ML19260B251
| ML19260B251 | |
| Person / Time | |
|---|---|
| Site: | Fort Calhoun  |
| Issue date: | 11/30/1979 |
| From: | William Jones OMAHA PUBLIC POWER DISTRICT |
| To: | |
| Shared Package | |
| ML19260B250 | List: |
| References | |
| NUDOCS 7912070464 | |
| Download: ML19260B251 (28) | |
Text
. i OMAHA PUBLIC POWER DISTRICT FORT CALHOUN STATION UNIT NO. 1 LOSS OF LOAD TO ONE STEAM CENERATOR INCIDENT ANALYSIS NOVEMBER 1979 1515 056 7912070 h
1.0 StDNARY OF RESULTS This report contains the results of the loss of load to one steam generator transient for the Fort Calhoun Station, to supplement the incidents covered in Reference A. The results of this analysis are summarized in table 1.1.
This transient is very similar to the turbine trip transient (see Section 3.2 in XN-NF-79-79, Loss of External Electric Load).
The pressurization of the primary system is slightly slower than in the turbine trip transient, where the thermal load on both steam generators is lost simultaneously. As the result in Table 1.1 shows, there is no significant reduction in margin to DNB.
From the analysis results it is concluded that the plant thermal response to this transient is acceptable for steady-state operation at 1500MW.
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TABLE 1.1 Summary of Results Maximum Maximum Core Average Pressurizer Power Level Heat Flux Pressure MDNBR Transient JPercent) (Btu /hr-ft2) (psic) (W-3)
Intiel Conditions 102 176,213 2053 1.87 for Transient Loss of Load to 103 177,941 2530 1.87 One Steam Gener-ator 1515 058
2.0 DISCUSSION OF RESULTS The loss of load to one steam generator, also called loss of heat sink in one loop, was analyzed using several bounding assumptions.
For the actual initiator to this transient, either a closure of the feed-water flow control valve or the steamline isolation valve would be sufficient.
For analysis purposes, in order to establish an enveloping set of conditions, both the steamline isolation valve and the feedwater flow control valve were closed within 4 sec. In addition, all steamline safety valves were blocked in both loops. Blocking these valves leads to conservatively high peak pressure values in both steam generators, and it terminates the heat absorbing capability of the isolated steam generator more quickly. With the steamline valves operating, the release of steam would enhance further steam generation and thereby serve as a heat sink. On the primary system, the pressurizer relief valve was blocked. Both BOC and EOC core conditions were analyzed, using high and low initial pressurizer pressure. The most limiting transient resulted from EOC conditions, starting from an initial pressurizer pressure of 2053 psia. This transient is similar to the loss of load (turbine trip) transient, where the load to both steem generators is lost. As in the loss of load transient, the reactor is tripped on a high primary system pressure signal, set at 2422 psia. For t loss of load in both loops (turbine trip) a reactor trip occurs between 8 see and 10 sec, whereas the loss of load in one loop generates a reactor trip at about 16.5 sec. The peak power is 103 percent, the pressurizer pressure reaches a peak 1515 059
value of 2530 psia at 16.6 sec, which is the safety valve setpoint. The pressurizer safety valve opens at 16.5 see and rescats in less than 2 sec.
Figures 2.1 through 2.6 show the traces for the plant response to this transient.
1515 060
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3.0 REFERENCES
- 1. XN-NF-79 Fort Calhoun Cycle 6 Reload Plant Transient Analysis Report, Exxon Nuclear Company, October 1979.
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c Question 1 The limiting crial pcuer shapes used in calculating the M/LP trip LSSS.
Respcnse The axial power distributions used in the generation of the TM/LP Lafety limits are shown in Table 6.1. The axial power profiles were calculated using EllC neutronics methodology for the Ef4C Ft. Calhoun reload fuel design.
Table 1.1 Axial Power Profiles for TM/LP Safety Limit Calculations F f X/L A _A 0.000 0.000 0.000 0.004 0.000 0.000 0.004 0.409 0.316 0.024 0.409 0.316 0.063 0.546 0.453 0.101 0.652 0.521 0.140 0.742 0.556 0.179 0.843 0.593 0.217 0.934 0.635 0.256 1.044 0.684 0.295 1.153 0.742 0.334 1.256 0.806 0.372 1.350 0.876 0.411 1.421 0.949 0.450 1.457 1.024 0.489 1.451 1.099 0.527 1.431 1.173 0.556 1.387 1.244 0.605 1.312 1.309 0.643 1.204 1.367 0 682 1.106 1.420 0.721 0.974 1.464 0.760 0.879 1.496 0.798 0.774 1.503 ('
O.837 0.682 1.467 0.876 0.570 1.325 0.915 0.425 n.978 '
0.934 0.425 0.978 O.934 0.000 0.000 1.000 0.000 0.000 1515 068
2 Question 2 List all bc=-to-core and pin-to-box peaking factors used for TM/LP calcula-tions.
Response
The box-to-core and pin-to-box peaking factors used in the analysis are shown in Figures 4.2 and 4.4, respectively. The box-to-core peaking factor is defined as the ratio of assembly power to core average assembly power. The pin-to-box peaking factor is defined as the ratio of fuel pin power to assembly average fuel pin power. The peaking factors given in Figures 4.2 and 4.4 were calculated using ENC neutronics methodology for the ENC Ft. Calhoun reload fuel design.
Question 3 Lict all hot channel factors and provide an c planation of hou and at uhat stage th3ll arc included in m calculation of the TM/LP trip.
Response
This response describes the generation of thermal margin / low pressure safety limit lines (Figure 3.1). The TM/LP safety limit lines are isobars on a plot of inlet coolant temperature versus core power and represent the loci of operating points which result in a calculated MDNBR 1.3 or in maxi-mum hot channel quality no greater than 15 percent, and otherwise satisfy the criteria set forth in Section 4.2 of XN-NF-507.
Generation of Til/LP safety limit lines entails two types of calculation:
- Core flow distribution calculation to determine the coolant flow to the hot assembly.
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540 f f f I I f f f f f I f I 60 70 80 90 100 110 120 Core Power, Percent of 1500 MW Figure 3.1 Thermal margin / low pressure safety limits four-pump operation 1515 070
4
- t4D!iBR calculation which employs the resulting hot channel flow factor to determine the mininum ceparture from nucleate boiling ratio.
Both types of calculations are performed in accordance with ENC methodology for predicting fuel assembly DNB performance (Xti-75-48).
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5 CORE FLOW DISTRIBUTION CALCULATIONS FOR TM/LP SAFETY LIMIT LINES PURPOSE:
Determine hot assembly flow factors for use in the thermal margin (MDNBR) calculations.
MODEL:
Quarter-core syn,me+ry is assumed 'see response to Question 4). Table 5.1 lists the component loss coefficients for both fuel types as used in the analysis. The axial locations of spacers and tie plates are noted in Figure 4.3. The radial power profiles employed for these calculations may be found in Figures 4.2 and 4.4. Additional model parameters used in the core flow calculations are given in the response to Question 4.
OPERATING CONDITIONS:
The operating conditions used in core flow distribution calculations are
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given in Table 3.1. A 5 percent plenum flow maldistribution is postulated.
Individual channel flow factors used to model the maldistribution appear below:
Channel No. (Figure) Flow Factor _
l-6, 8-12, 14-16, 20, 21, 26, 27, 38, 39 0.95 Realaining 1.067 1515 072
6 Table 3.1 Operating Conditions for Core Flow Distribution Calculations System pressure 2078 psia Coolant inlet temperature 547 F (nominal plus 2 measurement uncertainty)
Core average mass flux 6 2.097 x 10 lb/hr-ft 1515 073
7 MDNBR CALCULATIONS FOR TM/LP SAFETY LIMIT LINES PURPOSE:
Determine steady state operating conditions which preclude penetration of specified acceptable fuel design limits with respect to DNB.
MODEL:
The subchannel model is illustrated in Figure 4.4 and assumes 1/8 assembly symmetry. The hydraulic characterization and axial model for the limiting ENC fuel assembly is identical to that used in the core flow distribution calculation (see response to Question 4, Figure 4.3, and response to Ques-tion 5, Table 5.1) . The local peaking distribution appears in Figure 4.4 and the axial power profiles are given in the response to Question 1.
PROCEDURE:
(1) Select the nominal operating pressure for which a TM/LP safety limit line is desired, ano ca.1 it P . gProgram input system pressure is given by Pg - 22 psi, where 22 psi is the magnitude of pressure measurement uncer-tainty.
(2) Select an appropriate nominal core coolant inlet temperature and call it T .
g Program input core coolant inlet temperature is given by Tg +2F, where 2 F is the magnitude of temperature measurement uncertainty.
(3) The core coolant mass flux is determined as the product of the nominal mass flux (based on nominal active coolant flow as cited in the response to Question 6, Table 6.1) and the hot channel flow factor calculated in the core flow distribution calculation.
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8
- 4) The limiting hot assembly LHGR is determined by iterative calculations which employ the values of system pressure, inlet coolant temperature, and coolant mass flux froa Steps 1 through 3, varying only the hot assembly LHGR between calculations. Allowable core pcwer is obtained according to:
P = power fraction = LHGR j /LHGRref where LHGR is the hot assembly LHGR at which MDtlBR = 1.306. Sub-crit W-3 script W-3 indicates the ratio obtained via the W-3 correlation. The value 1.306 is determined from the 95/95 MD?tBR limit of 1.3 by applying a 1.03 local engineering heat flux factor and a 0.975 correction which adjusts the heat flux to reflect only that fraction of energy which is generated directly in the fuel:
MDilBR W-3 = 1.3 x 1.03 x 0.975 = 1.306 LHGR is given by:
ref (I)l0.346 for P > 1. 0 LHGR =4(2)6.3864 (P - 2.0125)' for 0.925 < P < l.0 ref
. - 0.625 .
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(3)l1.1124 for P < 0.925 The iteration procedure is complicated by the dependenge of the hot assembly flow factor on hot assembly LHGR. That dependence must be deter-mined using the core flow distribution model. Therefore, Step 4 may have to be modified by varying also the coolant mass flux.
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9 Steps 2 through 4 are repeated at appropriate temperatures to generate a single TM/LP safety limit line. Further TM/LP safety limit lines may be i generated by repeating Steps 1 through 4.
(1) 10.346 = 1.02 x 1.02 x 1.06 x 1.62 x 6.0076/0.988 x 1.05 T
1.06 is a measurement / calculational uncertainty o: F xy 1.62 is F xy at P = 1.0 6.0076 is nominal 100 percent power core average LHGR 0.988 is a calculational factor apnlied to account for reduction of nominal hot channel flow area due to allowed manufacturing tolerances 1.05 is conservatively estimated as the least possible maximum local peaking factor 1.02 is porter measurement uncertainty 1.02 additional conservatism (2) 6.384 = 10.246/Fx (P = 1.0)
F xy = (P - 2.0125)[- 0.625) f or 0.925 < P < l .0 (see figure 5)
T (3) 11.1124 = (10.346/Fxy(o = 1.0)) 1.74 1.74 =F xy(p < 0.925) where F is given as a function of core power level as shown in Figure 3.2.
1513 076
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1515 077
. u 11 Ouestion 4 Provide a schcmtic shouing the core and fuel asocmbly (axial and radial) modeling used for tha TM/LP calculations.
Response
The quarter core model used in the Cycle 6 analysis is shown in Figure 4.2 and 4.3. Figure 4.2 shows the radial layout of the core while Figure 4.3 shows the axial layout of the core including spacer type and location. The values of the parameters used to define the core geometry are shown in Table 4.1. The assembly layout used in the subchannel (MDNBR) analysis is shown in Figure 4.3. Assembly parameters also appear in Table 4.1. Table 4.1 Core and Assembly Geometrics CE ENC Rod pitch 0.580 in. 0.580 in. Rod 0.D. 0.440 ir. 0.412 in. Guide tube 0.D. 1.115 'n. 1.115 in. Assembly pitch 8.18 in. 8.18 in. Outer assembly to core shroud distance 0.08 in. Active fuel length 128 in. 128 in. Total fuel length 137.7 in. 137.7 in. 1515 078
12 CE CE CE CE CE CE 0.796 0.579 1.080 1.078 1.057 1.079 l 37 36 35 34 33 , 32 EriC l 0.885 CE CE CE CE CE CE 39. 1.216 0.979 1.298 0.976 1.123 0.856 7 31 30 29 ENC 28 27 26 g 0.675 CE CE I CE CE CE ErlC 38 1.071 1.302 1.166 1.125 0.954 1.084 I 25 24 23 22 21 20
! CE CE CE CE , CE Et;C l.102 0.979 1.119 1.070 0.875 0.898 19 ; 18 , 17 16 15 '
14 CE CE CE , CE CE ENC 1.077 1.125 0.943 0.871 0.902 0.625 13 12 11$ 10 9 8 CE CE Efic EilC EfiC 1.097 0.862 1.495 i 0.897 0.513 l 7 6g _ 5l 4j 3 EllC EtlC 0.893 0.679 2, 1
.I AAA Assembly Type Figure 4.2 Core model for core flow distribution analysis.
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13 3 - 0.999 2 - 0.863 2 - 0.741 2 - 0.619 2 - 0.497 2 - 0.374 2 - 0.252 2 - 0.130 1 - 0.001 X/L t Spacer and lowei tie plate 1 2 Spacer 3 Spacer and upper tie plate Figure 4.3 Axial stations of spacers and tie plates. 1515 080
14 w
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ey e e e a'_ 7 [TO O) O[9 O O[ t[h2' Rod Power I (.'lonna lized to i.009 9* 1.037 1.00,i 0. %0 assembly average roilpower) 9 1.045j 007 Control Rod 20 21 ' Guide Tube 1.05 1.011 2h h3 1.013 1.003 1 24 1.001
- Includes 1.03 engineering heat flux factor.
Figure 4.4 Ft. Calhoun subchannel model for MDNBR evaluation. 1515 081
15 Question 5 Provide informtion used for the hydraulic modeling of the E!/C fuel: grid friction factors (K) and lateral mixing coefficient B.
Response
The nominal component loss coefficients f< r both Et1C and Combustion Engineering fuels used in the thermal-hydraulic analysis for Cycle 6 appear in Table 5.1. These values were obtained by performing single-phase pressure drop testing (Ref) for both fuel types. The turbulent mixing correlation also appears in Table 5.1. Table 5.1 Fort Calhoun Single-Phase Hydraulic Loss Coefficients Exxon fluclear Combustion Engineering Bare rod friction 0.1987 Re
-0.20 -9.20 0.1987 Re
-0.067 Spacers 1.62 Re 1.019 Re -0.04 Lower tie plate 3.53 3.55 (includes lower core support structure)
Upper tie plate '7.53 7.45 (includes upper core alignment strJCture)
-0.1 Turbulent mixing coefficient, W O.0062 Re GD H
W' = turbulent crossflow G = mass flux (arithmetic average over interconnected subchannels) D H = hydraulic diameter (arithmetic average over interconnected subchannels) Re = Reynolds number (arithmetic average over interconnected subchannels) 1515 082
16 Question 6 Provide acowncd base ecolant conditions for Ft. Ca thow: Cycle G calculations. R,esponse The design operating conditions for Cycle 6 as used in the thermal-hydraulic analysis appear in Table 6.1. Table 6.1 Ft. Calhoun Design Operating Conditions (Nominal Values) Rated power 1500 MW h Inlet coolant temperature 545 F Active coolant flow 68.5 x 106 lb/hr Pressure 2l00 to 2075 psia 1515 083}}