ML19206A248
| ML19206A248 | |
| Person / Time | |
|---|---|
| Site: | Crane  |
| Issue date: | 07/31/1977 |
| From: | BABCOCK & WILCOX CO. |
| To: | |
| References | |
| BAW-1455, NUDOCS 7904180429 | |
| Download: ML19206A248 (38) | |
Text
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b BAW-1455 July 1977
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1 THREE MILE ISLAND, UNIT 2 I
i FUEL DENSIFICATION REPORT i,
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O BAW-1455 July 1977 THREE MILE ISLAND, UNIT 2 FUEL DENSIFICATION REPORT e
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- SM,2 C Centrcl # 7729DCF Date /e//7/77 ef ? cr.uS IES:!LLTC'!Y C3;;'T 7"_I BABCOCK & WILCOX Power Generation Group Nuclear Power Generation Division P. O. Box 1260 Lynchburg, Virginia 24505 Babcock & Wilcox
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Babcock & Wilcox Power Generation Group Nuclear Power Generation Division Lynchburg, Virginia Report BN4-1455 July 1977 Three Mile Island, Unit 2 - Fuel Densification Report Key Words:
Fuel. Densification Effects ABSTRACT In Nos m 1976, Babcock & Wilcox filed topical report BN4-100831, Rev. 1,I which de. tibes rc ' sed e.ethods to be used in analyzing fuel densification effects as required by the guidelines set forth in the AEC report, " Technical Report on Densification of Light '4ater Reactor Fuels," data Nover.ber 14, 1972.
This report presents an analysis of the effects of fuel densification on the fuel for Three Mile Island, Unit 2 and supports the safe operation of that unit at the rated core power level of 2772 }Mt.
Babcock & Wilcox
- 11 _
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CONTENTS Page 1.
INTRODUCTION.
1-1 2.
CONCLUSIONS 2-1 3.
RESULTS 3-1 3.1.
Power Spike Model 3-1 3.2.
Ther=al Analysis.
3-1 3.2.1.
Fuel Thermal Analysis 3-1 3.2.2.
Thermal-Hydraulic Design.
3-2 3.2.3.
DNBR Analysis 3-2 3.3.
Nuclear Analysis.
3-4 3.3.1.
Reactor Protection System 3-4 3.3.2.
Analysis of Power Distributions Before Densification...................
3-4 3.3.3.
Analysis of Power Distributions With Densification Effects 3-6 3.3.4.
Snema ry 3-8 3.4.
Safety Analysis 3-8 3.4.1.
Introduction.
3-8 3.4.2.
Reactivity Insertion Transients 3-8 3.4.3.
Loss of Coolant Flow.
3-10 3.5.
Mechanical Analysis 3-10
- 3. 5.1.
Cladding Collapse 3-10 3.5.2.
Cladding Streau 3-11 3.5.3.
Cladding Strain 3-12 4.
REFERENCES.
4-1 List of Tables Table Page
-~
3-1.
Minimum Fuel Melt Limits.
3-12 3-2.
Thermal-Hydraulic Design Conditions 3-13 3-3.
Modifications to Reactor Protection System Design 3-14 Parameters.
3-4.
Thermal Data Input for Safety Analysis.
3-15
- 111 -
Babcock & Wilcox v.'
' TG s
s List of Figures Figure Page 3-1.
Maximum Gap Size Vs Axial Position.
3-16 3-2.
Power Spike Factor Vs Axial Position.
3-17 3-3.
Fuel Melt Limit Vs Burnup 3-18 3-4.
Maximum Fuel Temperature Vs Linear Heat Rate at 100 mwd /mtU 3-19 3-5.
Average Fuel Temperature Vs Linear Heat Rate at 100 mwd /mcU 3-20 3-6.
Minimum DNBR Vs Reactor Inlet Temperature 3-21 3-7.
Minimum DNBR Vs Power-Maximum Design Conditions 3-22 3-8.
Trip Setpoint Vs Axial Imbalance Without Densificatica Effects 3-23 3-9.
Calculated Offset Limits Vs Power 3-24 3-10.
Pressure, Fower, and Flux Vs Time for Densified Fuel Rod Ejection Accident 3-25 3-11.
DNBR, Average Fuel Temperature, Centerline Fuel Teuperature, and Cladding Temperature Vs Time for Rod Ejection Accident 3-26 3-12.
Power, Flow and Flux Vs Time for Densification Fuel -
Four-Pump Coastdown 3-27 3-13.
DNBR and Centerline Fuel Temperature Vs Time for Four-Pump Coastdown 3-28 3-14.
Pressure, Power, Flow, and Flux Vs Time for Locked Rotor Accident 3-29 3-15.
DNBR, Maximum Fuel Temperature, and Maximum Cladding Temperature Vs Time For Locked Rotor Accident 3-30
- iv -
Babcock & Wilcox kY
1 1.
INTRODUCTION In Nosember 1972, the Atomic Energy Co= mission published a " Technical Report on Densification of Light Water Reactor Fuels," containing guidelines and criteria for including the effects of fuel densifice. tion in safety analysic reports.
In November 1976, Babcock & Wilcox filed topical report BAW-10083P, Revision 1,1 which established revised methods for implementing the guidelines
~'
and criteria of the AEC report. These methcds have subsequently been approved by the NRC.2 This report contains the specific analysis of the effects of fuel densifica-tion on the initial core for Three Mile Island, Unit 2 (THI-2).
The analysis supports the operation of IMI-2 at 2772 MRt for the first three fuel cycles.
The thermal analysis section considers protection of the fuel =elt and DN3R criteria.
The nuclear analysis section deceribes thermal design criteria, core offset limits, and core power distributions.
The mechanical analysis section contains the input sir-nry and results, cladding creep and collapse, cladding stresses, and fuel pellet irradiation swelling.
The safety analysis section addresses all postulated accidents analyzed in the TMI-2 FSAR3 except the loss-of-coolant accident (LOCA), assuming that densification occurs. The LOCA analysis is included in BAW-10103, Revision 2,4 which was approved in February 1977.
l-1 Babcock & \\Nilcox la
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2.
CONCLUSIONS Based on the analysis perforned for TMI-2, the following conclusions cre =ade regarding the effects of fuel densification:
1.
The cladding vill not collapse during the first three fuel cycles.
2.
The integrity of.he fuel rods will not be i= paired. Neither the central fuel te=perature nor =ini=u= DN3R li=1ts will be violated.
3.
The reactor can be operated safely at the rated power level of 2772 MWt with only =inor =odifications to reactor protection syste= (RPS) setpoints.
These modifications will ensure that the thermal design criteria are not exceeded.
2-1 Babcock & \\Vilcox
'l3(h
a 3.
RESULTS
' l.
Power Soike Model The power spike model utilized in this analysis is identical to that presented in BAW-10054, Rev 2,5 and in BAW-100556 except for two modifications. The modifications have been applied to F and F.7 These probabilities have been changed to reflect additional data from operating reactors and yield less severe penalties due to power spikes.
F was changed f rom 1.0 to 0.5.
F,& was g
changed from a Guassian to a linear distribution, which reflects a decreasing frequency with increasing gap size. These modificacions have been approved on previous B5W-related applications.7 The =aximum gap size versus axial position is shown in Figure 3-1, and the power spike factor versus axial position is shown in Figure 3-2.
These fig-ures also show the initial and final theoretical densities (TDI, TDF) used in the calculations.
3.2.
Thermal Analysis 3.2.1.
Fuel Thermal Analysis Fuel thermal aralyses have been performed with TACO,8 a fuel pin temperature and internal pressure analysis code that includes models for fuel and cladding te=perature distribution, t ime-dependent fuel densification and swelling, fis-sian gas release, and cladding creep. These analyses have been performed on the basis of specification fuel data -- both with densification to 96.5% TD assumed ard including the results of fuel pellet resinter tests.
Pertinent input data for these analyses include the following:
1.
Initial fuel pellet densities and assumed density reductions due to densi-fication (Table 3-1).
2.
A fuel pin power history based on radial power factors versus burnup that envelop the maximum radial power factors obtained from physics calculations.
3.
No fuel restructuring.
3-1 Babcock & Wilcox
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Assumptions 1 through 6 above were also used in FSAR analyses. Table 3-2 is a su= mary of the i=portant ther=al-hydraulic design conditions used for both the FSAR and this report.
Figure 3-6 depicts the minimum DNB ratio as function of both reactor inlet temperature and 4, stem operating pressure. The infor-mation on Figure 3-6 serves as the bas 's for the variable low-pressure trip function in the RPS to ensure that DNB does not occur.
Figure 3-7 shows the minimum DNBR as a function of core power.
3.3.
Nuclear Analysis 3.3.1.
Reactor Protection Svstem The safe operation of a reactor core requires an extensive analysis of power distributions because of the various modes of plant operation. The primary considerations of this analysis are as-follows:
1.
Assurance that thermal criteria are not exceeded; i.e.,
specified minimum DNBRs and centerline fuel temperatures may not be viciated.
2.
Definition of i= balance limits to prevelt adverse power peaks that would exceed the foregoing criteria.
3.
Definition of core operational limits and recccmended operating procedures to prevent unnecessary reactor trips.
The complete maneuvering study includes a combined nuclear-ther=al analysis of the power distributions. This section describes the metheds and criteria used in developing the RPS setpoints required to account for postulated densifica-tion effects.
3.3.2.
Analysis of Power Distribur. ions Before Densification The three-dimensional PDQ07 code with thermal feedbcck effects was used to analyze power distributions. This analysis determined power distributions for all modes of reactor operation except accidents and other rapid transients.
The design power transient (100-50% power and return to 100% at peak xenon) was analyzed throughout core life. The fuel cycle and transient analyses de-termine power distributions for normal equilibrium and transient conditions, respectively. The extremes of core operation, such as control rod bank inser-tion beyond normal limits and maloperation of axial power shaping rods, were also examined. The extreme control rod bank conditions defined the limits for a preliminary imbalance protection system.
3-4 Babcock & kVilcox
3.3.2.1.
Correlation of Power Peaks to Ther=al Design Criteria The power peaks from PDQ cases were corrected for calculational uncertainty and were analyzed to determine the margin to the ther=al criteria - centerline fuel melt and departure from nucleate boiling. The margin to centerline fuel melting was defined as Fuel melt =argin =
- 1 100%
max calculated peak The maximum allowable peak defined as the pointwise pcwer that yields center-line fuel melting was fue3 melt limit (kW/ft)
Max allowable peak = 6.105 kW/ft x 1.014 x FOP 6.105 kW/ft = average heat rate at 2772 MWt, 1.014 := hot channel factor, FOP = fraction of power.
The =axi=um calculated peak is the largest total peak frcs the PDQ power in-creased by a factor of 1.075 to account for calculational uncertainty.
The deter =ination of DNB sargin requires a more complex analysis. DNBR is a function of peak location, the =agnitude of the pcwer peak component parts (radial and axial), and other core parameters. To arrive at true DNB condi-tions, each power distribution was analyzed explicitly.
From the PDQ power distribution, the maxi =um calculated total peak was obtained and adjusted for uncertainty. Tite DNB cargin was then defined as DNB sargin =
- 1 100%
, max calculated total peak The basis for the allowable total peak was a 1.30 DNBR associated with the protection system envelope, or a quality limit based on the CHF correlation applicability, whichever was most limiting.
3.3.2.2.
Offset-Margin Relationship Core offset, a measure of the axial power imbalance, is defined as the frac-tion of total core power in the top half of the core minus the fraction of total core power in the bottom half of the core:
3-5 Babcock & Wilcox o.'
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= Power (too) - cover (botton)
Offset power (top) + power (bottom)
The relationship between hot channel power peaks (i.e.,
thermal =argins) and core offset defines the protection system setpoints. Power imbalance is the primary signal to the protection system for flux shape protection:
Imbalance = ? "*" (' ")
- " "*" (b *)
full power The maneuvering analysis defines the relationship between core imbalance and thermal margin.
Limiting offsets are determined to prevent violation of thermal criteria for all operating conditions and power levels. To yield the imbalance trip enve-lope, the limiting offset values are corrected for potential instrumentation errors, imbalance detection bias, and calibration. The imbalance trip enve-lope defined the range of allowable operational imbalance and ensured that the 1.30 DNBR and/or the central fuel melting limit would not be exceeded.
Figure 3-8 presents the trip setpoints based on those criteria.
3.3.3.
Analysis of Power Distributions With Densification Effects Providing for possible fuel densification requires modification of the imbal-ance trip system for two reasons - the fuel melt (kW/ft) criterion change and the.aclusion of an additional power spike in the reactor power distributions.
Since the power spike factor is a function of axial position, the appropriate power spike factor is used to increase each peak to account for potential den-sification.
The modified offset limits with fuel densification effects included are shown in Figure 3-9 and are compared with the previous of fset limits. The primary, differences between the two-sets of calculated limits are as follows:
1.
D:iBR was based on BAR-2 including rod bcw penalties (section 3.2).
2.
The central fuel melt limits were based on TACOd (section 3.2).
3.
Average linear heat rates were based on as-built data and fuel column shortening from densification (section 3.2).
4.
The local power spikes were applied to the calculated total power peaking distributions for fuel =elt margin calculations.
5.
Local radial power distributions were calculated with PDQ07 in a two-dirensional, pin-bypin representation.
3-6 Babcock & Wilcox e
6.
Tbc previously limiting PDQ07 power distributions were reanalyzed with FLAME 310,11 to set the final limits consistent with present design tech-niques.
The trip setpoints are obtained from the calculated offset limits by adjusting for potential electronic errors and offset measurement bias by the out-of-core detectors.
The final error adjusted limits, i.e.,
trip setpoints, are included in the Technical Specifications.
The imbalance trip points and overpower trip nrovide operating flexibility with assurance that thermal criteria are not ex-ceeded.
Furthermore, potential relaxation of these limits may be realized from the physics tests, in which the offset bias instrumentation oehavior and the measured flow in the RC system will be determined.
3.3.1.1.
ECCS Considerations Power peaks during normal operation are mair.tained less than the LOCA kW/f t limits by placing restrictions on certain core operating parameters. These operating restrictions include the following:
1.
Axial power imbalance limits.
2.
Control rod position limits.
3.
Quadrant power tilt limits.
4.
Operational restrictions, as required, to limit transient xenon effects.
These operating restrictions are selected with respect to maximum peaking con-ditions that occur during steady-state depletion or power maneuvering.
3.3.3.2.
Normal Fuel Cvele Operation A summary of the typical fuel cycle calculations for the first cycle of a B&W 2772-MWe, nonrodded plant is presented in Table 3.3-2 of B&W-1393.12 A 7.5% uncertainty factor is included in the radial and axial components, and the total peaks inc.lude the 1.014 hot channel factor.
The maximum total hot spot peak is presented with the corresponding peak linear power density (kW/ft).
The kW/ft values are presented for the hot spot in a balanced core and normal control rod positions.
3.3.3.3.
Transient Data A 100-50-100% pcwer transient is the design transiest for the core. The core is designed to recover from 50 to 100% power at max. mum xenon, which occurs approximately 6 to 8 hours9.259259e-5 days <br />0.00222 hours <br />1.322751e-5 weeks <br />3.044e-6 months <br /> after the power reduction.
After recovery to full 3,/
Babcock r. Wilcox 5 1. 2 $ 8
O power, the transient xenen reactivity and undershoot are controlled by adding 12 sum =arize or remnving soluble boron. Tables 3.3-3 through 3.3-5 in BAW-13?3 the calculated data for three transients at various times in core life.
The uncertainty and hot channel factors are included in the components and tetal peaks, respectively.
3.3.4.
Sum =ary Luel densification, rod bow, and associated design limit changas produce mod-ifications to the RPS and operational changes. Revising the RPS with control operating restrictions for the emergency core cooling system (ECCS) will allow operation at 100% pcwer with assurance that thermal criteria, with all densi-fication effects included, are not exceeded. The design modifications are summarized and compared with the previous system in Table 3-3.
3.4.
Safety Analysis 3.4.1.
Introduction The effects of fuel densification are an increase in caximum fuel temperature, an increase in average heat flux due to shrinkage of the pellet sta k length, and spikes in the neutron power due to postulated gaps in the fuel. Ice de-creases in fuel length due to densification will lead to decreased initial DNBR for the accident calculations presented in the TMI-2 FSAR.3 The FSAR accidents, with more severe initial conditions, were assessed, and some were recalculated assuming the changes in fuel geometry and the higher fuel temper-ature dze to fuel densification.
The original va3ues of the moderator and
'ippler coef ficients were conservatively used. All calculations were made for conditions. The B&W-2 CHF correlation 9 was used for all the DNBR calcu-lations. This correlation provides a more realistic prediction of the DNB phenomenon than the W-3 correlation, which was used in the FS AR. 3 3.4.2.
Reactivity Insertion Transients The rod withdrawal accident was not recalculated since, for all ecmbinations of parameters including the simultaneous withdrawal of all rods in the core, the peak thermal power attained during the transient is alwrys less than the 112% design thermal power level. Therefore, the 1.30 limit on DNBR is main-tained for this transient.
3-8 Babcock & kVilcox 19h[
Se startup of an inactive locp was not recalculated since the maxi =u= ther=al pcwer achieved during the transier.t is =uca less than 100% and cccurs after full ficw is reached.
S e rod drcp accident results in an initial decrease in pcwer, which is folicwed by a return to 100% pcwer and it has been shcwn that neither the withdrawal ner the drop of a single control element will perturb the flux shape encup. to exceed design conditiens at 112%. Se =cderator dilu-tien accident results in reactivity insertion rates that are very sicw, and the accident is ter=inated by the high-pressure trip well before pcwer reaches the 112% design thermal pcwer level.
Ther* fore, the 1.30 li=it on OGR is raintained. This accident was not recalculated.
"'he ejecticn frc= the core of the red of =aximu= Technical Specification value (0.6544 k/k) has been analyced censidering the effect of fuel densification.
'"he basic assu=ptiens for calculating the plant parameters are the sa=e as those presented in TMI-2 FSAR.3 Figure 3-10 shows the neutron pcwer, 13 pressure, and heat flux as cc=puted with CACD for the ejection of a 0.654 ak/k centrc1 red at beginning cf ccre life. The neutron pcwer reaches abcut 7C0%
pri_ _ to inward red =otion, which eccurs at abcut 0.4 second, af ter which the pcwer decays to value of about 40%.
The pressure increa:: 2s to about 2445 psia due to the increased energy transfer to the ecolant and decreases later in the transient. Table 3-4 shews the i=por ant assu=ptions for the ther=al analysis. Figure 3-11 shews the fuel and cladding te=peratures at the peint of maxi =u= temperature during the transient.
'"he fuel te=perature reaches centerline =olting about 0.1 seccnd after peak tautron pcwer.
Figure 3-11 also shows the cladding temperature and O!BR as functions of time.
The CNBR reached 1.30 at about 0.2 seccnd, after which the =aximu= cladding temperature reached was 1570F, a value well belcw the assu=ed l'~i t of 23COF.
~~
The cladding te=perature decreases after abcut 2.1 seconds due to the decreased neutron pcwer. A parametric study was perferred to deter =ine the percentage of fuel pins that would experience a CNER less than er equal to 1.30.
It was determined that fer the red worth analyzed (0.65% Ak/k), about 30.5% of the pins would exhibit a CNBR of 1.3 cr lcwer. As further evidence of the accept-ability of the the. mal effects during this transient, the peak enthalpy of the hottest fuel red was calculated to be 231.6 cal /g=.
This is well belew the safety 14 i t of 280 cal /g=.
Secondary system accidents resulting in pcwer increases cccur at or near end of life (ECL) when a highly negative =cderater ccefficient exists. Since =cre CND =argin exists at ECL, these secondary accidents, such as stea= line breaks, will not cause thermal ~dts that are =cre severe than these presented in the 3-9 OCk3 W COX
FSAR.
The FSAR analysis of secondary system accidents, such as steam genera-ter tube rupture and loss of electric power, is unchanged since the thermal power re=ains the same or decreases during the transients and therefore does not increase the potential for reaching design limits.
3.4.3.
Loss of Coolant Flow A 1.5 cosine axial flux shape has been used for the loss-of-coolant-flow acci-d ents.
E. Jure 3-12 shows power, flow, and average heat flux as calculated ~by CADD for densified fuel for a four-pump coastdown initiated from 1027. power.
For this slow transient, the ther=al-hydraulic hot channel calculations were performed with RADAR.14 Figure 3-13 shows the calculated DN3R and the center-line fuel camperatures as functions of ti=e.
The minimus DNBR is above the 1.30 limit and film boiling will no'. occur.
The locked rotor accident has been analyzed using the assu=ptions presented in Table 3-4.
Figure 3-14 shows the power, flow, pressure, and average heat flux as calculated by CADD for densified fuel.
The initial pcwer level for this accident was 102% of 2772 MWt.
Trip occurs at 0.23 second from the be-ginning of the coastdown.
RADAR was used to calculate the hot channel tem-perature; Figure 3-15 shows the results of the calculation.
Film boiling does occur; however, the marimum cladding te=perature is 1440F, and the =axi-eum fuel temperature is lower than 4600F.
The DN3R reaches the criterion value of 1.33 at 0.6 seconds, after which the cladding te=perature increases to a peak value of 1440F at 4 seconds frem the beginning of the transient.
3.5.
Mechanical Analvsis 3.5.1.
Cladding Collapse Creep collapse analyses were performed for die most conservative three-cycle power history. The history conservatively enveloped the proj ected design con-ditions and was used to determine the limiting collapse ti=e as described in BAW-10084P, Revision 1.15 Measured power distribution data obtained during Cycle 1 will confirm the ac-curacy of the Cycle 1 design calculations used for the collapse analysis.
3.5.1.1.
Conservatis=s 1.
The CROV computer code was used to predict the time to collapse.
CROV conservatf.vely predicts collapse times, as de=custrated in reference 15.
3-10 Babcock & Wilcox c
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2.
No credit is taken for fission gas release. Therefore, the net differ-ential pressures used in the analysis are conservatively high.
3.
A conservative cladding thickness and c conservative initial ovality were assured. Previous as-built cladding dimensions were the bases f or these values.
4.
The design power history envelope is conservatively high for three-cycle operation.
3.5.1.2.
Results The predicted time to collapse for the conservative conditions chossa is more 4
than 30,000 effective full power houts (EFPH).
3.5.2.
Cladding wtiass The fuel red cladding stress analysis selected individual loads representing limiting conditions and used them in combination. The conservative appt. wh increased the actual =argics of safety over the calculated margins.
SecLion III of the ASME Boiler and Pressure Vessel Code is used as a guide in classi-fying the stresses into various categories, assigning appropriate 11mics to these categories, and co=bining the stresses to detersine the stress intensity.
Present design criteria require that the internal pressure does not exceed system pressure during normal operation, ensuring that cladding stresses due to the pressure differential are always compressive during nor=al operatic 1.
The 'ollowing stresses were addressed in the analysis:
1.
Pressure differential stresses.
2.
Ovality bending stresses.
3.
Ther=al stresses.
4.
Grid load' stresses.
Differential fuel rod growth and flow-induced vibration stresses were analyzed and did not affect the worst-case stresses.
Internal and external pressures, ther=al gradients, and grid loads were determined and analy:ed simultaneously; conservative cladding dimensions were used.
The worst-case conditions for this analysis were found to occur at the beginning of life (BOL).
Long-term creep ovality stresses are addressed in the creep collapse analysis.
The primary membrane stress was less than two thirds of the minimum specified unirradiated yield strength, and all stresses were less than the mini =um
'-11 Babcock & \\Vilcox
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specified unirradiated yield strength.
In all cases the margin is more than 30%.
3.5.3.
Claddinz Strain The fuel design criteria specify a limit of 1.0% on cladding circumferential plastic strain. The pellet design is set so that the plastic cladding strain is less than 1% at 55,000 mwd /mtU. The conservatisms in this analysis are listed below:
1.
The maximum specification value for the fuel pellet diameter was used.
2.
The maximum specification value for the fuel pellet density was used.
3.
The cladding ID used was the icwest permitted specification tolerance.
4.
The =aximum crpected three-cycle local pellet burnup is less than 55,000 mwd /mtU.
Table 3-1.
Mini =um Fuel Melt Limits Fuel-cladding heat transfer Min fuel coeff at hot U o ao melt limit, kW/ft spot, Btu /h-ft2_ay Design values 92.5 4.0 20.5 1300 Resintered data, 93.02 2.43 21.0 1790 batch 1(8)
Resintered data, 93.74 2.15 22.1 2240 batch 2(a)
Resince ed data.
92.99 2.25 21.6 1680 batch 3 a)
(a)The method used to determine these values is discussed in BAW-10083P, Rev 1.1 c.
3-12 Babcock & Wilcox Os? - gy]
9 9
Table 3-2.
Thermal-Hvdraulic Desien Conditions Densif'n TMI-2 FSAR report Design power level, l'We 2772 2772 System pressure, psia 2200 2200 RC flow, gpm 369,600 369,600 Vessel inlet coolant 557 557 temperature - 1002 power, F Reference design radial-local 1.783 1.783 power peaking factor Reference design axial flux 1.5 cos 1.5 cos shape Hot channel factors Enthalpy rise 1.011 1.011 Heat flux 1.014 1.014 Flow area 0.98 0.98 Active fuel length, in.
144.0 141.7 Average heat flux - 100% power, 185,090(*)
137,722
)
Btu /h-ft2 CHF correlation W-3 BAW-2 h imum DNBR - 112% power 1.39 1.62 (a) Based on the active fuel length and cold fuel pin diameter.
(b) Based on the densified active fuel length and hot fuel pin diameter.
I*s 32 % ~ $_T! 9 Babcock & Wilcox 1
3
Table 3-3.
Modifications to Reactor Protection Svstem Design Paraneters Previous Modified (a)
Parameter system system Imbalance Svstem Fuel melt limit, kW/ft 22.2 21.0 min Nominal heat rate, kW/ft 6.11 6.19, 6.20 Overpower, of 2772 MWt 112 112 Offset limits at rated power Positive offset
+38 40.8 Negative offset
-52
-49.2 Spike factor None 1.00-1.08 Nu.:1 car power peak uncertainty 1.075 1.075 ECCS Considerations Resttictions on operating parameters are included in the Technical Specifications to ensure that LOCA kW/ft limits are not exceeded.
(a) Includes the effects of densification, rod bow, and as-built data.
Bgggogk & Wii n 3-14 0,,1 AuU
Table 3-4.
Thermal Data Input for Safety Analysis Active fuel length, in.
141.7 Fuel pellet diameter, in 0.370 Fuel cladding thickness, in.
0.0265 Hot channel factors Overall power factor (F )
1.0107 q
Local heat flux f actor (F)
1.0137 q
Flow area reduction factor 0.98 Assumed DNB for film boiling 1.30 DNB correlation used B&W-2 Errors Inlet te=p, F
+2 Pressure, psi
-65 Flux trip setpoint, %
+6.5
~
=-
~
4f's 3-15 Babcock & kVilcox
Figure 3-1.
Maximum Cap Size Vs Axial Position 2.4 2.2 2.0 1.8 5
1.6 1.4
- n
(-
1.2 e
1.0 a
wx 3
C.8 TDF = 96.5%
0.6 TDI = 94.0%
~
enrichment = 3.2%
0.4 0.2 O
f 9
t e
I t
0 20 4u eu eu 100 120 140 Axial Location, in.
Babceck & Wilcox 3-16 o.., i.
,r 4<
Figure 3-2.
Power Spike Factor Vs Axial Position 1.09' 1.08 1.07 1.06 t
1.05 b,)
1.04 u
1.03 s
TDF = 96. 5%
M = 94.07.
1.02 enrichment = 3.2*.'
1.01 1.00 i
0 20 40 60 80 100 120 140 Axial Location, in.
~0'.'3 Babcock & Wilcox 3-17
Figure 3-3.
Fuel llelt !.i mi t Vs Iturntip 23.8 J'
23.4 Hesinte real Dat a ( Ha t c h 1},.
23.0
/ *'
22.8 22.6 o
s 3
22.4
/
22.2
~
/
u g
/
Design Data d
y 3
22.0 l-7 5
M 21.8 L
/
~
v
\\
21.6 [
21.4 t-
/
I I
21.2,f 21.0 20.8 20.6 to i
20.4 g
0 4
8 12 16 20 24 28 32 36 40 n
[,,
10-3 t%I/mt11
- s cs
=
O I,
x s
Figure 3-4.
Maximum Fuel Temperature Vs Linear Heat Rate at 100 M'Jd/mtU 5400 Fuel Melt 4600 4200 i
5 5
C 3800 I,
au i--
2 3400
~
uv w
i 3000 a
~
="
2600
- 2200, 1800 1400 6
8 10 12 14 16 18 20 22 Linear Heat Rate, kW/ft Babcock & Wilcox 3-19 r,
oe -nin
.s..Q
Figure 3-5.
Average Fuel Te=perature Vs Linear Heat Rate at 100 Pid/mtU 3000 2800 2600
-+
a 2400 4w 5..
2200
- )
9 5
2000
+
z 2
"4 5
1800 1600 1400 1200 1000 t
6 8
10 12 14 l6 18 20 22 Linear Heat Race, kW/ft 3-20 Babcock & Wilcox G y - ~,, D
2=
amu 2
'u i
=.
CNN zN
- . N E=w
- . O W e4
=N e -.
%.M
=m a
A u
-0
=
=
r a
z c
w u
c
=
s e
L N
A i-a N
y v
A e
u=
=
=
z w
w t
=
- e u
= ?
u
=
N c H u
N
=
av z
=
1
=
r c
=
c
=.
m E
=
=
z s
w
=.
u
=
c
=
e
^ :
=
=
N y
n u
o x
1 M
4 O
v m
m u
a e
=
=.
2 O
C m
I f
f f
N c
c c
c c
= c M
M N
N 4
4 C-Mil
'dENG snsTuTA' 3-21 Babcock & Wilcox 7.,
es < ~ m,,.,
e y
Figure 3-7.
Mlai=um DNBR Vs Power-Maximum Design Conditions 2.1 2.0 Four Pu=ps Operating 1.9 t2 i
1.8 5:
E E
1.7 5
=.
1.6 1.5 1.4 100 102 104 106 106 110 112 Core Power - ; of 2772.Y,it 3-22 Babcock & Klp'ox
,y o
a
Figure 3-8.
Trip Setpoint Vs Axial Imbalance '41thout Densification Ef fects 120 110 100 Overpower Trip Setpoint a
2 A
90 e,
n t%
F4 w
o a
80 u
3o O.
70 l
60 f
I f
f f
-60
-40
-20 0
20 40 Core Imbalance, %
Babcock & Wilcox 3-23 l -' s
') * ~
'{}
.r
Figure 3-9.
Calculated offset Limits Vs Power 120 Prior to Densification 110 100 h
~41th Densification and Other Effects y
R N
90 wc w
0v 80 70 t
i t
t
-60
-40
-20 0
zo 40 Core Offset, *:
3-2
,',Ba bcock rWilcox
Figure 3-10.
Pressure, Power, and Flux Vs Time for Densified Fuel Rod Ejection Accident
/
- 2440 s
- 2420 s
7.0 2400 0
- 2380
.3-an 6.0
- 2360 4
f 5
e Pressure
~
a 0
- 2340 7
}
[ Power
}c 5.s 2320 w
3e 1
Flow = Constant 2300 3
e e
U 4.0 2280 i
=
~
2260 3
?
Initial Core Avg a
y 3.0 -
Heat Flux = 159451 Btu /hr-f t2-2240 b
C E
x
/
2220 8
a 2.0
- 2200 3
n Core Avg Heat Flux
.y 2180
/
t
/
/
/
- amen '
1.C d 2160
, ' * ~ ~ ~
2140
(~
0 o
1.' 0 2.' 0 a.o a.o 5.'0 e.0 2120 Time, s
,;'; - n r 4
~
.> r 3-25 Babcock & Wilcox
G 3 'd=al Tang auTTaa2ua3 xrg
=
=
=
c
=
=
=
=
=
=
=
=
=
e x
e-e D
r
,c a
4 a
i e
3 'dmal Tang assaary xrg C
o
=
c o
e C
C
=
0 0
=
x
-r n
=
,o x
sc m
m m
e i
i M
i /
g ;
= w
- =
p
- a u n
~J u
j ee u o
= =.
uE f
N O au f
=
. e-u e,~y w
2-
= -U j
w ~3 U
(
m,<
w e f
. ~.
v-=
l
- c. O =
e uma i
H EY e J 2
ei
.a l
/*
3
- J 3
aW, a
y u
- u
w uax
's e
a
~
D 21 6
s nu u y
\\
u u u=
- V
<?-
n
<.uu u
u
. =.
i
=
.ss
\\l u -
=
e-a z
6
=3 a
ad-U e
a
=
=2
=:
- e.o
= wn u
u. -
=
=
=
U 1'
=
x u
y
~
N i
N N
u N
k m
s I
w t
t e
O C
C O
Q O
O C
C C
C C
~r o
e m
e.
n.
3 *dsal 3r7ppeta xrd i i i,,, i t
i 1
1
- o. m. x. r-c. e. 4. m
. c.
N
~ ~. -. ~
maura v,
. y,d
'A
. c, Babcock & Wilcox 3-2
O Figure 3-12.
Power, Flow and Flux Vs Ti=e for Densi-fied Fuel - Four-Pump Coas tdown 1.0 Flow Initial Core Avg Heat d.9 - Flu @ 102% Power =
193240 Btu /hr-ft2 1
0.8
-1.00
~~
,,~~s j-Power N
Core Avg g
-o,99 5
Heat Flux N
2
\\
g 0.7-N
-0.98 e
\\
Initial Power = 102% of 2772 YJt 3o
- 0. 9,e
-w l
0.6
\\
- 0.96 9
\\
d u
~
\\
M u
c
\\
- 0.95 C
3 x
x
\\
0.5
_ o.94 3
\\
- 0.93 0.4 -
0.92 0.91 0.3 -
- 0.90 i
i i
0 0.2 0.4 0.6 0.6 1.0 1.2
- 1. 4 1.0 1.0 2.0 Time, s
!af m <.J, Babcock & Wilcox 3-27
e,ar Figure 3-13.
DNBR and Centerline Fuel Temperature Vs Time for Four-Pump Coastdown 4600 2.0 4550 Cen t e.rline
- 1.95 Fuel Temp
- 1.90 4500 u.
E
- 1.85 4450 s
5z n.
E 4400 DNBR
- 1.80 s
i U
i,;
u 5v
- 1.75 4350 x
5 4300
- 1.70 4250 1.65 1.60 4200 0.0 0.5 1.0 1.5 2.0 2.5 3.0 Time After Loss of Primary Power, s Babcock & Wilcox 3-28 l> '
'y p
Figure 3-14 Pressure, Power, Flow, and Flux Vs Time for Locked Rotor Accident Initia1 Core avg Heat J1ux 3102% Power = 193240 Stu/ hr-g Core Avg Heat Flux 1.0 g
C Pcwcr
's e
- 2175
~
n s
g 5
o8
' N..
N i
' - s, 3
Flcw /
7165 w
/
3
.Cx 0.6 Pressure g
I_
- =
c 5
Initial Power =
-2155 i
102". o f 2772 MWt
.=
a 3
p
- t 0.4 s
3 5
~
- 2145 5m "J
L I
0.2 d
- 2135 x
~
m 0.0 O
1.0 2.0 3.0 4.0 5.0 Time, s t
O T -^ 4 < -
.v.4 9 Babcock & Wilcox 3-29
O Figure 3-15.
DNBR, Maximum Fuel Temperature, and Maximum Cladding Temperature Vs Time For Locked Rotor Accident 4700-1500 2.6 4600-1400 2.,
s 4500 130C
- 2.2 w
8
{120C Fue1 w
4400 d
U Temp w
c.
=
a a
u 3
4300 H 1100
- 1.8 E 8
5 5'
a Cladding Temp I
with Densification 5
4200 -
5 1000 Power Spike 1.6 E
w 3
w
=
~
M "3
m
- c 4100-m 900
- 1.4 u
M 4000 -
300
- 1.2 Maximum DNBR 3900 -
700
\\:
1.6 60C Time, s v
7 im _, h Babcock & Wilcox 3-30
4.
REFERENCES 1
B. J. Buescher and J. W.
regram, Babcock & Wilcox Model for Predicting In-Reactor Densification, BAW-10083P, Rev 1, Babcock & Wilcox, Lynchburg, Virginia, November 1976.
2 S. A. Varga (NRC) to J. E. Taylor (B&W), Letter, " Evaluation of BAW-10083P, Revision 1," May 16, 1977.
3 Three Mile Island Unit 2 Nuclear Station, Final Safety Analysis Report, USNRC Docket No. 50-320.
R. C. Jones, et al., ECCS Analysis of B&W's 177-TA Lowered-Loop SSS, BAW-10103, Rev 2, Babceck & Wilcox, Lynchburg, Virginia, April 1976.
5 R. A.
Turner, Fuel Densification Report, BAW-10054P, Rev 2, Pabcock &
Wilcox, Lynchburg, Virginia, June 1973.
6 R.
A.
Turner, Fuel Densification Report, BAW-10055, Babcock & Wilcox, Lynchburg, Virginia, June 1973.
7 K.
E.
Suhrke (B&W) ra S. A. Varga (USNRC), Letter, "Densification Power Spike," Dece=ber 6.
1976.
a R. H. Stoudt, et a l., TACO -- Fuel Performance Analysis, BAW-1008 7P, Rev 1, Babcock & Wilcox, Lynchburg, Virginia, May 1976.
9 Correlation of Critical Heat Flux in a Bundle Cooled by Pressuriced Water, BAW-10000A, Babcock 6 Wilcox, Lynchburg, Virginia, March 1970.
10 C. W.
Mays and M. Furtney, FLAME 3 - A Three-Dimensional Nodal Code for Calculating Core Reactivity and Power Distributions, B AW-10124 A, Babcock &
Wilcox, Lynchburg, Virginia, August 1976.
11 C. W. Mays, Verification of the Three-Dimensional FLAME Code, EAW-10125PA, Babcock & Wilcox, Lynchburg, Virginia, August 1976.
12 Rancho Saco Unit 1 Fuel Densification Report, BAW-1393, Babcock & Wilcox, Lynchburg, Virginia, June 1973.
7 Babcock & Wilcox 4-1
13 R. H. Stoudt and S. E. Busby, CADD - Cocputer Application to Direct Simula-tien of Transients in Water Reactors, BAW-10076?A, Rev 2, Babcock & Wilcox, Lynchburg, Virginia, December 1974.
C. D. Morgan, e t al., RADAR -- Code to Analyze Slow Reactor Transients in 1"
Water Reactors, BAW-10069A, Rev 1 Babcock & Wilcox, Lynchburg, Virginia, October 1974 A. F. J. Eckert, et al., Program to Determine In-Reactor Performance of 15 B&W Fuel Cladding Creep Collapse, B AW-10084P, Rev 1, Babcock & Wilcox, Lynchburg, Virginia, October 1976.
,r;r
,1 Babcock & Wilcox 4-2