.2 USAR, Chapter 14, Safety Analysis Event.s Not Considered in Reload Core Analyses 3 2.2.1 Malpositioning of Group N CEAs !
(formerly Part Length CEAs) 3 2.2.2 Idle Loop Startup Event 4 2.2.3 Turbine Generator Overspeed Event 4 2.2.4 Loss of Load Event 4 2.2.5 Malfunctions of the Feedwater System 5 2.2.6 Steam Generator Tube Rupture Accident 6 2.2.7 Loss of Coolant Accident 7 2.2.8 Containment Pressure Analysis 7 2.2.9 Generation of Hydrogen in Containment 8 2.2.10 Fuel Handling Accident 8 2.2.11 Gas Decay Tank Rupture 8 2.2.12 Vaste Liquid Event 8 2.3 USAR, Section 14, Events Considered in a Reload Core Analysis 8 3.0 TRANSIENT AND ACCIDENT ANALYSIS AND TECHNICAL SPECIFICATIONS 9 4.0 TRANSIENT AND ACCIDENT ANALYSIS MODELS 10 4.1 Plant Simulation Model 10 !
4.2 DNBR Analysis Models 12 t 4.3 Application of Uncertainties 13 5.0 TPASSIENT AND ACCIDENT METHODS 14 5.1 CFA Vithdrawal Event 16 5.2 Boron Dilution Event 23 5.3 Control Element Assembly Drop Event 28 l 5.4 Four Pump Loss of Flow Event 32 5.5 Asymmetric Steam Generator Event 36 5.6 Excess Load Event 40 5.7 RCS Depressurization Event 47 5.8 Main Steam Line Break Accident 50 OPPD NA 8303 NP Rev. 02 11
Table of Contents (Continued)
,~
(
Section Eagg 5.0 TPANSIENT AND ACCIDENT MFTHODS (Continued) 5.9 Seized Rotor Accident 60 5.10 CEA Ejection Accident 64 5.11 Loss of Coolant Accident 67 5.12 Loss of Load to Both Steam Cenerators Event 68 5.13 Loss of Feedwater Flow Event 73 6.0 TRANSIENT ANALYSIS CODE VERIFICATION 79 6.1 Ine souction 79 6.2 Comparison to Plant Data 80 6.2.1 Turbine Reactor Trip 80 6.2.2 Four-Pump Loss of Coolant Flow 83 6.3 Comparison Between OPPD Analyses and Independent Analyses Previously Performed by the Fuel Vendors 85 6.3.1 Dropped CEA 86
()
6.3.2 Hot Zero Power Main Steamline Break 87 6.3.3 Hot Full Poser Main Steamline Break 88 6.3.4 RCS Depressurization 90 6.4 Summary 92
7.0 REFERENCES
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OPPD NA 8303 NP, Rev. 02 111
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LIST OF TABLES j TABLE TITLE 26EE i 501 Reactor Protective System and Safety Injection 15 5.1 1 Initial Conditions Assumed in CEAW Event Analysis 20 5.3.4 1 Key Parameters Assumed in the Full Length CEA Drop 31 Analysis 5.4.4 1 Key Parameters Assumed in the Loss of Coolant Flow 34 Analysiv i 5.5.4 1 Key Parameters Assumed in the LL/1SG Event 39 i
5.6.4 1 Key Parameters Assumed in the Excess Load Event Analysis 45 5.7.4 1 Key Parameters Assumed in the RCS Depressurization Event. 49
) Analysis ,
j 5.9.4 1 Key Parameters Assumed in the Seized Rotor Analysis 62 [
', 5.12.4-1 Key Parameters Assumeo n the Loss of Loa'd to Both Steam 71
?
Generators Analysis l 5.13.4 1 Key Parameters Assumed in the Loss of Feedwater Flow 77 Analysis ,
6.3 1 Comparison of Parameters Including Uncertainties Used in 93 the CEA Drop Analysis for Cycles 6 and 8
', 6.3-2 Comparison of Parameters Including Uncertainties Used in 94
^
i the HZP Main Steamline Break Analysis for Cycles 6 and 8 ,
6.3 3 Comparison of Parameters Including Uncertainties Used in 95 l the HFP Main Steamline Break Analysis for Cycles 6 and 8
]
6.3 4 Comparison of Parameters Including Uncertainties Used in 96 the RCS Depressurization Analysis for Cycles 6 and 8 I i
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-- - ..--- ---- --.- - - --,- - - - -.- ----,- -- - -- - - J
A +- =-h
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LIST OF FIGURES FIGURE TITLE E6EE 11 Full Power Turbine. Trip Nuclear Power vs Time 99 1-2 Full Power Turbine Trip P essurizer Pressure vs Time 100 13 Full Power Turbine Trip Pressurizer Level vs Time- 101 <
14 Full Power Turbine Trip RCS Temperatures vs Time 102' .
1-5 Full Power Turbina Trip Steam Generator Pressure vs Time 103 16 Full Power Turbine Trip Main Feedwater Flow vs Time 104 17 Full Power Turbine Trip Steam Flow vs Time 105 r
21 4 Pump Loss of Flow Total RCS Flow vs Time 106 i
22 4 Pump Loss of Flow Pressurizer Pressure vs Time 107 23 -4. Pump Loss of Flow Pressurizer Level vs Time 108 24 4. Pump Loss of Flow Core Power vs Time 109 l 25 4 Pump Loss of Flow Steam Generator Pressure vs Time 110 26 4 Pump Loss of Flow RCS Temperatures vs Time 111 !
27 4 Pump Loss of Flow Main Feedwater Flow vs Time 112 28 4 Pump Loss of Flow Steam Flow vs Time 113 ,.
1 31 CEA Drop Incident Core Power vs Time 114 i 32 CEA Drop Incident Core Average Heat Flux vs Time 115 ,
33 CEA Drop Incident Coolant Temperature vs Time 116 l i
34 CEA Drop Incident Pressurizer Pressure vs Time 117 ,
f 41 Zero Power Steam Line Break Incident Core Power vs Time 118 42 Zero Power Steam Line Break Incident Core Average Heat 119 Flux vs Time l 43 Zero Power Steam Line Break Incident Total Reactivity 120 vs Time O
OPPD NA 8303 NP, Rev. 02 v
i LIST OF FIGURES o (Continued)
EICURE TITLE E6.GI 4-4 Zero Power Steam Line Break Incident Coolant System Pres- 121 sure vs Time 4-5 Zero Power Steam Line Break Incident Steam Generator 122 Pressure vs Time 51 Full Power Steam Line Break Incident Core Power vs Time 123 52 Full Power Steam Line Break Incident Core Average Heat 124 Flux vs Time 53 Full Power Steam Line Break Incident Total Reactivity Heat 125 Flux vs Time 54 Full Power Steam Line Break Incident Coolant System Pressure 126 vs Time 5 5A Full Power Steam Line Break Incident Reactor Coolant Temper- 127 atures vs Time: Cycle 6 5 5B Full Power Steam Line Break Incident Reactor Coolant Temper- 128
-s atures vs Time: Cycle 8 56 Full Power Steam Line Break Incident - Steam Generator Pres- 129 sure vs. Time
)
OPPD NA 8303 NP, Rev. 0?
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OMAIM PUBLIC P0b'ER DISTRICT RELOAD CORE ANALYSIS MET 110DOLOGY O TRANSIENT AND ACCIDENT MET 110DS AND VERIFICATION REVISION DATE 1 00 September 1983 01 November 1986 02 April 1988 l 9
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O l OPPD NA 8303 tiP, Rev. 02 vil
f Omaha Public Power District Reload Core Analysis Methodology O, Transient and Accident Methods and Verification t
i 1.0 1NTP.ODUCTIOF AND
SUMMARY
! I This repors discusses the methodology the Omaha Public Power District utili:es to analyze transients and accidents for reload cores. In addi-tion, the report discusser the District's verification of the combustion i t
j Engineering System Excursion Code (CESEC) for Fort Calhoun Station j transients. The purpose of this verification is to demonstrate the I l
l District's ability to properly utilize the CESEC code. ;
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{ The District's transient and accident analysis methodology for reload l i,
I cores is based upon the reanalysis of those Updated Safety Analysis
) t t
Report (USAR), Chapter 14 events whose consequences may be adversely F 1
I affected by changes in parameters associated with any reload core. The ,
- USAR Chaprer 14 events which must be considered during a reload core j analysis are discussed in Section 2.0. Section 3.0 discusses the trans-f
- ient analyses which determine certain parameters specified in the Tech- !
j nical Specifications. The District's transient analysis models are dis- l j cussed in Section 4.0. The District's application of these transient
- analysis models to the various Chapter la events is discussed in Section !
4 5.0. The verification of the NSSS simulator model used by the District !
l I is discussed in Section 6.0. References are provided in Section 7.0. j j
1 2.0 CHAPTER 14 EVENTS CONSIDERED IN THE RELOAD CORE ANALYSES l l
} This section discusses the criteria utilized to determine if a Chapter
} 14 event need be considered in reload core analyses. Each event which 1
j is not formally considered in a reload core analysis is discussed and i
l the reasons given for not normally including the event in the reload 1
a core analyser,. The methodology applied to these events will not be dis.
4 cussed in this report.
OPPD NA 8303 NP, Rev. 02 Page 1 of 129
2.0 CHAPTER 14 EVENTS CONSIDERED IN THE RELOAD CORE ANALYSES (Continued)
\- 2.1 Criteria The criterion used to determine the events' considered in reload core analyses is that changes in various neutronics parameters ad-versely affect the safety analyses of these events. The core para-meters considered are the pin peaking factors, FR and Fxy, the Moderator Temperature Coefficient (MTC), the Fuel Tenperature Coefficient (FTC) or Doppler Coefficient, the boron concentration, the inverse boron worth, the neutron kinetics parameters, the CEA reactivity worth and the cooldovn teactivity associated with a steam line break. If these parameters change such that the previ-ously reported results for a Chapter 14 event are no longer conser-
- er.t ive , then this event must be reanalyzed. If these parameters are conservative with respect to the values assumed in the refer-enced safety analyses, the criteria of 10 CFR 50.59 are met and this event is not reanalyzed. If a change in some of the para-meters may cause the results of a safety analyses to be nonconser-v vative, the event is reanalyzed. If the criteria for the event are still met, then the requirements of 10 CFR 50.59 are satis-fied. The event is reported as being reanalyzed and that it has been determined that no unreviewed safety question exists for the event. In some cases it may be possible that an event is reanal-yzed and it is determined that an unreviewed safety question exists. In these cases the analyses for these events are submit- '
ted. In addition, any safety analyses which are performed as a re-sult of a chana in the Technical Specifications are reported as part of the supporting docuaentation for a Facility License Change.
Criteria not directly associated with the reload core but which may be considered in a reload analysis are changes to plant sys-tems which would take place during a refueling and would first be utilized during the operar. ion of the subsequent core. In cases where either physical modifications or modifications in operating OPPD NA 8303 NP, Rev. 02 Page 2 of 129
2.0 C11 APTER 14 EVENTS CONSIDERED IN THE RELOAD CORE ANALYSES (Continued) 7 'x
( )
s/ 2.1 Criteria (Continued) procedures are made that do impact the safety analyses, the re-suits of the revised safety analyses are reported in a reload core analysis. This methodology report does not consider the methodol-ogy that is required to analyze all events which could be affected by this criteria, rather, if submic,tals are made which require analyses of events other than those discussed in this report, revi-sions to this methodology report will be made to incorporate the methodology used for those events.
2.2 USAR. Chapter 14. Safety Analysis Events Not Considered in Reload Core Analyses This section discusses the USAR, Section 14, safety analyses which are not normally considered in a reload core analysis. The USAR
(% section is discussed and the reasons for not including it in the
(/)
s scope of these analyses is discussed. Typically, the reasons for not analyzing these events are that the operating modes considered in the events are no longer allowable at Fort Calhoun Station, the event is not associated with any core parameters or the event is analyzed by a fuel vendor for the District.
2.2.1 Malcositionine of Grouc N CEAs (formerly Part Lencth CEAs)
This event is not analyzed in the reload core analysis because the Group N part length CEAs are replaced with full length CEAs starting with Cycle 11. The use of Group N during power operations will still be prohibited by the Technical Specifications. The drop of a Group N CEA is considered in the full length CEA analysis.
(
OPPD NA 8303 NP, Rev. 02 Page 3 of 129
2.0 CilAPTER 14 EVENTS CONSIDERED IN THE RELOAD CORE ANALYSES (Continued)
( ,) 2.2 USAR. Chaerer 14 Safety Analvmir Events Not Considered in Rel2ad Core Analvoe- (Continued) 2,2.2 Idle Loon Startuo Event This event is nor analyzed because part loop operation is not permitted by the Fort Calhoun Technical Specifications.
2.2.3 Turbine Generator Oversoced Event This event is an analysis of the consequences of a turbine wheel failure and is unrelated to any reload core changes 2.2.4 Loss of Load Event A. The loss of load to both generators is assessed to determine if:
f O The pressurizer safety valves limit the reactor cool-
- ant system pressure to a value below 110% of design pressure (2750 psia) in accordance with Section II?, of the ASME Boiler and Pressure Vessel Code, and suffi-cient thermal margin is maintained in the hot fuel assembly to assure that Departure from Nucleate Boll-ing (DNB) does not occur throughout the transient.
This event is not analyzed with respect to the first i
criteria since the relief capacity of the pressurizer safety valves does not change and the initial energy contained in the reactor coolant system will not change unless power level is raised above 1500 MV or the reactor coolant system inlet temperature is sig-nificantly increased. Section 14.9 of the USAR re-ports that the DNBR for the loss of load transient gs never decreases below the initial value considered in i
OPPD NA 8303 NP, Rev. 02 Page 4 of 129
1 l
J 2.0 CllAPTER 14 EVENTS CONSIDERED IN Tite RELOAD CORE ANALYSE 0 (Continued) l i
O l
2.2 USAR. Chanter 14 e s fe E Analysis Events Not Considered in Reload Core Analyses (Continued) l 2.2.4 Loss of Lead Event (Continued) the analysis nerefore, it is concluded that any change in a parameter which could effect the DNBR for this event would much more significantly effect other events and that it is nat necessary to analyze this event with respect to DNBR criteria.
Steam generator tube plugging performed during a re-fues ng outage has the potential for altering the heat transfer characteristics assumed in Seccion 14.9.1 of the USAR. Section 5.12 of this' document addresses the methodology to be employed should the steam generator f tube pluggf.ng exceed or be expected to exceed the cur-
~
rent USAR analysis assumptions.
B. The loss of load to one steam generator is discussed in this methodology report as one of the asymmetric steam generator transients.
2.2.5 Malfunctions of the Feedwater Systgg The analyses which are reported in USAR Section 14,10 Mal-functions of the Feeowater System, are the total loss of feedwater flow and the loss of feedwater heating. The re-sults of the total icas of feedvater flow show that the min-inwn DNBR does not decrease below its initial steady state value and that no safety limits are approached during the event. Therefore, this even. is not reanalyzed in a reload core analysis.
O OPPD NA 8303 NP. Rev. 02 Page 5 of 129
i 2,0 4
CHAPTER 14 EVENTS CONSIDERED IN THE RELOAD' CORE ANALYSES (Continued) 2.2 USAR. Chanter 14. Safety Analysis Events Not Considered in Relond Core Analyses (Continued) !
t
{
7 2.2.5 Malfunctions of the Feedwater System (Continued) i i
The loss of feedwater heating is the most adverse feedwater L malfunction eras of cooling on the RCS. This event. [
like the excess load event, is more limiting at EOC. This !
event has the same effect on the primary system as a small ,
- increase in turbine demand which is not matched by an in-
.j crease in core power. As a result, the DNRR degradation i
associated with this event is less severe than that for the [
excess load where a largo effective increase in turbine de-l mand is analyzed. The excess load event analysis is report-ed in Section 5.6 in this document, l
O Steam generatcer tube plugging performed during a refueling outage has the potential for degrading the heat transfer f
i j
characteristics assumed in Section 14.10.1 of the USAR for i j
,t the Loss of Feedwater Flow Event. Section 5.13 addresses ;
the methodology to be employed should steam generator tube plugging exceed or be expected to exceed the assumptions of
- the current Loss of Feedwater Flow Event. Reduced heat j
transfer for the Loss of Feedwater Heating Event does not #
g require reanalysis, since it is an overcooling event and the j i
increase in plugged tubes reduces the consequences of the !
event. ;
)
3 2.2.6 Steam Generator Tube Ruoture Accident
'i l
The steam generator tube rupture accident is analyzed to determine if the offsite dose acceptance criteria of 10 CFR 1
Part 100 is met. The analysis is a radioactive material 4
b i
4
{ OPPD NA 8303 NP, Rev. 02 Pye 6 of 129
2,0 CllAPTER 14 FNENTS CONSIDERED IN THE RELOAD CORE ANALYSES (Continued) ,
i 2.2 USAR. Chanter 14. Safety Analvris Events Not Considered in Reload f Egge Analyses (Continued) !
1 2.2.6 Steam Generator Tube Ruoture Accident (Continued) I 1
i ,
- release analysis based upon 14 failed fuel within the core.
- It is not dependent upon any reload core analysis related ,
j parameters, therefore, it is not analyzed in the reload core ,
analysis. In the future, the steam generator tube rupture accident analysis may be verified for high burnup fuel and/
or a change in heat transfer cheracteristics for an !ncrease i in the number of plugged tubes in the generators, i [
l I 1
j t
2.2.7 Loss of Coolant Accident !
) i The loss of coolant accident as reported in USAR Section
] 14.15, is analyzed for the District by CE. The large and 1
small break analyses were performed by CE. The Diserfct j
, confirms the assumptions used in these analyses are valid for each reload core. If reanalysis is required, the reanal- l ysis is done by a nuclear fuel vendor. The District does I not perform any loss of coolant accident analyses. .
!4 !.
}
} 2.2.8 Containtrent Pressure Analysis
[
d n !
Containment pressure analysis is dependent upon the initist f j liquid mass and energy contained ir. the primary or secondary f cystem. Since these parameters do not change when the cors !
j is refueled, the containment pressur. analysis is not done j i
j in a reload core analysis. I
- i I f l
i !
! l i i 1
1 h j OPPD NA 8303 NP, Rev. 02 I
! Page 7 of 129 !
i i
0 l
2.0 CilAPTER 14 EVENTS CONSIDERED IN THE RELOAD CORE ANALYSES -(Continued)
O 2.2 USA 2. Chanter 14. Safety Analysis Events Not Considered in Reload i
Core Analyses (Continued) 2.2.9 Generation of Hydreren in Containment The generation of hydrogen in containment analysis is inde-pendent of any reload core parameters, therefore, the anal-ysis is not performed during the course of a reload core analysis.
2.2.10 Fuel Handline Accident The fuel handling accident is a function of the isotopic inventory contained in the fuel pins. This is not normally considered in a reload core analysis, however, it may be necessary to reconsider this analyses for high burnup fuel.
2.2.11 Gas Decay Tank Ruoture The gas decay tank rupture is independent of any parameter associated with refueling the core. Therefore, the analysis is not performed during a normal reload core analysis.
2.2.12 Waste Lieutd Event l The waste liquid event analysis is not affected by refueling !
the core. Therefore, the vaste liquid event analysis is not !
porformed in the course of a normal reload core analysis.
2.3 USAR. Section 14. Events Corisidered in a Reload Core Analysis The reload core analysis consists of analyzing several events which are considered in the USAR and two events which previously were not O
OPPD-NA 8303 NP, Rev. 02 Page 8 of 129
2.0 CilAPTER 14 EVD;TS CONSIDERED IN Tile RELOAD CORE ANALYSES (Cone! aued)
) 2.3 IfSAR. Section 14. Events Considered in a Reload Core Analysis (Continued) l analyzed in the USAR. These events are analyzed in accordance with 8 the criteria discussed in this report and to determine if an unre- ,
viewed safety question would exist for a reload core. The USAR Chapter 14 events considered in a reload core analysis are the >
Control Element Assembly Vithdrawal (CEAV) event, the boron dilution event, the Control Element Assembly (CEA) drop event, the loss of coolant flow event, the excess load event, the steam line break accident, the CEA ejection accident and the seized rotor accident.
In addition, analyses are performed for incidents resulting from the malfunction of one steam generator and for the RCS depressurization event. The analysis for each of these events will be discussed in 4
detail in Section 5.0 of this report.
3.0 TRANSIENT AND ACCIDENT ANALYSIS AND TECHNICAL SPECIFICATIONS l Results of transient and accident analyses are used in the Technical Spec-ifications in two ways. The first way is that values from the Technical Specifications are included in the initial conditions of the transient analyses. These Technical Specifications guarantee thet the various trans-
, ient and acciosnt analysis acceptance criteria will not be exceeded if the i reactor is operated within the bounds of these Technical Specifications.
Technical Specifier tons of this typt. include the limits on FR, F xy, the PDIL and the Moderator Temperature Coefficient.
The second type of values factored into the Technica). Specifications are
] those that are determined by transient analysis. These parameters consist of tt.e transient response term applied to the TM/LP equation, the minimum required shutdown margin, the linear heat rate LCO and the DNBR LCO. The transient response term applied to the TM/LP equation in the Technical Spec- ]
j ifications is a result of the analysis of the RCS depressuri:ation event or
- O I
orFD NA 8301 NP, Rev. 02 Page 9 of 129 j 4
3.0 TRANSIENT AND ACCIDENT ANALYSIS AND TECHNICAL SPECIFICATIONS (Continued)
O <
b excess load event. The minimum required shutdown margin at hot shutdown conditions is determined by the steam line break accident. This value is also confirmed for the boron dilution event. The minimum required shutdown margin for cold shutdown and refueling shutdown conditions is determined by the boron dilution event or the five percent suberiticality requirement for i refueling, The values used in the linear heat rate LCO are typically deter-mined by the loss of coolant accident. These values are also confirmed for the dropped CEA event, The LCO on DNBR margin is calculated based on re-sults from the dropped CEA event, the loss of four pump flow analysis or the CEA withdrawal analysis. !
l 4.0 TRANSIENT AND ACCIDENT ANALYSIS MUDELS I 1
, The District utilizes the latest vetsion of the CESEC code (CESEC II. and hereafter referred to as CESEC) in the simulation of plant response to non. I LOCA initiating events. The District; utilizos the CETOP and TORC computer codes for calculation of DNBR during these events, 4
O t
! 4.1 Plant Simulation _Model i
r l
The District utilizes the CESEC digital e iputer code, References 4 1 (
and 4 2, to provide tha simulation of the Fort Calhoun Station nu- !
i clear steam supply system. Ine program calculates the plant response ;
) to non LOCA initiating events for a vide range of operating condi- [
- tions. Additional information on the model is provided in Reference I
- 4 3. The CESEC program, which numerically integrates one dimensional l r
l-mass and energy conservation equations, assunes a node / flow path net- y
] vork to model the NSSS. The primary system components considered in f the code include the reactor vessel, the reactor core, the primary (
]l coolant loops, the pressurizer, the stema generators ano the reactor coolant pumps. The secondary system ensponents include the secondary ;
side of the steam generators, the main steam system, the feedvater f j system and the various steam control valves. In addition, the pro-l gram models some of the control and plant protection systems.
- \
i i
OPPD t'A 8303 NP, Rev. 02
- l iage 10 of 12o j f
4.0 TPANSI2yr AND ACCIDENT ANA1.YSIS MODELS (Continued) s/
m 4.1 Plant simulation Model (continued)
The code self initializes for any given, but constant, set of reactor power level, reactor coolant flow rate and steam generator power shar-ing. During the transient calculations, the time rate of change in the system pressure and enthalpy are obtained from solution of the conservation equations. Those derivatives are then numerically inte.
grated in time under the assumption of thermal equilibrium to give the system pressure and nodal enthalp'.es. The fluid states recog-nized by the code are subcooled and saturated; superheating is 81<
loved in the pressurizer, Fluid in the reactor coolant system is assumed to be homogenous, Reference 4 1 provides a description of the CBSEC code, including the major models, and the input, output and plot packages.
The pressurizer model is described in Reference 4 1 and further dis-
) >
cussed in Reference 4 2. The District utilizes the vall heat trans.
1 for model to permit simulation of voiding in any node in which steam formation occurs. Voiding may occur in events such as a steam line break or steam generator tube rupture. Nodalization of the closure head, described in Reference 4 1 and further discussed in Reference 4 2, allows for the formation of a void in the upper head region when a the pressurizer empties. Flov to the closure head is terminated in simulations of those events in which natural circulation occurs and in those events such as the steam line break where this action delays safety inp etion, 4 ,
I The capabilities and limitations of the CESEC code are discussed in References 4 1 and 4 2. The District's CESEC model of Fort Calhoun Station is valid as indicated in Reference 4 3 for the transients discussed in Section 5 of this report, with the exception of the CEA Ejection Analysis and LOCA Analysis. The CESEC model is also valid for analysis of the loss of load, malfunctions of the feedvater sys- !
tem and the steam generator tube rupture incidents.
OPPD NA 8303 NP, Rev, 02 Page 11 of 129 l l
4.0 TRANSIENT AND ACCIDENT ANALYSIG MODELS (Continued) 4.1 Plant Simulation Model (Continued)
The CESEC code is maintained by CE on the CE computer system in Windsor, Connecticut. The District accesses the code through a time sharing system. CE maintains all documentation and quality assurance programs related to this code.
4.2 DNBR Analysis Models i
The DNBR analysis is currently performed usiag either the TORC code, Reference 4 4, or both the TORC and CETOP codes, Reference 4 5. The 1 TORC code is used as a benchmark for the CETOP code model. TORC solves the conservation equations, as applied to a three dimensional i representation of the open lattice core, to determine the local cool-ant conditions at all points in the core. Lateral transfer of mass and energy between neighh, ring flow channels (open core effects) are accounted for in the ec,1cu',atf on of local coolant conditions. These coolar,t conditions are then used with a Critical Heat Flux (CHF) cor-
', relation supplied as a code subroutine to determine the miniaws value i s
of DNBR for the reactor cora. The CE 1 CHF correlation (References j 4 6 and 4 7) is used for the Fort Calhoun reactor as approved in ;
i Reference 4 8. *he Detailed TORC code is used directly in the seised l rotor analysis.
j i j The CETOP code has been developed to reduce the computer time needed 1
y for thermal hydraulic analyses while retaining all of the capabili. i ties of the TORC design model. The CETOP model provides an addi-tional simplification to the c9nsetvation equations due to the spec.
9 ific geometry of t).e mode'. A complete description of the CETOP code is contained in Reference 4 5 and a descsiption of the District's application of the CETOP code is contained in Reference 4 9. '
't OPPD NA 8303 NP, Rev. 02 Page 12 of 129
4.0 TRANSIENT AND ACCIDENT ANALYSIS MODELS (Continued)
/"' _
(' ,
4.2 DNBR Analysis Modals (Continued)
The fraction of inlet flow to the hot assembly in r,he CETOP model is adjusted such that the model yields appropriate MDNBR results when compared to the results of the TORC analysis for a specified range of operating conditions.
The CETOP code is used to calculste DNBR for all transient analyses discussed in Section 5 with the exception of the seized rotor analy-sis, 4.3 Acclication of Uncertainties '
Uncertainties are taken into account either by deterministic or stat-istical methods. 'The deterministic method applies all uncertainties adversely and simultaneously when calculating the approach to a limit.
l Uncertainties in DNBR calculations are taken into account by statis-tical methods. The statistical method takes into account the like-lihood that the uncertainties will all be adverse. The statistical j method is discussed in Rofererce 4 10. In this method the impact of component uncertainties on DNBR is assessed and the DNBR SAFDL is in-creased to include the effects of the uncertainties. Since the uncer- ,
tainties are accommodated by the increased DNBR SAFDL in the statis-tical method, engineering ftetors are not applied to the DNBR analy-sis model. The statistical method of applying uncertainties is ap. i plied to the CEA withdrawal, CEA drop, loss of RCS flow, excess load, seized rotor and asymmetric steam generator event DNBR calculations.
I l
t b
f-~s t d
j OPPD NA 8303 NP, Rev. 02 1 I Page 13 of 129 t
5.0 TiW;SIENT AND ACCIDENT ANALYSIS METHODS b
\_,/ This section addresses the evaluation of the various transients and acci-dents that are performed during a reload core analysis. Specific methods are described for each transient and accident. For each accidene. or trans-ient the following material is described:
A. Definition of the Event - A brief description of the causes, conse-quences, and RPS trips involved in the incident.
B. Analysis Criteria A brief description of the classification of the event and the Specified Acceptable Fuel Design Limit (SAFDL) or the
, offsite dose criteria which must be met.
J C. Objectives of the Analysis A brief description of the methods that are used to assure that the criteria of the analysis are met.
D. Xey Parameters and Analysis Assumptions A description of the key parameters and assumptions used in the analysis.
i
)
O E. Analysis Method A description of the methodology employed by the !
+
District to analyze the event, f
F. Analysis Results and 10 CFR 50.59 Criteria The expected results of l the analysis and a discussion of the methods used to determine if the event meets the criteria of 10 CFR 50.59.
- l C. Conservatism of Results A description of the conservatism of the analysis. l The values of the trip setpoints and trip delay times used in these anal-yses are shown in Table 5.0 1.
l 1 I 1
l 1
OPPD NA 8303 NP Rev. 02 Page 14 of 129
i O O O l
Table 5.0-1 REAC10R HUIECTIVE SYSIDI 'IRIPS AND SAFEIY INIBCTION Used in Arnlysis Idp Setpoint Uncertainty Delav Time (sec) 6etpoint Illgh Rate-of-Qunge of Ptwer 2.6 dec/ min 10.5 dec/ min 0.4 2.1 dec/ min l
I liigh Power Invel 107% 5.0% 0.4 112%
Variable High Power Ieml 9.1% above set 0.9% 0.4 10% above power level to initial power a low of 19.1% level low Reactor Coolant Flow 95% f2% 0.65 93%
liigh Pressurizer Pressure 2400 psia f22 psi 0.9 2422 psia
'Ihermal MargirVIow PressureIII 1750 psia 122 psi 0.9 1728 psia Iow Steam Generator Pressure 500 psia 122 psi 0.9 478 psia Iow Steam Generator Water Isvel 31.2% of narrow 110 in. (5.7% of narruw 0.9 25.5% of span ,
range span range span)
Steam Generator Differential Pressure 135 psid f40 psi 0.9 175 psid (bntairment Pressure High 5 psig 10.4 psi 0.1 5.4 psig Illgh Pressure Safety Injection 1600 psia 122 psi 12(2) 1578 psia (1) Values represent the low limit of the thermal margirVlow pressure trip. 1he setpoinc of this trip is dime:ed in Reference 5-3.
l I2) DLmp start - loop valve opening time.
OPPD-!E-8303-P, Rev. 02
l 5.0 TRN;SIENT AND ACCIDENT ANALYSIS HETHODS (Continued) j O 5.1 CEA Withdrawal Event l.
P 5.1.1 Definirion of the Event A sequential CEA Group Vithdrawal Event is assumed to occur f
as a result of a failure of the contral element assembly i drive mechanism control system or by operator error. The l CEA Block System eliminates the possibility of an out of [
l sequence bank withdrawal or single CEA withdrawal due to a ,
6 single failure, i i
Any controlled or unplanned withdrawals of the CEAs results j in a positive reactivity addition which causes the core ;
power, core average heat flux and reactor coolant system j temperature and pressure to rise and in turn decrease the :
DNB and Linear Heat Rate (LHR) margins. The pressure in- !
crease, if large enough, activates the pressurizer sprays [
which mitigate the pressure rise. In the presence of a positive Moderator Temperature Coefficient (MTC) of reactiv. ,
ity, the temperature increase results in an additional pos.
itive reactivity addition further decreasing the margin to the DNB and DiR limits.
Withdrawal of the CEAs causes the axial power distribution to shift to the top of the core. The associated increase in the axial peak is partially compensated by the corresponding decrease in the integrated radial peaking factor. The mag- j nitude of the 3 D peak change depends primarily on the ini- f tial CEA configuration and axial power distribution.
The withdrawal of the CEAs causes the neutron flux as mea-
] sured by the excore detectors to be decalibrated due to CEA l
motion, i.e., rod shadowing effects. This decalibration of I t
- O r i OPPD NA-8303 NP. Rev. 02
) Page 16 of 129 i i l l
5.0 TRAFSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.1 CEA Withdrawal Event (Continued) 5.1.1 Definition of the Event (Continued) excore detectors, however, is partially compensated by neutron attenuation rising from moderator density changes l (i.e., temperature shadowing effects). i As the core power and heat flux increase, a reactor trip on t
high power, variable high power, or Thermal Margin / Low Pres-sure may occur to terminate the event depending on the ini- i i
tial operating conditions and rate of reactivity addition.
Other potential trips include the axial power distribution and high pressurizer pressure trips. If a trip occurs, the
- CEAs drop into the core and insert negative reactivity which ;
l quickly terminates further margin degradation. If no trip occurs and corrective action is not taken by the operators, the CEAs fully withdraw and the NSSS achieves a new steady t state equilibrium with higher power, temperature, peak i
linear heat rate and lower hot channel DNBR value. l i
d 5.1.2 Ansivsis Criteria j
The CEA Withdrawal (CEAW) event is classified as an Antici-pated Operational Occurrence (AOG) for which the following i
criteria muct be met: !
t A. The transient minimum DNBR is greater than the 95/95 '
1 confidence interval limit for the CE 1 correlation, i
and 1
B. the Peak Linear Heat Generation Rate (PLHCR) does not exceed 22 kw/ft (Reference 5 1).
l I
OPPD-NA 8303 NP Rev. 02 l Pare 17 of 129 l 1
i 5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued)
'- 5.1 CEA Withdrawal Event (Continued) 5.1.3 Obiectives of the Analysis The objectives of the analysis performed for the "at power" 4
CEAV event is to calculate the Required Overpower Margin i (ROPM) which must be factored into the setpoint analysis. l 1
The objective of the analysis for the hot zero power CEAV event is to demonstrate that the Variable High Power Trip (VHPT) is initiated in time to insure that the analysis 4 criteria are met, t 1
5.1.4 Kev parameters and Analysis Assuentions l The initial conditions assumed in the CEAV analysis are shown in Table 5.1 1. The reactor state parameters of !
(}
primary importance in calculating the margin degradation l
f are:
A. CEA withdrawal rate * (i.e., reactivity insertion rate),
l l
]
B. Cap thermal conductivity (Hg ,p),
i I
4 1
C. Initial power level, !
l D. Flux power level determined from the excore detector response during the transient, 1
) E. The moderator temperature coefficient reactivity, and
interchangeably in this report, i
4 O l OPPD NA-8303 NP Rev. 02 Page 18 of 129 j
l l
l 5.0 TRANSIENT AND ACCIDENT ANALYSIS ME7110DS (Continued) i 5.1 CEA Vithdrawal Event (Continued) 5.1.4 MeV Parameters and Analysis Assumptions (Continued) l F. Changes in the axial power distribution and planar and integrated radial peaking factor during the transient, ,
} The excore responses for each initial power level analyzed are based on the CEA insertions allowed by the Power Depend- l ent Insertion Limit (PDIL) at the selected power level, the f changes in CEA position prior to trip, and the corresponding !
L 1
rod shadowing and tempe'ature attenuation (shadowing) fac- i tors.
For the CEAW cases where combinations of parameters result in a reactor trip, the scram reactivity versus insertion ,
s characteristics are assumed to be those associated with the core average axial power distribution peaked at the bottom i
) of the core. The scram reactivity versus insertion charac-teristics associated with this bottom peak shape minimize ,
) the amount of negative reactivity inserted during initial j portions of the scram following a reactor trip.
1 l
All control systems except the pressurizer pressure control system and the pressurizer level control system are assumed j to be in a manual mode. These are the most adverse operat-l ing modes for this event. The pressurizer pressure control '
system and pressurizer level control system are assuned to j be in the automatic mode since the actuation of these sys-1 I ,
d i
i I
OPPD NA 8303 NP, Rev. 02 l Page 19 of 129 [
l t
Table 5.1 1 I
nitial Conditions Assumed in CEAW Event Analysis l'a rme t e r Units Value Initial Core Power MWt 1(HZP)/1530(HTP)t i
Initial Core Inlet Coolant 'T 532(HZP)t Temperature Maximum allowed by Tech. Specs.
Moderator Temperature Coefficient x10* hap /'r Tech Spec. Range Initial RCS Pressure psia Minimum allowed byt Tech. Specs.
Fuel Temperature coefficient x10ap/'F Least Negative Predicted
, During A Cycle Initial Core Mass velocity x106 lbm/hr Minimum allowed byt Tech. Spees.
Fuel Terp. Coeff. Uncertainty n 15.0 j CEA Withdrawal Speed in/ min 46.0
- Radial Peaks Maximum Allowed by
.( Tech. Specs, for a Given Initial 1 Power Level Scram Reactivity 4 Minimum Predicted During a Cycle I
- High Power Trip Analysis Setpoint n of 1500 MVt 112.0 Variable High Power Trip Analysis t Above Initial 10.0 Setpoint Power Level 1 1 1 l 4
1 i For DNBR calculations, effects of uncertainties are combined statistically.
i l
I i
O
! OPPD NA 8303 NP, Rev. 02 Page 20 of 129 i
5.0 TRMISIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.1 CEA k'ithdrawal Event (Continued) l -
5.1.4 MeV Parveters and Annivsis Assumotions (Continued) tems minimizes a rise ir. the coolant system pressure. The net effect, is to delay a reactor trip until a high power trip is initiated. This allows the transient increases in power, heat flux and coolant temperature to proceed for a longer period of time. In addition, minimizing the pressure increase is conservative in the margin degradation calcula. I tions since increases in pressare would offset some of the DNB margin degradation caused by increases in the core heat flux and coolant temperatures, i
5.1.5 Analysis Methodolorv 1
The methodology used for analysis of the CEAW event is de-scribed in CEN 121(B) P. Reference 5 2. The District does lt not perform all parametric analyses discussed in Reference [
5-2 for Fort Calhoun Station. Rr.ner, the District utilizes ' !
the analyses performed in Refe,ence 5 2 to limit the number of analyses necessary for Fort Calhoun Station.
i The rod shadowing factors for the Fort Calhoun Station full !
I power case with Bank 4 inserted are the inverse of the rod !
shadowing factors used in Reference 5 2 (The rod shadowing l factors for Fort Calhoun Station are such that the excore detectors see more flux when the rods are withdrawn than ,
when they are inserted.) The analysis at intermediate power 1 I
levels is the same as doewtented in Reference 5 2. l O ,
OPPD NA 8303 NP Rev. 02 I Page 21 of 129 I
5.0 TRANSIENT AND ACCIDENT ANALYSIS METil0DS (Continued)
O
'd 5.1 CFA Vithdrawal Event (Continued) l 5.1.5 Analysis Methodology (Continued)
The hot zero power CEAV event is analyzed assuming the vari-able high power trip is initiated at 29.14 (19.14 plus 10%
uncertainty) of rated thermal power. In addition, the anal.
ysis assumes that the uaximum CEA withdrawal rate is com-bined with the maximum differential rod worth. This case is i analyzed using CESEC and the minimum DNBR is calculated us-ing CETOP using the assumptions discussed in Reference 5 2. l The CEAV event analyzed to determine the closest approach to l the fuel centerline melt SAFDL assumes those values of the CEAV rate and H g ,p discussed in Reference 5 2. This com- l' bination of CEAV rate and Hg ,p was used to determine the 1
PillCR at all power levels, l
5.1.6 Tveical Analysis Results and 10 CFR 50.59 Crit.crip The results of the analyses of the CEAV event for Fort Cal.
houn Station at full power and at intermediate power levels are expected to be similar to those presented in Peference 5 2. TheresultsofthehotzeropowerCEAwithdrawalanal-l ysis are expected to be similar to those discussed in the Cycle 8 reload submittal and the 1983 update of the USAR.
The 10 CFR 50.59 criteria are met if the analysis for the full power and intermediate power level CEAV events shows that the required overpower marSin for these events is less than the available overpower margin required by the current Technical Specification DSB and P1JICR LCOs. The 10 CFR ;
50.59 criteria is satisfied for the hot zero power CEA= l event if the minimum DNBR is greater than that reported in O the latest submitted analysis.
(O OPPD NA 8303 NP, Rev. 02 Page 22 of 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued)
V) 5.1 CEA k'i thdrawal Event (Continued) 5.1.7 Conservatism e, Results Conservatism of the results of the CEAV incident analyses is discussed in Reference 5 2 for the full power, intermediate 4
power level and hot zero power cases.
5.2 Bnron Dilution Event 5.2.1 Definition of Event I
Boron dilution is a manual operation, conducted under strict procedural controls which specify paraissible limits on the rate and magnitude of any required change in boron concentration. Borota concentration in the reactor coolant f
s systen can be d treased by either controlled addition of l l unborated makeup water vlth a corresponding removal of reactor coolant or by using the deborating ion exchangers.
To effect boron dilution the makeup controller mode selector of the chemical and volume control system (CVCS) must be set to "dilute" and then the demineralized water batch quantity selector set for the desired quantity, khen the specific ;
amount has been injected, the domineralized water control I valve is shut automatically. An inadvertent boron dilution r can occur only if there is a combination of operator error l and a CVCS malfunction occurring at the same time. No RPS trips are assumed to terminate this incident. 3 J
5.2.2 Analysis Criteria J
1 The boron dilution event is classified as an A00 for which the following criteria cannot be exceeded:
O OPPD NA 8303 NP. Rev. 02 Page 23 of 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS MET 110DS (Continued) 5.2 Poron Dilutien Event (Contir.ue d) 5.2.2 Analysis Criteria A. DNBR greater than the 95/95 confidence interval limit using the CE 1 correlation, and B. The PUICR less than 22 kv/ft.
1 5.2.3 Obiectives of the Ansivsis The DNBR and PUICR criteria are met by showing that suffi-j eient time exists for the operator to take corrective action to terminate the event prior to exceeding the SAFDLs. This is accomplished by calculating the time interval in which the minimum Technical Specification shutdown margin is lost.
, The acceptable time interval for the operator to take correc. '
tive actions before shutdown margin is lost are 15 minutes foe Modes 2, 3 and 4 and 30 minutes in Mode 5.
l 5.2.4 Kev Parameters and Analysis 1,ssumetions l The boron dilution event at power (Mode 1) is bounded by the l faster rectivity insertion rate of the CEA withdrawal event and it lacks the local power peaking associated with the j I i withdrawn CEA. For the boron dilution event in Modes 2 {
1 through 5, it is assumed that all three charging pumps are j operating at their maximum capacity for a total charging l rate of 120 gpm. For the dilution at hot standby (Mode 2) ;
)' the event is assumed to be initiated at the Technical Spec-ification hot shutdown margin requirement at 532'F. The ;
y reactor coolant system is 5,506 cubic feet.
OPPD NA 8303 NP, Rev. 02 l Page 24 of 129
I 5.0 TPv\MSIENT AND ACCIDENT ANA1.YSIS HETHODS (Continued) !
5.2 Boron Dilution Event (Continued) 5.2.4 MeV Parameters and Analysis AssuPotions (Continued)
The boron dilution event at hot shutdown (Mode 3) is assumed to be initiated from the Technical Specification shutdown margin requirement at 210'F. The boron dilution event at cold shutdown (Mode 4) is initiated from the Technical Spec- l
] ification minimum shutdown margin requirement at 68'F. The 3 analysis is conducted for two RCS volumes, one of 5,506 cubic feet and the other of 2,036 cubic feet, which corre-sponds to the volume for a refueling operation condition.
The analysis for the lower volume cold shutdown condition assumes that shutdown groups A and B are withdrawn from the ;
core and all regulating groups are inserted in the core with the exception of the most reactive rod which is assumed to '
be stuck in its fully withdrawn position. These assumptions O~ are consistent with the Technical Specifications for cold shutdown conditions. The boron dilution event during refuel- l ing is analyzed assuming that reactor refueling has just been completed and the head is in place but the coolant vol.
- use is sufficient t.o only fill the reactor vessel to the bot-i tom of the piping nozzles (2,036 cubic fect) and the minimum '
permissible boron concentration allowed by Technical Specifi- ;
I cation for refueling exists. All CEAs are withdrawn from
- the core.
i These assumptions represent shutdown conditions for the var- '
j ious modes wherein the core reactivity is greatest, the ,
I vater volume and total boron content is at a minimun, and l l the rate of dilution is as large as possible. Hence, these i l
conditions, represent the minimum time to achieve inadvertent 1
criticality in the event of an uncontrolled boron dilution, i
! OPPD NA 8303 NP Rev. 02 i Page 25 of 129
l 5.0 TRR;SIENT AND ACCIDENT ANALYSIS METHODS (Continued) !
p u
\' 5.2 Bnron Dilutten Event (Continued) l 5.2.5 Analysis Methods '
The method used to calculate the dilution time to critical-icy from Modes 2 through 5 is through the use of the follow-
, ing equation:
i ACterit " 'BD . In C3 + SDM*IBW l
C 3
Where 'BD - boron dilution time constant, which is a function of RCS volume and temperature (sec)
C 3 - critical boron concentration (ppm)
SDM - shutdown margin (top)
O IBW - inverse boron worth (ppm /ap) !
) l l As can be seen from this equation, the dilution time to ;
criticality is minimized with a greater critical boron concentration, a smaller inverse boron worth, or a smaller
'ED' 5.2.6 Analysis Results and 10 CFR 50.50 Criteria l
i i
The analysis results are similar to those reported in the !
Cycle 11 safety analysis report and in the 1987 update of l the USAR. The criteria of 10 CFR 50.59 are satisfied if the !
Technical Specification requirements on shutdown margin and f the refueling boron concentration are unchanged as a result of this analysis. i.
l 3
o ;
i OPPD NA 8303 NP, Rev. 02 !
Page 26 of 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS METl!0DS (Continued)
- ) O 5.2 Boron Dilution Event (Continued) !
5.2.7 Conservatism of Re sults l
! Because of the procedures involved in the boron dilution and the numerous alarm indications available to the operator, the probability of a sustained or erroneous boron dilution is very lov. There is usually a large interval between the calculated time and the time limit for the boron dilution at hot standby and hot shutdown modes. Therefore, the results show considerable margin to che limit. The calculated time
. 1 l to critical for the boron dilution at cold shutdown with the j minimwn RCS volume is reasonably close to the acceptance i criteria; however, the event is analyzed with only shutdown i
groups A and B being fully withdrawn from the core.
Cold
, shutdown is normally achieved with the shutdown groups A and j B fully inserted in the core and, therefore, the core has a much lower k,gg than assumed in the analysis, The boron I dilution at refueling is conservative since it is improbable J i i that more than a few CEAs vill be removed at any one time ;
j during a refueling and the approach to critical following j refueling is done under strict administrative control with !
) only onu bank of CEAs removed at a tine. The analysis (
} assumes that all CEAs are withdrawn from the core. l l
l I
l O l OFFD.NA 8303 NP, Rev. 02 Page 27 of 129
. - - _ _ _ . _ - - . -- ..- _. .~. _ .--
t i
) 5.0 TRANSIENT HID ACCIDENT A!!ALYSIS METHODS (Continued)
O 5.3 contiel Element Assemb1v Dron Event 1
-) 5.3.1 Definition of Event ;
The control element assembly (CEA) drop event is defined as the inadvertent release of a CEA causing it to drop into the li i
reactor core. The CEA drive is of the rack and pinion type f I
j with the drive shaft running parallel to and driving the j j rack through a pinion gear and a set of bevel gears. The !
drive mechanism is equipped with a mechanical brake which i
4 due to an inadvertent interruption of power to the CEA drive f i magnetic clutch or an electrical or mechanical failure of i I !
i a
the mechanical brake in the CEA drive mechanism when the CEA >
i is being moved. ;
i ;
t l The full length CEA drop event is classified as an AOO which j does not require an RPS trip to provide protection against !
< i
]
exceeding the SAFDLs, The CEA drop results in a redistri.
] bution of the core radial power distribution and an increase in the radial peaks which are not directly monitored by the l RPS and which are not among those analyzed in determining I
the DNB and LRR LCOs and LSSSs. As such, initial steady f state margin must be built into the Technical Specification i
LCOs to allow the reactor to "ride out" the event without ;
exceeding the DNBR and LRR SAFDLs. !
a I
i ,
I I I t
l i
O l OPPD NA 8303 NP, Rev 02 Page 28 of 129
'____-_ _ __ _ _____-____-_-_--_ -__ l
i 5.0 TRANSIENT AND ACCIDENT ANALYSIS MrTHODS (Continued) 5.3 control Elerent Assembly Dron Event (Continued) l 4 4 1
5.3.2 Analysis Criteria i i
] The full length CEA drop event is classified as an Antici. l 4
l pated Operational occurrence for which the following cri.
i teria must be met:
l A. The transient minimum DNBR must be greater than or l
J equal to the 95/95 confidence interval limit, using ;
] the CE.1 correlation, and 4
I i
(
B. ')
The Peak Linear Heat Rate (PLHR) must be less than or e
i equal to 22 kv/ft. ;
1 '
\
1 5.3.3 Obicetives of the Analysis
- 4 i
- L The objective of the analysis is to determine the Required i Overpower Margin (ROPM) which must be built into the LCOs to l i
assure e.he DNBR and LHR SAFDLa are not exceeded for the CEA drop which produces the highest distortion in the hot chan.
nel power distribution. Since the ROPM is dependent upon I
) initial power level, rod configuration and axial shape s
index, an analysis parametric in these variables is per. (
,' formed, i
)
i i
5.3.4 Key Paraweters and Ansivsis Assumotions I .
t
! Table 5.3.4 1 contains a list of the key parameters assumed 1
i in the full length CEA drop analysis. Assumptions used in j
the analysis include:
4 1
} i j
OPFD.NA 8303.NP. Rev. 02 Page 29 of 129 l
i
5.0 TRANSIENT AND ACCIDENT ANALYSIS HETHODS (Concinued)
[^')s
\'-
5.3 Control Element Assembly Dron Event (Continued) 1
] 5.3.4 MeV Parameters and Analysis Assuretions (Continued) f A. The rod block system is assumed to prevent any other rod motion during the transient.
I B. The turbine admission valves are maintained at a con. I
- stant position during the transient. This is because
) the turbine admissicn valve position is set manually ,
at Fort Calhoun Station and, therefore, the turbine 1 {
]
admission valves will not automatically open in re-j sponse to a reduced electrical generation output. !
i 1
5.3.5 Analysis Method i
The analysis methods utilized by the District to analyze the i j
( CEA drop event are discussed in Section 8 of Reference 5 3. .
I l 4 i j 5.3.6 Annivsis P.esults and 10 CFR 50.5o Criteria t 1
j Typical analysis results are contained in Section 8 of i j Reference 5 3 and in the 1987 update of the Fort Calhoun 1
ll Station Unit No. 1 USAR. The criteria of 10 CFR 50.59 are !
met if the required overpower margin calculated for this incident is less than the overpower margin being main:ained j 1
by the current Technical Specifications, t l
i '
l 1
)
I 1
1 OPPD NA 8303 NP, Rev. 02
-1 Page 30 of 129
),
I Table 5.3.4 1
- KEY PARAMETERS ASSUMED Ill T11E ITLL LENCT11 CEA DROP ANALYSIS
- Parareter Units Value Initia! Core Power MWe 1500t I
l Initial Core Inlet 'r Maximum allowedt Temperature by Tech. Specs.
Initial RCS Pressure psia Minimum allowedt by Tech. Specs.
Initial Core Mass Flow Rate x1061bm/hr Miniaws allowedt l by Tech. Specs.
Moderator Temperature x10ap/'r Most negative l Coefficient allowed by Tech.
Spees.
CEA Insertion n Insertion Maximum allowed by Tech. Specs.
Radial Peaking Distortion Maximwn value predicted Factor during core life
!O t For DNBR calculations, the effects of uncertainties on these parameters are combined statistically.
i f
l OFFD NA 8303 NP, Rev. 02 '
Page.31 of 129 !
l
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued)
V 5.3 Control Element Assembiv Dron Event (Continued) 5.3.7 Conservstism of Results i
j The following area of conservatism is included in the anal.
ysis:
A. The moderator temperstare coefficient as used in the l analysis is the most negative value alleved by the Technical Speelfications. The actual end of life value, including roast.rement uncertainty, is less negative.
j l
- 5.4 Four Punn loss of Flov Event '
i i
j 3.4.1 Definition of the Event ,
The four pump loss of coolant flov event !.s initiated by the ,
i simultaneous loss of electrical power to all four reactor coolant pumps. The loss of AC power to reactor coolant ,
pumps may result from either the complete loss of AC power ;
1 to the plant, or the failure of the fast transfer breakers 1 -
1 to close after a loss of offsite power. !
} i
( :
Reactor trip for the loss of coolant flow is initiated by a [
low coolant flow rate as determined by a reduction in the i j sum of the steam generator hot to cold lag pressure drop.
This signal is compared to a setpoint which is a function of j the number of reactor coclant pumps in operation (which ;
current Technical Specifications require to be four). A reactor trip would be initiated when the flow rate drops to ,
) 934 of full flow (956 minus 2n uncertainty).
I iO ,
OFFD NA 8303 NP. Rev. 02
- Page 32 of 129 l
i 5.0 TRMISIENT AND ACCIDENT ANALYSIS METHODS (Continuedi 5.4 Four-Pumn Loss of Flow Event (Continued)
~ 4.2 Analysis Criteri; 3 The four pump loss of flow event is classified as cn ADO for which the transient minimum DNBR must be greater than the 95/95 percent confidence interval limit.using the CE 1 cor-relat. ion.
5.4.3 obiectives of the Annivsis ,
The objective of the analysis is to determine the required overpower margin that must be built into the DNB LCOs such that in conjunction with the low flow trip the DNBR SAFDL is not exceeded. Since the required overpower margin is de-pendent upon both axi.al shape index and the CEA rod config- l uration, an analysis parametric in these parameters is per-formed, l
, 5.4.4 Kev Parameters and Analysis A qumotions !
The closest approach to the DNBR SAFDL occurs for a loss of flow event initiated from the full power conditions. Table 5.4.4 1 gives the key parameters used in this analysis. The flow coast down is calculated in the CESEC code.
5.4.5 Analysis Meth m The analysis method used by the District to analyze.the four-pump loss of coolant flow is discussed in Section 7 of Refer-ence 5 3. The District utilizes the CESEC TORC method to analyze axial power distributions characterized by both nega-tive and positive shape indices. The STRIKIN TORC method is not utilized by the District because of the high rotational energy of the pumps (N - 1185 rpm. I - 71,000 lb ft / pump).
OPPD NA 8303 NP Rev. 02 Page 33 of 129
i .
rL t
Table 5.4.4 1 KEY PARAMETERS ASSUMED IN THE LOSS OF C001 ANT / LOW ANALYSIS
.g-Parameter h Value ,
Initial Core Power MWt 1500t a
Moderator Teraperature x10 op/'F Maximum allowed l Coefficient by Tech. Specs. ,.
, Fuel Temperature x10'0 Ap/'F Lease negative l.i 4
Coefficient predicted during .
] core-life. !
Low Flow Trip Delay Time see Maximum l CEA Drop Time sec' Maximum allowed-
' l-'
by Tech. Specs.
q Scram Rese.tivity Worth op Minimum predicted ,
during core lifetime Scram Reactivity Consistent with axial i
- Curve shape of interest O
v S
i l
I i t For DNBR calculations, effects of uncertainties on these parameters were i combined statistically.
i 1
t l
i l
i
, l
]
i l
l 1
I OPPD NA 8303 NP, Rev. 02 i PATe 34 of 129 4 !
t 5.0 TRAF.SIENT AND ACCIDENT ANALYSIS METHODS (Continued) l d- 5.4 Four-Pumo Loss of Flow Event (Continued) 6 5.4.6 Analysis Results and 10 CFR 50.59 Criteria Expected enalysis results are presented in Section 7.1 of ;
Reference'5 3. The main difference between these results l l and the results for Fort Calhoun Station is that the ROPM will be significantly reduced for Fort Calhoun Station. ;
This f s because of the higher rotational energy of the Fort Calhoun' reactor coolant pumps.
The criteria of 10 CFR 50.59 are met.if the required over- ]
power margin calculated for the four pump loss of coolant
- flow event is less than the overpower margin'being main-I tained by the current Technical Specifications. I J
l Q 5.4.7 Conservatism of Results i d
Q i The conservative nature of the DNBR R0PM valves calculat-ed for the four pump loss of flow event is demonstrated
- by the following conservative assumptions
- !
A. Field measurements of the CEA magnetic clutch decay l i is more rapid than assumed in the safety analysis.
B. The available scram worth is higher than assumed in l the safety analysis.
C. The MTC at full power is more negative than the l t
value assumed in the -afety analysis.
lO j OPPD NA.8303 NP, Rev. 02 Page 35 of 129
]
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) ,
t rs 5.4 Four Pumn Loss of Flow Event (Continued) 5.4.7 Conservatism of Results (Continued)
D. The actual CEA drop time to 906 inserted is faster l than that assumed in the safety analysis.
E. The conservatism of the CETOP calculations is dis- l cussed in Section 7 of Reference'5 3.
5.5 Asymmetric Steam Generator Event 5.5.1 Definition of the Event The asymmetric transients arising from a secondary system malfunction in one steam generator result in changes in '
core power distribution which are not inherently covered by the TM/LP or APD LSSS. Consequently, these events must be analyzed to determine the initial steady state thermal margin which is built into and maintained by the Technical Specification LCO such that assurance is pro- i vided that the DNBR and peak linear heat rate SAFDLs are ;
not exceeded for these transients. The four events which !
effect the steam generator are:
A. Loss of load to one steam generator. !'
i B. Loss of feedwater to one steam generator. l C. Excess feedwater to one steam generator. l' i
D. Excess load to one steam generator l' O :
OPPD NA 8303 NP, Rev. 02 Page 36 of 129 l I
1
4 e
5.0 TRANSIENT AND ACCIDENT ANALYSIS.METil0DS (Continued) w' 5.5 Asymmetric Steam Generator Event (Continued) 5.5.1 Definition of the Ev;ng (Continued)
The possible RPS trips which can occur to mitigate the ;
consequences of these e.>/ats include the low steam gen-erator level, TM/LJ, low steam generator pressure, and the asymmetric steam gene ator transient protection trip i function (ASCTPTF). The particular trip which intervenes ')
! is dependent u'Jon the event initiator and the initial-
\ ,
] operatin8 conditions. !
The ASGTPTF trip was installed in the Fort Calhoun Sta- ),
- tion RPS prior to operation of Cycle 9 to reduce the mar.
gin requirements associated with these asymmetric events and to insure that these events do not become a limiting A00 for establishing initial margin which must be main-1 tained by the LCO. A system description of the ASGTPTF is presented in Appendix 8 of Reference'5 3.
5.5.2 Analysis Criteria I
+
The asymmetric steam generator events are classified as i A00s for which the following criteria must be met: !
A. The transient minimum DNBR must be greater than or ;
equal to the 95/95 confidence interval limit using the CE 1 correlacion, and l l B. The peak linear heat must be less than or equal to 2
1 22 kw/ft.
i
!o
, OPPD.NA 8303.NP, Rev. 02 j Page 37 of 129
. . = _ .
5.0 TRANSIENT AND ACCIDENT ANALYSIS MET 110DS (Continued)
\
5.5 Asymmetric Steam Generator Event (Continued) i 5.5.3 Obiectives of the Analysis The objectives of the analysis are to determine the re-quired overpower margin that must be built into the LCOs sut.h that in conjunction with the ASGTPTF the DNBR and l PIEGR SAFDLs is not exceeded.
i 5.5.4 Kev Parameters and Analysis Assumn 4gna Section 7 of Reference 5 3 demonstrates that the loss of load to one steam generator (LL/1SG) is the limiting asym.
.(i metric steam generator transient for establishing initial steady state thermal margin which must be maintained by the Technical Specification LCO. Therefore, information is only provided for this asymmetric steam generator event. The key parameters used in the analysis of the LL/1SG event are given in Table 5.5.4 1. The charging ;
pumps and proportional heater systems are astumed to be ;
inoperable during the transient. This maximizes the pres-l sure drop during the event. The turbine admission valves !
are assumed to maintain a constant position throughout the event since f.he turbine control system at Fort Cal- j houn utilizes manual setting of the turbine admission {
valves.
I 5.5.5 Analysis Method l
i' The method utilized by the District to analyze the LL/1SG is discussed in Sec. tion 7 of Reference 5 3.
t OPPD NA 8303 NP Rev. 02 i Page 38 of 129 !
f J
n
Table 5.5.4 1 V KEY PARAMETERS ASSUMED IN Tile LL/1SC EVENT Paramoter Units Value Initial Core Power MW e 1530t Initial Core Inlet 'F Maximum allowedt Temperature by Tech. Specs.
Initial Reactor Coolant psia Minimum allowedt System Pressure by Tech. Specs.
Core Average H gap BTU /hr-ft2,.* Maximum value predict-ed during core life.
6 Initial Core Mass x10 1bm/hr Best estimate flowt Flow Rate Scram Reactivity Vorth top Minimum predicted during core life, t For DNBR calculations, effects of uncertainties on these parameters were combined statistically, J
i
)
A I
I l
(0 1
v
/
l i
j OPPD NA 8303 NP, Rev. 02 l Page 39 of 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS METi!0DS (Continued)-
(~M N s/ 5.5 Asymmetric Steam Generator Event (Continued) 5.5.6 Analysis Results and 10 CFR 50.59 Criteria The _sults of the analysis for the LL/1SG event are dis-cussed in Section 7 of Reference 5-3.
t The results for Fort Calhoun Station are expected to be similar. The criteria of 10 CFR 50.59 are satisfied.if the required overpower margin calculated for the'LL/1SG ,
event is less than the overpower margin being maintained 'l by the current Technical Specifications.
i 5.6 Excess Load Event '
i l
4 5.6.1 Definition of Event i i
An excess load transient is defined as any rapid increase in the steam generator steam flow other than a steam line break. Such a rapid increase in steam flow results in a power mismatch between the reactor core and the steam gen-erator load demand. In addition, there is a decrease in i the reactor coolant temperature and pressure. Under these
- conditions the negative moderator temperature coefficient j reactivity causes an increase in core power.
4 !
] The rapid opening of the turbine admission valves or the 1
1 steam dump bypass to the condenser causes an excess load event. Turbine valves are not sized to accommodate steam
. flow for powers much in excess of 1500 Mut. The steam dump valves and steam bypass valves to the condenser are i
i O
1 l
3 OPPD NA 8303 NP Rev. 02 Page 40 of 129
.. ~ . . . .- - . ,
1 5.0 TRANSIENTANDACCIbENTANALYSISHETHODS(Continued) l O 5.6 Excess Load Event (Continued) l 5.6.1- Definition of Event (Continued) sized to accommodate 334 and St. respectively, of the .l steam flow at 1500 MW. Therefore, the following load !
increase events are examined:
3 A. Rapid opening of the turbine control valves at power: The maximum increase in the steam flow due 1
to the turbine control valves opening is limited by the turbine load limit control. The load limit con. :
trol function is used to maintain load, so unlesc ;
- valve failure occurs, the control valves will remain [
where positioned.
! B. Opening of all dump and bypass valves at power due to steam dump control interlock failure: The cir-cuit between the steam dump controller and the dump ,
valves is open when the turbine generator is on line. Accidental closing of the steam dump control interlock under full load conditions, according to
, g the temperature program of the controller, causes full opening of the dump and bypass valves. Since the reactor coolant temperature decreases during the f event, these valves will be closed again after the 1
average reactor coolant temperature decreases to l
$ 3 5 ' F.
J C. Opening of the dump and bypass valves at hot standby
- I conditions due to low reference temperature setting in the steam dump controller
- When the plant is in hot standby conditions the dump valve controller is OPPD NA 8303 NP, Rev 02 i Page 41 of 129 ,
a
5.0 TRANSIEN1' AND ACCIDENT ANAINSIS METHODS (Continued)
C% 5,6 Excess Load 1 m (Continued) l 5.6.1 Definition of Evert (Continued)
C. (Continued) s operative but does not act because the hot standby temperature is lower than the lowest value required to open the valves. At hot standby the reactor cool - t ant temperature is 532*F, which is 8'F below the min-imum temperature required to open the dump and by-I pass valves (540'F). The maximum error that can be ;
introduced in the referenced temperature setting is limited to 17'F since a narrow range instrument is used for this purpose. Reducing the dump valve con-troller reference setting from 532' to 515' would re-sult in a partial opening of the 141ves but as soon as the reactor coolant temperature dropped to 518'F the valves would again be completely closed.
D. Opening the dump and bypass valves at hot standby due to steam dump controller malfunction: The most severe event at hot standby would occur in the event ,
the steam dump valve controller yields an incorrect '
signal and causes the steam dump and bypass valves l to open completely. This case is considered to be ;
much less probable than case C above but represents the most limiting event under hot standby condi-i tions.
l The possible RPS trips that might be encountered dur-
! ing this event are:
1
- 1. Variable high power trip (V11PT).
OPPD NA 8303 NP, Rev. 02 4 Page 42 of 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.6 Excess Lond' Event (Continued) 5.6.1 Definition of Event (Continued)
D. (Continued)
- 2. TM/LP trip.
9 3, Low steam generator water level trip.
! 4. Low steam generator pressure trip, d The RPS trip initiated to mitigate the consequences of the event will depend upon the initial conditions and the rate of reactivity insertion due to moderator feed-back effects.
5.6.2 Analysis Criteria The excess load event is classified as a AOO for which the following criteria must be met:
A. The transient minimum DNBD. must be greater than or equal to the 95/95 confidence internal limit using the CE-1 correlation.
B. The peak linear heat rate (PIllR) must be less than or equal to 22 kw/ft.
5.6.3 Obiectives of the Analysis The objectives of the analysis are to calculate a term which.
when incorporated in the TM/LP equation will ensure that the SAFDLs are not exceeded for those excess load events which require a TM/LP trip for protection and to ensure that the DNBR and LJIR SAFDLs are not exceeded for excess load events for which the TM/LP does not provide protection.
OPPD NA.8303 NP. Rev. 02 Page 43 of 129
5.0 TPMSIENT AND ACCIDENT ANALYSIS METHODS (Continued)
(N 5.6 Excess Load Event (Continued) 5.6.4. Kev Parnmeters and Analysis Assumotions As discussed in Section 5 of Reference 5-3, sensitivity studies performed by CE have demonstrated that the maximum calculated term for the excess load event occurs at hot full ,
power conditions. District sensitivity studies show similar results. Therefore, only the hot full power case is anal-yzed. The key parameters used in the analysis of the excess load event are given in Table 5.6.4 1. The remaining assump-tions are the same as those discussed in Reference 5-3.
5.6.5 Analysis Method The steps used for determining the value and calculating the largest term for all excess load events which rely on the ;
TM/LP trip for DNBR protection are given in Section 5 of Reference 5 3. The minimum transient DNBR value for excess load events protected by the Variable High Power Trip is I
calculated .ing the procedure discussed in the same
- t Section. l i
The PLHR is calculated by obtaining the core average linear heat rate at time of peak core power and multiplying it by i the appropriate peaking factors and associated uncertain-ties.
l l
I i
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- Page 44 of 129 l i !
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t '
Table 5.6.4-1 '
i KEY PARAMETERS ASSUMED IN THE EXCESS LOAD EVENT ANALYSIS Parameter Units Value f
Initial Core Power MWt 1530t ;
i Initial Core Inlet F At Power Maximum allowedt !
Temperature. by Tech. Sp,ecs.-
Initial Reactor Coolant psia Minimum allowedt System Pressure by Tech. Specs.
Initial Core Mass 6 x10 1bm/hr Minimum allowedt (
Flow Rate by Tech. Specs. ;
t
{ CEA Drop Timo sec Maximum allowed l by Tech. Specs. ;
Scram Reactivity top Minimum predicted ,
Vorth during core life. l Moderator Temperature x10*4ap/*F Negative values up to '
i Coefficient the most negative j value allowed by Tech. ,
1
(::) => -
I t For DNBR calculations, effects of uncertainties on these parameters were combined statistically.
4 l
l l i !
)
i l.
1 l
l l OPPD.NA.8303.NP, Rev. 02 l
! Page 45 of 129
5.0 TPANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 0, 5.6 Excess Load Event (Continued) 5.6.6 Analysis Results and 10 CFR 50.59 Criteria The results of the excess load analysis are similar to those presented in Section 5 of Reference 5-3. The cri-l teria of 10 CFR 50.59 are met if the term is less than or equal to the value used in the' current TM/LP trip equa-i tion.
5.6.7 Conservatism of Results The following points demonstrate the conservatism of the overall results for the excess load event:
A. Field measurements demonstrate that the CEA magne- l' tic clutch decay time is less than that assumed in the analysis.
B. The actual scram worths are higher than those in l the analysis, i
C. L'here the most negative MTC is used, the value is l[
more negative than that measured during plant oper- l ation. ,
i j D. The actual Doppler reactivity is more regative than l' 1'
assumed in the analysis.
I !
.l 1
y 1
\ :
OPPD NA 8303 NP, Rev. 02 Page 46 of 129 '
5.0 TPANSIENT AND ACCIDENT ANALYSIS ME1110DS (Continued)
(~T i
5.6 Excens Lond Event (Continued) 5.6.7 Conservnrism of Results (Continued)
E. Field data demonstrates that the actual CEA drop time is less than that assumed in the analysis.
F. The conservatism of the term is discussed in Section 5 of Reference 5 3.
5.7 RCS Depressurization Event 5.7.1 Definition of Event The RCS depressurization event is characterized by a rapid decrease in the primary system pressure caused by either the inadvertent opening of both power operated relief valves (PORVs) or the inadvertent opening of a single primary safety valve operating at rated thermal '
power. Following the initiation of the event, steam is discharged from the pressurizer steam space to the quench tank where it is condensed and stored. To compensate for the decreasing pressure the water in the pressurizer flashes to steam and the proportional heaters increase the heat added to the water in the pressurizer in an attempt to maintain pressure. During this time the pres-surizer level also begins to decrease causing the letdown l control valves to close and additional charging pumps to start so as to maintain level. As pressure continues to l
drop, the backup heaters enerrize to further assist in maintaining primary pressure. A reactor trip is initiat-ed by the TM/LP trip to prevent exceeding the DNBR SAFDL.
i a
OPPD NA 8303 NP, Rev. 02 Page 47 of 129
4
)
J 5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.7 RCS Depressuri ntion Event (Continued) i 5.7.2 Analysis Criteria L
i The RCS depressurization event is classified as an ADO l I
for which the transient minimum DNBR must be greater than i
- or equal to the 95/95 percent confidence interval limit
) using the CE 1 correlation. ;
I f
! 5.7.3 Qbiectives of the Analysis l I '
1 r i
This event is classified as an A00 for which there must l 1
- be sufficient margin built into the TM/LP trip such that l j the DNBR SAFDL is not exceeded. The objective of this l
analysis is to calculate a conservative term for incorpor- -
] ation into the TM/LP equation. [
]
1 5.7.4 Kev Parameters and Analysis Assumotiong ,
The key parameters for the RCS depressurization event analysis are given in Table 5.7.4 1. Additional assump-
, tions are discussed in Section 5 of Reference 5 3.
l 5.7.5 Analysis Method i t
1 2
! The methods used by the District to analyze the RCS de. i
! pressurization event are contained in Section 5 of Refer- l i ence 5 3.
E i i f
I, l
i j
j i
OPPD NA 8303 NP, Rev 02 j Page 48 of 129 J
i i
I l
Table 5.7.4 1 i) KEY PARAMETERS ASSUMED IN THE RCS DEPRESSURIZATION EVENT ANALYSIS Q
Parnecter Units "alue Initial Core Power Mut 1530t Initial Core Inlet 'F Maximum allowed byt Temperature Tech. Specs.
Initial Reactor Coolant psia Upper limit of norma '
System Pressure operating range Moderator Temperature x10op/'F Host negative allow <
Coefficient by Tech. Specs.
Fuel Temperature x10*'op/'F Most negative predic Coefficient during core life.
Core Average H gap BTU /hr ft2,.F Minimum predicted during core life.
Total Trip Delay see 1.4 Time m t For DNBR calculations, effects of uncertainties on these parameters were combined statistically.
l l
l N.
OPPD NA 8303 NP, Rev. 02 Page 49 of 129
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6 i
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5.0 IRANSIENT AND ACCIDENT ANALYSIS' HET110DS (Continued)
, sg R
- i. s- / 5. 7 RCS Deoressurization Event (Continued) l ,
5.7.6 Analysis Results and 10 CFR 50.59
-l Results of the RCS depressurization transient are discussed '
in Reference 5-3 and in the 1984 update of the Fort Calhoun l Station Unit No. 1 USAR. The criteria of 10 CFR 50.59 are ;
satisfied if the term is less:than or equal _to the value 7 used in the current TM/LP trip equation, j 5.7.7 Conservatism of Results i
s P
The conservatism of the calculated pressure bias term is obtained by using the combination of the following conserva. l tive key parameters:
1
, A. Conservative scram reactivity characteristics are used in the analysis.
B. Conservatively slow RPS response times are used. l l Conservatively hi8 h primary relief or safety valve C.
l j areas are used.
l D. The RCS pressure is initially assumed to be in its l j upper limit as opposed to the normal operating pres- ,
j sure. l
, i
, 5.8 Main Steam Line Break Accident !
i
)' 5.8.1 Definition of the Event i I
i i !
] A large break of a pipe in the nain steam system causes a {
rapid depletion of steam generator inventory and an in- l creased rate of heat extraction from the primary sy tem. '
l 1
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OPPD NA 8303 NP, Rev. 02
, Page 50 of 129 1
1 I
- . < - , - - - - - - n --- ----v, n .. ---m,--,., - - . - - . . , , ..,m w ww w ,-m m m _ , - , - - - ,~r,,,r--v,
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued)
O 5.8 Main Steam Line Break Accident (Continued) 5.8.1 Definition of the Event (Continued)
The resultant cooldown of the reactor coolant,' in the pre-sence of a negative maderator temperature coefficient of reactivity, will. cause an increase in nuclear power and trip ,
I the reactor. A severe decrease in main steam pressure will l also initiate reactor trip and cause the main steam isola-1 tion valves to close. If the steam line rupture occurs be- -
j tween the isolation valve and the steam generator outlet nos-i zie, blowdown of'the affected steam generator will continue, i l (However, closure of the check valve in the ruptured steam h line, as well as closure of the isolation valves in both j ctons lines, will terminate blowdown from the intact steam generator). The 1,wcast blowdown, and therefore, the most rapid reactivity addition, ocev.cn yhan the break is at a steam generator nozzle. This break locacies ty traueed for i che cases analyzed.
Both full power and no load (hot stand'o-f ) initial condition cases were considered for two loop operation (i.e., four reactor coolant pumps). (
! f Since the steam generators are designed to withstand reactor j coolant system operating pressure on the tube side with at- I j mospheric pressure on the shell side, the continued inte-grity of the reactor coolant system barrier is assured.
1 The most probable trip signals resulting from an MSLB in.
clude low steam generator pressure, high power, low steam generator water level, TM/LP, and high rate of change of power (for the no load case).
1 l OPPD NA 8303 NP, Rev. 02
) l Page 51 of 129 l 4
5.0 TPANSIENT AND ACCIDENT ANALYSIS MET 110DS (Continued)
V 5.8 Main Stenm Line Break Accident (Continued) 5.8.2 Analysis Criteria The steam line break accident event is classified as a postulated accident for which the site boundary doses must be within the 10 CFR 100 criteria. Acceptable site boundary doses are demonstrated by showing that the crit-ical heat flux is not exceeded.
5.8.3 Obieetives of the Analysis The objectives of the analysis are to demonstrate that the margins to DNB for the reload core no load two loop and full load two-loop main stJam line break cases are greater than that for the Cycle 1 cases given in the ori-ginal FSAR. This is accomplished by demonstrating that the return to power during the event for the reload core is less than the return to power calculated for Cycle 1.
5.8.4 Kev Parameter and Analysis Assumotions The MSLB accident is assumed to start from steacy state conditions with the initial power being 1530 MVt (1024) for the full power case and 1 MVt for the no load case.
The reactor coolant system cooldown causes the greatest i 8
positive reactivity insertion into the core when the Moderator Temperature Coefficient (MTC) is, the most j negative. For this reason the Technical Specification j negative MTC limit corresponding to the end of cycle is !
l assumed in the analysis. Since the reactivity change i associated with moderator feedback varies significantly I O
OPPD NA 8303 NP, Rev. 02 Page 52 of 129
{
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHODS (Continued) 5.8 Main Steam Line Break Accident (Continued) ;
i 5.8.4 Kev Parameter and Analysis Assumotions (Continued) t over the temperature range covered in the analysis, a curve of reactivity insertion versus temperature rather 1
than a single value of MTC is assumed. This curve is j derived on the basis that upon reactor trip the most reac- !
tive CEA is stuck in the fully withdrawn position thus 1 f i yielding the most adverse combination of scram worth and 9 -
l reactivity insertion. Although no single value of MTC is
- assumed in the analysis, the moderator cooldown reactiv.
4 ity function is calculated assuming an initial MTC equal to the most negative Technical Specification limit.
Reactivity feedback effects from the variation of fuel !
temperature (i.e., Doppler) are included in the anal-
} ;
, ysis. The most negative Doppler defect function, when used in conjunction with the decreasing fuel temperature
{ causes the greatest positive reactivity insertion during !
1 i
the MSLB event. In addition to assuming the rnost nega- (
tive Doppler defect function, an additional 154 uncer-tainty is assumed, i.e., a 1.15 multiplier. This multi-1 plier conservatively increases the suberitical multiplica- l t
l tion and results in a larger return to power. '
I i
The delayed neutron precursor fraction, B, assumed is j the maximum absolute value including uncertainties for end of cycle conditions. This is conservative since it i also maximizes suberitical multiplication and thus, en-j hances the potential for a return to power.
I i
a 1
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) OPPD NA 8303 NP, Rev. 02 i Page 53 of 129 i
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 0 V 5.8 Main Ste.m Line Break Accident (Continued) f 5.8.4 M g3 meter and Analysis Assumotions (Continued)
- The steam generator low pressure trip, which occurs at 478 psia (including a 22 psia uncertainty below the nominal trip i retting of 500 psia), is the trip assumed in the analysis.
No credit is taken for the high power trip which occurs at f approximately the same time for the full power case. For l the cases analyzed, it is assumed that the most reactive CEA j l,
is stuck in the fully withdrawn position. If all CEAs in- f sert (no stuck CEAs), there is no return-to-critical and no
- power transient following trip. i
. The cold edge temperatures are used to calculate moderator 1
1 reactivity insertion during the cooldown, thus maximizing the return to critical and return to. power potentials, t
i The Emergency Operating Procedures incorporate the Trip 2/ Leave 2 RCP operating strategy as indicated in Reference 5 8. For a steam line break, the Trip 2/ Leave 2 strategy i will result in tripping two RCPs (at 1350 psia). If the l I
event was misdiagnosed as a LOCA, all four RCPs would be j tripped. As discussed below, for a main steam line break the consequences of Trip 2/ Leave 2 is bounded by the loss of l
] offsite power and the loss of offsite power case is bounded j I
by tripping no RCPs. Consequently, the limiting main steam l line break accident occurs with all RCPs operating. [
., l f
The MSLB case with the RCPs tripped is similar to the HSLB l
case with a loss of offsite power since the RCPs coastdown in both events. As discussed in Reference 5 4, the loss e offsite power delays safety injection due to the time delay ,
l '
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OPPD NA.8303.NP, Rev. 02 j Page 54 of 129 i
i
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued)
)
5.8 Main Steam Line Break Accident (Continued) 5.8.4 Kev Parameter and Analysis Assumotions (Continued) for the emergency diesel generators to restore power to the safety injection pumps and causes a coastdown of the RCPs.
f, !
The coastdown affects the degree of overcooling and in.
creases the time for safety injection borated water to reach ;
l the core midplane. Because manual tripping of the RCPs f results in a later coastdown of the RCPe and because safety _[
injection is not delayed since offsite power is available (
(i.e., the diesel generator startup and pump loading delays i are not present), the injected' boron will arrive at the core {
l midplane sooner for a MSLB with the RCPs tripped than for a ,
MSLB with a loss of offsite power. Therefore, the r, . 'v- '
5 ity effects of a MSLB with the RCPs tripped are less severe ;
! than for the MSLB with a loss of offsite power.
! Reference 5 4 states that the MSLB case with a loss of off-j site power results in the injected boron being dominant over j the RCS cooldown and concludes that the reactivity effects j l of a MSLB accident would be reduced in severity with a con-
! current loss of offsite power when compared to the same I event with offsite power available and the RCPs operating.
l Because the reactivity effects of a MSLB with tho'RCPs !
j tripped after SIAS are loss severe than a MSLB with a con- l
! current loss of offsite power, it is concluded that the f 4 reactivity effects for the MSLB case with the RCPs tripped i
! utilizing Trip 2/ Leave 2 at 1350 psia are less severe than for a MSLB with offsite power available and RCPs operating. f i
1 l
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OPPD NA.8303.NP, Rev. 02 Page 55 of 129 l
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5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) !
l 5.8 Main Steam Line Break Accident (Continued) !
5.8.4 MeV Parameter and Analysis Assumotions (Continued)
The reactor coolant volumetric flow rate is assumed to be constant during the incident. The LCO flow rate (197,000 ;
gpm) was used in order to obtain the most adverse re-suits.
1 A lower flow rate increases the initial fuel and average primary coolant temperatures and consequently results in a higher steam generator pressure and a greater steam gen-erator mass inventory. These effects cause a longer blow-down, a Sreater blowdown rate and a greater decrease in average primary coolant temperature. After MSIV closure h the lower flow rate decreases the rate of reverse heat j transfer from the intact steam generator, thereby increas- l in5 the heat extracted from the primary steam by the rup- l tured steam generator. The overall effect is that the ;
potential for a return to power is maximized.
Maximum values for the heat transfer coefficient across I the stearn generator are used for the no load initial con.
dition case, while nominal values are used for the full- ;
load initial condition. These ' neat transfer coefficients i i
result in the most severe conditions during the incident ,
because of the shape of the reactivity versus moderator i temperature function and the difference in average moder. I I
ator temperature for the maximum and minimum values of ;
the steam generator heat transfer coefficients. l l
l l
l i
OPPD NA 8303 NP, Rev. 02 Page 56 of 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS MET 110D (Continued)
[ \
U 5.8 Main Stearn_ Lirie Break Accident (Continued) 5.8.4 Kev Parameter and Analysis Assumptions (Continued)
The fast cooldown following a MSLB results in a rapid shrink-ing of the reactor coolant. After the pressurizer is em-ptied, the reactor coolant pressure is assumed to be equal to the saturation pressure corresponding to the highest temperature in the system.
Safety injection actuation occurs at 1578 psia (i.e., 1600 psia minus the 22 psia uncertainty) after the pressurizer empties. Additional time is required for pump acceleration.
I valve opening, and flushing of the unborated part of the safety injection piping along with the requirement that the RCS pressure decrease below the shutoff head of the safety injection pumps (1376 psia for high pressure safety injec-tion (HPSI) pumps and 201 psia for low pressure safety inje::-
tion pumps (LPSI) pumps). The analysis takes credit for one HPSI pump, one LPSI pump, and the safety injection tanks.
The boric acid is assumed to mix hornogeneously with the reac-tor coolant at the points of injection into the cold legs.
Slug flow is assumed for movement of the mixture through the i
piping, plena, ati core. After the boron reachei the core l
midplane, the concentration within the core is assumed to increase as a step function after each loop transit inter- !
val.
The boron concentration of the safety injection water is <
assumed to be at the Technical Specification minimum limit.
The values of the inverse boron worth are conservatively chosen to be large to minimize the negative reactivity in-sertion from safety injection.
OPPD NA 8303 NP, Rev. 02 Page 57 of 129
5.0 TPJJiSIENT AND ACCIDENT ANALYSIS METHOD (Continued)
C 5.8 Main Steam Line Break Accident (Continued) 5.8.4 Kev Parameter and Analysis Assumotions (Continued)
Since the rate of temperature reduction in the reactor coolant system increases with rupture size and with steam ,
pressure at the point of rupture, it is assumed that a circumferential rupture of a 26 inch (inside diameter) steam line occurs at the steam generator main steam line nozzle, with unrestricted blowdown. Critical flow is assumed at the point of rupture, and all of the mass leaving the break is assumed to be in the steam phase.
This assumption results in the maximum heat removal from the reactor coolant per pound of secondary water, since the latent heat of vaporization is included in the net heat removal. A single failure of the reverse flow check valve in the ruptured steam generator is assumed; so that the intact steam generator will have steam flow through the unaffected steam line and back through and out the ruptured line. Based on sensitivity analyses performed by the District, this is the most severe single failure for the steam line break event. The analysis credits a choke which is installed in each steam line immediately above the steam ganerator and assumes the stearn flow frors the intact stears generator is through a 50t area reduc-tion choke installed in a 24 inch stears line. This flow will be terminated upon MSIV closure. ,
The feedwater flow at the start of the MSLB corresponds to the initial steady state operation. For the futi load initial condition, it is automatically reduced in accord-ance with the prograra used in the valve controller. For the no load initial condition, feedwater flow is assumed O
OPPD NA 8303 NP, Rev. 02 Page 58 of 129 I
I
I 5.0 TRANSIENT AND ACCIDENT ANALYSIS METH00 (Continued)
~
\- / 5.8 Main Stenm Line Break Accident (Continued) 5.8.4 MeV Parameter and Analysis Assumptions (Continued) to match energy input by the reactor coolant pumps and the 1 MVt core power. Feedwater isolation upon the re-ceipt of a low steam generator pressure (at 478 psia) is credited for both the full load and no load cases. A valve closure time of 30 seconds was used.
5.8.5 Analysis Method The analysis of the main steam line break accident is per-formed using CESEC which models neutron kinetics with fuel and moderator temperature feedback, the reactor pro-tective system, the reactor coolant system, the steam gen-(N erators and the main steam and feedwater systems.
5.8.6 &nalXti s Results and 10 CFR 50.59 Criteria The results of the analysis for the Fort Calhoun steam i
l line break event are discussed in Section 14.12 of the 1983 update of the Fort Calhoun Station Unit No. 1 USAR.
The criteria of 10 CFR $0.59 are met if the calculated return to power is less than the return to power reported I for the Cycle 1 analysis, using the current Technical Specification limit on shutdown margin and moderator tem- I l
perature coefficient.
l 5.8.7 Conservatism of Results Con:arvatism is added to the analysis by inclusion of un-i certainties in moderator and fuel temperature coeffi-cients of reactivity, by taking no credit for void reac-(G)
OPPD.NA 8303 NP, Rev. 02
) Page 59 of 129 j
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) ,
/
(%I
\- 5.8 Main Steam Line Break Accident (Continued) 5.8.7 Conservatism of Results (Continued) tivity feedback, by taking credit for only 1 HPSI pump, by assuming all RCPs operate instead of manually tripping two pumps and by taking no credit for the stuck CEA worth.
5.9 Sei ed Rotor Accident 1
5.9.1 Definition of Event The seized rotor accident is assumed to be caused by a mechanical failure of a single reactor coolant pump. It is assumed that the rotor shears instantaneously, leaving a low inertia impeller attached to a bent shaft. This latter combination comes to a halt immediately causing a sharp drop in the flow rate. The rapid reduction ... core flow will initiate a reactor trip on low flow within the first i
few seconds of the transient.
5.9.2 Analysis Criteria i
A single reactor coolant pump abaf t seizure is classified
'I as a postulated accident for which the dose rates must be within 10 CFR 100 guidelines.
5.9.3 Obiective of the Analysis i
i
)
The objective of the analysis is to demonstrate that the 1 j
radiological releases are within a small f action of 10 CFR 100 guidelines. This objective is met if it can be shown that less than in of the pins fail during the avent.
i j
OPPD NA 8303.NP Rev. 02 Psge 60 of 129 1
l 5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued)
/
\'- ,)
5.9 Seired Rotor Accident (Continued) 5.9.4 MeV Parameters and Analysis Assumntions The key parameters used in the analysis of the seized rotor event are given in Table 5.9.4 1. The saized rotor is con-servatively assumed to result in a 0.1 second rampdown of the core flow from its initial value to the 3 pump value.
5.9.5 Analysis Method Two methods of analyzing the seir . r 6o. . vent are dis-cussed in this section. Section . i 5.1 discusses a method which does not require t ansient analysis input. Section 5.9.5.2 discusses a met. hod which u.Alices transient anal-ysis input.
5.9.5.1 Analysis '4ethod Vithout Transient Analysis Resconse input This method calculates the number of pin fail-ures assuming that the core flow instantan-eeusly decreases to the 3 pump flow rate. This I
method util'zes the TORC analysis with a 3 purp inlet flow distribution. The initial RCS pres-sure and core inlet temperature are used as input to TORC and the core average heat flux is conservatively assumed to remain at its initial l
l j OPPD NA 8303 NP, Rev. 02 l Page 61 of 129 l
l
Table 5.9.4 1 r%
) KEY PARAMETERS ASSUMED IN THE SEIZED ROTOR ANALYSIS Parareter h Value Initial Core Power MWt 1$30t l Initial Core Inlet 'T Maximum allowedt Temperature by Tech. Specs.
Initial Reactor Coolant psia Minimurn allowedt System Pressure by Tech. Specs.
Initial Core Mass 6 x10 1bm/hr Minimum allowedt l Flow Rate by Tech. Specs.
Moderator Terperature x10*'ap/'F Most positive l Coefficient allowed by by Tech. Specs.
Fuel Temperature x10*4ap/'F Least negative pre- l Coefficient dicted during core life.
Core Average H gap BTU /hr ft2 'T Minimurn predicted l during core life.
V CEA Drop Time see i
Maximurn allowed by l Tech. Specs.
1 Scram Reactivity gap Minimurn predicted Vorth during core life.
t Uncertainties on these parameters are combined statistically, i
i
)
)
s/
l OPPD NA 8303 NP. Rev. 02 Page 62 of 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued)
(
J tb -
5.9 Seired Rotor Accident (Continued) 5.9.5 Analysis Method (Continued) 5.9.5.1 Analysis Method Without Transient Analysis Response Incut (Continued) i value. The maximwn value of FR T is com-bined with a conservatively flat power distri-i bution. The TORC calculation determines the I
number of pins that have failed.
5.9.5.2 Analvsjs Methods Usine Transient Analysis i
f This method utilizes the CESEC code to calcu-late the transient response for the seized j ) rotor event. The CETOP code is then used to i determine the time of minimum DNBR.
The TORC
] code utilizes the 3 pump inlet flow distribu-tion, 3 rump cort. flow rate, and the RCS pres.
1 sure, core inlet temperature and core heat flux calculated at the time of minimum DNBR by I CESEC. !
i 5.9.6. Analysis Aesults and 10 CFR 50.54 Criteria The results of the seized rotor analysis are contained in Section 14.6.2 of the Fort Calhoun Station Unit No. 1 l {
USAR. The criteria of 10 CFR 50.59 are met, if the nurnber f
of pin failures is less than one percent. I I
l I
J 1
.J OPPD NA 8303 NP. Rev. 02 Page 63 of 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) 5.9 Sei: cd Rotor Accident (Continued) 5.9.7 Conservatism of Results Conservatism in the calculated number of fuel pins pre-dicted to experience DNBR is added through the use of the following assumptions:
A. The most positive MTC is assumed in the analysis. l l
The actual MTC is more negative and would limit core power and heat flux rise. .
B. A relatively flat pin census is assumed in the anal. l ysis. A more peaked pin census distribution would lower the number of pins predicted to experience DNB.
O C. For the case without transient analysis, no credit l is taken for the pressure increase during the trans-ient and calculating the minimum transient DNBR.
5.10 CEA Eiection Accident 5.10.1 Definition of Event i 1
A CEA ejection accident is defined as a mechanical failure i of a control rod mechanical pressure housing such that the coolant system pressure would eject the CEA and the drive shaft to a fully withdrawn position. The consequences of this mechanical failure is a rapid reactivity insertion which when combined with an adverse core power distribu-tion potentially leads to locali::ed fuel damage. The CEA O
OPPD NA 8303 NP, Rev. 02 Page 64 of 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS ::ETHOD (Continued)
(GV) 5.10 CEA Election Accident (Continued) 5.10.1 Definition of Event (Continued) t ejection accident is the most rapid reactivity insertion that can be reasonably postulated. The resultant core and thermal power excursion is limited pr!.marily by the Doppler reactivity effect of the increased fuel temperatures and is terminated by reactor trip of the remaining CEAs activated by the high power trip or variable high power trip.
5.10.2 Analysis Criteria i
The CEA ejection event is classified as a postulated acci-dent. The design and limiting criteria are:
1 A. Tuel cladding and enthalpy thresholds (Reference 5 5) are:
Clad Damage Threshold Total Average Enthalpy - 200 cul/ gram l
Centerline Melting Threshold Total Centerline Enthalpy - 250 cal / gram 5
Fully Holten Centerline Threshold Total Centerline Enthalpy - 310 cal / gram ;
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. .. -, . _ ~ - . . _ - - - ~ . - . . . . . - . . . . .-
l I
i 5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued)
- 5.10 CEA Eiection Accident (Continued) 5 5.10.2 Analysis Criteria (Continued) f B. The peak reactor pressure during a portion of the l l 1 l transient will be less than the value that will cause .j j stress to exceed the emergency conditions stress limits [
l as defin1d in Section 3 of the ASME Boiler and Pressure f Vessel Code. This objection is achieved if the peak j i RCS pressure does not exceed 2750 psia. !
t i l i
- C. Fuel melting will be limited to keep the offsite dose. l l l consequences well within the guidelines of 10 CFR 100. I i
) 5.10.3 Obiectives of the Analysis I J
l De objective of the analysis is to demonstrate that fuel
! failures are less than those reported in Section 14.3.4.1 of j I
the Fort Calhoun Station Unit No. 1 USAR of that site boun- !
r q dary doses are within the 10 CFR 100 limits. (
) !
l 5.10.4 Analysis Method i !'
l The District utilizes the CEA Ejection Accident Analysis of !
3 our current fuel vendor, Combustion Engineering. nis anal. l 4
ysis methodology is documented in Reference 5 5 and is per. l
{
i formed by Combustiren Engineering. This methodology utilizes l j physics parameters, calculated by the District in accordance j with the methods outlined in Reference 5 6. i i l l
l i
!O d
- OPPD NA 8333 NP Rev. 02 i Page 66 of 129 i
5,0 TRANSIENT AND ACCIDENT ANALYSIS METliOD (Continued) ,
5.10 CEA Eiection Accident (Continued) 5.10.5 Analvmis Results and 10 CFR 50.5o Criteria The results of the CEA Ejection Analysis are reported in Sec- !
tion 14.13 of the Fort Calhoun Station Unit No. 1 USAR, Cri-teria of 10 CFR 50,59 are satisfied if fuel failures are less !
than those assurned for input to the Radiological Consequences
, portion of the analysis, i j 5.10.6 Conservatism of Results 1 ,
The major area of conservatism is the calculation method used to obtain the ejected CEA worth and the ejected radial pe.2k.
The ejected worth and the ejected radial peak are calculated
! without any credit for Doppler or Xenon feedback. In addi-tion, the hot full power ejected worth and ejected peak are calculated assuming the no load temperature of $32'F. The lower temperature is more adverse since this causes a power r roll to the core periphery which also happens to be the loca- l l tion of the ejected CEA. Also, the ejected worth is calcu-j lated assuming the CEAs are fully inserted for hot full power j case regardless of PDIL, Thus, the ejected worth is conser.
vative, J
5,11 Loss nf Coolant Accident l
l The 'sistrict does not perform the Loss of Coolant Accident Analysis. ;
The large and small break loss of coolant analyses were performed by l Combustion Engineering (CE). The large break topical is mentioned in
] Reference 5 7. The small break analysis shows the closest approach l I to the 10 CFR 50.46 criteria for ECCS analysis. The actual differ- ;
J OPPD NA 8303 NP Rev. 02
) Page 67 of 129 l
5.0 TRANSIENT AND ACCIDENT ANALYSIS HETHOD (Continued) t 5.11 Loss of Coolant Accident (Continued) ence is less than 10'F in PCT. The District verifies that the physics input assumptions and the maximum rod burnup are within the bounds assumed in the CE large break analysis.
5.12 Loss of Load to Both Steam Generators Event 5.12.1 Definition of Event A total loss of load to both steam generators usually results from a turbine trip due to a loss of external electrical load or to abnormal variations in electrical network frequencies.
Other possible causes include the simultaneous closure of all turbine stop valves or main steam isolation valves. All ini-tiating mechanisms result in a corresponding reduction in
, f heat removal from the reactor coolant system due to the loss b of secondary steam flow. Although a Reactor Protective Sys-
} tem trip signal would normally result from c turbine trip, no credit is taken in the analysis of this event for the turbine trip signal.
5.12.2 Analysis criteria 4
The loss of load to both steam generators event is classified as an Anticipated Operational Occurrence (A00) for which the l following criteria must be met:
OPPD NA-8303 NP, Rev. 02 Page 68 of 129 J
l l
\
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued)
/~"x i t
\~2 5.12 Loss of Load to Both Steam Generators Event (Continued) 5.12,2 Analysis Criteria (Continued)
A. The peak RCS pressure dces not exceed 2750 psia l (110% of design pressure).
B. The transient minimurn DNBR is greater than the 95/95 l confidence interval limit for the CE.1 correction limit.
C. The Peak Linear Heat Generation Rate (PwCR) does l not exceed 22 kw/ft. ,
Criteria B. and C. are not of major concern since DNBR in- l creases during the event and the P GCR margin required is much less limiting than other A00s. Therefore, criterion A. is the main concern in analyzing this event. The loss l
of load to both steam generators event is the limiting A00 event with respect to peak RCS pressure.
5.12.3 Orlectives of the Analysis The objective of the analysis is to demonstrate, for modi-fications to the plant which potentially degrade RCS heat removal capability (including steam generator plugging) that the peak RCS pressure stays within 110% of the design pressure in accordance with Section III of the ASME Pres-sure Vessel Code. This objective is achieved if the peak RCS pressure does not exceed 2750 psia.
OPPD NA 8303 NP. Rev. 02 Page 69 of 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued)
O 5.12 Loms of Load to P.oth Steam Gene-ators Event (Continued) 5.12.4 Kev Parameters and Annivsis Assumptions The key parameters used in the loss of load (to both steam generators) event are given in Table 5.12.4 1. Assumptions '
used in the analysis to maximico heat up of the RCS and con-sequently peak RCS pressure include:
i A. The event is initiated by a sudden closure of the tur- l j bine stop valves without a simultaneous reactor trip.
- B. No credit is taken for operation of the PORVs, pressur- l izer sprays, and the turbine steam dump and bypass sys-tem, i.e., the pressurizer pressure control system is assumed to be in MANUAL. '
!o j
C. The rod block system is assumed to prevent rod motion (other than scram) during the transient.
l 4
l
} D. Maximum charging flow and zero letdown flow are as- l sumed. l E. Termination of the event occurs as a result of a high l
7 pressuriser pressure trip, i i
a 5.12.5 Analysis Method '
The analysis methods utili:ed by the District to analy:e the loss of load to both steam generators event consists of simu-1 lating the event using the CESEC computer code, utili:ing the analysis assumptions listed in Section 5.12.4 (above) as in-put, and extracting the peak RCS pressure for comparison with the 2750 psia upper limit. !
OPPD NA 8303 NP, Rev. 02 Page 70 of 129 1
Table 5.12.4 1
,a
( REY PARAMETERS ASSUMED IN THE LOSS OF LOAD TO BOTH STEAM CENERATORS ANALYSIS Paramerg Units Value Initial Core Power MWt 1530 l Initial Core Inlec *F Maximum allowed Temperature by Tech. Specs.
Initial RCS Pressure psia Minimum allowed by Tech. Specs initial Stearn Generator psia Minimum value cor-Pressure responding to core inlet temperature operating range.
Initial Core Mass x106 lbm/hr Minimum allowed Flow Rate l by Tech. Specs.
Moderatec Temperature x10*0ap/'F Most positive Coefficient l allowed by Tech.
Specs.
Fuel Temperature x10'0ap/'F Least negative Coefficient l
predicted during
( core life.
Fuel Temperature Coefficient Multiplier 0.85 1
CEA Drop Tirte see Maximura allowed l by Tech, Specs.
Scram Reactivity Vorth tas Minimum predicted i
during core lifetime
! Scram Reactivity Consistent with most
! Curve l positive axial shape
- (bottom peaked)
I allowed by Tech.
Specs.
Core Average H g3p BTU /hr ft2 'F Maximum predicted l during core lifetirme.
Minetics Parameters EOC parameters {
(minimum absolute '
S).
i i
S RPS Response Time sec 1.4 l OPPD NA 8303 NP. Rev. 02 Page 71 of 129
5.0 TRANSIENT AND ACCIDENT ANALYSIS HETHOD (Continued)
(D
!' 'l 5.12 1.o s s nf Load to Both Stenm Generators Event (Continued) 5.12.6 Analvmis Results and 10 CFR 50.59 Criteria The results of the loss of load to both steam generators are contained in the Fort Calhoun Station Unit No.1 USAR. The criteria of 10 CFR 50.59 are met, if the peak RCS pressure is less than 110% of design pressure in accordance with Sec-tion III of the ASME Pressure Vessel Code. This objective is achieved if the RCS pressure does not exceed 2750 psia.
5.12.7 Conservatism of Results The following areas of conservatism are included in the anal-ysis to obtain a conservatively high peak RCS pressure:
() A. Field measurements demonstrate that the CEA magnetic clutch decay time is less than that assumed in the l
analysis.
B. The actual scram worths are greater than those assumed in the analysis, i
C. The actual MTC is more negative during power operation l ,
than assumed in the analysis, l
i D. The steam dump and bypass system and the pressurizer l <
pressure control system (PORVs and sprays) are oper-ated in the AUTO mode rather than MANUAL as assumed in the analysis.
I
,0
' N-] ,
OPPD NA 8303 NP Rev. 02 Page 72 of 129
l l
5.0 TRANSIENT AliD ACCIDENT ANALYSIS METHOD (Continued) 5.12 Loss of Load to Both Steam Generators Event (Continued) i 5.12.7 Conservatism of Results (Continued) i E. Actual secondary. pressure is higher which results in l {
earlier secondary safety valve opening and earlier ;
alleviation of the primary system temperature and l pressure rises, f i
I F. The maximum pressurizer safety valve capacities are l assumed to be "On of the ASME rated values.
- I C. A one percent pressure uncertainty is applied to the l h
primary and secondary safety valve setpoints, i.e., a j 1.01 multiplier.
5.13 . Loss of Feedwater Flow Event O,
i I
l 5.13.1 Definition of Event ,
?
A total loss of main feedwater flow event is defined as a '
loss of feedwater flow when operating at power without a corresponding reduction in steam flow from the steam gen-erators. The most likely causes for this event are the loss of all feedwater or condensate pumps or the inadver.
{
tent closure of either the main feedwater regulating valves or the feedwater isolation valves due to a feedwater con. I troller malfunction or manual positioning by the operator.
The result of this mismatch in which turbine demand remains at 100t, is a reduction of the steam generator liquid inven-tories and a degrading RCS heat removal capability. As the heat removal capability is lost, through decreasing steam generator inventor.tes (i.e., levels) the RCS temperacutes O
OPPD NA 8303 NP, Rev. 02 Page 73 of 129 l
l 5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued)
I
- 5.13 Loss of Feedwater Event (Continued) !
5.13.1. Definition of Event (Continued) ,
, and pressure increase. Normally the event would be ter. I f
minated by a reactor trip on low steam generator level.
Since no credit is taken in the analysis for the steam l l
l generator low level trip, a high pressurizer trip even- i tually results, l
Automatic actuation of the auxiliary feedwater (AW) sys+
tem will also eventually occur (after reactor trip) if r
i either main feedwater is not restored or manual actuation ^
j of the AW system is not performed by the operator. The !
- AW system actuation ensures the maintenance of a second-ary heat sink. I i '
5.13.2 Analysis Criteria '
a +
.i t
The loss of feedwater flow event is classified as an Anti- ,
j cipated operational Occurrence (A00) for which the follow-1 ing criteria must be met: !
/ ,
l i 1
l A. The peak RCS pressure does not exceed 2750 psia
' l i, j (110% of design pressure). l i
i B. The transient minimus DNBR is greater than the l l 1
95/95 confidence interval limit for the CE 1 cor- j f relation limit. I i
l
(
1 I C. The Peak Linear Heat Generation Rate (PUICR) does j t
not exceed 22 kw/ft. l
) r t :
1 1 t
! OPPD NA 8303 NP, Rev. 02 l
} Page 74 of 129 I l' l 1
I {
5.0 TRANSIENT AND ACCIDENT ANALYSIS METil0D (Continued) i t-5.13 1.oss of Feedwate' Flow Event (Continued) i 5.13.2 Analysis Criteria (Continued)
Criteria B. and C. are not of major concern because DNBR
] does not decrease below the initial steady state value [
! and the P1JICR margin required is much less limiting than f 1
other A00s. Therefore, only criterion A. requires re- l !
evaluation should plant modifications (such as stews gen-
- f erator tube plugging) be made which result in degraded i
, secondary heat transfer capability beyond that of this event. For Fort Calhoun Station, this event is bounded l by the loss of load incident. !
1 i r
5.13,3 Obieetives of the Analysis !
- i !
1 -
l The objective of this analysis is to demonstrate, for l l
1 plant modifications which potentially degrade RCS heat ;
71 removal capability (including steam generator tube plug- j l ging), that the peak RCS pressure stays within 110% of ;
- the design pressure in accordance with Section III of the f
ASME Pressure Vessel Code. This objective is achieved if l
)
1 the peak RCS pressure does not exceed 2750 psia. l r
I 5.13.4 Kev Parameters and Annivsis Assumotions {
- The key parameters used in the loss of feedwater flow i event are given in Table 5.13.4 1. Assunptions in the analysis to daximize heat up of the RCS and consequently the peak RCS pressure include
3 l
A. The event is initiated by an instantaneous loss of j main feedvater. No credit is taken for the low steam generator level trip.
1 j OPPD NA 8303 NP. Rev 02 j Page 75 of 129 I
J
5.0 TPANSIENT AND ACCIDENT ANALYSIS HETHOD (Continued)
O 5.13 1.oss of Feedvater Flow Event (Continued) 5.13.4 rey Parameters and Analysis Assumntions (Continued)
- 8. The steam dump and bypass system is assumed to be in MANUAL (i.e.. inoperative).
C. The pressurizer pressure control system is in l MANUAL (i.e., PORVs and sprays are inoperable).
D. The pressurizer level control system is in MANUAL l with maximum charging and zero letdown flows.
E. The rod block system is assumed to prevent rod l motion (other than scram) during the traasient.
5.13.5 Analysis Method The analysis methods used by the District to analyze a loss of main feedwater flow event consists of using the CESEC computer code to simulate the event, utilizing the analysis assumption itsted in Section 5.13.4 (above) as input. and extracting the peak RCS pressure for compari-son with the 2750 psia upper limit.
5.13.6 Analysis Results and 10 CF1t 50. 5o criteria _
The results of the loss of feedvater flow are contained in the Fort Calhoun Station Unit No. 1 USAR. The cri-teria of 10 CFR 50.59 are met. if the peak RCS pressure is less than 110% of the design pressure in accordance with Section III of the ASME Pressure Vessel Code. This objective is achieved if the peak RCS pressure does not exceed 2750 psia.
OPPD NA 8303 NP, Rev. 02 Page 76 of 129
Table 5.13.4 1 i
O)
(v KEY PARAMETERS ASSUMED Ill THE LOSS OF FEEDWATER FLOV A!1ALYSIS Parnmeter Units Value l Initial Core Power MVt 1530 l 1
Initial Core Inlet *F Maximum allowed Temperature by Tech Specs.
l Initial RCS Pressure psia Minimum allowed i
by Tech. Specs Initial Steam Generator psia Minimum value cor-Pressure responding to care inlet temperature operating range.
Initial Core Mass x106 lbm/hr Minimum allowed l Flow Rate by Tech. Specs, i
Moderator Temperature x10ap/*F Most positive Coefficient l ,
allowed by Tech.
Specs.
l Fuel Temperature x10Ap/'F Least negative O Coefficient predicted during l
b core life.
k Fuel Temperature Coefficient Multiplier 0.85 CEA Drop Time see Maximum allowed by Tech. Specs.
Scram Reactfvity Vorth tar Minimum predicted during core lifetime Scram Reactivity Consistent with most Curve positive axial shape (bottom peaked) l allowed by Tech, i Specs.
Core Average H E8P BTV/hr ft2.*p Maximum predicted l during core lifetime Kinetics Parameters EOC parameters (minimum absolute 4).
RPS Response Time see 1.4 l [
OPPD tiA 8303 t;P. Rev. 02 l Page 77 of 129
l l
5.0 TRANSIENT AND ACCIDENT ANALYSIS METHOD (Continued) m
! )
(s / 5.13 Loss of Peedvater Flow Event (Continued) ,
5.13,7 Conservatises of Results A. Field measurements demonstrate that the CEA magnetic clutch decay time is less than that assumed in the i analysis.
B. The actual scram worths are greater than those as- l sumed in the analysis. !
l C. The actual MTC is more negative during power oper- l l
{ ation than assumed in the analysis. !
\
l l D. The steam dump and bypass system and the pressurizer 1 l s
] pressure control system (PORVs and sprays) are oper- l ated in the AUTO mode rather than the KANUAL mode as Ik s assumed in the analysis.
3 E. Actual secondary pressure is higher which results in ,
f 1
earlier secondary safety valve opening and earlier I i
1 alleviation of the primary system temperature and ;
] pressure rises.
I I
I !
- F. No credit is taken for a steam generator low level i
- trip. !
.i e
k i
4 r O
)' OPPD NA 8303 NP. Rev. 02 Page 78 of 129 i
t
6.0 TRN;SIENT ANALYSIS CODE VERIFICATION
( 6,1 Introduction
. The District utilizes the CESEC III computer code to calculate the transient response of the NSSS during events discussed in this doc-l unent. Combustion Engineering has provided overall verification of the CEfca III code in References 6 1 and 6 2. The purpose of the i work reported here is to demonstrate the District's ability to cor- <
j rectly utili e the CESEC III code.
1
, In order to demonstr=?e Omaha Public Power District's ability to cor- ,
t rectly use the CESEC III coeputer code, verification work has been performed by benchmarking both actual plant transient data and in-i dependent safety analyses previously accepted by the NRC. The plant transients which were benchmarked were the Turbine Reactor trip and Tour Pump I.oss of Coolant Flov events. The independent safety anal-yses which were benchmarked were the Dropped CEA, Main Steamline Break, and RCS Depressurization events. Each of the comparisons
- will be addressed below, !
l f
6.2 Co carison en Plant Data j i
A prerequisite for beginning performance of transient analyses is verification that the code will stabilize with the correct system
- parameters when simulating steady state operation, this step was !
! performed following setup of the CESEC III code and correct results '
] were obtained. I i
i Tor plant transient benchmarking, the type of transients that have '
i occurred and both the quality and quantity of data existing for each I
- is very limited. In nearly all cases, operators take actions which l
I
, reduce the consequences of the event, introducing complicated pertur-bations in system response which cannot be easily modeled, because the actions taken and the time at which they are performed ase not OPPD NA 8303 NP, Rev. 02
] Page 79 of 129 t
6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued)
( )
L 6.2 Comonrison to Plant Data (Continued) recorded. Scrip chart recordings on an extremely compressed time scale are generally the only form of data available. This com-pressed time scale (with graduations typically of 10 minutes) does not permit adequate comparisons to CESEC III modeling in which seconds are of major concern. The only source of plant transient data in which system parameters were measured with high spead strip chart recorders and no operator action taken, was during the cycle 1 startup testing. Good data existed for a nominal full power turbine reactor trip and a 354 power total loss of RCS flow event. The CESEC III computer code was set up to model Cycle 1 in a best estimate mode to permit accurate comparisons to the actual measured plant responses for both of the above cases. A summary of each of these comparisons follows.
O 6.2.1 Turbine Reactor Trin For the turbine reactor trip case, the plant comparison data were obtained from the Cycle 1 startup testing per-l formed May 10, 1974 The event was initiated from 974 of full power, all rods out, and equilibrium xenon. The plant response data used in the CESEC III comparisons were obtained from vendor te9t recorders. No operator action was taken following the manual generator turbine trip (which provided the RPS "loss of load" trip). Prior to the trip the main feedvater, the pressurizer pressure, and pressuri:er level control systems were all in the l automatic mode, and the letdoJn backprussure control valve was in the manual mode. ith . the exception of adjusting ths letdown backpressure contrcl valvo at 20 l secondo, r.o operator action w s taken for 60 seconds l folly.ini *be trip 1O Q
. key. C2
.29 l
6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued) 6.2 Cnmenrison to Plant Data (Continued) 6.2.1 Turbine Reactor Trio (Continued)
Figures 1 1 through 1 7 show plots of the comparisons be.
tween the measured plant responses and the CESEC.III pre.
t dicted responses. It should be noted that this test was performed based on a rated power level of 1420 MVt rather than the current limit of 1500 MVt (the design power fe,r which licensing was obtained in Cycle 6). -
l 4
Figure 1 1 shows the nuclear power response following the turbine reactor trip. The CESEC.III prediction follows (
the same power decay rate, however, the endpoint residual ,
r power is slightly higher, i.e., conservative. It should :
j be noted that trip delays included in the CESEC III model.
ing prevent the immediate power drop observed in the 1
plant data; again this is conservative. The pressurizer !
l pressure response predicted by CESEC.III and shown in t
] Figure 1 2 shows very good agreement with the plant re- '
j sponse. The CESEC.III case was initiated 10 psia above
} the plant data and remained slightly above the plant re.
! sponse for the duration of the transient. The difference between the predicted and measured pressurizer pressures ,
increased slightly due to the higher residual power after
!i trip as shown in Figure 1 1. This difference between the !
1 predicted and neasured pressurizer pressures at 60 se. f q conds is only 19 psia, a value which is less than the pressure measurement uncertainty. Figure 1 3 shows the pressurizer level response. The comparison between the ;
j measured and predicted values shows excellent agreement.
Figure 1 4 shows the RCS cold leg and hot leg temperature responses for each steam generator loop for the plant l l
OPPD.NA 8303.NP. Rev. 02 Page 81 of 129
i l
1 l
6.0 TRANSIENT ANAL.YSIS CODE VERIFICATION (Continued) 6.2 Comearison to Plant Data (Continued) .;
- 6.2.1 Turbine Reactor Trio (Continued) i data and the CESEC III predicted average cold leg and hot- I leg temperatures. The diffennees in the transient. re- I
- sponse of the two steam generater loops for the plant [
data is attributable to the differences in the main feed.
i 1
water flow rate rampdown after trip (see Figure 1 6). [
The CESEC III responses lead the loop measurements be. {
s cause of the measurement delays associated with the re. I
- sponse time of the RTDs (resistance temperature devices) )
providing the temperature signals. Figure 1 5 shows the >
l
- measured and predicted steam generator pressure re.
sponses. These two plots show very good agreement with !
each other with only minor differences. The predicted j pressure is slightly higher early in the event due to a i
combination of the greater heat residual as shown in Fig- e t
ure 1 1, a quicker turbine stop valve closure, and quick- !
1 :
er steam dump bypass operation assumed in the CESEC III l 1
analysis. The latter two effectr , which are shown in the
- steam flow of Figure 1 7, would show better agreement if ;
! the CESEC III input were modified, however, the overall i i
1 j differences are small enough not to warrant the reanaly-
, sis.
?
i !
! j q In conclusi,on, the CESEC III predicted parameters for the j
, turbine reactor trip show very good agreement with those measured in the Cycle 1 startup testing performed at nom- '
inal full power conditions, i
l I i l !
i !
\ i i I OPPD NA 8303 NP, Rev. 02
) Page 82 of 129 !
i [
i !
6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued)
(O
(_-) 6.2 Comparison to Plant Data (Continued) 6.2.2 four Puro toss of Coolant Flow For the four pump loss of coolant flow case, the plant com-parison data were obtained from the Cycle 1 startup test performed March 6, 1974 This event was initiated from 354 power by manually a.d simultaneously tripping all four reac-tor coolant pumps. At the time of trip the pressurizer ,
pressure, pressurizer level, main feedwater, and steam dump and bypass controllers were in the automatic mode. At ap- ;
proxt.sately 20 seconds after the trip, the operators took i
manual control of feedwater in order to preclude overfeed-ing of the steam genetators and too rapid of a cooldovn for the following natural circulation test. ,
The behavior of the various RCS and secondary parameters
) that were measured and the CESEC III predictions for the first 30 seconds following the RCP trips are shown in Fig-l ures 2 1 through 2 8. These comparisons show excellent
! agreement. The minor differences that exist are discussed below.
Figure 2 1 shows a plot of the measured total RCS flow ver-sus time and that predicted by the CESEC III code which in-i corporates explicit modeling of the reactor coolant pumps.
These data show excellent agreement with the predicted flow
, being slightly conservative. Figures 2 2 and 2 3 show the pressurizer pressure and level response comparisons which 2
also show excellent agreement. Figure 2 4 shows plots of core nuclear power versus time. As in the turbine reactor trip case. CESEC III shows a slightly higher residual power after trip. The predicted and measured steam generator OPPD NA S303 NP Rev. 02 Page 83 of 129
6.0 TPA';SlisT ";'> CODF VERIFICATION (Continued) 6.2 32s,.. a to Plant Data (Continued) l 6.2.2 Four Pumn Loss of Coolant Flow (Continued) pressure responses as plotted in Figure 2 5. also show very good agreement. The response of the hot leg and cold leg temperatures, as shown in Figure 2 6, is con-l sistent with the data obtained from the turbine reactor
, trip case. Again the delay associated with the RTD re-sponse causes the predicted temperatures to lead those that were measured. Figure 2 7 shows that the main feed-water input function used in CESEC III was acceptable in terms of the actual feedwater system response. It should 3
be noted that the operator action of assuming manual con- l j trol of the main feedwater system at approximately 20 seconds had little effect on any of the other system para-f meters examined, and that following a several second re-
\
l duction in flow the previous flow rate was reestablished.
i Figure 2 8 shows that turbine stop valve closure rate l
l assumed in the CESEC III analysis was quicker than the L j actual valve response. The fi5ure also shows a steam j flow rate mismatch between the two steam generators for ,
J the plant data. This is something one vould not expect '
and raises the question of the validity of the measure.
ment or its uncertainty for this steam generator steam i
rate flow, because the two corresponding feedwater flow j rates (in Figure 2 7) are consistent. !
j In cenclusion, the CESEC III predicted parcaeters for the 354 power total loss of coolant flow show very good agree-ment with those measured during Cycle 1 startup testing. !
f !
t I
OPPD NA 8303 NP. Rev. 02
, Page 84 of 129 e
i i
6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued)
(
N- 6.3 Coenarisons Between OPPD Analyses and Independent Analyses Previously Perforced by the Fuel Vendors Of the transients analyzed by OPPD for reload core licensing (using CE methodology) no plant data existed, so comparison of the limiting l events to previous independent analyses performed by either Advanced Nuclear Fuels, formerly Exxon Nuclear Company (ENC), or Combustion l
Engineering (CE) was done. Since Exxon Nuclear Company performed some analyses in this section (used for comparison) prior to becoming ANF, all references to this company will be to ENC. For the compar-ison cases, the assumptions used in the analyses were similar to :
those used by the District, i.e., the core physics parameters did not vary significently between fuel cycles. The events chosen for compar.
ison vere:
(1) The Dropped CEA event is dependent upon the initial avail- l
/~ able overpower margin to prevwnt exceeding the SAFDLs. The goal of the analysis is to determine the DNBR required over-power margin (ROPM).
(2) The Hot Zero Power (HZP) Main Steamline Break which deter.
mines the minimum required shutdown margin.
(3) The Hot Full Power (HPP) Main steamline Break which dete- !
raines the most negative moderator temperature coefficient of reactivity allowed.
l l
(4) The RCS Depressurization event which is used in the deter- I mination of the term. The term accounts for DNBR margin degradation in the thermal margin / low pressure (TM/LP) trip.
i O- %
OPPD NA 8303 NP, Rev. 02 1
Page 85 of 129
\
l
d i f I
6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued) l O 6.3 Coronrisons Between OPPD Analyses and Indeoendent Analyses Previousiv l
Perforced by the Fuel Vendors (Continued) 6.3.1 Droceed CEA f
j pared to the previous analysis, contained in the Updated i
) Safety Analysis Report (USAR). .The USAR analysis was per. f
] formed by ENC for Cycle 6. Table 6.3 1 summarines the para. [
meters and their values for Cycles 6 and 8. Plots of core f power versus time for the OPPD (Cycle 8) and ENC (Cycle 6) [
1 l analyses are found in Figure 3 1. The curves show a very l
l similar prompt drop, to 694 versus 704, respectively, and [
i t both cases show a return to a nominal 1004 power. Both j j cases assumed that the turbine admission valves opened to
(
)
their full open position in an attempt to maintain full load {
during the event (i.e., the turbine control system was
- placed in the load set mode which is not used at Fort Cal. ;
l houn Station). The core heat flux plots are contained in l 3 ,
j Figure 3 2. Both are very similar, as was the case in the l
j core power cases. Figure 3 3 contains pints of the coolant !
l i
j average temperature versus time. Both figures are in good t I
agreement showing a drop in avetage coolant temperature to
\
567'F. Plots of the inlet and outlet temperatures for Cycle l 8 are also included. Figure 3 4 shows plots of the pres. i surizer pressure versus time. The minimum pressures pre.
- dicted at 160 seconds are 1957 psia and 1945 psia for Cycle f
j 8 and Cycle 6, respectively. This difference is small I
- enough to be less than the pressute measurement uncertainty.
i s f
f In summary, the primary system responses between the ENC and l t
j OFFD analyses show excellent agreement with each other which 1
is consistent with reioad cores having similar core physics i parameters.
l i f OPFD.NA 8303.NP, Rev. 02
] Page 86 of 129 1
j '
6.0 TPANSIENT ANA1.YSIS CODE VERIFICATION (Continued)
\ '
6.3 Coevarisons Petween OPPD Analyses and Independent Ansivses Previousiv Perforced by the Fuel Vendors. (Continued) 6.3.2 Hot Zero Power Main Steamline Break The hot zero power (HZP) Main Steanline Break, which is the basis for detertaination of the required shutdown margin, was analyzed by OPPD for Cycle 8. The results of this analysis
- have been compared to those of ENC in their Cycle 6 analysis and to those obtained by CE in their Cycle 6 control grade auxiliary feedvater (AW) system analysis. Table 6.3 2 -
shows comparisons of the psreinent input values for each of '
the analyses, j Figure 4 1 shows plots of core power for the Cycle 8 OPPD l analysis and Cycle 6 ENC analysis, respectively. The. max.
imum return to power is Aess for Cycle 8 than for Cycle 6 and occurs later due to the use of a higher shutdown mar.
gin. The Cycle 6 CE AW analysis power is not included because there was no return to. critical and no return to.
power. Figure 4 2 shows plots of the core average heat flux
! for OPPD, ENC and CE, respectively. Both the OPPD and CE f analyses, which were performed using CE!EC.!!! and CESEC I.
I respectively, show a slight heat flux increase at approxi. ,
raately 12 seconds. This is due to suberitical multiplica. ,
tion. Otherwise, the heat flux curves within the specific analyses are essentially the same as the core power curves with a slight decay. Figure 4 3 shows the total reactivity
- versus time for each of the analyses. With very similar A
I moderator cooldown curves, the peak reactivities occur chron.
olot.ically with increasing shutdown margin as expected:
4 1.e., for increased shutdown cargin (CEAs) it takes longer 1' i to be offset by the positive moderator cooldovn reactivity l insertion.
j OPPD NA 8303 NP. Rev. 02 Page 87 of 129
6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued) 6.3 Cgynarisons Between OPPD Analyses and Independent Ansivses Previous 1v Perforced by the Fuel Vendors (Continued)
! 6.3.2 Hot Zero Power Main Steamline Break _(Continued) i Figure 4 4 shows plots of RCS pressure versus time for Cycle 8 (OPPD) and Cycle 6 A W (CE). Also included in Figure 4 4
]
j is the Cycle 1 (CE) results. All three of these curves show
- excellent agreement. The Cycle 6 A W (CE) analysis shows a
.l lower end point pressure than the Cycle 1 (CE) and Cycle 8
) (0 PPD) analyses due to the assumption of auxiliary feedwater I
addition. The ENC data available did not include the RCS pressure respr se.
rigure 4 5 m. plots of the steam generator pressures for Cycle 8 (OPPD) and Cycle 6 A W (CE), respectively. These plots sho.r reasonable agreement between prcssures and times.
The increase in the intact steam generator's pressure is due 1 l
j to MSIV closure; i.e., failure of the reverse flow check valve on the intar.t steam generator was chosen as the most !
' s adverse single failure. Following dryout of the ruptured i steam generator, the pressure drops to atmospheric. The j
times of dryout are slightly different due to the increased normal water level value used in the Cycle 8 analysis. i i
In summary, the HZP Main Steamline Break analysis for Cycle 8 shows trends similar to those in Cycle 6 as analyzed by j
6.3.3 Hot Full Power Main Steaalina Break l The hot full power (HFP) Main Ster.mline Break prov W s en 4
l acceptance criteria for the most negative moderator tempera-OPPD NA 8303 NP Rev. 02 Page 88 of 129 i
l
6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued)
(,,)
N' 6.3 Comparisons Between OPPD Analyses and Independent Analyses Previousiv Performed by the Fuel Vendors (Continued) 6.3.3 Hot Full Power Main Steamline Break (Continued) ture coefficient (MTC) of reactivity. If a return to crit-ical occurs, the goal of the reload analysis is to show that the return-to-pouet is bounded by the most limiting case which, for the Fort Calhoun Station, is the Cycle 1 analy-sis. The Cycle 8 HFP analysis of this event was compared to the previous analyses performed by ENC in Cycle 6 and by CE in their Cycle 6 control grade AFW system analysis. Table 6.3 3 shows a comparison of the important input parameters for each of the analyses.
Figures 5-1, 5 2, and 5 3 show plots of core power, core
(~'j average heat flux, and total reactivity for Cycle 8 (OPPD).
Cycle 6 (ENC), and Cycle 6 AFW (CE). Within each cycle's analysis, the core average heat flux slightly lags the core power which peaks at a time several seconds after the peak reactivity is reached (for the return to critical cases).
The return to power peaks occur at different times due to the different scram worths used, as explained for the shut-down margin in the HZP Steamline Break analysis section.
Figure 5 4 shows plots of the RCS pressure versus time for the Cycle 8, Cycle 6 AFW, and Cycle 1 analyses. These plots are very similar and show excellent agreement. Figures 5 A and 5 B show plots of the RCS temperatures for Cycle d and Cycle 6 AFV. Again good agreement exists to approximately 180 seconds. At this time, the C?cle 6 AFW analysis assumed runout flow from both AFW pumps to the ruptured steam gen-p)
(
L/
OPPD NA 8303 PP, Rev. 02 Page 89 of 129
l
}
I 6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued) r%
i) 6.3 Comonrisons Between OPPD Analyses and Indeoendent Analysis Previously l
Performed by the Fuel Vendors (Continued) 6,3,3 Hot Full Power Main Steamline Break (Continued) erator which resumed the RCS cooldown. This additional cool-down caused by the AFW system is prevented from occurring in Cycle 8 by the logic of the newer safety grade AFW system.
Figure 5 6 shows plots of steam generator pressures versus time for Cycle 8 and Cycle 6 AFV (CE). These results are very similar except that the intact steam generator pres-sure, in the CE analysis, begins to drop after 180 seconds ,
due to the AFW induced RCS cooldown, 6,3,4 RCS Deoressurization O The RCS Depressurization analysis is performed to calculate a term for the TM/LP trip which accounts for the DNBR margin degradation.
Because no figures from previous cycle analyses exist, com-parison was made between the transient analysis training manual sample analysis and the figures generated by OPPD for j Cycle 8. Pertinent input parameters are summarized in Table 6.3 4 t
1 I
4 O ,
CPPD NA 8303 NP, Rev, 02 I Page 90 of 129
6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued) i r~x l
( ) I
6.3 Comparisons Between OPPD Analyses and Independent Analyses Previously l
EtL{ormed by the Fuel Vendors (Continued) 6.3.4 RCS Deoressurization (Continued)
A manual trip is next simulated at the time of maximum mar-gin degradation, i.e., at the time the maximum RCS Depres-surization rate occurs. The maximum RCS Depressurization rate occurs in approximately the first 20 seconds and is constant. T.-Tore, the time at which a manual trip should occur is arbA..ary but must be in the first 20 seconds. A trip time corresponding to a 100 psia drop is adequate to perform the analysis.
In the CE example, the initial pressure was 2300 psia, a value which corresponds to the maximum pressure before which
(~') the pressurizer sprays will be activated in a 2700 MW(th) class plant (whose normal RCS pressure is 2250 psia). In the Cycle 8 analysis, a value of 2172 psia was used for the initial RCS pressure, since the normal operating RCS pres-sure at Fort Calhoun is 2100 psia. The Fort Cal nun pres-surizer sprays are fully closed at 2175 psia and fully open at 2225 psia.
The comparison of the figures show good agreement in the trends for the core power, core asarage heat flux, and RCS pressure.
A OPPD NA 8303 NP, Rev. 02 Page 91 of 129
6.0 TRANSIENT ANALYSIS CODE VERIFICATION (Continued)
O 6.4 Summary The CESEC III computer code was developed for the analysis of FSAR transient and accident events for the two by four loop combustion Engineering plants. OPPD's engineering staff was trained for tb proper use of CESEC-III and a close consultive relationship has been maintained with the CE development team. Extensive man-hours were involved in setting up the plant model, testing the input, validating and verifying the output and quality assuring the base-deck to ensure the applicability of the use of CESEC-III for steady state plant operation. Benchmarking against Cycle 1 plant data for the Turbine Reactor Trip and the Four Pump Loss of Cool-ant Flow was performed and excellent agreement between the predict-cd and observed responses was obtained.
For transients in which plant data were not available, comparisons were performed between the OPPD Cycle 8 analyses of the limiting '
transients and the Cycle 6 analyses of the fuel vendors (CE and ENC) and, in one case, the transient analysis training manual exam-ple. In all cases, these benchmarking comparisons showed very good agreement.
The District continues to maintain a quality assured model and pro-vides an update basedeck for each cycle, contained in an Opera-tions Support Analysis Report (OSAR).
O OPPD NA 8303 NP, Rev 02 Page 92 of 129
,__. TABLE 6.3 1
\
COMPARISON OF PARAMETERS INCLUDING UNCERTAINTIES USED IN THE CEA DROP ANALYSES FOR CYCLES 6 AND 8 Parameter Units Cvele 6 Cvele 8 Initial Core Power Level KWt 102% of 1500 102% of 1500 Core Inlet Temperature 'F 547 547 Pressuri::or Pressure psia 2053 2053 RCS Flow Rate gpm 190,000 197,000 Moderator Temperature Coeff. 10 ap/*F 2.3 2.7 Doppler Coeff. Multiplier 1.20 1.15 CEA Insertion at Full Power t Insertion 0.0 25.0 Dropped CEA Vorth top 0.34 0.28 bi V
1 l
l
)
I l
i V
OPPD NA 8303 NP. Rev. 02 Page 93 of 129
~. .
I 1
TABLE 6.3 2
+
COMPARISON OF PARAMETERS INCLUDING UNCERTAINTIES j
USED IN THE HZP MAIN STEAMLINE BREAK ANALYSIS FOR CYCLES 6 AND 8 l
cycle 6 i Parmeter Units Cvele 6 AW Cvele 8 {
t 4
Initial Core Power Level MWt 0.0 1.0 -1.0 .
i i
Core Inlet Temperature 'T 532 532 532 ;
- Pressurizer Pressure psia 2053 2175 2172 f RCS Flow Rate gpm 190,000 190,000 197,000 Effective Moderator Temperature 10~0ap/'F 2.3 -2.3 2.5 [
Coefficient l
Doppler coeff. Multiplier 0.8 .1.15 1.15 l E
, Minimum CEA Scram Worth g op 3.0 -4.2 4.0 -
(Shutdown Margin) t Initial Steam Generator Pressure psia N/A 900 895 '
Initial Steam Generator Mass t Narrow 63 63 70 Inventory (Level)
! Range Scale j
t i i
4 !
I l
1 !
1 ,
e ;
i i l 1
i !
l i i :
i l
A OPPD.NA 8303 NP, Rev. 02 l
4 Page 94 of 129 l
1
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TABLE 6.3 3
/--T COMPARISON OF PARAMETERS INCLUDIt'C UNCERTAINTIES USED IN Tile IlZP MAIN STEAMLINE BREAK ANALYSIS FOR CYCLES 6 AND 8 Cycle 6 Parameter Units Cvele 6 AW Cvele R Initial Core Power Level MWt 102% of 1500 102% of 1500 102% of 1500 Core Inlet Temperature 'F 547 547 547 Pressurizer Pressure psia 2078 2175 2172 RCS Flow Rate gpm 190,000 190,000 197,000 Moderator Temperature 10Ap/'F 2.3 2.3 2.5 Coefficient Doppler Coeff. Multiplier 0.8 1.15 1.15 Minimum CEA Scram Worth 4 op 5.81 -5.81 6.68*
Initial Steam Generator psia N/A 880.5 890 Pressure ON V Initial Steam Generator Mass Inventory (Level)
% Narrow Range Scale 63 63 70
- Reduced to -6.57 to account for axial shape.
' (
OPPD NA 8303 NP, Rev 02 Page 95 of 129
.f m s TABLE 6.3 4 e ;
COMPARISON OF PARAMETERS INCLUDING UNCERTAINTIES USED IN THE RCS DEPRESSURIZATION ANALYSES FOR CYCLES 6 AND 8 Parameter Units Example Case
- Cvele 8 Initial Core Power Level Mk't 102% of 1500 102% of 1500 Core Inlet Temperature 'F 547 547 Pressurizer Pressure psia 2300 2172 RCS Flow Rate gpm N/A 209,796 Moderator Temperature 10'04/' F -2.5 2.7 Coefficient Doppler Coeff. Multiplier 1.15 1.15 Example case input data consistent with 2700 MVt plant operating
' characteristics.
U(3
[
\
OPPD NA 8303 NP, Rev. 02 Page 96 of 129
7.0 REFERENCES
(j Section 4 References 41 CESEC, Digital Simulation of a Combustion Engineering Nuclear Steam Supply System December, 1981, transmitted as Enclosure 1-P to LD 82 001, January 6, 1982.
4-2 CEN 234(C)-P, Louisiana Power and Light Company, Wateriord Unit 3 Docket 50 382, Response to Questions on CESEC, December, 1982.
4-3 Letter from A. E. Scherer (CE) to F. J. Miraglia (NRC), "Applica-bility of CESEC III to the Fort Calhoun Station," February 27, ,
1987.
44 CENPD 161 P, TORC Code, A Computer Code o f' r Determining the Ther-mal Margin of a Reactor Core," July.1975.
45 CEN 191(B) P, "CETOP.D Code Structure.and Modeling Methods for Calvert Cliffs 1 and 2," December, 1981.
46 CENPD-162 P A, "CE Critical Heat Flux, Critical Heat Flux Correla-
. tion for CE Fuel Assemblies with Standard Spacer Crids Part 1 Uni-form Axial Power Distributions," September 1976.
47 CENPD 207-P "CE Critical Heat Flux, Critical Heat Flux Correla-tion for CE Fuel Assemblies with Standard Spacer Crids Part 2 Non.
() 48 uniform Axial Power Distributions," June, 1978, 1 '
Letter from E. C. Tourigny (NRC) to W. C. Jones (OPPD) dated March 15, 1983.
49 OPPD NA 8301, Rev. 03, "Reload Core Analysia Overview", April, 1988. >
4 10 CEN 257(0) P, "Statistical Combination of l'ncertainties",
November, 1983.
Section 3 References 51 CEN 347(0) P, Rev. 01, "Omaha Batch M Reload Fuel Design Report", i j January, 1987. '
57 CEN-121(B) P, "CEAW, Method of Analyzing Sequential Control Ele-ment Assembly Croup Withdreyal Event for Analog Protected Sys-tems", November, 1979.
53 CENPD 199 P, Revision 1.P. "0E Satpoint Methodology", April,1982.
J OPPD NA 8313 NP, Rev. 02 l Page 97 of 129 2
i
7.0 REFERENCES
(Continued)
-s .
Section 5 References (Continued) 54 _ Fort Calhoun SER on Automatic Initiation of Auxiliary Feedwater, contained in tht letter to V. C. Jones from Robert A. Clark, dated-February 20, 1981. ;
5-5 CENPD-190 A "CE Method for Control Element Assembly Ejection Anal-ysis", July 1976.
56 OPPD NA 8302, Rev 02, "Nuclear Design Methods and Verifications", ,
April, 1988.
5-7 - Letter from D. M. Crutchfield (NRC) to A. E. Scherer. (CE). "Safety Evaluation of Combustion Engineering ECCS Large Break Evaluation Model and Acceptance for Referencing of Related Licensing Topics Reports". July 31, 1986.
58 Fort Calhoun SER on Generic Letter 86-06 (TMI Action Item II.K.3.5, "Automatic Trip Reactor Coolant Pumps During Loss of Coolant-Acci-dent"), contained in the letter to R. L. Andrews from Anthony '
Bornia, dated March 25, 1988.
Section 6 References 61 "CESEC - Digital Simulation of a CE NSSS", Enclosure 1-P to O LD 82 001, January 6,1982. ,
62 Letter from A. E. Scherer to F. J. Miraglia, LD 87 013 dated February 27, 1987. ;
h l
l 1
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OPPD NA 8303 NP, Rev. 02 Page 98 of 129 l
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OPPD.NA.8303.NP. Rev. 02 Page 112 of 129
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PLANT DAT A: TEST PERFORMED MAACH 6, 197a o
G I 0:anaPublicPowerDistrict Figure 4-Pu.coLossOfFlow FcrtCalhounStation-Unitflo.1 2-8 Stes:FlowvsTir.e OPPD NA 8303 NP, Rev 02 Page 113 of 129
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Page 114 of 129 l
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CoreAveraaeHeatfluxvsTite Fort Calhou.n " ' n-Unit tio. i 3-2 1 I -
4 NA.8303 NP. Rev. 02-t'G' 115 Of 129
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i O :
CEADropIncident OmahaPublicPowerDistrict Figure
! CoolantTe:ceraturevsTite FortCalhounStation-UnitNo.i 3-3 ;
4 l 0FPD.NA.8303.NP. Rev. 02 e I
i Page 116 of 129
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Figure ;
CEADrcoIncicent 0:anaPublicPowerDistrict Fort Calhoun Staticn-unit No i I 3-4 PressurizerPressurevsTite OPPD NA-8303 NP Rev. 02 Page 117 of 129
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- CYCLE 8
- OPPD ANAL.YSIS WITH SDM= 4.0%sf i
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) Zero Power Stesa t.ine Break Incident OmahaPublicPowerDistrict Figure l 1 CorePowervsTime FortCalhounStation-UnitNo.i A-i j
' OPPD NA.8303 NP Rev 02 l i
Page 118 of 129 j
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-,CoreAveraceHeatFluxvsTi:e OPPD NA 8303 NP, Rev. 02 Page 119 of 129
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CYCLE 6 AFW: CEAtlALYSISWITHSOM=4.2%af l 4
CYCLE 0: OPPDA!!ALYSISWITHSOM=4.0%af l
O lero Power Steas une Break Incicent CtanaPublicPowerDistrict 'Fjqure l TotalReactivityvsTite C3 Fcrt Calhoun Station-Unit !!0.,1 y
' OPPD NA 8303.NP Rev, 02
! Page 120 of 129
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OsanaPublicPowerDistrict figure j ZeroPowerSteasLineBreakIncident 4-5 i Steas Generatar Pressure vs lite Fort Calhoun Station-Unit ib.1 OPPD NA 8303 NP. Rev. 02 Page 122 of 129
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CYCLE 6 AFW: CE ANALYSIS (CEA WORTH = 5.81hf) l c CYCLE 8: OPPD ANALYSIS (CEA WORTH = 3.57hp) l i
O Figure lI i FullPowerSteasLineBreakIncident OmahaPublicPowerDistrict CorePowervsTite FortCalhounStation-UnitNo,i 5-1 i OPPD NA 8303 NP, Rev. 02 Page 123 of 129
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l;ill Power Stes: Line Break Incident 0:anaPublicPowerDistrict figure '
CoreAveraceHeatFluxvsTire FortCalhounStation-UnitNo,i 5-2 OPPD.NA.8303.NP. Rev. 02 Page 124 of 129
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! l OPPD NA 8303 NP. Rev, 02 Page 125 of 129 [
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O l Full Power Sten Line Break Incident 0:ana Public Power District ' Figure I Ccolant Systen Pressure vs Ti::e FortCalhounStation-Unittio.i 5-4 _ l OPPD.NA.8303.NP, Rev. 02 Page 126 of 129
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Full Power Steam Line Break Inciden 0:ana Pub',ic Power District figure i !
Peact:r Ccolant Te:ceratures vs Tite I Fcrt Calhoun Station-Unit tb. ii s-s'. i OPPD NA.8303 NP, Rev 02 Page 127 of 129 ;
l
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Full Power Steam Line Break Incident OmahaPublicPowerDistrict Figure i P.eactcr Coolant Te:ceratures vs Tite fortCalhounStation-Unit!!0.i s-se OPPD NA 8303 NP Rev. 02 Page 128 of 129
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! FullPowerSteanLineertakIncident OmahaPublicPowerDistrict Figure i StesaGeneratorPressurevsTite Fort Calhoun Station-Unit !!o.1 5-6 l ..
i OPPD.NA 8303.NP, Rev. 02 i
Page 129 of 129
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