ML20213E683
| ML20213E683 | |
| Person / Time | |
|---|---|
| Site: | Seabrook  |
| Issue date: | 11/30/1986 |
| From: | Kenton M, Lutz R, Plys M FAUSKE & ASSOCIATES, INC., WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML20213E656 | List: |
| References | |
| FAI-86-39-DRFT, NUDOCS 8611130256 | |
| Download: ML20213E683 (88) | |
Text
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~ 1 a. r as j SEABROOK STEAM GENERATOR INTEGRITY AN ALYSIS Martin G. Plys Marc A. Kenton Robert E. Henry Fauske & Associates, Inc.
Burr Ridge, Illinois Robert Lutz Peter Kirby Westinghouse Electric Cort, oration Pittsburgh, Pennsylvania November,1986 8611130256 PDR 861107 F ADOCK 05o00443 PDR
- - . - - - - . - - - - - -,c . - ..,,.. . - . ._..-
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TABLE OF CONTENTS Page
1.0 INTRODUCTION
. . . . . . . . . . . . . . . . . . . . . . . . . . 1-1 1.1 Background . . . . . . . . . . . . . . . . . . . . . . . . 1-1 1.2 Method . . . . . . . . . . . . . . . . . . . . . . . . . . 1-1 2.0 ACCIDENT SEQUENCES STUDIED . . . . . . . . . . . . . . . . . . . 2-1 2.1 Sequences Without Operator Actions . . . . . . . . . . . . 2-1 2.2 Sequence With Operator Action . . . . . . . . . . . . . . . 2-1 2.3 Uncertainty Analyses . . . . . . . . . . . . . . . . . . . 2-1 3.0 SEABROOK SPECIFIC INFORMATION , . . . . . . . . . . . . . . . . 3-1 3.1 NAAP 3.0 Parameter Fil e . . . . . . . . . . . . . . . . . . 3-1 3.2 Operator Actions . . . . . . . . . . . . . . . . . . . . . 3-1 4.0 RESULTS . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-1 4.1 No Ope ra to r Ac ti on s . . . . . . . . . . . . . . . . . . . . 4- 1 4.2 Operator Actions . . . . . . . . . . . . . . . . . . . . . 4-5 4.3 Influence of Uncertainties . . . . . . . . . . . . . . . . 4-13
5.0 CONCLUSION
S . . . . . . . . . . . . . . . . . . . . . . . . . . 5-1
6.0 REFERENCES
. . . . . . . . . . . . . . . . . . . . . . . . . . . 6-1 APPENDIX A: MAAP 3.0 Seabrook Parameter File . . . . . . . . . . . . . A-1 APPENDIX B: Steam Generator Tube Integrity Analysis . . . . . . . . . B-1 B.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . B-1 B.2 Tube Degradation . . . . . . . . . . . . . . . . . . . . . . B-1
\
TABLE OF CONTENTS (Continued)
Page B.2.1 Tube Properties . . . . . . . . . . . . . . . . . . . B-2 B.2.2 Thinning / Cracking Type Defects . . . . . . . . . . . . B-2 B.2.3 Tube Denting . . . . . . . . . . . . . . . . . . . . . B-5 B.3 C reep Ru ptu re . . . . . . . . . . . . . . . . . . . . . . . . B-9 B.4 Suma ry . . . . . . . . . . . . . . . . . . . . . . . . . . . B-12 B.5 References . . . . . . . . . . . . . . . . . . . . . . . . . B-14 APPENDIX C: Estimation of Steam Generator Tube Wall Temperatures . . . C-1 l
[ b
- iii -
1 .
LIST OF FIGURES Figure No. Page
\
1-1 Hot leg and steam generator natural circulation flow . . 1-2 4-1 Base case primary system pressure . . . . . . . . . . . . 4-3 4-2 Base case primary system structure temperatures . . . . . 4-4 4-3 Base case and seal LOCA case primary system CsI re-tention . . . . . . . . . . . . . . . . . . . . . . . . 4-6 4-4 Base case and PORV case primary system pressure com-parison . . . . . . . . . . . . . . . . . . . . . . . . 4-8 4-5 Base case and PORV case steam generator inlet plenum gas temperature comparison . . . . . . . . . . . . . . . 4-9 4-6 Base case and PORV case steam generator average tube temperature compari son . . . . . . . . . . . . . . . . . 4-10 1 4-7 Base case and PORV case in-vessel hydrogen production compa ri son . . . . . . . . . . . . . . . . . . . . . . . 4-11 4-8 Base case and PORV case primary system CsI retention . . 4-12 4-9 Base case and high eutectic temperature case hottest core node temperature comparison . . . . . . . . . . . . 4-14 4-10 Base case and high eutectic temperature case steam generator inlet plenum gas temperature comparison . . . . 4-15 4-11 Base case and natural circulation uncertainty cases steam generator inlet plenum gas temperature compa ri son . . . . . . . . . . . . . . . . . . . . . . . 4-17 4-12 Base case and natural circulation uncerta-nty 'ases steam generator tube temperatrire comparisons . . . . . . 4-18 4-13 Base case and steam generator PORV case primary sys-tem pressure comparison . . . . . . . . . . . . . . . . 4-21 4-14 Base case and steam generator PORV case pressurizer water level comparison . . . . . . . . . . . . . . . . . 4-2i 4-15 Base case and steam generator PORV case steani gen-erator inlet plenum gas temperature comparison . . . . . 4-23 4-16 No blockage case core-uoper plenum flow . . . . . . . . . 4-24
- iv -
LIST OF FIGURES Figure No. Page f
4-17 No blockage case upper-plenum gas temperature . . . . . 4-25 4-18 No blockage case steam generator plenum gas temperature . . . . . . . . . . . . . . . . . . . . . . 4-26 B-1 High temperature tensile properties of annealed (1600*F/1 hr.) hot-rolled plate (B-1) ..........B-3 B-2a Burst data for 0.875 x 0.050 unifonn thinning speci-mens - defect length variation (B-1) . . . . . . . . . . . B-4 B-2b Burst data for 0.875 x 0.050 uniform thinning speci-mens - defect depth variation (B-1) . . . . . . . . . . . B-4 B-3a Burst data for 0.875 x 0.050 EDM slot specimens -
defect depth variation (B-1 ) . . . . . . . . . . . . . . . B-6 B-3b Burst data for 0.875 x 0.050 EDM slot specimens -
defect length variation (B-1) . . . . . . . . . . . . . . B-6 B-4 Burst pressure cata of 0.875 x 0.050 uniform thinning specimens with and without denting (B-1) . . . . . . . . . B-7 B-5 Creep and creep-rupture comparisons (B-4) . . . . . . . . B-10 B-6 Master creep rupture curve for 316 stainless steel, taken from Ref. (B-7) . . . . . . . . . . . . . . . . . . B-13 C-1 Natural circulation flows on the inside and outside of a tube carrying fluid from the inlet to the out-let plenum . . . . . . . . . . . . . . . . . . . . . . . . C-2 1
-v-LIST OF TABLES Table No. Page 4-1 Blackout Base Case Key Event Times, Without and With Seal LOCA . . . . . . . . . . . . . . . . . . . . . . . . 4-2 4-2 Operator Action Figures of Merit and Event Times .... 4-7 4-3 Key Event Time Comparison for Steam Generator PORV Open Case . . . . . . . . . . . . . . . . . . . . . . . . 4-19 B-1 Comparison of Burst Pressures of Elliptically Wasted
.875 00 x .050 Wall Tubin (B-1) . . . . . . . . . .g With................
and Without Denting B-8 i
- B-2 Creep Rupture Data for Inconel-600, Hot Rolled, and Annealed at 1600 F ................... B-ll
vi -
1-1
1.0 INTRODUCTION
I
1.1 Background
A frequently-studied postulated severe LWR accident is the station blackout sequence (TNLB'). In this sequence, all off-site and on-site AC power are assumed to be lost. When analyses of this sequence are performed, high reactor vessel upper plenum temperatures are computed (1). This is caused by two factors, the heating of the upper plenum due to core-upper plenum natural circ.ulation, and that caused by volatile fission products (chiefly iodine isotopes) which are released from the core and deposit in the upper plenum.
l Experiments performed at Westinghouse (2_) in a one-seventh scale test facility have shown that these high upper plenum temperatures will cause natural circulation to initiate between the upper plenum and steam genera-tors. This process, which is shown schematically in Figure 1-1, would result in an increase in steam generator inlet planum temperatures. Conse-l quently, concern has been expressed that temperature-induced failures in the tt.bes or in the tube-to-tubesheet welds could result in a discharge of fission products to the secondary sides of the steam generators. Continued blowdown of the primary system to the secondary system through sur lures could lift the steam generator safety valves, bypassing the containment.
The principle purpose of this study was to assess the likelihood of these failures in the Seabrook plant, given a station blackout sequence.
1.2 Method Several variations on station blackout sequences were simulated using the Modular Accident Analysis Program (MAAP) version 3.0. This code is an integrated thermal-hydraulic and fission-product analysis code for severe accidents developed in the IDCOR program. Earlier version have been used by utilities, NSSS vendors, and consultants for approximately 4 years. The code has been verified and has been extensively benchmarked against experi-mental data, actual plant transients, and detailed code calculations (3_). .
1-2
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1-3 ,
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MAAP is particularly suited for this study since it contains fully integrated models for: l l
- a. Core overheating, oxidation and melting.
- b. Natural circulation between the core and upper plenum, between the upper plenum and inlet steam generator plena, and between the inlet and outlet plena of the steam generators. Calculaticas using the natural circulation models ha:' compared well to the Westinghouse one-seventh scale tests (1,2_).
- c. Fission oroduct release, transport, deposition, and revapori-zation.
In addition, MAAP allows arbitrary operator actions to be applied so that the efficacy of the likely operator responses can be studied.
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2-1 2.0 ACCIDENT SE0VENCES STUDIED In all the sequences presented below, it is assumed that off-site and on-site AC power remains lost indefinitely. In such a case, the emergency response procedures (ERPs) (4_) instruct the operators to initiate auxiliary feedwater (AFW) using the turbine-driven pump. As long as sufficient AFW flow can be maintained, decay heat would be removed and the accident would not progress further. For the cases studied here, it has been further assumed that all AFW has been lost. This reflects a major conservatism in the analysis, i
2.1 Sequences Without Operator Actions Two sequences were studied in which no operator actions were credited.
In the base case, no additional failures were assumed. In the second case, loss of cooling to the main coolant pump seals was assumed to result in a leak area corresponding to a flowrate of 50 gal / min of water per pump. This leak area is the same as that assumed in the IDCOR program (5,).
2.2 Sequence With Operator Action The Seabrook ERPs (6_,) specify that when AC power is available and all other measures have been attempted, if core thermocouple temperatures exceed 1200'F, the pri ary system should be depressurized using the pressurizer PORVs. This enables the accumulators to be used to inject coolant into the core. While this action is not invoked for blackout sequences, it was considered that this action might well be recommended by the Technical Support Center. Accordingly, this action was simulated in one case.
2.3 Uncertainty Analyses Several variations on the base case were studied to investigate the impact of uncertainties in phenomenological parameters and accident pro-gression. In these cases, parameters were varied in such a way as to l increase the potential for high steam generator temperatures:
2-2 l
- a. A high (3000*K) core melting temperature was assumed compared to the nominal value (2500*K). This delays the onset of core geome-try degradation and thus enhances the potential for the core to 4
heat the rest of the primary system.
- b. Low values of axial and cross-flow friction factors were assumed in the core modelling. This also tends to maximize convection of heat out of the core.
- c. Lower values of steam generator natural circulation flow W3g (see Figure 1-1) relative to hot leg natural circulation flow W HL were assumed to minimize cooling of the steam generator inlet plenum by flows from the outlet plenum. This was accomplished by choosing lower limit values of the number of steam generator tubes partic-ipating in the flow from the inlet to the outlet plena, guided by observations in the Westinghouse experiments (2_).
- d. A run was made where it was assumed that the steam generator PORVs stick open. Depressurizing the secondary side tends to decouple the tubes from the shell after the steam generator dries out.
This could potentially increase tube temperatures,
- e. A run was made in which it was not assumed that coolant channel blockage occurs in a core node when melting begins in that node.
MAAP ordinarily makes such an assumption to represent the rapid reduction in flow area that would occur after melting commences.
In this run, complete blockage of the flow area in a node was only assumed to occur if the node completely filled with molten ma-terial. By allowing continued core oxidation and core-upper plenum flow, this assumption maximizes the potential for the core to heat the rest of the primary system.
3-1 j 3.0 SEABROOK SPECIFIC INFORMATION ;
l i i 3.1 MAAP 3.0 Parameter File ;
Since MAAP 3.0 was used for this analysis, it was necessary to develop a Seabrook-specific parameter file. The major part of this effort was completed by using the MAAP 2.0 Seabrook parameter file which had been developed in support of the Emergency Planning Zone study (_7_).
Additional changes were made to the Seabrook MAAP 2.0 deck to reflect the model revisions in MAAP 3.0. The major additions and changes were in
! the areas of reactor vessel heat sinks, peaking factors, fuel rod ballooning data, model parameters, and auxiliary building data.
The new MAAP 3.0 data for this analysis was derived from Seabrook plant drawings, data deleted or modified from MAAP 2.0, the Westinghouse IMP data base, and other generic and Seabrook specific documents. A copy of the Seabrook MAAP 3.0 parameter file is included in the Appendix.
3.2 Operator Actions In order to get a complete picture of steam generator tube response during a station blackout transient, operator actions were also considered.
The potential operator actions were derived from the Seabrook ERPs (4,6).
In the event of a station blackout, the operators are instructed to initiate auxiliary feedwater (AFW) operation using the turbine-driven pump.
This action was not modelled in the MAAP 3.0 runs, since it would result in the removal of decay heat and would effectively prevent core uncovering from i
occurring.
In sequences with AC power available, the symptom-oriented procedures -
(i) call for depressurizing the primary system using the PORVs if core temperatures exceed 1200*F and all other means for cooling the core have
3-2 been attempted. While this action is not invoked in the procedures if AC power is not available, it was considered that such an action might well be recomended by the Technical Support Center in a blackout in order to obtain flow from the accumulators. For this reason, this action was simulated in one of the comparison cases.
4-1 4.0 RESULTS 4.1 No Operator Actions Two cases were run with no operator action, the base case and the base case with a pump seal LOCA at 45 minutes after initiation. Since the results of these sequences are quite similar, the base case will be de-scribed here and only differences will be noted. Major events are listed in Table 4-1. After the blackout is initiated, the steam generator inventory begins to be depleted. As the water level in the steam generators de-creases, heatup and expansion of the primary coolant leads to the pressur-izer " going solid", i.e., completely full of water. The quench tank rupture disk then breaks after discharge from the pressurizer overpressurizes the tank. The steam generators dry out completely soon afte:vards.
The seal LOCA case has slightly different timing until steam generator dryout because of the loss of primary system inventory through the failed seals. Thus, in this case, the pressurizer goes solid later because there is less water which can expand. Steam generator dryout, on the other hand, is determined simply by integrated decay power and the available secondary side inventory and differs little from the base case.
Loss of primary coolant leads to eventual core uncovery, meltdown, and vessel failure. The primary system remains at high pressure until vessel failure as dictated by the pressurizer relief valve setpoints (Figure 4-1).
Strong natural circulation occurs between the core and upper plenum after the core uncovers; this in turn sets up circulation between the upper plenum and the steam generators. Fission products leave the core during the heatup and can migrate through the upper plenum to the hot leg. The circulation and fission product transport are affected by temperature differences, and feed back to influence region temperatures (Figure 4-2).
After vessel failure, the primary system blows down, the pump bowls clear, and the accumulators empty. Most core debris and water in the lower plenum at that time are entrained into the lower compartment, which is
4-2 Table 4-1 BLACK 0UT BASE CASE XEY EVENT TIMES.
WITHOUT AND WITH SEAL LOCA Event Base Case Seal LOCA Initiation 0.0 0.0 Seal LOCA -
2714.
Pressurizer Solid 4829. 5920.
Quench Tank Disk Rupture 5486. 5944.
Steam Generator Dryout 5527. 5517.
Steam in Pressurizer 6319, 6269.
Core Uncovery 7280. 7124.
Vessel Failure 11648. 11054.
Accumulator Depleted 11700. 11102.
End of Simulation 20000. 20000.
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4-5 connected to the reactor cavity. Airborne fission products are swept from the primary system at that time. With the pump bowls clear, natural circu-lation can occur throughout the primary system, and some long-term revapor-ization of fission products is possible. In the case of the seal LOCA, slightly more Csl leaves the primary system due to leakage through the pump seals and the lower system pressure (Figure 4-3).
4.2 Operator Actions Operator action to depressurize the primary system by opening the pressurizer relief valves was assumed to occur when the maximum core temper-ature reached 1200*F, as discussed earlier. This resulted in complete blowdown of the primary system before bottom head failure. Thus, the time during which steam generator tubing was exposed to both high pressures and high temperatures was significantly reduced by operator action.
Key events and figures of merit are sumarized in Table 4-2 and com-pared to the base case. It can be seen that not only is the peak steam generator inlet plenum yas temperature lower after the operator action, but the primary system pressure is lower as well. The action occurs at 8000 seconds, about 15 minutes after core uncovery, and leads to accumulator dis-charge well before bottom head failure. This discharge occurs over a period of several thousand seconds, as seen by the gradual decrease in primary system pressure (Figure 4-4). The steam generator inlet plenum gas tempera-ture (Figure 4-5) is high only after depressurization, and the tubes them-selves are relatively cool (Figure 4-6). While, as shown in Figure 4-6, tube temperatures continue to increase after vessel failure, simulations continued for a longer time than those presented here show that the tempera-tures do not increase much beyond 700*K due to heat losses to containment.
Since the differential pressures are much lower after vessel failure, these moderate temperatures are not limiting.
The accumulator discharge causes only slightly more hydrogen production (Figure 4-7) through availability of steam. However, the large flows caused by steaming of accumulator water flush fission products from the primary system into the containment (Figure 4-8).
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A 4-7 Table 4-2 OPERATOR ACTION FIGURES OF MERIT AND EVENT TIMES Base PORV Seal Case 0 1200*F LOCA
- 1. Peak SG inlet plenum gas tem- 858. 682. 865.
perature while primary system is at high pressure (*K)
- 2. Primary system pressure at 17.1 1.83 14.9 bottom h(ad failure (MPa)
- 3. Peak SG inlet plenum gas 858. 843. 865.
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- 4. Primary system pressure at 17.1 4.23 16.9 time of #3 above (MPa)
- 5. Time of peak T #3 (sec) 9230. 9290. 9000,
- 6. PORVopen(sec) --
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- 7. Time accumulators depleted (sec) 11701. 10643. 11102.
- 8. Vessel failure (sec) 11648. 12601. 11054.
- 9. Approximate peak SG tube wall 750 low 760 temperature (see Appendix C)
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- Figure 4-8 Base case and PORV case primary system Csl retention.
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4-13 4.3 Influence of Uncertainties Five sensitivity cases are presented here which show the influence of high core melting temperatures, enhanced core to upper plenum natural circulation, reduced steam generator tube circulation, blowdown of the steam generator secondary sides, and neglect of core blockage phenomena. Each of those factors individually can increase hot leg and steem generator plenum temperatures relative to nominal cases.
A high UO2 -Zr eutectic temperature of 3000*K (versus the nominal 2500*O was used for uncertainty analysis of the base case presented above.
In the base case, vessel failure occurred at about 11,600 seconds (Table 4-2), while in this case vessel failure was delayed until 12.300 seconds.
This is because more time is required to reach the higher eutectic tempera-ture and cause melting. Other thermal-hydraulic behavior and event ti ling is similar to the base case, with exceptions relating to higher core temper-atures. The driving potential for natural convection can be compared between this and the base case by considering the hottest core node tempera-ture (Figure 4-9), which shows the effect of the input eutectic temperature.
Only a small impact on the steam generator inlet plenum gas temperature is caused by this parameter change (Figure 4-10), with a slight delay in reaching the peak values. Thus, considering an uncertainty of 500*K in the core melting temperature, a difference of only tens of degrees in steam generator temperatures results. This is because the heatup rates in the core are very much more rapid than the heatup of the steam generator plena onme temperatures are high enough to cause rapid Zircaloy oxidation.
Natural circulation between the core and upper plenum was enhanced by lowering the axial and cross-flow core friction factors in one sequence.
Circulation between the steam generator tubes and the inlet plenum was reduced by lowering the fraction of tubes carrying flow out of the plenum in another case. Each case acts to increase the steam generator inlet plenum temperature over the base case value. There is a small effect of these changes on sequence timing: vessel failure occurs at 11600 seconds in the base case, 12100 seconds in the high core circulation case, and 11200 seconds in the low steam generator circulation case. In the high core
l
! SEABROOK UNCFRTAINIY COMPARISON m , ON AHI TDJ I
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i Figure 4-9 Base case and high eutectic temperature case hottest core node temperature
- comparison.
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] Figure 4-10 Base case and high eutectic temperature case steam generator
] inlet plenum gas temperature comparison, i
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1
4-16 circulation case, more heat transfer occurs to the steam generators and vessel failure is slightly delayed. In the low steam generator circulation case, less heat transfer from the plenum to the tubes occurs, and vessel ,
failure is slightly hastened. The steam generator inlet plenum gas tempera-ture (Figure 4-11) is significantly higher (about 150*K) only for the steam generator circulation case ("SG NC"), while it is only slightly higher for the core circulation case (" CORE NC"). Tube temperatures (Figure 4-12) follow the same trend. It should be noted that the low steam generator circulation case, which assumed that only 10 percent of the tubes carried flow in the "out" direction, is in all likelihood unrealistically severe.
More typical extremal values of the flow split observed in the Westinghouse experiments (about 20 percent of the tubes carrying out flow) result in only l about a 900*K peak inlet plenum gas temperature.
A blackout case with stuck open steam generator relief valves on all the units was run to minimize the heat transfer capability of the tubes, and l thus increase the steam generator inlet plenum temperature. In the base case, steam on the secondary side serves as an efficient heat sink for the l small recirculation flow through the primary side, and also couples the tube l mass with the steam generator shell heat sink. When the secondary side is blown down, and this heat capacitance is impaired, the tube outlet tempera-ture is higher and thus the mixed mean steam generator inlet plenum tempera-ture is higher.
The early behavior of this sequence differs dramatically from that of the base case due to enhanced cooling by the secondary side early in the transient. Comparing events (Table 4-3), the steam generators go dry much earlier in the steam generator PORV case, and the primary system cools down enough that the pressurizer is drained due to contraction of the coolant.
Later, of course, with the heat sink lost, primary system fluid reexpands and the pressurizer goes solid. Core uncovery and vessel failure occur slightly earlier in the steam generator PORV case because the overall integrated heat removal is lower with the steam generators blown down.
(
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SFABROOK tlNCFRTAINTY COMPARISON OBASE A CGIE NC ySC NC a.... .ig. .... iig iiiiiiiiiiiiiiiiiig . iiigiiiiiiiiig .
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- generator inlet plenum gas temperature comparison.
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SEABROOK UNCFRTAINTY COMPARISON e onaSE A CORE E USC NC g ; i . .. gi m i . .. iiiiii iiiiiii m o ii g . . . . . g ii iiii iiii iiii iiii g i s i iiii iiii iiii i . . i g iiiiiii n t- : :
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Figure 4-12 Base case and natural circulation uncertainty cases steam generator tube temperature comparisons.
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4-19 i
Table 4-3 KEY EVENT TIME COMPARISON FOR STEAM GENERATOR PORY OPEN CASE Event SG PORV Base Case p
Initiation 0.0 0.0 SG PORV Open --
0.0 SG Dryout 5527. 2738.
Pressurizer Drained --
350.
Pressurizer Refilling --
2000.
Pressurizer Solid 4829. 4205.
Quench Tank Disk Rupture 5486. 4937.
. Pressurizer Has Steam 6319. 5920.
Pressurizer Empty 7000. 7400.
Core Uncovery 7280. 6626.
Vessel Failure 11648. 11045.
I
4-20 Primary system pressure (Figure 4-13) shows the cooling achieved while the secondary side blows down, resulting in a 10 MPa transient over 4000 seconds. The pressurizer water level (Figure 4-14) illustrates the con-traction of primary system coolant during the cooldown, and later reexpan-sion. Since the core uncovers earlier in the steam generator PORV open case, high gas temperatures in the steam genera'or inlet plenum are shifted to slightly earlier times (Figure 4-15). The initial high peak, prior to degradation of core geometry, is the same in each case both in magnitude and rise behavior. Thereafter and before vessel failure, the inlet plenum gas temperature is about 50*K higher for the steam generator PORV case. How- '
ever, it is still below 800*K for most of the high temperature period.
The last uncertainty case assumed that nodal core flow channel blockage does not occur at the onset of melting in the node. MAAP normally assumes a
such blockages to represent the rapid reduction in flow area that would occur after melting begins. In this case, complete blockage was credited only when the geometry would allow no flow, i .e. , when the node was com-pletely full of refrozen eutectic. As shown in Figure 4-16, a rapid reduc-tion in core-upper plenum flow still occurs after the beginning of accel-ersted oxidation. While the flow continues at relatively low values, heat removal from the upper plenum due to the hot leg flows about equals heat convected frcm the core and that due to fission product heating, resulting in a stabilization in upper plenum gas temperature (Figure 4-17). Steam generator inlet plenum gas temperatures (Figure 4-18) also stabilize. The peak sustained plenum gas temperature, which occurs just before vessel failure (12,000 secs) is about 1060*K. Thus, even when blockage in the core is essentially neglected, only relatively moderate increases in plenum gas temperature are seen over the base case.
. _ . ,___..___m..___,.._ _ , _ , _ _ , _ _ . , - , . . . _ . _ _ _ _ _ - . , , _ _ , _.,-__ -
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- SEABROOK THLB COMPARISON
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e l SEFER00K TMLB COMPARISON O cl ASC N I O l''''l''''P'''l''"I 'l''''l''''l''''l''''l''''-
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Figure 4-15 Base case and steam generator PORV case steam generator i inlet plenum gas temperature comparison.
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5-1
5.0 CONCLUSION
S Various blackout sequences have been analyzed for Seabrook which cover a range of potential operator actions and phenomenological uncertainties.
An important result of the base case is that for the period when the primary system is at high pressure, the peak steam generator inlet plenum gas temperature is only s 850'K, and average tube temperatures are well below 700'K. Operator action to depressurize the system can successfully reduce the system pressure prior to the time high temperatures are reached in the steam generator. When uncertainties in circulation between the core and upper plenum or in the core melting temperature is considered, negligible increases in the steam generator temperatures are observed. Similarly, only slight increases in temperature were obtained in a case where the secondary sides of the steam generators were depressurized. For an unmitigated case, considering a large variation in the steam generator inter-plenum circula-tion indicated that gas temperatures can briefly reach about 1000'K in the steam generator plenum. Similar temperatures (s 1060*K) are achieved when the blockage of core node coolant channels due to melting is neglected.
i Doth of these latter cases are considered unrealistically extreme.
s MAAP does not contain a detailed model for the change in tube wall temperature as one leaves the inlet plenum and moves toward the outlet 4
plenum of the steam generator (i.e., only average wall temperatures are computed). However, by using the average secondary side gas temperature and the temperature of the primary side gas entering the tubes, the wall temper-atures can be estimated by knowing the value of the heat transfer coeffi-cients on the primary and secondary sides. As shown in Appendix C, these heat transfer coefficients are approximately equal. By equating the heat flux convected to the tubes by the primary side gas to that convected away from the tubes by the secondary side gas, the analysis in Appendix C leads to the conclusion that the peak tube wall temperature will be the average of the inlet gas temperature and the average secondary side temperature. In the worst case discussed above, this results in a peak tube wall temperature of approximately 850'K and is only 750*K in the best estimate case. As
5-2 shown in Appendix B, these are considerably less than the temperature values which would challenge the integrity of the tubes. Therefore, steam gener-ator tube rupture is judged to be very improbable for the sequances ex-amined.
I I
l j
{
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l
6-1
6.0 REFERENCES
(1) Fauske & Associates, Inc., " Technical Support for Issue Resolution",
IDCOR Report 85-2 (July,1985).
(2) W. A. Stewart, et al., " Experiments on Natural Circulation Flows in Steam Generators During Severe Accidents", Proc. Inter. ANS/ ENS Topical Meeting on Thermal Reactor Safety, San Diego, California (1986).
(3_) R. E. Henry, " Benchmarking of Severe Accident Codes: tiow Should It be Done and How Should it be Used?", Proc. Inter. ANS/ ENC Topical Meeting on Thermal Reactor Safety, San Diego, California (1986).
(4) New Hampshire Yankee Station Emergency Operating Procedure ECA-0.0,
" Loss of All AC Power" (May 16, 1986).
(5) Coninonwealth Edison Company. " Zion Station Integrated . Containment
- Analysis", IOCOR Report 23.12 (1985).
(6,) New Hampshire Yankee Station Emergency Operating Procedure FR-C.1,
" Response to Inadequate Core Cooling", (May 16,1986).
( 7_) Pickard, Lowe, and Garrick, Inc., "Seabrook Station Risk Management and Emergency Planning Study", PLG-0432 (December,1985).
i i
A-1 APPENDIX A MAAP 3.0 Seabrook Parameter File
i A-2 l l
- MAAP WHICH
.08 ARE MARKED WITHHAVE 88888M EN ADDED OR REDEFINED SINCE THE ISSUANLE OF
- 88888888888888888888888888888888888888888888888888888888888888888888 st
- GENERAL INFORMATION:
88
!!1. FDR THE GEDMETRICAL RELATIONSHIP BE1 WEEN VARIOUS CONTAINMENT AND FRIMA
- SYSTEM NUDES: SEE THE REGION SUBROUIINE WRITE-UPS IN VUL 2 0F USER'S
- MANUAL
- FOR LARGE, DRY CONTAINMENTS, SPECIFY ZERO VULUME FOR ICE-CONDENSER AND
- UPPER PLENUM-fHIS CAUSES ALL OfHtR UPPER PLENUM AND ICE CONDENStR
- PARAMETERS TO BE IGNORED 882. NOTE OUTER WALLS IN COMPARTMENIS A AND D SEPARATE THE CONI AINMEN1 FROM
- THE ENVINUNMENTI THE UUIER WALL IN COMPI B SEPARATES CUMPTS B AND Di
- THE OUTER WALLS IN COMPTS I AND U (ICE CONDENSENS ONLY) AkE NOT MUDELED
- SINCE THESE WALLS ARE INSULAIEDI THE UU1ER WALL IN COMPT C IS ASSuntD
- THE INNtR FACE) 88 TO REPRESENT A FREE STANDING STEEL CONTAINMENr WITH A SHIELD BUILDING, 88 TREAT IHE SHIELD BUILDING AS THE WALL AND THE CONfMT PRESSURE BOUNDARY
- PROPERTIES (THERMAL CONDUC ETC.) 0F IN1ERIOR WALLS IN A AND B ARE
- ASSUMED THE SAME AND ARE ENTERED IN THE LOWER COMPT SECTION
- IF (AS IS USUALLY IHE CASE) YOU MUST LUMP WALLS OF StVERAL THICKNESSES
- TOGETHER. YOU SHOULD LUMP DNLY RELATIVELY THILA WALLS (EG OREATER THAN
- ABUUT 1 FOOT OR .3 MEftR IN IHICKNESS) AND ENIER IHE 1HICKNtSS
- OF THE THINNEST WALL CREDITED 84
- 4. DECK REFERS TO THE FLOOR (AND VERTICAL WALLS IN ICE CONDENSER PLANIS)
- IHAT SEPTRA 1ES THE UPPER CUMPARTMENT FROM IHE COMPARTMENTS LOWER IN
- THE CONTAINMENT
- 5. TWO WAYS TO HANDLE CONTAINMENI FAILUkE:
- A. MECHANISITIC MUDEL:
88 ENTER 0 FOR THE FAILURE PRESSURE (ACOMPT NO. 36) ALSO SUPPLY:
- (1) CUNCREl'E: SUPPLY ALL THE MATERIAL DATA: CUNCRElE PARAMS 13-22 ETC.
48 (2) FREE STANDING STEEL SHELL: ENTEN THE WALL THICKNESS IN THE UtFER
- AND ANNULAM COMPARIMENTS IN THE ' LINER' THICKNESS EN1 RIES
- AND SUPPLY ONLY THE LINER MATERIAL FRUFENTIESI THE NUMrER OF
- IENDONSs AND AMOUNT UF REBAR SHOULD BE SLT TO 4ERO IN THIS CASE
- B. SIMPLE MODEL:
- SUPPLY ACOMPT NO. 36 AND 3/4 FAILURE AREA ENTERED AS MUDEL PARAMAfER 88 NO.28 MEED NOT SUPPLY IHE OTHER PARAMtTERS 886. SEDIMENTATION AREA
- IS THE TOTAL UFWARD-FACING AREA IN A
- GIVEN COMPARTMENT UPON WHICH FISSION PHODUCT AER05ULS CAN SETTLEI THIS 88 SHOULD INCLUDE (WHERE APPROPRIATE): FLOORS CABLE THAYbe EQUIFMENI ETC 88
- 7. AS DESCRIBED IN THE 8 CON 1ROL SECTION: THE AUXILIARY BUILDING MUDELS ARE 88 ACTIVA 1ED BY buPPLYING A NONZERO NO. UF AUX. NUUtS TO BE MODELLED.
- THE MODEL CAN BE RUN SIMULTAMEOUS WITH A RUN OF THE CONIMI AND FAINARY
A-3
- SYSTEM MODELSe GR, BY SUPPLYING A NUNZERO INPUT FILE NO., THE
$8 AUX MODELS ONLY CAN BE RUN USING AN INPUT FILE OF T/H DATA FROM AN 88 EARLIER MAAP RUN.
!!8.FISSIONPRODUCTREMOVALBYINERTIALIMPACTIONISMODELLEDONLYIN
$$ UNE CONIAINMENT COMPARTMENT. IN LARGEe DRY'S SUCH PARAME1ERS 88 SHOULD CHARACTERIZE GRATES WHICH ARE ASSUMED TO BE IN THE ANNULAR st COMPARTMENT. IF MORE IHAN UNE LEVEL OF GRA1ES EXISISe SUPPLY IHE TOTAL ss IMPACTION AREA 0F ALL THE GRATESe AND THE MAXIMUM FLOW AREA AT ANY
- OF THE GRA1E ELEVATIONS 88 IN ICE CONDENSER PLANIS, THESE PARAMETERS (EVkN THOUGH LOCATED
- IN THE 8 ANNULAR COMPARIMENT DATA SECTION) SHOULD REFLECT IMPACTION AND
$8 FLOW AREAS AND STHAP WIDTHS IN THE ICE BOX--SEE EG POS1MA'S REPORI 489. THE UNITS FOR PARAMETER INPUIS ARE SPECIFIED BY EITHER A 851 (METRIL) 88 OR *BR (BRITISH) UNITS CARD. ALL PARAMETERS FULLOWING SUCH A CARD
$$ ARE ASSUMED TO HAVE THESE UNITS UNTIL THE NEXT UNITS CARD IS INLLUDED.
- THUS A PARAMEftR FILE CAN HAVE SECTIONS WITH DIFPERENT UNITSe IF 88 DESIRED. THE LAST UNITS CARD IN A PARAMETER FILE CONTROLS THE UNITS
- UF UTHER PROGRAM INPUTS IN TAPE S (EG START AND FINAL IIMES ETC.) AND
- THE UNITS TO BE OUTPUT IN THE DU1PUT FILE AND PLOT FILES.
88 MkTRIC UNITS ARE M-KG-SEC-DEGREE KELVIN-PASCALS-M883/SECeETC.
48 BRITISH UNITS ARE FEEL-LBM-HOURS-DEGREE F-PSI-GPM 88 EXAMPLESI 88 IN METRIC UNITS, FLOWRATES SPECIFIED TO BE VOLUMEIR1C SHOULD BE
- M883/SECl OfMER FLOWRATES IE ALL THOSE NOT EXPLICIILY STATED TO BE
- VOLUMETRIC SHOULD BE KG/SECl HEADS SHOULD BE IN Mi PRESSUMES IN pal
- IN ENULISH 1HE UN11S ARE RESPECTIVELY UPMeLBM/HRe>Te PSIA--
- NOTE TO MAAP/BWR USERS--GPM IS USED IN MAAP/PWR INSTEAD OF FT883/HR
- IHE UNLY EXCEPT!UN TO THIS PRUCtDURE IS THAT THE TIME S1EPS ARE ENTERED
- IN THE 8 TIMING SECTION ALWAYS IN SECONDS 8810.IN LARGEe DRY CONTAINMENIS ' FANS' REFERS TO FAN COOLERS WHILH TAAE
- PLAMIS AME USED TO CHARACIERIZE IHE AIR RETURN FANS.
88 4888*****888888888888888888888888888888888888888888888888888888888888884 85 88R 55 ss8:3888888888888888888888888888888888848888*sts******888888888888888888
$ UPPER COMPARTMENT (OR 'A' COMPT)
- 8888488888888888888888888888888888888888888888888488848*88888488884488 01 2.138D6 FREE VULUME 02 I415. AREA 0F REFUELING POOL
. 03 141. HEIGHT OF CUNTAIMMENT SPRAY HEAD ABUVE B0110M OF COMFARTMtNT
! 04 3437. FLOW AREA FROM UPPER COMPARTMEN1 INIO ANNULAR COMPT 05 15394 CFMAACTERISTIC CROSS-SEC AREA 0F COMPT FOR BURN TIME
- CALCS--EG THE BURN TIME IS THE SOUAkt ROOT OF THIS 48 AREA DIVIDED BY THE BURN VELOCIlY 06 0.0 CURB HEIGHT IN REFUELING POOL TO ALLOW OVERFLOW--NORMALLY 48 0 UNLESS YOU ASSUME REFUELING POUL DRAINS ARE BLOCKED (A
- CLASSICAL ICE CONDENSER SEQUENCE), THEN MAKE If LARGE 07 72122. SURFACE AREA 0F UUTER WALLS IN UPPER COMPARIMENT 08 .0357 LINER THICKNESS ON OUTER WALL
, 09 0.28 UUTER WALL LINER UAP RESISTANCE--SEE NOIE IN SLOWER COMPT i ** FOR HOW TO MODEL FhEE STANDING STELL CONTMTS WITH A SHIELD i
I
l A-4 1
11 WALL 10 4.1 OUTER WALL TOTAL THICKNESS 11 0.92 IHERMAL CONDUCTIVITY OF QUTER BALL (FOR CONCNt1E SfRUCTURES
- WITH A LIERe THIS REFERS TO THE CONCRETE PANT) 12 0.157 SPECIFIC HEAT OF UUTER WALL l 13 144 DENSITY OF OUTER WALL 14 0 ENTER A 1 IF THE OUIER WALL IS SOLID STEEL (IE A SIEEL CONTM1
- WITH NO SHIELD BUILDING)e 0 FOR CONCREIE WIIH OR W/0 LINER 15 11116. HALF AREA (WALLS MODELED AS 1-D SLASS) 0F INIERNAL WALLS i
16 0 LINER 1HICKNESS UN INTERIOR WALLe IF ANY 17 0 !
LINER GAP RESISTANCE IN IN1ER10R WALL '
18 4. IHICKNESS OF INIERNAL WALLS 19 13990. DECK AREA 20 0 LINtR THICKNESS ON DECK 21 0 LINER GAP RESISTANCE ON DECK 22 4.2 DECK THICKNESS 23 0.92 THERMAL CONDUCTIVITY OF DECK 24 0.157 SPECIFIC HEAT OF DECK 25 144. DENSITY OF DECK 26 0 ENTER A 1 IF THE DECK IS SOLID STEEL 0 FOR CONCRE1E 27 3.434D6 METAL E0PT MASS
' 28 2.0603D5 E0PT HEAT TRANSFER AREA 29 0. NUMBER OF IGNITION $00RCES IN UPPER COMPT ( A COMPARTMEN.')
30 0. AVERAGE DISTANCE OF THESE FNOM THE CEILING OF A
- FULLOWING PARAME1ERS ARE UStD TO DETERMINE WHICH IGNI1ERS OR IGNITION
- THEIR RESPECTIVE CUMPARTMtNTS WHICH CAN THEN PROPAGA1E INTO INE UPPER 88 COMPARTMENT--IF NO IGNITERS IGh0RE 31-33 31 0. NO. OF IGNI1ERS/IGN SOURCES IN B WHICH CAN BE STEN FROM A 32 0. NO. OF IGNITERS /IGN SOURCES IN D WHICH CAN DE SEEN FROM A 33 0. DISTANCE FROM THE 10P OF A 'fD THE DECK 34 .90 FRACTION OF UFFER COMPT SPRAY WATER THAT RUNS IN10 THE
- REFUELING POUL (VS. CUNTINUING ON DIRECTLY INTO LOWER COMPT) 35 .35 FRACTION OF WATER DRAINING OU1 0F REFUELING FOOL 1HA1 88 GOES INTU LOWER COMPT (REHAINING FRACTIUN RUNS INTO THE
- CAVITY) 48!NPUTS FOR SIMPLE (FAILURE PRESSURE SUPPLIED) OR DETAILED (LONIMI STkAINS
- CALCULA1ED) MUDELS FOR CONTAIMtNT FAILURE--SEE GENERAL NOTES ABOVE 36 187. FAILURE PRESSURE OF CONTAINMENT OR 0 TO USE DETAILED MODEL 37 0 ENTER A 1 IF CUNfMT FAILS IN UPPER COMPfl0 FOR 88 FAILURE IN THE ANNULAR COMPT (USED ONLY FOR THE SIMPLE MUDEL)
- 888**888888884***88NEW8888888888848:
38 64.4 CONTAINMENT RADIUS FOR STRESS CALCULATIONS 39 1 67D-4 EUUIVALENT AREA TO CALCULA1E CONTAINMENT NORMAL LEAKAGE--
- NORMAL LEAKAGE IS ASSUMt0 TO COME FROM THE ANNULAR COMPT 40 0.D0 MASS OF WA1ER IN NEU1RUN SHIELD BAGS--WHEN BAGS RUPluRE
- THEY DROP THEIR CONIENTS int 0 REFUELING FOOL 41 10263. StDIMENTATION AREA FOR FISSION PRODUCT SETTLING
- THE REST OF THESE ARE NEW 8488 THESE ARE ONLY REQUIRED IF THE DETAILED CONTAINMENT FAILUKE MUDEL Ib USED.
- 888888888888888***NEW8888888884888 42 425. NUMBER OF TENDONS IN HOOP DINECTION IN THE LENGTH OF WALL
- GIVEN IN ITtM 43 43 .04590 VOLUME OF REPAR PER UNIT AREA 0F OUTER WALL (EQUIV THICKNEbS)
- RUNNING IN THE HOOP DIRECTION 44 .0512 VOLUME OF REBAR PER UNIT AREA 0F OUIER WALL (EQUIV THICKNESS)
- NUNNING IN IHE Z DIRECTION 45 .1969 DIAMETER OF HOOP TENDONS 46 164. HEIGHT OF fHE CYLINDNICAL PART OF THE CONTAINMENT WALL ABOVE
- THAT PART OF THE WALL REPRESEN1ED IN DCOMPT ITtM NO. 5
i 1
A-5 I
88 (EG APPROX IHAT ABOVE THE OPERATING DECK) 47 16.4 HEISHT OF INTERNAL WALLS 48 .984 DISPLACEMENT IN AXIAL DIRECTION WHICH IS SUFFICIENT TO TEAR i
48 THE CONTM1 WALL (EG Al A PENETRATION) 49 .984 SAME AS 48 FOR INE RADIAL DIRECTION 488888888888888888888888888888888888888888888888888888888888888888888888 SLOWER COMPARTMENT (OR 'B' COMPT) 88888888888888888888888***8888888888888888888888888888888888888888888888 01 45. DISTANCE FRUM FLOOR TO 10P OF 8 C0hPARTMENT 02 3375. AREA 0F CORIUM FOOLI THIS MUST BE LLb5 IHAN THE AREA Of
- THE FLOUR (ENBERED BELOW) 03 2.162 HEIGHT OF CURB ON FLOOR (OVER WHICH WATER OVERFLOWS TO C) 04 1840 CHARAC. CROSS-SEC AREA 0F LOWER COMPT FOR BURN TIME CALCS 05 2.523E5 FREE VOLUME 06 40. VERTICAL DISTANCE FROM THE CAVITY BYPASS FLOW AkEA 88 (EG AREA ARUUND VEShEL NUZZLES BUT SEE DEFINITION 88 IN CAVITY SECTION BELOW) TO THE CENTEN OF THE CAVITY END 88 UF fHE IUNNtl FLOW AREA 07 25. DISTANCE FROM THE FLOOR OF A TO THE OPENING FROM B IN10 D 08 0.D0 FOR CASES WHERE l'HE OUIER BOUNDARY OF CONfMT IS A
$8 STEEL SHELL SEFERATED FROM A CONCRETE SHIELD WALLS 48 ENTER DISTANCE BElWEEN THE TWO AND IREAT 1HE STELL
- $ SHELL AS A LINEN (ACOMPT AND DCOMPT 001ER WALLS)--
- EN1ER 0 OfHERWISE 880 UTER WALL OF B DIVIDES IT FROM COMPT D 09 1125 0 AREA 0F OuitR WALL 10 0.0 OUftR WALL LINER THICKNESS 11 0.0 GAP RESISTANCE OF OUTtR WALL LINER 12 4. THICKNESS OF OUTER WALL 13 0.92 THERMAL CONDUCTIVITY OF OUTER WALL 14 0.157 SPECIFIC HEAT OF OU1EN WALL 15 144. DENS!fY OF OUIER WALL 16 0 ENIER 1 IF THE OUIER WALL IS SOLID STEEle 0 FOR CONCREIE 88NUIE THAL CORIUM IN B IS ASSUMED TO SEE ONLY ONE FACE OF THE INFERIOR 48 WALL FOR RADIATION CALCULATIONS 17 16290. HALF SURFACE AREA 0F INTERIOR WALL 18 0 INTERIOR WALL LINER THICKNESS 19 0 OAP RESISTANCE OF BUILDING INftRIOR WALL LINER 20 4.0 THICKNESS OF IN1ERIOR WALLS 21 0.92 IHtRMAL CONDUCTIVIfY OF INTERIOR WALLS 22 0.157 SPECIFIC HEAT OF INIERIOR WALLS 23 144. OtNSITY OF IN1ERIOR WALLS 24 6751. AREA 0F FLOOR (USE WATEN FOOL AREA IF LESS) 25 O. FLOOR LINER IHICKNESS 26 0. GAP RESISTANCE OF FLOOR LINEN 27 4.0 IHICKNESS OF FLOOR 28 0.92 THERMAL CONDUCTIVITY OF FLOON 29 0.137 SPECIFIC HEAT OF FLOOR 30 144. DENSITY OF FLOOR 31 1.26E5 MASS OF EQUIPMENT--THIS REFERS TO EUPT IN1ENNAL TO THIS
$$ REGIONI
- THE FRIMARY SYSTEM MASS SHOULD N01 BE INCLUDED SINCE IT 88 HAS A SPECIFIC IREAIMENT tLSEWHtRE 32 21422. HEAT TRANSFER AREA 0F EGPT 880UANT!IY J3 IS USED FOR ALL EXTERNAL WALLS 33 8.8 HEAT TNANSFEN COEFFICIENT TO BE USED ON THL OUIER SUhFALE
- OF lHE CONTAINMENI UUIER WALLS (EG IN A AND D) 8834 NOT USED
A-6 35 0.D0 FRACTIONAL AREA AVAILABLE FOR REVERSE FLOW ON B-1 FLOWFATH ss COMPARtD 10 THE FORWAND DIRECTION (E0 DOE TO ICE
- CONDENSER DOOR (S) SHUTTING)--THIS NO. MUST l
- s BE NUNZER0 AND POSITIVE IN ICE CONDtNSER PLANTS--IGNORED IN as LARGE, DRY CONINTS 36 1.00 FRACTIONAL AREA AVAILABLE FOR REVERSE FLOW ON A-D FLOWFATH ss (E0 AIR RtfuRN FAN FLOW DAMPERS IN ILE CONDENSLRS) l ss ENTER 1 IF NO DAMPER 37 384. FLOW AREA FROM B INTO D 38 1092. FLOW AREA FkOM B TO A 39 0. NUMBtR OF IUNITERS/ IGNITION SOURCES IN B 40 0. AVG DISTANCE OF THESE FROM THE CEILING OF B 41 27. HtIGHT OF FLOOR OF B ABOVE FLOOR OF C 42 4500.- SEDIMENTATION AREA sessssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssst
- CAVITY (CCOMPT) assasssssssssss**ss***sssssssssssssssssssse 88888888ssssssssssssssssssst aslHE CAVITY INCLUDES ALL THE VOLUME bELOW THE REATOR N0ZZLES INSIDE asTHE BIOLOGICAL SHIELD AND ALL THE VOL OUT TO WHkRE lHL TUNNEL SLOPES UF ss ssNOTE THAT THE CAVITY HAS TWO FLOWPATHS- ' TUNNEL' REFERS TO A WATER ssAND PERHAFS CORIUM FLOW PA1H THAT ENIERS NEAR THE BASE OF tHE CAVIIYi ss' BYPASS' REFERS TO A FLOWPATH HIGHER IN THE CAVITYI THIS CUULD BE THE ssAREA AROUND THE RV N022LES, OR IN fHE CASE OF SOME PLANTSe BLOWOUT PANELS
- HIGHER IN THE CAVITf--1HL BYFASS AREA IS ASSUMLD TO EMPTY INf0 B ss as!N SOME PLANTS WATER CAN FLOW DOWN FNOM THE RESUELING FOOL TO THE CAVITY ssAND IN SOMEe CURIUM CAN BE EN1 RAINED UP TO THE W PER COMPARTMENT AROUND asTHE RV ANNULUS--AT FRESENT GAS IS NOT EXCHANGED BETWEEN C AND A HOWEVER ss asIN MANY SEQUENCESe NAT. CIRC. IS SET UP WHEREBY COLD GAS ENTERS THE
- sCAVITY IHROUGH THE BUNNELe IS HEALED BY PASSING OVER CORIUMe AND LEAVES asTHROUGH THE BYFASS AREA 01 7.74 BYFASS (NON-TUNNEL) FLOW AREA COUPLING CAVITY TO LOWER /UFFEN 85 CUMPARTMENTSI THIS SHOULD BE 1HE LIMITING FLOW AREAe EG ss THE AREA AROUND THE N0ZZLES AS idtY FENETAATE 1HL BIOLOGICAL
- SHIELD OR THE ANNULAR FLOW AREA BETWEEN THE RV AND THE SHIELD 02 476.4 AREA 0F CAVITY FOOL--THIS INCLUDES KEYWAY EIC WHERE AFFLIC 03 233.0 CHARAC. CROSS-SEC AREA 0F COMPT FOR BURN TIME CALCULATION 04 17. HEIGHT OF VtSSEL ABOVE BOTTOM OF CAVITY 05 158.2 (UNNEL CROSS-SECTNL AREA 06 253 LARGEST CHARAC CROSS-SECTNL AREA 1 HAT CORIUM MUST ss TFAVERSED UN IIS WAY TO THE OPENING WHERE IT MAY BE as ENIRAINLD OR FLOODED TO COMPTS A OR B--IN FLAN 15 WITH
- bO1 TOM HEAD PiNtlRATIONSe (HIS WILL TYPICALLY BE THE as 'ktYWAY' AREA (THIS IS USED TO CALCULATE THE MINIMUM
- VELOC1fY WHICH CAN LNIRAIN IHE CORIUM AND WAftR) 07 16935 CAVITY FRET VOLUMt 08 22. HEIGHT OF TOP OF TUNNEL ABOVE CAVITf FLOOR (MLASUkEu AT ss CAV!fY END OF THE TUNNEL IF IT SLOFtS) 09 2670. AREA 0F CAVITY CUTEk WALLS 10 0.0 LINER THICKNESS 11 00 LINLR GAP RESISTANCE 12 10. THILANESS OF WALL (UR UtP1H TO BE HUDtLLED FOR HEAT ss TRANSFER IF IT IS VERY DEEP) 13 0.92 fHERMAL CGNDUCTIVIlY OF WALL 14 0.157 SPECIFIC HtAT OF WALL 15 144. DENSITY OF WALL 16 0. NUMBER OF IGNITION SOURCES IN C
A-7 17 0. AVG DISTANCE OF THESE FROM IHE CEILING 18 796. SEDIMENIATION AREA 19 43 1 MINIMUM FLOW AREA WHICH CONNECTS CAVITY TO LOWEN LOMPT 84 Tl TROUGH TUNNEL 48 8 CONCRETE AND CONTAINMENI SHELL SSI
- FIRST 12 GUANTITIES ARE USED FOR ALL CONCRE1E DECOMPOSITION CALCS,
- UNLESS OTHERWISE STAIEDe CONCRt1E PRUPERTIES ARE FOR ' PURE'
- (UNREINFORCED) CONCRETE 01 1300. AVERAGE SFECIFIC HEAT OF CONCRETE (UP TO hELT POIN'f) 02 1b03. MELTING TEMPERATURE OF CONLRE1E (BETWEEN SOLIDUS AND LIQUIDUS) 888888 MEW: DEFINITIONS OF FOLLOWING HAVE CHANGED 8888888
- ALL IHE CONCREIE MASS FRACS SHOULD ADD UP TO ROUGHLY 0.9 TO 1.8 88(THE DIFFERENCE BETWEEN THE SUM AND 1 IS DUE TU NOT ACCOUNIING FOR 88SMALL PERCENTAGES UF NELATIVELY INtRT MAIERIALSe EG AL203 AND MGO) 03 .029 MASS FNACTION OF CONCRETE THAT IS FREE WATER 04 .0 MASS FRACTIUN OF CONCRE1E IHAT IS C%MICALLY BOUND WATER
- ITEM $'0SHOULD K SMALL FOR ' BASALTIC' TYPE CONCREfES 05 .015 MASS FRACTION OF CUNCHETE IHAT IS CO2 06 2.74E5 ENERGY ABSORBED IN ENDOTHERMIC CHEMICAL REACTIONS
- DURING CONCHE1E DECOMPUSITION 07 5.5E5 LATENT HEAT OF MELTING 08 1.8E-2 MASS FRACTION OF CONCRE1E THAT IS NA20 09 5.4E-2 SAME FOR K20 10 0.55 SAME FOR SIO2 11 0.30 SAME FOR CAO + OTHER LESS VOLATILE CONCREfE COMPONENTS 8888888888888888888 MEW 48888888888888888 12 183. REBAR DENSITY ssDCSRCN (MASS OF REBAR FEN UNIT VOLUME OF
- REINFORCED CONCRETE) a KG SIELL / M883 SIEEL + CONCRETE
$8 RELATED TO (KG STEEL / KO CONCktTE)ssR BY:
88 DCSRCNs R8DCN0/(1 + R8DCN0/DCS) WHERE DCN0 IS THE
- VIROIN CONCRETE DENSITYas2300 KU/M883 AND DL5==VIR0!N
- SIELL UtNS11Yss0000 K0/M883 -- CONSIDLR IHOSE VALUES 88 HARD-WIRED BECAUSE THEY WILL BE USED BY MAAP 1NIEhMALLY.
- REMAINDER OF IHE QUANTITIES ANE USED IN IHE CONTAINMENT FAILURE MODEL
$8AND NEED NOT BE SUPPLIED IF THE ' SIMPLE' MODEL IS USEU (SEE GENERAL
$8N0lES SECTION) 88 NOTE: FOR FREE-STANDING STEEL CONTAINMENIS, YOU NEED SUFPLY ONLY THE
- ' LINER' PROPERTIES (WHICH ARE TAktN TO DESCRIBE lHE STLEL SHELL) 88AND THE STEEL THICKNESS (STEEL THICKNESS IS INPUI AS
- ' LINER' 1HILKNESS IN fHE UPPER AND ANNULAR COMPARTMENT SECTIONS)--
- SEE GENERAL NOTES SECTION 13 3.E11 ELASTIC YOUNGS MODULUS FOR TENDONS 14 1 99E11 ELASTIC YOUNGS MODULUS FOR REUAR 15 3.9/E9 PLASTIC YOUNGS MODULUS FOR TENDONS 16 1.4E9 FLASTIC YOUNGS MODULUS FOR REUAR 17 9. '/ E 8 PRESTRESS ON HOOP IENDONS 18 1.01E9 PRESTRESS ON AXIAL TENDONS 19 1.53E9 TENDON YILLD STRESS 20 4.137E8 21 1.65E9 REBAR YIELD SSTNES'IRESS TENDON ULTIMATE 22 6.2E8 REBAR ULTIMATE STRESS 23 1.99E11 ELASTIC YOUNGS MODULUS FOR LINER 24 1.4E9 FLASTIC YOUNGS MODULUS FOR LINER 23 4 137E8 LINER YIELD SIRESS 26 6.2E8 LINER FAILUkE STRESS 88 88
_ . _ , _ . - - _ _ - . , , - _ _ . . . - -----.- _~ . . ~ , , _ . , - . .
A-8
- BR
- sssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssages CONTROL CARDS itell'5i'lil'iiii'lil!A'tiltil'ii'lll'Illinittiil'!Al'litil'tfit!!'1111' 01 1 ENTEN A 0 TO USE FAST STEAM TABLES IN PRI SYS WHEN FUSSIBLE 02 1 ENTER A 0 T0 USE FAST STEAM TABLES IN CONIMT WHEN POSSIBLE 01 1 INTEGRATION METHOD 1 RUNGE-KUTTA DRDER (1 OR 2)I 1 IS
- RECOMMENDED 04 29 UNIT NUMBER i TAPE' NO. IN CDC JAkGON) sa fu WRiiE PESTART FILES FOR MAIN PROGRAM FROM THIS RUN 06 30 UNIT NUMBER TO WRITL RESTART FILEb FOR HEATUP FKOM THIS RUN ss07 NOT USED 08 10 UNIT NUMBER TO FUI FRI SYSTEM OUTPUf DN 09 10 UNIT NUMBER TO PUT CONTAINMENT GUTPUT ON (MOST USERS PUT as IN THE SAME NO. WHICH APPENUS THE TWO FILES) 10 31 UN!f NUMBER FUR THE FIRST PLOT FILE (UfHERS SEQUENTIAL) 11 39 UNIT NUMBER FOR SCENARIO FILE
- sNEXT 3 OUANTITIES CUNTROL IHE PLOT POINT STORAGE FREQUENCY (SLE VOL 1 0F ssUSER'S MANUAL) 12 250 NON-S?IKE NUMWER OF POINTS (AVENACE BEHAVIOR) S10 RED 13 15 NUMBLR OF POINTS Sf0 RED DURING A SPIKE (TO RESOLVE FAST
! 17 0 EN1ER A 1 FUR B AND W PLANTSe 0 OIHERW15E (NOTE THAT MAAP as WAS WHITTEN FOR B AND W FLAN 15 WHOSE OTSG LOWLR TUWESHEE1S
- LIE BELUW THE LEVEL OF lHE PHIMARY SYSitM NUZlLES--NOT as YET TESTED FOR THE HIGH TUWESR Ef CONFIGURATIONS) 18 13 FILE NUMBER 10 WR11E AUX DATA ON FM LA1ER STAND-ALONE ss AUX RUNS (OR 0 NOT 10 WNITL DAIA) assassssssssssss**EWassasssas**sssastsal 19 0 FILE NO. TO READ AUX DATA FROM (IF THIS NUMBER IS NONZERO, ONLY
, as THE AUX BUILDING MODLLS ARE RUN, IHE INPUT T/H DAIA FROM THE as CONTAINMENT HAVING BELN RECORDED FROM A FREVIOUS RUN) 20 4 NUMBER OF NODES IN IHE AUX BUILDING (MAX =$i!F Os IHE AUX BLDNG ss PODELS ARE NOT RUNE PUI A FILE MAY S1!LL BE CREATED FOR l
ts SUBSEUUEhi STAND-ALONE AUX BUILDING ANALYSES BY SUFPLYING s* A NONZERO NO. FOR ITEM 18)
ssessssss**sssssssssssssssssssssssssssssssssssssssssssssssss44*ssssses
- CORE assessssssssssssssssssssssssssssssssssssssssssssssssss**ssas*******ssess 01 .031167 FUEL PIN UUTER DIAMETER 02 46920 INTIAL ZIRCALLOY MASS 03 50952. NUMPER OF FUEL PINS 04 222739. TOTAL UO2 MASS ss! TEM 5 MUST BE ABOVE THE ELEVATION SUFPLIED FOR THE TOP OF THE RV HLAD ss!N IHE PRIMARY SYS'IEM SECTION 05 10 132 ELEVATION OF BOTTOM OF ACTIVE FULL ABOVE BOTTOM OF VESSEL 06 22.132 ELEVATION OF TOP OF ACTIVE FULL AB0VE 80Tf0M OF VESSEL 07 8766. TIME OF IRRADIATION 08 1 16E10 FULL POWER a:THE CORE NODALIZATION ADMITS Ur TO 70 NODESI IN ADDITION NO MORE THAN as20 RuWS MAY BE USED AND NO MOKE IHAN 7 RINGS OR COLUMNS ssWHATEVER NODALIZATION IS USED, INSERT FEAKING F ACTORS INio AFFROFRIATE
A-9 88kNTRY NUMBERS (EG SECOND RING FROM INSIDE RADIAL PEAKING FACTOR IS
$$ALWAYS ITEM 32 NO MATTER HOW MANY AXIAL NODES) 88f0P NODE IS UNFUELED (FISSION GAS FLENUM ETC) AND MUST HAVE ZERO FEAKING
$8 FACTOR 09 7 NUMBER OF RINGS 10 10 NUMBER OF RUWS
$$ TOP ROW IS STRUCTURE (UFFEN FLENA ETC): S0 FEAKaZENO. THE FULLOWING
$8 9-ROW VALUES HAVE BEEN QBTAINtD BY AVERAGING 10-ROW VALUES.
48 THIS MEANS THAT ROW 5 HAS A SLIGHILY LOWER FEAK THAN OTNENWISE-~
88 REEVALUATE 1HESE FEAKING FACTORS IF DESIRtDe BUT DIFFERtNCES ARE SMALL.
11 0.498 AXIAL FEAKING FACTOR POTIUM 12 0./34 AXI AL PEAKING FACTOR 13 1 126 AXIAL FEARING FACTOR 14 1 402 AXIAL FEAKING FACTOR 15 1.480 AXIAL FEARING FA: TOR 16 1.402 AXIAL FEAKING FACTOR 17 1.126 AXIAL FEAKING FACTOR 18 0./34 AXIAL PEAKING FACTOR 19 0.499 AXIAL FEAKING FACTOR 8888888888888888888NEW88888888888888888 20 0.000 AXIAL PEAKING FACTOR TOP 88 ENTRIES 21-30 AXIAL FEAKING FACTORS NOT USED IN 1HIS NODALIZATION 31 1.09 RADIAL FEAKING FACTOR INSIDE 32 1.11 RADIAL FEAKING FACTOR 33 1.10 RADIAL FEAKING FACTOR 34 1 115 RADIAL PEAKING FACTOR 35 1.096 RADIAL FEAKING FACTOR 36 1.01 RAUIAL PEAKING FACTOR 1 37 0.75 RADIAL FEAKING FACTOR OUISIDE 38 0.047 AREA OR VOLUME FRACTIONS INSIDE 39 0.062 AREA OR VOLUME FRACTIONS 40 0.145 AREA OR VOLUME FRACTIONS 41 0.124 AREA GR VOLUME FRACTIONS 42 0 207 AREA OR VOLUME FRACTIONS 43 0.166 AREA OR VOLUML FRACTIONS 44 0.249 AREA OR VutunE FRaciluNb UUISIDE
$8FOLLOWING OUANTITIES CONTROL ANSI DECAY HLAT CALCULATION 45 32000 FUEL EXPOSURE AT SCRAM (ALWAYS IN
$$ MEGAWATT-DAYS /METkIC TON NO MAffEN WHAT UNITS SELECTEU) 46 .39 FUEL ' ALPHA
- AT SHU100WN (FISSILE ISOT0FE
$$ CAFTUhES/ FISSION) 47 .032 INITIAL tNNICHMtNT OF FUEL IN ATOM FRACTION 48 .642 CONVERSION RATIO (PRODUCTION RAlt OF U-239/ ABSORPTION RAT 88 IN FISSILE ISUTOPES) AT SHUTDOWN 49 .487 FRACTION OF FISSION F0WER MADE DUt TO FISSIONS IN U-235
$$ AND FU-241 AT SHUTDOWN 50 .443 SAME AS 49 FOR FU-239 51 .069 SAME AS 49 FOR U-238 (FAST FISSIONS) 52 6.5E-4 FRACTIONAL ZR021145S (COMPARED TO ZN MASS) AT TIFE O 88888888888884LL THE REMAINDER IN IHIS SECTION ARE NEW8888 53 .01344 FUEL FELLET RADIUS 54 3.185 CORE FLOW AREA IN THE BYFASS AREA BE1 WEEN THE CUKE BAFFLE
$$ AND THE CUNE BARRtL (ENSURE fHIS IS CONSISfENT WITH FRI 88 SYSTEM CORE FLOW AREA FARAMtTER NO. 5) 88 PARAMETERS b5-60 ARE USED FOP CALCULATING BALLOONING (DATA SHOWN IS 88MOSTLY FROM TMI REPORTS) 55 1.875E-3 CLAD THICKNESS 56 GAS VOLUME FER FUEL FIN b7 450. AS-BUILT ROOM TEMP FUEL FIN FILL OAS FRESSUNE 58 CURE SUFFORT PLA1E MASS--1HIS FLAIE IS MELTED BY IHE DEBRIS
--..,n -- - - - - - - - - - - . - - . - - - e. -, - . - - , , , . - - -- - - - - - - , - , -
A-10 ss AS If LEAVES THE ORIGINAL CORE BOUNDARY 59 FRACTION OF THE TOTAL FUEL PIN GAS VOLUME WHICH IS
- CONTAINED IN THE LOWER GAS FLENUM OF THE FIN 60 SAME IN IM UPPER GAS PLENUM 61 RADIUS OF CORE BAFFLE (2sPISTHIS ENIRY SHOULD BE THE
- CIRCUMrERENCE OF IHE BAFFLE) 62 0. ' FLOW AREA FEk ROW' IN CORE BAFFLE (IMPOR1AN1 ONLY IF
- IN-VESSEL NAIURAL CIRC %ATION RETURN LEG IS IN BAFFLE-CORE
- BARREL ANNULUS--SEE SMOKL)--THIS REPRESENIS THk AFFROXIMATE as FLOW AREA AVAILABLE AS lHE FLUW IURNS SIDLWAYS AND PENETRATES ss THE C0fr 63 0. FOR TMI-TYFE GEOMEIRIES THE FLOW AREA THROUGH EACH CORE
- FURMtR PLATE IN AXIAL DIRECTION 64 0. FOR TMI-TYFE CORESsNUMBER OF CORE FORMLR PLATES IN THE
- BAFFLE-CORE ANNULUS ss as satsssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssas s!CE CONDENSER ('I' COMPARTMENT) ses***sssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssa 01 0.E0 TOTAL VOLUME INCLUDING 1HE ICE 8802 EXIT GAS TEMPERATURE--THIS IS THE TLMFENATUNE OF GAS LEAVING THE
- ICE BOX (SEE WRIIE-UP FOR SU8 ROUTINE HTICE IN VOL 2 0F USER'S MAN) 8803 INITIAL TLMFERATURE OF THE ICE 8804 SPECIFIC VOLUME OF ICE--N01E THE TOTAL VOLUME MINUS IHE ICE MASS as TIMES THE SPEC VOL SHOULD BE THE FREE VOLUML
$308 FLOW AREA BETWEEN LOWLR COMFARTMENT AND THE ICE CONDENSEN ss09 SEDIMENTATION AREA as 4*sssssssssssssssssssss***ssssstes***ssssssssssssssssssssssssssssssssssa sVPLENUM (UFPER PLENUM OF ICE LUNUtNSER- 'u' COMPARTMENT) asssas****sssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssa 01 0. VOLUME--ENIER 0 VULUME FOR LARGE, DRY CUNTAINMENTS 4802 CHARACTERISTIC CROSS-SEC AREA 0F THE COMPT FOR BURNS 3803 HEIGHT OF UPPER PLENUM 3:04 LIMITING FLOW AREA WHICH COUFLES THE ICE CONDLNSER TO THE UFFER
- COMPARIMENf--!E USE 1HE LESStR OF 1HE UPPER PLLN TO UPPER COMPT sa FLOW AREA DR THAT COUFLING THE UFFEH FLEN TO THL ICE CONU.
8805 NUMBER OF IGNITERS IN U ss06 AVLRAGE DISTANCE OF IGNITERS BELOW THE CEILING OF U 8307 AVG DISTANCE FROM THE TOP OF UP FLEN TO IHE PORTION OF THE ss CEILING OF THE UPFEN COMPT WHICH IS JUST OVtR THE EAIT 0U1 0F UI as THIS IS USED TO CALCOLA1E LOCAL BURNING IN THE UPPER COMPT 88 INTIATED BY FLAME PRWAGATION OUT OF U 8308 SEDIMENTATION AREA IN U se es sesssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssa 84NNULAR COMPARTMENT ('D' COMPARIMtNT) essssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssa as as!N LARGE DRY CONTAINMENTS:
asTHIS COMPARTMENT REPRESENTS THE VOLUME BEfWELN THE CRANL WALL (IF ANf) asAND 1HE CON 1MT WALLe AND BEfWEEN fHE DECK AND IHE LOWER COMPT FLUOR--
ss!F NO CLEAR DISTINCTION: ARBITRARILY DIVIDE THE SFACE BELOW THE UFFER
A-11 tsCOMPT AND USE LAN6E FLOW ANEAS TO KLEP THE GAS WELL MIXtD--AT FRESENT, asCORIUM sa IS ASSUMED NOT TO GE) INf0 THIS COMPARTMEN1 ss!N 01 ICE CWDENMRS:
2.968E5 FREE VOLUME THIS VOLU5E REFLECTS THE ' DEAD-END' COMPARTMLNTS 02 4888. AREA 0F WATER FOOL 03 0. DISTANCE THE FLOOR OF D IS AB0VE THE FLOOR OF B I). A h*E ObALL 06 0 03125 WALL LINER THICKNESS 07 0 28 GAP RESISTANCE OF WALL LINER 08 43 THICKNtSS OF WALL 09 0.92 fHtRMAL CONDUCTIV! Y OF WALL 10 0.157 SPECIFIC HtAT OF WALL 11 144. DENSIlY OF WALL 12 0. ENTER A 1 IF THE OUTER WALL (CONfMT OUIEN BOUNDARY) 88 IS MADE OF STELL 13 0.
14 0. HEIGHT OF CUNB SEFERATING D AND 8 MEASUALD FNOM B'S FLOOR 15 0. NUMBER UF IGNITEMS OR IGNITION SOURCES IN D AVG DISTANCE OF THESE FNOM THE CEILING 16 9776. SEDIMENTATIM AREA assasssssssasALL THE REMAINLIER IN THIS SECTION ARE NEWasas 8:1HE NtXT THREE PARAMtTENS ARE USED TO DEFINE THE EFFICIENCY OF ss!NERTIAL IMFACTION asIN LARGE, DRT'S THESE PARAMETERS SHOULD CHARACTERIZE s*0 RATES WHICH ARE ASSUMtD TO BE IN IHE ANNULAR COMPARTMENT ss
- IN ICE CONDENSER PLANTS, THtSE PARAMETERS (EVEN THOU @ LOCATLD
- s!N THE ANNULAR CDMFARIMENT DAIA SECTION) SHOULD REFLECT IMPACTION AND ssFLOW AREAS AND STRAP WIDTHS IN THE ICE BOX--SEE EG FOSTMA 17 733.2 18 0.010 IMPACTION AREA (ANtA 0F BANS IN GNAIES l HAT INftRCEPT FLOW)
WIDTH OF GRATE PARS 19 4154.8 sa FLOW AREA THROUGH GRATES (DEFINES FLOW VtLOCITY OF AEN050LS TRAVtRSING GRATES) asNOTE! IF MORE THAN ONE LEVLL OF GRATES EXISTSe USE THE TOTAL INFACT!0N AREA ss0F ALL THE GRAIE$e AND THE MAXIMUM FLUW AkEA AT ANY OF THE GRA1E ELEVATIONS ss sSI s USEDDETAILED CONTAINMLNr FAILURE MODEL INFUIS--IGNORE IF SIMrLE MODEL 20 130.
ss NUMBER OF TENDONS IN HOOP DIRECTION IN THE FART OF THE WALL WHOSE AREA IS O! vin IN ITtM 5 ABOVE 21 216.
22 NUMBER OF TENDONS WHICH RUN IN THE AXIAL (VERTICAL) DIRECTION as .01399 VOLUME OF RtBAR PER UNIT AREA 0F QUTER WALL (EQUIV THICKNESS)
RUNNING IN THE HOOP DIRECTION 23 .06 DIAMtitR OF HOOP TtNDONS 24 .06E0 DIAMETER OF THE AXIAL TENDONS k
26 .3 RhNN G N EAfL REC!f$N !
sa O!SPLACEMENT IN AXIAL DIRECTION WHICH IS SUFFICIENT TO TEAR THE CONTNT WALL (EG AT A FENElRATION) 27 .3 SAME AS 24 FUR THE RADIAL DIRECTION ss sFR ssessssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssssses sENGINEERED SAFEGUARDS ssessssssssssssssssssssssssssssssssssssssssssssssssssssssssses**8488444s
- IN BRITISH UNITS, asFLOWkATES SPECIFIED TO BE VOLUMETRIC SHOULD BE M843/SECl OTHtR FLOWNATLS ss!E ALL THOSE NOT EXFLICIfLY STATLD TO PE VOLUMETRIC l
{
A-12
- SHOULD DE KG/SECl HEADS SHOULD BE IN MI FRESSURES IN FAI IN ENGLISH THE
- UNITS ARE RESPECTIVELY GPMsLBM/HRerie PSIA--
- NOTE TO MAAP/BWR USERS--GFM IS USED IN MAAP/PWH INSTEAD OF Fis*3/HH
- IN THE FOLLOWING,' FANS' REFER TO FAN C00 LENS--(AIR RETUNN FANS IN
- CONDENSER PLANTS)
- FOR BETTER ACCURACYe YOU MAY ELECT TO INrVI ' SYSTEM' FUMP HEAD CURVEb WHICH
- INCLUDE thE EtFECTS OF FRICTION IN IHE INLti AND OUTLLT PIPING (WHICH IS
- IGNORED IN MAAF)1 IF YOU DO S0e BE SUkE THE ASSUM*TIONS ON STATIL HtAD
- WHICH ARE USED IN IHEIR CALCULATION ARE CONSISIENT WIIH 1HE PUMP ELEVATIONS
- ETC. WHICH ARE INPUI BELOW--THIS IS GENERALLY A FACTOR ONLY IN CRITICAL 4 ** APPLICATIONS SUCH AS FLED AND BLELD WHERE 1HE CHARGING PUMP FLOW IS
- BARELY (OR NOT) ADEQUATE TO MATCH DECAY HEAT 01 0 833 ACCUMULATOR FIFE DIAMETER 02 195. FRESSURE SETPOINT FOR LPI 03 1336. FRtSSURE StTPOINT FOR HPI 04 615. INITIAL FRESSURE OF ACCUMULATORS 05 86. 1EMPERA10RE OF REFULLING WATER STORAGE TANK (RWST)--IE
- THE TANK FROM WHICH THE CHARGING, HPIs LPle AND SFRAYS
- DRAW 1HE!R WATER DURING 1HE INACTION PHASE 06 100. TEMPERATURE OF ACCUMULATORS 07 2.91E6 INITIAL MASS IN RWST 08 5.28E4 INITIAL MASS FER COLD LEG ACCUMULATOR 09 1319.5 AREA UF 8ASE OF RWST 10 18.3 LENGTH OF AN ACCUMULATOR PIFE 11 35 1 FRESSURE SETPOINT OF BLDG SPRAYS 12 300. PRESSURE SETFOINT OF BLDG FANS 13 0 NUMBER OF OPERATING FAN COOLERS OR FANS 14 0.0 VOLUMETRIC FLOW THROUGH ONE FAN COOLEN OR FAN 15 2.165E-3 NOMINAL DIAMETtR OF CONTAINMtNT SPRAY DROPLETS AS THEY N
17 4 1350. bE O bACCUMULATOR NUMWER OF 0FERATIONAL COLD LEG ACLUMULA10RS 18 1 NUMBtR OF OPtRATIONAL HPI PUMPS 19 1 NUMBER OF OPERATIONAL LFI FUMPS 20 5 NUMBER OF ENTRIES USED IN HPI PUMP-HD CURVE TABLE (5 MAX) 21 3600. HIGHEST HLAD IN TABLE (UNITS ARE MElENS) 22 3200. NEXT HIGHEST HEAD IN HPI PUMP-HEAD CURVE TABLE 23 2900. NtXT HIGHtST HEAD IN HPI FUMP-HLAD CURVE TABLE 24 1650. NEXT HIGHEST HEAD IN HPI PUNF-PEAD CURVE TABLE 25 0.0 LOWEST HEAD IN HPI FUMP-HEAD CUkVE TABLE 26 00 VOLUMEIRIC FLOWRATE CORESPONDING TO FIRST ENTRY IN
- THE FRESSURE TABLE 27 325. NEXT VOL. FLOWRATE 28 425. NEXI VOL FLOWNATE 29 650. NtXT VOL. FLOWRATE 30 650. NEXI VOL. FLOWHATE 31 5 NUMBER OF ENTRIES USED IN LFI TABLE 32 470. HIGHEST HEAD IN LPI TABLE 33 425. NEXT HEAD 34 390. NEX1 HEAD 35 325. NtXT HEAD 36 0.0 NEXT HEAD 37 00 FIRST VOLUMETRIC FLOWRATE IN TABLE 38 2000. NEXT VOL. FLOWkATE 39 3000. NEXT VOL FLOWRATE 40 4500. NEXI VOL. FLOWAATE 41 4500. NtXT VOL. FLOWRATE 42 2687. CHARGING FUMP FRESSURE SEff0!NI
A-13 43 1.0 NUMBER OF WCRKING CHARGING PUMPS 44 5 NUMUER OF ENIRIES IN CHARGING PUMP HEAD CUKVE TABLE 45 6000. FIRST HEAD 46 5800. NEXi HEAD 47 4800. NEXT HEAD 48 2000. NEXf HtAD 49 0.0 NEXT HEAD 50 0.0 FIRST VOL. FLOWWA1E 51 150. NkXT VOL. FLOWRATE 52 300. NEXT VOL. FLOWHATE 53 b50. NEXT VOL. FLOWRATE 54 550. NEXT VOL. FLOWNATE 55 160. AREA 0F BASE OF CONIMT SUMP 56 8 DEPTH OF CONTMI SUMP
- HEADe THkN A CONSTANT FLOW MODEL WILL BE USED 57 1 NUMBER OF USED LNIRIES IN SPRAY PUMP HEAD CURVES (5 MAX) 58 1000 FIRST ENTRY IN SFRAY FUMP HEAD TAWLE
- HEADS 59-62 NOT USED IN 1HIS SCHEME 63 1.65D-1 FIRST VOLUMLTRIC FLOW ENIRY IN SPRAY PUMP TABLE
- ASSuntD TO CORRESPONG T0 IHE NPSH HEADS GIVEN 68 28. NFSH (UNITS OF LENGTH) RE0'D FOR CHARGING PUMP sa AT FIRST FLOW IN TABLE 69 28. NEXT NPSH ENTRY FOR CHARGING FUMPS 70 28. NEXT NFSH kNTRY FOR CHARGING PUMPS 71 28. NEXf NFSH ENfRY FOR CHARGING FUNFS 72 28. NEXT NPSH ENTRY FOR CHARGING PUMPS 73 13.5 FIRST NFSH ENfRY FOR LPI 74 13.5 NEXT EN1RY FOR LPI 75 13.5 NEXf ENTRY FOR LFI 76 13.5 NtXT LNTRY FOR LPI 77 13.5 NEXf ENTRY FOR LPI 78 25. FIRST NPSH ENTRY FOR HPI 79 25. NEXf ENIRY FOR HFI 80 25. NEXT LNTRY FOR HPI 81 25. NEXl ENTRY FOR HFI 82 25. NEXT LNTRY FOR HPI 83 19. FIRST MPSH EN1RY FOR SFRAY FUMFS 84 3.05 NEXT EN1RY FOR SPRAY FUMPS 85 3.05 NEXf ENTRY FOR SFRAY FUMFS l
86 3.05 NEXT ENTRY FOR SPRAY PUMPS 87 3.05 NEX1 ENTRY FOR SPRAY FUMFS 88 1 NUMBtR OF OPERATING SPRAY PUMPS FOR UPPtR COMFARTMENT 89 0 NUMBER OF OPERATING SPRAY FUMPS F0k LOWEN COMFAATMLNI 90 77.5 HEIGHT OF B01 TOM 0F RWST ABOVE IHE ENG SAFE FUMPS 91 22.3 HEIGHT OF BOTTOM OF CONTAIN SUMr AWOVE THE ENb 5AFE FUMr$
92 38. ELEVAT!uN OF IHE RV INJECT!UN N0ZiLES ABOVE THE SI PUMFS 93 7577. FLOW THNOUGH OWt SFRAY FUMP WHEN ITEM 94 IS MEASUKEu 94 40. DIFPERtNTIAL PRESSURE ACROSS IHE SFRAY N04ZLES 95 0.0 MASS FLOWWATE OF EXIERNAL RWST REFLACEMLNf WATEke IF ANY 96 .00278 TIME DELAY FOR HFI (IE TIME BtTWtEN THE ACTUATION AND WHEN
- ACTUAL OPERATION BEGINS) 97 .002778 TIME DELAY FOR LPI 98 .002778 TIME DELAf FOR CHARGING PUMPS 99 .00833 TIME DELAY FOR UPFER COMPAR1 MENT SFRAYS 100 .00833 TIME DELAf FOR LOWER COMPARTMENf SFRAYS 101 5.0 TIME DELAY FOR FAN COOLERS 102 NUMBER OF TUBES IN A FAN COOLER
A-14 103 OUTSIDE AREA 0F ALL IUBES IN A FAN COOLER 104 AREA 0F ALL FINS IN A FAN COOLEN 105 FAN C00LtR FIN EFFICIENCY 106 FAN COOLER INSIDE FOULING FACTOR 107 FAN COOLER FIN DIAMETER 108 FAN COOLER TUBE THICKNtSS 109 FAN C00LtR fuBE THERMAL CONDUCTIVITY 110 MINIMUM FLOW AREA THROUGH FAN COOLEN 111 FAN C',0LER 1UBE ID 112 5 NUMsER OF NODES USED TO MODEL FAN COOLER (5 MAX) 113 INLtf COOLING WAftR fLMP TO FAN COOLER--NOTE IHIS IS
- ALSO USED AS THE COOLING WATEN TEMP FOR ALL OTHtR
- SAFEGUARDS HEAT EXCHANGERS 114 INLET COOLING WATER FLOW TO A FAN COOLEN 115 1 NUMBtR OF LFI FUMFS UStD FOR RHR SPRAYS WHEN VALVE OPEN 116 0 ENTER A 1 IF FANS /C00 LENS DISCHARGE TO BIO TO D
- sCALCULATIONS CONTROLLED BY HEAT EXCHANGER TYPE 88 HEAT EXCHANGER TYPE:
st -1 SET OUTLET TEMP OF HX TO RWST TEMrERATURE
- 2 U-1UBE HX 88 88!MPORTANT NOTE!
- 0R fHE N1U (NUMBLR OF TRANSFER UNITS) FER HX--ALL RNOWN USERS DO 88THE LATTER--NTUS ARE AVAILABLE BY CONSULTING NAMtFLATL DATA ANU
- USING GRAPHS INe FUR EXAMPLE, HOLMAN, HEAT TRANStER
!2! 0.D0 SkRA HfTU!kfk!CKNESS 122 0.00 IUBE TO fuse SEF4 RATION IN SPRAY HX 123 0.00 SHtLL LENGTH IN SFRAY HX 124 0.D0 fHtRMAL CONDUCTIVIlY OF SPRAY HX TUBES 125 0.00 LARGEST FERP DISTANCE FROM SHELL 10 BAFFLE (' BAFFLE CUf')
126 V.00 SHELL (0 lube CLEARANCE AT GUTSIDE OF SFHAY HX IbBE BDL 127 2.35E6 SFRAY HX COOLING WATER MASS FLOWNAIE 8:128 NOT USED 129 2 TYFE OF HX FOR RHN 130 0.D0 NUMBtR OF LUBES IN RHR HXS 131 0.D0 NUMBER OF BAFFLES IN RHN HXS 132 0.00 TUBt ID IN RHR HXS 133 0.D0 TUBE THICKNLSS IN RHR HXS 134 0.D0 fuBE TO fuBE SEFERATION IN RHR HXS 135 0.00 SHELL LENGTH IN RHN HXS 136 0.D0 (UBE lHtRMAL CONDUCTIVilY IN RHR HXS 137 0.00 BAFFLE CUT DISTANCE IN RHk HXS (SEE 125) 138 0 00 SHtLL TO lusE CLEARANCE AT OUTSIDE OF RHR HX IUBE BUNDLE 139 2.475E6 RHR HX COOLING WAlEN MASS FLOWRATE 140 0.996 SFRAY HX NIU 141 1.416 RHR HX NTU 142 0 00 SHELL ID OF SFRAY RECIRC HX 143 0 00 SHELL ID OF RHR RECIRC HX 88tNTER ZERO VOLUME FOR ITEM 148 IF NO UH1 SYSTEM
l A-15 l
l 1
- 144 INITIAL MASS IN THE UNI WATER ACCUMULATOR
- 145 LtN01H OF 1HE UHI PIPE TO IHE RV
- 146 DIAMETER OF THE UHI PIPE l
- 147 INTIAL PRESSURE OF 1HE UHI ACCUMULATOR 148 0.D0 TOTAL (WATER t gab) VOLUME IN THE UNI ACCUMULATORS
- 149 FAILURE DIFFERENTIAL PPESSURE OF 1HE UN! PIPE RUP10RE DISK
- THE ' CAVITY INJECTION SYSTEM' IS (RARELY) USED TO SIMULATE A
- USER HAS THE OPTION TO THROTTLE ESF SYSTEMS AT LESS THAN
- IHEIR FULL FLOW GIUtN 1HE CONDITIONS tXISTING--TO DO THIS,
- GENERATOR SECTION) A TOTAL FLOWRA1E DESIREDI IHE CODE WILL USE
- THE MINIMUM OF THIS FLOW AND THAT CALCULATED FROM THE HEAD CUKVES
- AND 1HE NO. OF OPERATIONAL PUMPSIIF OPERATOR ISN'T 1HRO1TLINGs
- ENTER A LARGE NO.11F HE CHANGES THE DEGREE OF THNOTTLINGs ENIEN
- PARAMETER CHANGES USING INIERVENTION N0. 1000 IN CONTROL CARDS 152 7.936E9 THROTTLED FLOW FOR LPI SYSTEM (TOTAL) 153 7.936E9 SAME FOR HPI 154 7.936E9 SAME FOR CHARGING PUMPS 155 7.936E9 SAME FOR UPPER COMPT NORMAL SPRAYS 156 7.936E9 SAMt FOR UFFER COMPT RHR SPRAYS (WHEN AC1IVAlED) 157 7.936E9 SAME FUR LOWtR COMPT SPRAYS
- t************************************************************
- INITIAL CONDITIONS
- tst38*******************************************************
01 591.8 NOMINAL FULL POWER PRIMARY SYS1EM WATER 1EMPERA10RE 02 2250. NOMINAL FULL POWER PRIMARY SYSTEM PRESSUkE 03 28.05 PRESSURIZER WATER LEVEL (ABOVE BOTTOM OF PZR HEAD) 04 14.7 CONIAINMENT BUILDING PRESSUKE 05 120 LUWER CONTAINMENT BUILDING COMPARTMENTS (ALL BUT I
l
- UPPER COMPT ANU ICE CONDENSEN) TEMPERATUNE 06 0. ILE CONDENSER GAS lEMPERAIUREe WHERE APPLICABLE 07 1. LOWER CONIAINMENT BUILDING COMPARTMENIS REL. HUMIDITY (0-08 0 INITIAL ICE MASS 09 99253. INITIAL MASS OF WATER ON SECONDARY SIDE OF EALH S/G
- VALUE TAKEN FROM MODtl F SG TlH DATA FOR MILLSTONE 10 120. INITIAL TEMPERA 1UNE OF CONIAINMENr CONCRETE ANU
- METAL STRUCfuRES 11 964. INITIAL PRESSURE ON SEC SIDE OF S/G'S
- UPPER COMPT CONDITIONS COULD BE DIFFERtNT IN ICE CONDENSERS 12 1 00 UPPER COMPARTMENT REL HUMIDITY (0-1) 13 100. OPPtR COMPARTMENT 1EMPERATURE 14 591 8 INITIAL PRIMARY SYSTEM WATER TEMrERATUKE FOR THIS RUN 15 2250 INITIAL PRIMARY SYSIEM PRESSURE FOR 1HIS RUN 16 0.0 AMOUNT OF SUPERHEAT AT EXIT N0Z OF AN OTSGi IGNOKED FOR
- U-TUBE STEAM UENERATORS
- PRIMARY SYSIEM
- UNLESS OTHERWISE NOIEDs ALL ELEVATIONS IN THIS SECTION SHOULD BE
- REFERENCED TO THE LOWEST FOINI 0F THE INSIDE OF THE RV HEAU
- WHEN A PARAMETER SUCH AS IHE VOLUME OF IHE DOWNCOMER IS CALLED FOR,
- THE ACTUAL DOWNCOMER VOLUnt SHOULDs OF COURSE, BE USED EVEN lHOUGH THE
- MAAP NODALIZATION LUMPS UTHER VOLUMES WI1H 1HE DOWNCOMER VOLUME (IHE
A-16
- LUMPING IS DONE INTERNALLY IN THE CODE) 01 4 NUMBER UF COLD LEGS 02 2.42 INNLR DIAMETER OF A HOT LEG FIFE
- NtWS****************1 03 INSIDE RADIUS OF THE CYLINURICAL PANT OF THE REACTOR VEbSEL
- NEWS ****************1 04 VOLUME WHICH IS INSIDE THE CORE BARREL AND LIES BETWELN
- lHE BulTOM OF IHE CORE AND 1HE LINE WHICH DENOTES IHE TOP
- OF THE RV HEAD (IE THt BOTTOM OF THE RV CYLINURICAL SECTION) 05 51.325 FLOW AREA 0F CORE PLUS CUNE BYPASS AREA 06 86 01 VOLUME OF HORIZONIAL RUN OF PIPE IN ONE COLD LEG FNUM
- 1HE REAll0R VESStL UUT TO 1HE MAIN COOLANT PUMP 07 .0623 RADIUS OF VESSEL FENtlRATION--IF NO VESSEL FENElRATION
- (tG SUME CE PLANTS) USE 1HE ASSUMED INITIAL RADIUS OF
- FAILUKE WHEN THE RV HEAD FAILS DUL TO CORIUM ATTACK ANU
- SUPPLY 1 FOR 1HE NO. OF FAILED PENETRATIONS IN *MODtl 08 1.2E7 ENERGY INFUf FROM ONt FRIMARY SYSTEM PUMP (WHtN RUNNING) 09 0.0 IUTAL MAKEUP FLUW TO 1HE PRIMARY SYS1EM--UNDtR NORMAL
- OPERATION SHOULD EQUAL LEfDOWN FLOW BELOWITHIS IS USED
- MAINLY IN lHE TMI SCENARIO AND MOST USERS WILL INPUT ZEROI
- THIS WATER IS NOT SUBTRACTED FROM THE RWST AND CONTINUES
- (IF POWER IS AVAILABLE) UNTIL MANUALLY SHUT OFF 10 560. TEMPERATURE OF MAKEUP WATENe IF ANY, GIVLN IN 09 11 2.2917 INNER DIAMEfER OF A COLD LEG PIPE 12 28 35 ELEVATION OF THE N0ZZLE WHICH ATTAChtS THE SUNGE LINE
- PUTTING IHE BREAK ELEVATION IN THE VICINITY OF THE SURGE LINEi
- ARTIFICIALLY INCREASING THE
- ELEVATION OF THE SURGE LINE 0.5-1 METER OR SU ABOVE THE BREAK IS SUGGESTED
- FUR 1HER, IT IS HELPFUL TO AVOID PUITING BREAKS NEAR 1HE ELEVATION OF IHE
- TUBESHtET IN U-TUBE TYPE S/G PRIMARY SYSTLMS--B0fH OF THESE MEASUkES
- HELP AVOID WATER SLOSHING INTO AND OUT OF NODES (WHICH CRANKS IHE TIME STEP
- DOWN) AND WILL GREATLY DECREASE RUNNING TIME Af NEbLIGIBLE LUSS OF ACCUNACY
- NEW*****************;
13 6 ENIER BROKtN LOOP BREAK LOCATION KEY (NODE NO.):
- 3--BROKEN HOT LEG NODE
- 4--BROKtN HOT LEG ' TUBE' NODE (B AND W ONLY)
- 6--8ROKEN INiERMEDIATE LEU NODE (BLIWtEN PUMP AND COLD SIDE OF
- S/G)
- 7--BROKEN COLD LEG NGDE (HORIZ FART OF COLD LEG)
- IWO COLD LEGS PER HOT LEG) WHICH IS UNutR A HORI20NTAL
- LINT DRAWN THNOUGH THE BOTTOM OF A COLD LtG NOZilt 17 MAX VOLUME OF WATtR IN ONE COLD LEG WHICH WILL STILL ALLOW
- GAS TRANSFER TO OCCUR PAST THE LOWEST FART OF THt CULD LEG 18 13TAL VOLUME OF ONE COLD LEG 19 TOTAL VOLUMt OF ONE HOT LEG 20 TOTAL FLUID VOLUME OF fHE RX VESStLe IE IHE VOLUME NOT
- INCLUDING THE CORE ITSELF OR INILRNAL STRUCTUKES
- 21 GAS FLOWHA1E OF REACTOR HIGH POINT VENT (S),IF ANY, AT
- NOMINAL SYSTEM PRESSURE
- DOWNCOMER IS M00tLLED AS LNDING AT IHE POINT WHERE IHE LOWER HEAD
- 0F THE RV MEETS THE CYLINURICAL SECTION -NOTE THE CORE BARREL IS
- ALSO ASSUMtD TO STOP AT 1HIS FOINT 22 TOTAL VOLUME OF DOWNCOMLR 23 PORTION OF DOWNCOMER VOLUME WHICH IS BELOW IHE
A-17
- ELEVATION OF THE BOTTOM OF THE COLD LEG N0ZZLES 24 3
- tNTER A 3 FOR PZR TO BE IN BRDKEN LOOPl 9 TO BE IN UNBROKEN
- LOOP FOR U-TUsE GEONETRIESI USE 4 ANU 10 RE5FECTIVELY Fuk 25 4 B AND W PLANTS (NODE NO. OF PRIMARY SYSTEM SURDE LINE N0Z)
NUMBER OF HOT LEGS 26 0.10 VOID FRACTION AT WHICH REACTOR C00LANf FUMPS TNIF OR FAIL
- SCRAM SEIPOINTS1 IF A GIVEN IRIP DOES NOT EXIST, INFUT A VALUE WHICH THE
- NUMBER OF FOINIS IN MAIN C00LANI FUMP COAS1-DOWN CUkVE (5 MAX) 35 3.55E7 FIRST MASS FLOWNATE IN MCF COAST-DOWN CUNVE(MUST BE THE
- ONE Ful+P FLOW UNDtR NOMINAL CONDITIONS) 36 3.23E7 SECOND FLOWNATE 37 2.48E7 NEXT FLOWRATE 38 1.77E7 NEXT FLOWRATE 39 1.10E7 NEXT PLOWRATE 40 0.00 FIRST TIMt IN COAST-DOWN CURVE--MUST BE 0 41 2.778E-4 NEXT TIME 42 1.111E-3 NEXI TIME 43 2.5E-3 NEXT TIME 44 5.833E-3 NEXT TIME 45 33.13 ELEVATION OF BOTTOM OF S/G TUsESHtET ABOVE BUTTOM OF RV
- (IGNORED IN B AND W PLANTS) 46 .50 THICKNESS OF RV HEAD 47 26.20 ELEVATION OF IHE BASE OF lHE COOLANT LOOP N0ZlLES
- (DISTANCE FROM BOTTOM OF N0ZZLES TO B01r0M OF RV HEAD) 48 66 VERTICAL DISTANCE FROM LOWEST POINT OF A COLD LEG TO THE
- ELEVATION OF THE BASE OF THE COLD LEG NUZZLt ON THE RV 49 VOLUME OF THE HORIZONIAL RUN OF A HOT LEG PIPE 50 0.0 TOTAL LETDOWN FLOW--SEE NOTE NEAR MAKtur FLOW ENikY ABOVE 51 35.5 NORMAL DIFFERENTIAL PRESSURE FROM CORE INLET TO HOT LtG
- ALLSIDE OF OUTLET N0ZZLES WHEN MAIN C00LANI FUMFS ARL ON THE REMAINDER IN 1HIS SECTION ARE NEWa***
- MOST USERS WILL USE THE ' UNBROKEN
- LOOP BkEAK UNLY FOR PUMP SEAL LOCAS
- SIMULATION)
- THIf BREAK, ALONG WITH THE BROKEN LOOP BREAK IS CONTROLLED Bf EVENI CODE ,
- 20VI ONE CAN TURN 1HE BREAKS ON AND UtF SEPERAIELY BY USING A FARAMETER '
- CHANGE-TYPE INTERVENTION (CODE 1000---SEE VOL 1 0F USER'S MANUAL) 52 12 LOCATION KEY FOR UNBROKEN LOOP BREAKS IF ANY
- 9 --UNBROKEN HOT LEG NODE
- 10--UNBROKLN HOT LEG '10BE' NODE (B AND W ONLY)
- 12--UNBROKEN INIERMtDIATE LEG NODE--
- DOWNCOMtR NOT ALLOWLD AT THE FRESENf 53 0. AREA 0F UNBROKEN LEG BRtAK--PUT IN ZERO IF NONE
- 54 35 ELEVATION OF UNBROKEN LOOP BREAK (SEE NOTES FERTAINING
- TO BREAK ELEVATION ABOVE
- THE ' DOME' REFERS TO THE REGION ABOVE THE UFFEN FLENUM '
- lHE ' DOME PLATE' IS THE PERFORAIED PLATE fHAT DIVIDES IHE UPPER PLENUM '
- FROM 55 THE DOME--SEE DRAWINGS IN THE PRISYS SECTION OF THE USEK'S MANUAL I ELEVATION OF THE RV DOME PLATE l
A-18 56 ELEVATION OF THE INSIDE OF THE RV HtAD 57 ELEVATION OF INE RV FLANGE (CLOSURE STUDS) at (NOTE THAT THIS ELEVATION IS N OU DPANA# M EMe m lx?tT0lWP 59 MASS OF fHE CORE BARREL BELOW 1HE ELEVATION OF 1HE TOP OF
- THE CORE (' LOWER CORE BARREL') FROM INF DATA + MAAP2 FILE 60 MASS OF fHE CORE BARREL ABOVE 1HE LLEV UF IHE TOP OF THE CORE
- (*UFFER CORE BARREL') -- MAAP2 ZION CALC NOTES 61 MASS OF UPPER PLENUM INTERNALS -- MAAP2 ZION CALC NOTES 62 MASS OF THE RV DOMt PLATE -- MAAP2 ZIDN CALC NOTES 63 MASS OF ikE WALL FORMING IHE EXTERIOR OF lHE DOME (IE
- INCLUDES THE RV CLOSUNE HtAD) FNOM MAAP2 FILE 64 TOTAL MASS OF ONE HOT LEG + HOT INLET PLENUM WALL OF THE S/G+
- THE TUBESHtET MASS ASSOC WITH THE INLET FLENUii 65 TOTAL MASS OF ONE COLD LEO + COLD QUTLET PLENUM OF IHE S/G
- FLUS THE TUsESHtET MASS ASSOCIAltD WITH THE OUILEf PLENUM
- NOTE: FOR PLANTS WI1H 1WO COLD LEGS PER OUILET PLENUM,
- ADD ONLY HALF THE OUTLE1 FLENUM MASS--THE OlHtR HALF IS
- IHEN ASSOCIATED WIIH IHE UlHER COLD LEG IN 1 HAT LUDP 66 MASS OF THE RV WALL (BELOW THE RV FLANGEi THE DUMt WALL
- ENTERED ABOVE STARTS AT IHE PLANGE) FROM MAAP2 FILE 67 WATER LINE AREA IN THE UFFER PLENUM (ABOVE THE CDNE AND SELOW
- 1hE DOME PLA1E) -- ESTIMATE WIlH 0=12Fre 1/2 WATER 68 HYDRAULIC DIAMETER IN THE UPPER FLENUM -- INP 69 TOTAL HEAT TRANSFtR AREA 0F lHE UPPER PLENUM INTERNALS '
70 CONVECTIVE (NON-RADIA1IVE) HEAT LOSSES UNDER NOM CONDITIONS
- FNOM S1EAM UENERATORS PRESSURIZER: AND REST OF PRIM. SYS.
- NOTE: DETAILED CALCULATIONS INDICATE THAT UNDER NORMAL
- OPtRATIONe IHE PRIMARY SYSTEM HEAT LOSS IS DDE VIRTUALLY
- ENTIRELY TO UNINSULATED FARTS OF THE SYSTEM (LOSS THKUUGH
- INSULATION IS NEGLIGBLE)I 1HUS 1HIS NUMBER
- SHOULD BE AFFROXIMATELY THE TOTAL NOMINAL PRIMARY
- SYSTEM HEAT LUSS (SEE IDCOR REPORT 85-2 FOR DISCUSSION) 71 NO. OF FLATES IN PRIMARY SYSTEM REFLECTIVE INSULATION ORI
- EN1ER 0 FOR CALCIUM SILICATE BULK INSULATION OR
$* ENTER -1 FOR ROCK WOOL INSULATION--IF YOU HAVE A
- DIFFERENT lYPE OF INSULATION YOU SHOULD CONSIDER MODIFYING
- FUNCTION THCBUL WHICH SUPPLIES THE THtRMAL CONUUCTIVITY 72 TOTAL 1HICKNESS OF INSULATION 73 ELEVATION OF THE BASE OF THE CYLINDRICAL PART OF THE RV 74 VOLUME OF lHE LOWER HEAD OF THE RV 75 TOTAL HEAT TRANSFER AREA 0F LUWER CORE BARREL / THERMAL
- ts***********sttsstats*stsst*t****t**t*********t*********************
- PRESSURIZER
- sts*********st*S*tsts*stststst**stsassssststats**t****t***3***********
01 1800. PRESSURIZER VOLUME 02 FRESSUNIZER CROSS-SECTIONAL AREA 03 2235. PRESSURIZER HEATER PRESSURE SETPOINT 04 2325. FRESSURIZER SPRAY PRESSUkE SEfFOINI 05 7.8 WATER LtVEL BELOW WHICH PZR HEATERS TRIP 06 6.14E6 FRESSUKIZLR HtATER TOTAL OUTPUT--IN MAAF THE HEATERS
A-19 09 2500. LOWEST SETFOINI 0F A SAFETY VALVE (UFENING FRES5UNE) 10 2500. HIGHEST SEIFOINT OF A SAFE 1Y VALVE (OFtNING FRESSURE) 11 0.93 DIAMtTER OF THE SURGE LINE 12 46.25 ELEVATION OF SFRAY HEAD ABOVE BOTTOM OF FZR 13 64. LENGTH OF THE SURGE LINE 14 3 NUMBER OF SAFELY VALVES 15 3.281E-3 NOMINAL PZR SFRAY DROFLET 16 2450. LOWEST SET FOINT OF FORV (OPtNING PRESSURE) 17 2350. HIGHEST SET FDINI 0F FORV (OFENING FRES5 uke) 18 2 NUMBER OF PORVS 19 2.1E5 NOMINAL FLOWNATE OF A FORV AT ITS SETFOINI 20 1.656E5 EMPlY MASS OF FZR STEEL 21 0 ENTER A 1 IF THE SURGE LINT HAS A LOOP SEAL (EG TMI)I
- 1HIS FREVENTS COUNTER-CURRENT DRAINING OF FNESSURIZER
- THkOUGH SUNGE LINE WHEN THL FRIMANY CUOLANI LOOP SIDE
- IS VOIDED (SEE WRilEUP FOR SUBNOUTINE DRAIN) 22 38.5 SEDIMENIATION ARLA
- ALL IHE RtMAINDER IN IHIS SECTION ARE NEWa***
- FRESSUNIZER RELIEFS ARE ASSUMED TO CLOSE AT FRESSUNE FSET-FDEAD WHENE
- PSET IS 1HE OPENING PRESSURE DEFINED AB0VE AND PDEAD IS GIVEN BELOW 23 100 DEADBAND ON FRESSUNIZER SAFETY VALVES 24 100 DEADBAND ON FRESSURIZER PORVS l
- STEAM GtNERATOR (VALUES REFER TO ONE UNIT) 01 5900. TOTAL SECONDARY SIDE FREE VOLUMES EG OUT TO THE MSIV'S 02 DOWNCOMER CROSS-SECTIONAL FLOW AREA 03 IU8E BUNDLE (SECONDARY SIDE) PLOW AREA 04 0.D0 B AND W ONLY--ELEVATION OF AUX FEED SFRAY HEAD ABOVt
- 1HE TOP OF THE LOWER IUBESHEET 05 1.158E6 INITIAL MASS IN CONDENSATE STORAGE TANK--OR A LARbE
- INVENTORY SUPPLitD IN THE
- INITIAL CONDITIONS SECTIONI
- THIS IS USED TO ADJUST THE VOID FRACTION DISTRIBUfl0N
- SIMPLIFICATIONS IN THE MAAF MODEL i THE CURRECTION
- SHOULD MOST IMPACT LOSS UF FEED SEQUENCES 07 440. MAIN FEEDWATER TEMPERATURE 08 1199.7 LOWEST SETPOINT OF SECONDARY SAFE 1Y VALVES 09 1269.7 HIGHEST SETFOINI 0F SEC SAFETr VALVth 10 5 NUMBER OF SAFETY VALVES PER S/G 11 9.19E5 NOMINAL FLOWNATE OF A SAFETY VALVE AT THE SETF01NI 1135.
12 SETPOINT OF SEC RELIEF VLV (ASSUMtB SAME FOR ALL RtLIEFS) !
- IF NO ' RELIEF VALUES *--SUFFLY A SET FOINI PRESSURE HIGHtN THAN THE l
- SAFETIES AND USE THE RELIEFS AS MANUALLY CONTROLLED STEAM DUMPS 13 1 NUMBER OF RELIEF VALVES FER S/G 1 14 4.0E5 NOMINAL FLOWRATE OF A RELIEF VALVE '
15 3.97E6 MAX FEEDWATER FLOWHATE FER S/G l
- DEFINITION CHANGED FOR OTSG'S***********1 '
- INCLUDE THAT FORTION OF THt TUBE VOLUME WHICH IS NOT COOLED (IE 15 INSIDE
- 1HE 10BESHEET(S)) IN I1EM 16 l 16 318.8 TOTAL (BOTH FLENA FOR OTSG'S) FRIMARY HtAD(5) VOLUMt-- !
- UNLESS DEFEAIED WIIH APPROPRIAIE LVtNT CUDES 17 .0028 TIMt DELAY FOR ACTIVATION OF AUX FELD AFTER SCRAM 18 .00138 TIME NEQUIRED FOR MSIVS TO LINEARLY RAMP FROM OPEN
A-20 tt TO CLOSED 19 966 1 TOTAL PRIMARY SIDE VOLUME OF ONE STEAM GENERATOR 20 1.78E5 MAXIMUM AUX FEED FLOWRATE PER S/G 21 100. AUX FEED TEMPERATURE 22 5626 NUM8ER OF 10BES IN A STEAM GENERATOR 23 0.0033 THICKNESS OF STEAM GENENATOR TusES 24 0.05067 ID OF STEAM GENERATOR TUBES 25 9.35 THERMAL CONDUCTIVITY OF STEAM GENtRATOR TUDES 26 7.936E9 IHROTfLED FLOW PER STEAM GENERATOR FOR AFW SYSTEM OR LARGE st NUMBER IF FLOW NOT THROTTLED (SEE DISCUSSION AFIER ENGIN.
35 SAFEGUARDS ITEM 151) 27 1. FRACTIONAL AREA USED FOR STEAM DUMPS IN BROKEN LOOP S/G 28 1. FRACTIONAL AREA USED FOR STEAM DUMPS IN UNBKN LOOPS S/GS
- STEAM GENERATOR WATER LEVEL CONTROL (SbWLC) SYSTEM PARAMETERS:
29 34 8 DOWNCOMER PROGRAM WTR LtVEL FOR STEAM GtNERATOR WATER
- LEVEL CONTROL SYSTEM IN BROKLN LOUP S/G 30 34 8 DOWNCMR PROG WTR LVL FOR SGWLC SYSTEM IN UNBKN LOOP S/GS 31 STEAM GENENATOR TUBESHEET DIAMtTER
- FOR 34 8 ACCIDENT SIMULATION IT WAS NECESSARY TO INCORPORATE A BANG-BANG
- SMODE OF S/G WATER LEVEL CONfROL--IE OPERATOR CONfMOLS THL WATER LEVEL stIN A34 8CILLATORY WAY WIfMIN A DEADBANDI MUST USERS WILL NOT WISH t*TO USE THIS MODE AND SHOULD LEAVE THE NEXI THNEE ENIRIES EQUAL TO O 32 0.0 B-LOOP SGWLC DEADBAND 33 0.0 A-LOOP SGWLC DEADBAND (NON4ERO VALUE ACTUATES 48 BATCH FEED MODE) 34 0.0 FOR BANG-BANG MODE, THE MINIMUM AFW FLOWNATE PER S/G TO
, St BE USED ON THE DECREASING CYCLE
- t**8888888ALL THE REMAINDER IN THIS SECTION ANE NtWtsts 35 0 MAIN STEAM LINE BREAKS CAN BE SIMULATED: ENTER 0 FOR N0 st MAIN STEAM LINT BREAKI 1 DIRECTS STEAM FROM BHOKEN
- $ LOOP S/G TO CONTMTl 2 DIRECTS STEAM FROM ALL S/GS TO
- CONIMT--TO MODEL A BROKEN MAIN STEAM LINE, FORCE OPEN 38 MSIVS OR S/G STEAM DUMP (S) AND USE NON4ERO VALUE OF
- THIS PARAMETER 36 TOTAL HEIGHT OF S/G SHELL ABOVE TUBESHEET 37 MASS OF S/G SHELL--DON'T INCLUDE MASSES ASSOCIATED WITH PRIMARY
- HLADS OR TUBESHEETS WHICH ARE LUMPED WIlH IHE ASSOCIATED COLD l ** AND HOT LEG MASSES IN THE PRIMARY SYSTEM SECTION l
38 12 NUMBER OF PLATES IN REFLECTIVE INSULATION ON S/G SHELLS OR CODE INDICATING OTHER INSULATION TYPE (SEE FRIMARY SYSTtM
- t INPUT NO. 71) tsststattsssssssssistsssssssssstssssssssssssssststittsststststStsstststs STIMING DATA
$s8$35888888888tt$88883583388888stS$8888888tststststiltsstsisstssstSttSt
- 01 NOT USED
$s02 NOT USED 03 20.0 MAX TIME STEP (ALWAYS INPUI IN SECONDS) 04 .005 MINIMUM TIME STEP (ALWAYS INPUT IN SECONDS) t* TIME SELECTION ALGORITHMS ARE EXFLAINFD IN THE WHITE-UPS FOR SUBROUf!NES
- INTGRT (T/H MODELS) AND INTGFP (FISSION PRODUCT MODELS) 05 .05 RELATIVE MASS CHANGE USED TO SELECT TIME STEP 06 .02 MINIMUM INTER-NODE FISSION PROD MASS TRANSFER CONSIDERED WHEN
- PICKING TIME STEP IN FISSION PRODUCT MUDELS 07 .02 RELATIVE GAS TEMPtRA1URE CHANGE USED TO SELECT TIME STEP 08 .1 REL MASS CHANGE FOR FISSION PRODUCTS USED TO SELECT TIME 88 STEP IN FISSION PRODUCT ROUTINES
- t l
A-21
- s sss:
- 0UENCH TANK ('GT' COMPT) s s s s s a s tsssssssss s ss s t att s S 8 ss as stst sts ss ssss ss sis t at s t s s t *
- s t
- s t s s s s t
- s 01 1800. TUTAL v0 Lunes INCLUDING 1 HAT OF 1HE INITIAL WATER MASS 02 83695. INITIAL WATER MASS 03 105.7 FAILURE DIFFERENTIAL PRESSURE OF RUPTURE DISK 04 12.5 HEIGHI 0F RUPTURE DISK ABOVL BCOMPT FLOOR
88 IN THE GAS SPACE OF THE QUENCH TANK IS NOT MODELLED ts sSI
- s**s***********************************sas*******************s********
- MODLL PARAMETERS
- tssssts**stsst*sstats***sts ststatstastattsstssttsS8sSSSSSS*3858tss*sts
- SEE DISCUSSION IN VOL 1 0F USER'S MANUAL FOR ALLOWABLE LIMITS ON
- MODEL PARAMtTER VALUES AND THE DIFFERENI SENSITIVITY ANALYSIS MODES
- SCALE FACTORS' MULTIPLY MODEL PREDICTIONS OF FLOWMATES ETC.--
- lHE BEST-ESTIMATE VALUE IS USUALLY 1
- stass*******sMANY OF THE FOLLOWING HAVE BEEN CHANGED 8***
01 .005 CORIUM FRICTION COEFFICitNT FOR VESSEL ABLATION HEAT
- TRANSFER (REYNOLD'S ANALOGY) CALCS 02 .002 LEAK-BEFORE-BREAK CONIMT LEAKAGE AREA (IF 1HE CON 1MT STRAIN s* MODEL IS NOT USEDs THIS IS THE AREA USED WHLN THE CUNfMr PRESS
- PRESSURIZATION IN A GROSS MANNER (IE BEFORE THE STMAIN
- CRIIERION IS REACHED AT PENETRATIONS)I USED ONLY IF DETAILED
- CONIAINMENT FAILUME MODEL ACTIVATED 05 2.0 MULTIPLIER OF NORMAL CLAD SURFACE AREA USED IN OXIDATION
- CALCS TO ACCOUNI FOR STEAM INGRLSS AFTER
- CLAD RUP1URE (MUST BE Bt1 WEEN 1 AND 2) 06 983.D0 CRITICAL FLAML TEMP AT ZERO STEAM MOLE FRACTION
- USED IF NO IGNITION SOURCESI 1HIS IS MULTIPLIED BY THE
- WESTINGHOUSE FLAME TEMPERATURE MULTIPLIER CURRELA110N
- FOR NONZERO STEAM MOLE FRACTIONS 07 1. SCALE FACTOR FOR FISSION PRUDUCT AND INERT AERO RELEASE RATES
- WHEN 2- OR 1- PHASE NATUKAL CIRCULATION IS OCCURING
- IN INE COOLANT LOOPS--NOTE IHAT COOLANT VELOCI1Y AND
- VOID FRACTION DISTRIBUlION ARE NOT COMVUILD UNDER THESE COND.
10 .5 FRACIION OF S/G TUBES CARRYING '0UT' FLOWS IN lHE HOT LEG )
- NATURAL CIRC MODEL (SEE SU8ROUlINt HLNC WHITE-UP)I 1F YOU j
- WISH TO FORCE THE FLOW OFFS USE O (1HIS REQUIRES BYPASSING !
s* PARAMATER CHtCKING BY USING THE SENSITIVITY ANAL OPTION l
- IBATCH=2)I PARAMETER DOES NOT AFFECT B AND W GEOMETRY UNLESS st 0 IS INPUI SINCE OTSG TUBES DON'T PARTILIPA1E IN FLOW 11 .1 8 AND W ONLY' FRACTION OF S/G TUBES STRUCK BY A>W 12 1.D3 HT. TRANSFER COEFF BETWEEN MOLTLN CURIUM AND A FROZEN CRUSTi
- USED IN DECOMP AND IN CALCULATIONS WIlHIN A MOLTEN POOL IN
- THE CORE 13 0 ENTER A 0 FOR ENTRAINMENT FROM C TO Bf 1 FOR C TO A--LATTER
- UENERALLY USED ONLY IF CAVI1Y HAS NO INSTRUMENT IUNNEL 14 0.D0 IF 13 IS NONZLRO: FRACTION OF THE ENfRAINtD MASS WHICH
- S1RIKES IHE MISSLE SHIELD (BEFORE SIGNIFICANILY INTERACTING
- WITH THE UPPER COMPARTMtNI GAS) l l
_ - _._ _ __. _ _ . . _ _ . . _ ,__ _ , _ _ . _ _ _ . _ . . _ _ _ _ _ _ _ ._ _ i
A-22 15 1.D0 DRAG COEFFICIENT OF RISING PLUME DURING BURNS IN OFFER
- COMPT-- LARGER VALUtS RESULT IN A SLOWER ANU FATTER PLUnt
- AND THUS INCREASE 1HE EFFICIENCY OF INE 10 NITERS l 16 1.D0 SAME FOR B COMPT 1 17 1.D0 SAME FOR C COMPT i 18 1.D0 SAME FOR D COMPT 19 1.D0 SAME FOR U COMPT 20 1 53 CHURN-TUNBULENI CRITICAL VtLDCITY CDEFFICIENI 21 3.7 DROPLET FLOW CRITICAL VtLOCI1Y COEFFICIENT 22 1. SFARGED FOOL VOID FRACTION CDEFFICIENr 23 2. V1 0ME1RIC STEAM DENERATION VOID FRACTION COEFFICIENT ,
24 .5 DE8RIS ENfRAINMLNI TIME CDNSTANI (UNITS OF TIME) )
25 .9 EMISSIV!1Y OF WATER 26 .85 EMISSIVITY OF WALLS 27 .85 EMISSIVIlY OF EQUIFMENT 28 .85 EMISSIVITY OF CORIUM SUNFACE 29 .6 EMISSIVITY OF GAS i I
30 .3 CORE HfDR0 DYNAMIC LIMIT KUTATELADIE NO. FOR REFLOUDING Hi
- AND OXIDATION CALCULATIONS 31 .33D0 NUMBER TO MULTIFLY KUI ATELADZE CRITERION BY TO REFRESEN1
- OUT OF CAVITY 32 3.0 FLOODING CHITICAL VELOCI1Y COEFFICIENT 33 .14 FLAT FLATE CHF CRITICAL VtLOCITY COEFFICIENI 34 1. NUMBER OF VESSEL PENEl' RATIONS IHAT FAIL 35 .75 DISCHARGE COEFFICIENr FOR PRIMARY SYSTEM BREAK (S)
- SCALE FACTORS' MULTIPLY MODEL PREDICTIONS--lHE BEST-ESTIMATE VALUE
- IS USUALLY I '
36 1.D0 SCALE FACTOR FOR BUNN VELOCITY CORRELATION 37 1.D0 SCALE FACTOR FOR HEAT TRANSFER COEFFICIENTS TO PASSIVE .
- HEAT SINKS l 38 2.5D0 GAMMA SHAFE FACTOR (TO ACCOUNI FOR NON-SFHERICAL SHAPES IN
- lHE C0AGULATION E00AT10N) USLD FOR AER0s0LS 39 1.D0 CHI SHAFE FACTOR (TO ACCOUNI FOR NON-SFHtRICAL SHAPES IN
- STOKES LAW) USED FOR AEROSOLS 40 3.D0 RATIO OF AIRBORNE AEROSOL MASS TO THE MASS WHICH WOULD LEAVt
- YOU IN STEADY-STATE WIlH lHE CURRtNT SOURCE STRtNGlHilHIS IS
- USED TO CONfkOL THE SELECTION OF DLLAY VS STLADY-STATE AtkOSOL
- SETILING CORRELATIONS 41 10.D0 DECONIAMINATION FACTOR ASSOCIATED WITH THE FASSAGE THkOUGH 1
- MtTER (REFERENCE LtNG1H USED FOR EIlHER SET OF UNITS) 0F uATERi
- ASSUME DF IS LINEAR FUNCTION OF DLPTH FOR OTHtR DLPTHS J 42 .02 CAPluRE EFFICitNCY OF CONIMT SFRAY FOR AEROSOLS--lHIS IS
- THE FRACTION OF THE TOTAL VOLunt SWLPT ff FALLING DbOPS WHICH
- IS CLEANSED OF AEROSOLS 43 1.D0 ABSOLurE VALUE OF THE DESIRED MULTIFLIER OF CbI AND
- CSOH VAPOR FKESSURE--ENTLR A NEGATIVE NUMBER TO SLLECT
- JANAF CSON FUNCTIONi FOS FOR SANDIA CURELLATION (BEST-EST) 44 .1D0 FRACTION OF CLAD OXIDIZED WHICH CAUSES CORE TO COLLAPSE ON
- REFLOOD (GIVES SMALLER KU FOR HtAl IRANSFER THAN INIAtt'
- MODEL) AND CAUSES CORE GEONElRY TO CHANGE 1 45 0.D0 FOR B AND W UNITS ONLYe FRACTION OF FtRFtCT CUNUENSATION !
- OF STEAM EN1ERING DOWNCOMtR IHkOUGH PLAFFER VALVES
- SI 46 2100. TEMFERATUkE AT WHICH CLAD FAILS IF IT HASN'T ALREADY RutTUAEDi
- IHIS Halls ruR1HER BALLOONING AND ALLOWS FIsS FROD RELEASE 47 2.5E5 LATENI HEAT OF U-IR-ZR02 EUIECTIC 48 .25 VUID FRACTION OF A COLLAPSED CORE
A-23 49 .3D-6 SEED RADIUS ASSUMED FOR HfGROSCOPIC AEROSOL GROWiH CALC 50 -2 ENTER A 2 FOR FISSION FR0 DUCT RELEASE TO BE COMPUTED
- BY THE IDCOR/tFRI STEAM OXIDATION MODELi 1 FGR
- NUREG-0772 MODEll NEGATIVE NOS. ACTIVATE lHE SAME MODEL
- AS POSITIVE NUMBERS BUT ALSU TUhN ON A BLOCKAGE MODEL '
- WHICH REDUCES IHE RELEASE OF NONVOLATILE FISSION PRODS
- WHEN THE NODE IS BLOCKED FOR GAS ThANSPOR1 1
i 51 0 ENTER A 1 IF TELLURIUM IS RELEASED IN-VESStLI O IF IT
- IS ASSUMED TO BE TOTALLY SOUNu Ur WITH ZIRCALLOY
- (0 IS 8EST-EST) 5' 2500. ASSUMED EUIECTIC MLLTING TEMP 53 .1 FRICTION COEF FOR AXIAL PLOW USED FOR UPVER PLtNUM-CORE FLCW
- CALCSITHIS CAN BE ESTIMATED By F=2.*bPakH0/Ga*2 WHtRE (ALL
- VALUES ANE FOR NORMAL OPERATION WIlH MCP'S ON) DP=C0KE
- PRESSUKE DROP, RH0= DENSITY OF PNIMARY SYSTEM C00LANI, Gm
- CORE AVERAGE MASS FLOW PtR UNIT AREA (IN BRIT UNITS,
- INCLUDE G0 ANU OTHER NECESSARY CONVtRSIGHS TO MAKt F
- DIMENSIONLESS)--USE GT 100 TO ARTIFICIALLY STOP PLOW
- (REQUIRES USING THE SENSTIVITY OPTION IBATCH=2) 54 0. INSERT 0 IF IN-VESSEL NAIURAL CIRCULATION PLOW REluRN LEG
- IS IN OUTER FULL ASSEM8 LIES (USUAL CASE)i1NSERI 1 IF RETUNN
- BE EXFECTED ONLY IF THERE WAS A LOT OF FLOW AkEA IN THE
- BYPASS (EG PERHAPS B AND W PLANTS) 55 .1 A VOID FRACTION, BELOW WHICH A CORE NODE IS Ab50MED BLOCKED
- LONG-TERM OPERATION (OFTEN CALLE3 IHE ' GAP RELEASE',1HIS
- IS USED IN CALCULATING THE PKEShuRE INSIDE THE FUEL PIN FOR
- SALLOONING CALCS--NUMEG 0772 SAYS OBSERVED VALUES ARE 0-0.25) 59 .35 VOID FRACTION IN PRIMARY SYSTEM A80VE WHICH THE PHASES
- SEPARAIE AND IWO-PHASE NA1 URAL CIRCULATION STUPS 60 1060. TEMPERATUKE OF H2 JEl ENTERING NON-INLRTED COMPARTMENT WHICH
- IS SUFFICIENT TO CAUSE A LOCAL BURN--FROM HEDL-lME 78-50
- FISSION PRODUCTS
- NEW*****************:
- FISSION PRODUCT GROUPING SCHEME:
- GROUP 1: NOBLE GASSES AND 'INtRT' (NON-RADIUACTIVt) AEr.0SOLS
- GROUP 2: CSI
- GROUP 3: TELLURIUM (TAKEN TO BE ELEMENIAL TE)
- GROUP 5: MOLYBDENUM (TAKEN TO BE N003)
- GROUP 6: CSOH
- STRUCTURAL MATERIAL GROUPING SCHEME
- USED IN CORE NODES (IdACKED IN CONIAINMtNT AS LUMrED GROUP 1 AEROSOLS)
- GROUP 1: CD
- GROUP 2: IN
- GROUP 3: AG
- GROUP 4: SN
- GROUP 5: MN
- NEW*****************:
A-24 01 .0428 FRACTION OF FISSION PRODUCT POWER IN GROUP 1 02 .222 SAME FOR GROUP 2 03 .0467 SAME FOR GROUP 3 04
- SAMt FmR GROUP 4 05 .0**5I
.07 SAME FDR GROUP 5 06 .0415 SAME FOR GROUP 6 07 433. INITIAL MASS OF FISSION FRODUCTS IN GROUP 1 (NOBLES CNLY) 08 23.49 INITIAL MASS IN GROUP 2 09 40.92 GROUP 3 10 181.5 GROUP 4 (TOTAL KG OF SR AND BA EXFRESSED AS OXIDES) 11 552.6 GROUP 5 (EXPRESS AS KG OF TRI-0XIDE) 12 250.0 GROUP 6 13 116.8 INITIAL MASS OF CD IN CORE (STRUC MATERIAL GRUUF 1) 14 305.4 INITIAL MASS OF IN IN CORE 15 1868.6 INITIAL MASS OF A6 IN CORE 16 308.6 INITIAL MASS OF SN IN CORE 17 6.68 INITIAL MASS OF MN IN CORE
- 18 NOT USED
- AUXILIARY BLUILDING
- CAN MODEL A MAXIMUM OF 5 SERIALLY-CONNECTED NODES
- THESE RECEIVE FLOW FROM THE CONIAINMENI FAILUKE ANU, IN V-SEQUENCES:
- FROM THE PRIMARY SYSTEM BREAK (S)
- YOU NEED SUPPLY INFORMATION ONLY FOR THE NO. OF NODES YOU SELECTED IN
- THE ECONfROL SECTION--NOIE 1HAL, FUR EXAMPLEe 1HE MASS OF WATER
- INITIALLY IN NODE 1 ALWAYS GOES IN INFUI NO. 6 NO MATTER HOW MANY
- NODES YOU ARE USING 01 421.4 VOLRB(I) VOLUME OF NODE 1 02 711.8 NODE 2 03 2638.2 NODE 3 04 1380 2 NODE 4
- 05 O. NODE 5
- MASS OF WATER CAN BE USED TO REFRESENT, FOR EXAMrLE, REFUELING FUULS
- IN BWR'S (NORMALLY NONE):
06 0.0 MWRB(I) MASS OF WATER IN NODE 1 07 0.0 NODE 2 08 0.0 NODE 3 09 0.0 NODE 4
- 10 0.0 NODE 5 11 0.0 AWATRB(I) SUKFACE AREA 0F WATER FOOL, IF ANT IN NODE 1 12 0.0 NODE 2 13 0.0 NODE 3 14 0.0 NODE 4
- 15 0 0 NODE 5 ,
- AT FRESENT, ONLY ONt EXIERIOR WALL FER NODE IS MODELED IN THE AUX CODEi 1
- IE YHE WALL HAS AUX CONDITIONS ON ONE SIDE AND IHE ENVIRONMENT ON lHE :
- APPROPRIATE MATERIAL PROPERTIES l 16 211.8 AHSRB(I) ONE-SIDED WALL AREA FOR NODE 1 i 17 357.8 NODE 2 18 1266.6 NODE 3 19 710.2 NODE 4
- 20 0. NODE 5 21 .4572 XHSRB(I) THICKNtSS FOR NODE 1 WALL 22 .4572 NODE 2 23 .4572 NODE 3 24 .4572 NODE 4
- 25 0. NODE 5
A-25 26 1.59 KHSR8(I) THtRMAL CONUUCTIVITY OF WALL IN NODE 1 27 1.59 NODE 2 28 1.59 NODE 3 29 1.59 NODE 4
- 20 0. NODE 5 31 656.7 CFHSRB(I) SFECIFIC HLAT OF WALL IN NODE 1 32 656.7 NODE 2 33 656.7 NODE 3 34 656.7 N0DE 4
- 35 O. NODE 5 36 2.65 ZHSRB(I) HEIGHT 0F WALL (FOR NC CALCS) FOR NODE 1 37 4.48 NODE 2 38 16.61 NODE 3 39 8.69 NODE 4
- 40 0. NODE 5 41 2308.5 DHSRB(I) DENSITY OF WALL IN NODE 1 42 2308.5 NODE 2 43 2308.5 NODE 3 44 2308.5 NODE 4
- 45 O. NODE 5
- A FORCED OUT FLOW AND/uk A FGHCED IN FLOW
- THIS FLOW IS ON UNflL THE FIRE DAneER SETPUINi(SEE Ett0W) IS
- REACHtD IN A COMPARIMkNT--lHIS SHUTS FLOW DOWN IN lHAT COMPT
- NOTE THE AC F0WER EVENf CODE DOES NOT AFFECT THE Sbis FLOWI TO
- VENTILATION FLOW OUT OF NODE 1 47 0.0 NODE 2 48 0.0 NODE 3 49 22.26 NODE 4
- 50 0.0 NODE 5 51 17.96 WVIRB(I) FORCED VOLUMETRIC VtNIILAIION FLOW INTO NOLL 1 52 0.0 NODE 2 53 4.30 NODE 3 54 0.0 NODE 4
- 55 0.0 NODE 5 56 200.6 ASEDRB(I) AEROSOL SETTLING AREA FOR NODE 1 57 200.6 NODE 2 58 200.6 NODE 3 59 200.6 NODE 4
- 60 4900. NODE 5
- AEROSOL IMFACTION DATA
- SEE DISCUSSION OF IMPACTION FARAMETERS IN
- ANNULAR SE TION AbOVt
- IF IMPACTION IS MODELLD IN A NODE, lHE InF ACTION AFEA, DI ANETER (EG GRATE
- THICKNESS), AND FLOW AkEA MUST ALL E.E GIVLN 61 0. AInFRF(I) IMPACTION AREA FOR NODE 1 62 0. NODE 2 63 0. NODE 3 64 0. NODE 4
- 65 0. NODE 5 66 0. XDIMNB(I) IMVACTION DIAMtTER FOR NODE 1 67 0.0 NODE 2 68 0.0 NODE 3 69 0.0 NODE 4
- 70 0.0 NODE 5 71 0. AGRARB(I) GRATE FLCu AREA FOR NODE 1 72 0.0 NOUE 2 l 73 0.0 NODE 3 1 74 0.0 NODE 4 l
l l
-r
A-26
- 75 0.0 NODE 5
- SPRAYS (EG FIRE SPRAYS)--THESE ARE TURNED ON AND OFF USING EVENS .
- CODE 240 NANUALLY--NO AUTOMATIC INITIATION 76 0.0 WSPRB(I) SPRAY MASS FLOW RATE FOR NODE 1 77 0.0 NODE 2 78 0.0 NODE 3 79 0.0 NODE 4
- 80 0.0 NODE 5 81 0.0 XHSPRB(I) SFRAY FALL HEIGHf FOR NODE 1 82 0.0 NODE 2 83 0.0 NODE 3 84 0.0 NODE 4
- 85 0.0 NODE 5
- INITAL CONDITION DATA 86 305. INITIAL TEMPERATURE OF AUX BUILDING 87 305. AUX BLDG SPRAY WATER TtMP 88 1.D-3 AUX BLDG SFRAY DROF DIAMtTER 89 .5 INIIIAL RtL HUMIDIlY OF AUX BUILDING COMPTS 90 311. ENVIRONNt'N TEMr 91 1.D5 ENVIRCNMENT/ AUX BLDG FRESSURE 92 355. FIRE DAMFER ACTIVATION TEMr (fur IN 0 TO SHUI DUWN SGTS (SEE
- A80VE)I PUT IN A VERY HIGH NO. IF NO FIRE DAMPERS)
- IF DESIRE TABULAR OUlFUI AND OFERATOR INiERVtNIIONS IN BNITISH UNITS
- INSERT A *BR HERE (EG FOR A PARAMEIER DUMP IN BKITISHe OPERATOR
- INTERVENfIONS, ANU TABULAR OUfFUI)
- SI
B-1 APPENDIX B Steam Generator Tube Integrity Analysis B.1 Introduction In the previous sections, the analyses show that the steam generator tubes can be subjected to temperature and pressure conditions beyond the design basis for certain severe accident sequences. After core uncovery, natural circulation in the primary system can bring hot steam and hydrogen into the steam generator inlet plenum and the tubes. Heatup of the tubes is retarded by heat transfer to the secondary side, so gas temperatures will exceed the tube temperatures. At this time, primary system pressure could be at the pressurizer safety or relief setpoints, while the secondary side could be at the steam generator safety or relief points. The combination of high pressures and potentially high tube wall temperatures thus raises the possibility of tube failure and containment bypass. The purpose of this section is to provide information related to the strength of the steam generator tubes at these conditions, from which conclusions related to their integrity can be drawn.
B.2 Tube Degradation During the course of operation, tubes may experience some degradation, and the effect of such degradation must be considered in evaluation of tube integrity. On the basis of operating experience, tube degradation may be separated into two distinct categories: 1) denting and 2) thinning and/or cracking. The effect of these two types of defects on tube integrity are discussed in the following sections, in light of the predicted severe accident loadings. l I
Numerous studies have reported on the various modes of tube degradation and the burst strength of steam generator tubing. The results discussed '
below are based on test data reported in Reference (B-1). All of the tests were conducted in a simulated steam generator environment at 600 degrees F
B-2 for 0.875 inch diameter and 0.05 inch wall thickness Inconel-600 mill-annealed tubing.
B.2.1 Tube Properties The following sections summarize results of tests conducted with steam generator tubes in research facilities to determine the burst strength of new and degraded tube at a temperature of 600 degrees F whereas the pre-dicted wall temperature for severe accidents may be substantially higher.
In this section, a method of relating these data to the accident case is presented. Theories on ductile failure indicate that the flow stress is approximately proportional to the sum of the material yield and ultimate strength (B-2). The strength properties of mill-annealed Inconel-600, at elevated temperatures, taken from Reference (B-3), are shown in Fig Are B-1 and are summarized below for a few selected temperatures.
Temperature Yield Ultimate Flow Stress Flow Stress (F) (ksi) (ksi) (ksi) Ratio 600 43 94 137 -
1000 41 82 123 0.90 1350 30 47 77 0.56 The correlation of material properties shows that the expected burst strengths at 1000 degrees F would be 90% of those reported in the tests; the expected burst strengths at 1350 degrees F would be 56% of those reported in the tests. The reported data at 600 degrees F can therefore be used and adjusted for the predicted higher severe accident temperatures.
B.2.2 Thinning / Cracking Type Defects In the case of thinning the defects were produced by machining uniform-ly on the tube outside diameter (OD). Combinations of three different penetration depths and four defect lengths were tested, in addition to the undefected virgin specimens. Results of these tests are shown in Figures B-2a and B-2b. As indicated by these results for a tube with a 1.5 inch
B-3 100 % , , , [ , , ,
T'%s %
3 ,n ,t , %
90 *
"#fh -
I e 80 -
- i
' ~
c 0
~
) '/0 -
[
l 5 '
i w .
l 60 - l -
. . I (SO -
(O hk3f..ry,ngt f e r) , _ _ = = - - - ., . . - . .-
g -. -..._-..._-%.
- = = ' '-
- 40 " Elongation ' ' i -
j ,
I
$ l .
- 30 - >
,---- -g-- -. -- - -
.n 20 - ; -
10 -
l O I I I l I l I I O 200 400 600 800 1000 1200 1400 1600 1800 Temperature. 'F Figure 8-1 High temperature tensile properties of annealed (1600 F/l hr.) hot-rolled plate (8-1).
B-4 to -
So : ;
% 9 37 -
e oo w
g5 -
=
, e
< e
['
g4 -
E 9 3 3 o UNDEFECTED o 25-30*. '
2 -
55 60'.
- 7 5 8 0.
1 -
0 ' ' ' ' ' ' _
0 0 25 0 50 0 75 1 00 1 25 1 50 DEFECT LENGTH tlNCHES)
Figure B-2a Burst data for 0.875 x 0.050 uniform thinning specimens - defect length variation (B-1).
10 -
b s -
O ~
q o 7 @
w
- g. -
=
- g
['
g4 -
l3 0 UNDEFECTED t
- a 2 -
o 3/16- TH'N NE O LE NGTH o 3, S T HIN N E O s i *.G T et \
e 3 / 4-* TMia.N E D L E *.G T H g a 1 1/2" TMINNEO LE NGTM 0 ' ' '
O 10 20 30 40 50 40 70 to 90 100 MAXIMUM DEGR AQ A TsON t'. WALU Figure B-2b Burst data for 0.875 x 0.050 uniform thinning specimens - defect depth variation (B-1).
B-5 long defect 50 percent of nominal wall thickness is required to withstand the postulated event loadings when the degradation in flow stress at high temperatures is considered. That is, 5 ksi
- 0.56 = 2.8 ksi which exceeds any possible AP across the tubes, given the rather high 1350*F temperature.
Considering a AP of 1250 psi present in the reactor, a good margin of safety exists. For shorter and/or nonuniform defects, the allowable degradation would be somewhat larger than 50 percent.
To study the effect of tight cracks the defects were simulated mechan-ically by EDM slots. Slot lengths of 0.25 inch, 0.5 inch, and 1.5 inch and slot depths in the range of 25-30 percent, 55-60 percent, and 85-90 percent '
of nominal wall were tested. The results of these tests are shown in Figures B-3a and B-3b. Again, from these results, it is seen that a tube with a 1.5 inch long and 50 percent through-wall crack would be able to withstand the postulated event loading without rupturing.
B.2.3 Tube Denting Because of the inherent mechanism of denting, namely the corrosion of carbon steel tube support plates, the dented region of tubes is confined within the support plate thickness. The denting was therefore simulated by hot-swagging a carbon steel ring onto the tube. The ring thickness was 0.75 inch, simulating the thickness of the support plate. The nominal denting depth of 0.04 and 0.05 inch.
To investigate the effects of tube thinning near a dent, tests were also performed on specimens with denting plus uniform thinning and denting plus elliptical wastage defects.
Results of burst pressure tests with various denting and thinning configurations are summarized in Figure B-4 and Table B-1. These results clearly indicate that denting has no degradation effect on the burst strength of a tube with either the nominal or wasted wall. A similar )
conclusion has been derived regarding the strength of a dented tube with l
superimposed cracks (B-5). Thus, the observations made previously regarding l l
B-6 r,
_____ l ll 8 .
S =
c 0
= %
y7 - a3 w
5*
o G g Es - %
t o o
,4 _
! o E3 0 UNDEFECTED 2 - 0 1/4" SLOT ,
o 1/2~ SLOT g ,
a 1 1/ 2" SLOT 0 ' ' ' ' i r e i t 0 10 20 30 40 50 so 70 80 90 100 MAXIMUM DEGR AD ATION I'. WALL)
Figure B-3a Burst data for 0.875 x 0.050 EDM slot specimens - defect depth variation (B-1).
n to -
g y c===I e ...__)
o U e -
o
!? %
le -
a
. O s e g4 .
l3 ,
o uwoareCTeo o as sos a - o as-sos a ss. sos
- 1 .
o > > > > - ,
o o as o so o 7s i oo i as i so OEFECT LENG7M NCHES Figure B-3b Burst data for 0.875 x 0.050 EDM slot specimens - defect length variation (B-1).
B-7 4
10 -
] *
.o.
e >
j .
O e .
8 -
n 9
- 7 5 w
g6 - e y*
i 0
9 e
g4 -
= v 3 3 0 UNOFFECTED
- O 25 30.
e n SS 609.
2 -
1
- 75 80%
i 5
a 1
aa ' '
0 .T3 ' ' '
O 0.25 0.64 0 76 1.00 1.25 1 SG DEFECT LENGTH (INCHES)
Figure B-4 Burst pressure data of 0.875 x 0.050 uniform thinning specimens with and without denting (B-1).
i 1
1
B-8 Table B-1 COMPARISON OF BURST PRESSURES OF ELLIPTICALLY WASTED
.875 OD x .050 WALL TUBING WITH AND WITHOUT DENTING (R-1)
Radius of Wrap Burst Pressures (psi)
Cutter Angle Depth (in.) (degrees) (% Wall) Without Denting With Denting 24 0 25-30 8155, 8150 7905, 8150 24 45 25-30 7615, 7940 9200, 8060 24 135 25-30 5200, 7780 7800, 7770 12 0 55-60 5550, 5635 5520, 5605 12 45 55-60 5680, 5840 5650, 5675 12 135 55-60 5610, 5455 5320, 5230 1
l
B-9 the required wall thickness to withstand the event loads are unaffected by the presence of denting.
B.3 Creep Rupture The potential for creep rupture must be evaluated given the possible corrbination of high temperatures and pressures. Pressure differences of 1230 to 1300 psi can exist across the tubes when the primary and secondary sides are both at the safety setpoint pressures. A pressure difference of 1215 psi can exist if both sides are at the respective PORV setpoints.
Given the tube dimensions, this translates to nominal hoop stresses between 9330 and 9980 psi. Creep rupture times for various tube temperatures can be estimated given these stresses.
Creep data for Inconel is presented in Figure B-5 and summarized below.
At temperatures of 1150'F and 1350*F, corresponding roughly to 900*K and 1000*K respectively, rupture times are seen to be on the order of a thousand hours or greater for the stresses considered.
Rupture Stress (ksi)
Temperature 1000 hr. 10,000 hr.
'F,(*K) Ruoture Rupture 1150 (894) = 15 = 12 1350 (10C6) 9 7 Additional creep rupture data for Inconel-600 from Re f. (B-8) is summarized in Table B-2. Data are provided here for shorter rupture times, corresponding to higher temperatures or stresses. The 1000 hour0.0116 days <br />0.278 hours <br />0.00165 weeks <br />3.805e-4 months <br /> rupture data are consistent with that of Figure B-5. According to this table, a temperature of 1500*F (1089'K) is required for a 10 hour1.157407e-4 days <br />0.00278 hours <br />1.653439e-5 weeks <br />3.805e-6 months <br /> period to rupture tubes with a 10 ksi stress. Thus, for the accidents considered, creep rupture is not a credible phenomena.
l
B-10 l
1000 hr - 410 31 +
io cco,n, '
fv,0 10 l l ,
.t... 304 ,ere e
, s s ,30, , , ;
j , i 309 '
4I0# ! r 316 3I6 i i 10 I
430' I, / . 3
\ i ?
o - 04y g ggg g ,
,g gg g g, , ,
8 ,
\\ \ ,_ ,,,,, , y , \ ,\\\ g _,ae='oy W
\. ,
- i. IX l I \ \\%p ,n ;,, ;
j '
I /' i inconely 43nj
! / . I Type 446j/
stainless Type 446 I l
steiniess l l l 800 iOOO 1200 1400 1600 1800 BCO iC00 1200 1400 16CO iSCO j Testing temperature, F Testing temperature, F 40 ,
01% Creep l
per iO O O ar
_ 304 A N N .
e Type 316 steintess l ko c
t
' ! \N\
b l
i l
i l
i i l
h I h ND \\ \l l i I j f4 a
Nickei \ /-430 A w
j-'nconel l :
l 1 N( N N / i l ; l
/ \ ' /
i "5
\ N~3 io '
)
700 800 900 iOOO 110 0 I200 i300 14C0 ISCO 1600 170 0 3;; 1 Tessing temce,ature, W Figure B-5 Creep and creep-rupture comparisons (B-4).
_. , _ _ _ - _ . _ . . _ _ _ _ . _ _ _. _ ____o
B-ll
. Table B-2 CREEP RUPTURE DATA FOR INCONEL-600, HOT ROLLED, AND ANNEALED AT 1600*F Rupture Stress (ksi)
Temperature 10 hr. 100 hr. 1000 hr.
'F,('K) Rupture Rupture Rupture 3 1100 (866) 48 35 22 1200 (922) 33 22 14 1300 (977) 22 14 8.6 1400 (1033) 13 8.6 5.8 1500(1089) 10 6.5 3.9 b
_=. ._ _- . _
B-12 Prediction of the failure time given a temperature and stress, for failure times below 1000 hours0.0116 days <br />0.278 hours <br />0.00165 weeks <br />3.805e-4 months <br />, is possible by correlating a parameter combining time and temperature with the stress. The Larson-Miller parameter is such a parameter [Ref. (B-6)]. For 316 stainless steel, it has the form LMP = T*(20 + log t )
- 10-3, where T is in temperature in degrees Rankine, r
t is the rupture time in hours, and log represents the base 10 logarithm. l r
A master rupture curve for 316 stainless steel is shown in Figure B-6, as !
presented in Ref. (B-7). The Inconel data from Figure B-5 maps onto the !
same curve as the ASME N-47 data for stainless steel, and from Figure B-5 it can be seen that these two materials exhibit very similar creep rupture behavior. Thus, for example, assuming stress of 10 ksi, a Larson-Miller parameter value between 40 and 41.5 is obtained. Given a temperature of 1500'F = 1960*R = 1090'K, a rupture time between 2.5 and 14.9 hours1.041667e-4 days <br />0.0025 hours <br />1.488095e-5 weeks <br />3.4245e-6 months <br /> is obtained. Checking the validity of this extrapolation, at 10 ksi and thus the same Larson-Miller parameter range, for tr= 1000 hours0.0116 days <br />0.278 hours <br />0.00165 weeks <br />3.805e-4 months <br /> a temperature between 1280*F and 1340*F is predicted. This agrees quite well with the value 1325'F from Figure B-5 and an interpolated value of 1275*F from Table B-2. Therefore, the extrapolation approach is seen to be valid, and temper-atures in excess of 1500*F would be required to fail tubes by creep rupture, since the time at elevated temperatures is well under an hour.
B.4 Summary Based on the material properties of Inconel-600 at elevated tempera-tures and the results of the test program on burst strength of steam genera-tor tubes, it can be shown that the combined pressures and temperatures at which the tubes can continue to maintain integrity are within the predicted conditions for severe accidents described elsewhere in this report. The results are applicable to conditions of tube degradation up to 50 percent uniform through-wall wastage as well as for crack profiles which represent the limits of operation.
Creep rupture is not considered to be a failure mechanism for steam generator tubes at the pressure, temperature and time conditions predicted for severe accidents. This conclusion is based on three independent data
b
! Master Rupture Curve for 316 Stainless Steel
- 10 0 --
i i i i
i i
n-4 hag g
',y ' x ,
ASME Code Case N-47 g Larson-Miller data / ,*
- q' m f or 18Cr- 8Hi S.S a Ei o 1
6 N j u 10 Ii - 700 K (800 ar) - a sy
-U X 755 K (900 *F) '
-s
, O 0 - 811 K (1000 *F) .'[' '
- () - 922 K (1200 *F) "'
- y
~
A - 1033 K (1400 "f)
W V - 1089 K (1500 "F) e
+*
i W
- j_ _ __ _ i __.__.l_.___ l_________ _1_ _ __ . . _ l __L _ . _ .
) 15 20 25 30 35 40 45 50 Larson-Miller parameter (T(20+1ogt r) x 10-a_
l Figure B-6 Master creep rupture curve for 316
- l stainless steel, taken from Ref.
jt (B-7).
B-14 sources, which indicate that temperatures in excess of 1500'F must be sustained for long periods for creep rupture.
B.5 References (B-1) M. Vagins, et al., " Steam Generator Tube Integrity Program - Phase I Report", NUREG/CR-0718, PNL-2937 (September,1979).
(B-2) R. J. Eiber, et al., " Investigation of the Initiation and Extent of Ductile Pipe Rupture (Final Report)", BMI-1908 (June,1971).
(B-3) " Source Book on Industrial Alloy and Engineering Data", American Society for Metals.
(B-4) " Metals Handbook", 8th Edition, A,nerican Society for Metals.
(B-5) D. L. Harrod and D. A. Kaminski, " Burst Strength of Dented and Cracked 1600 Steam Generator Tubes", Westinghouse R&D Memo 79-102-DEFIN-MI (January, 1977).
(B-6) F. R. Larson and J. Miller, "A Time-Temperature Relationship for Rupture and Creep Stress", Transactions of the ASME, pp. 765-775 (July,1952).
(B-7) V. Shah, Presentation to IDCOR/NRC Exchange Meeting on Direct Con-tainment Heating (April 23,1986).
(B-8) "Inconel-600", Technical Bulletin of the International Nickel Company, Inc., 1969, cited in Nuclear Systems Materials Handbook, Volume 1.
l i
l l
C-1 APPENDIX C Estimation of Steam Generator Tube Wall Temperatures MAAP 3 does not contain a detailed model for natural circulation on the secondary side of the steam generators. An overall heat transfer coeffi-cient is calculated in the code which is used to represent the loss of heat from the tubes to the secondary side gas, but this does not consider the detailed behavior in the vicinity of the hottest tube surfaces. This area lies just above the tubesheet on the inlet plenum side on those tubes which carry fluid from the inlet plenum out to the outlet plenum ("out" tubes).
The purpose of this appendix is to determine the relative magnitude of the heat transfer coefficients on the primary side of the inlet part of the out tubes to that on the secondary side. If, as will be shown to be the case, these heat transfer coefficients are comparable, the tubes will achieve a temperature about half-way between the inlet primary side gas temperature and the average secondary side gas temperature.
The procedure to be used is to calculate the natural circulation flows on the primary and secondary sides and to use these flows to calculate the heat transfer coefficients. Since both flows are driven by the high tem-peratures which exist near the inlets of the out tubes compared to the cold temperatures existing elsewhere, scaling arguments can be effectively used to compare the flows and heat transfer coefficients.
As the primary side flow Wp (see Figure C-1) enters the out tubes, it rapidly cools off from a temperature Tp to a temperature which is essential-ly the mixed average secondary side temperature T . If we assume a roughly s
linear drop in temperature occurs over some length t, the magnitude of the flow is given by equating the total frictional pressure drop required to move fluid a distance 2L on the primary side to the difference in hydro-static head created due to the high temperatures in the zone of length t at the inlet of the out tubes:
C-2 T
/ '
[ \
n I l i O 3
~
,, n "
Tp inlet outlet plenum plenum e
/ ,
w Figure C l' Natural circulation flows on the inside and outside of a tube carrying fluid from the inlet to the outlet plenum.
vy- - -+r_ . . - - . - - - . . - . - - - , , .-.~-----w..-,,---- ...-.,-----,---.---..-----,------c----- - - - - - -- ,-
C-3 g8 L (C-1) pDA p 2(,+(),,)2)"#p where, op = average density of fluid on the primary sides of the tubes, Dp = hydraulic diameter on primary side (tube diameter),
A = total tube flow area, t
a = fraction of tubes carrying out flow; (1 - a) carry the return flow back from the outlet plenum, f = friction factor, S = coefficient of volumetric expansion, g = acceleration of gravity.
The temperature decrease in the fluid aT p is given by aTp Wp cp =O gg (C-2) where, c = specific heat, I p
Q3g = total heat loss to secondary side.
Substituting l 2 '
A SG p t
! W =o p 98 j (C-3) p p fl[ j y + \
- (a (1 - a)2/
On the secondary side, we have by the same reasoning l
C-4 2~
W aT
" #s 98 1 (C-4) 2a 2 2s s 0 s
2 where, aTs Ns cp =Qg 3 (C-5) or W 3"D 98 OcG a Af ~
(C-6) s 2f
{p where A s is the flow area associated with the inlet half of the tube bundle.
For simplicity, we have ignored the typically small differences in volu-metric expansion coefficient, specific heat, and friction factor between the primary and secondary sides. If we divide Equation (C-3) by Equation (C-6) we obtain
[ ) }l/3 l+ a 1
/
For typical values of a s .2, the last term is very close to 1. l The heat transfer coefficients can now be calculated by using the Dittus-Boelter correlation for turbulent flow in passages. Since super-heated steam has a Prandtl number of 1, and if we neglect small differences in thermal conductivity and dynamic viscosity between the primary and secondary sides we have:
f
,. , _ - . _ . - _ _ _ . _ ~ ,. -_ _ _ _ - _ . .
C-5 g.8 hs (C-8)
A*8 02 or s2 # ( (C-9)
For Seabrook:
At"II" '
As = 2.95 m2 Dp = .0155 m Ds = .038 m L = 17 m t s 1-2 m (based on the rate at which the primary fluid loses its heat to the tube wall in the MAAP calculations).
At the time of peak inlet plenum temperature, the primary side pressure is about twice the secondary side value and has an absolute temperature which is approximately 1.3 times as high so that
- h=1.5 Thus
[h g (1.2)(1.2)(.5)(1)(1.3) s
% .9
C-6 or '
hsyh p It should be noted that this conclusion is a relatively weak function of the assumptions (e.g., on the value of t), given the small values of the expo-nents in Equation (C-9).
As noted previously, if the heat transfer coefficients are equal, in steady-state the tube temperature will te half-way between the primary and secondary side gas temperatures. Further, by substituting in Equation (C-7), we obtain W3 s5W p so that there will be a relatively small variation AT in s secondary side gas temperature compared to AT p ; thus, the average secondary side temperature can conveniently be used for estimating the tube temperature.
1
RAI 48
Most of the work pertinent to severe accidents has addressed plant behavior l at full power, on the assumption that this represents the major contribution to risk. Also, WASH-1400 assumed containment failure was probable following a core melt, making containment bypass sequences relatively less important.
Therefore:
j a. Please address the possibility of accidents inside the containment building while in Modes 2-6 (Startup, Hot Standby, Hot Shutdown, Cold Shutdown, and Refueling) insof ar as these accidents could impact upon risk. In particular,
' consider the ef fect of reduced safety equipment availability and containment integrity requirement permitted by technical specifications while shutdown li or refueling.
l b. Event V and steam generator tube rupture provide a direct path from the RCS to the environment during severe accidents. Please describe the Seabrook
! work which identifies any other direct path.
- c. Please provide f urther information and/or specific references pertinent to I- release of radioactive material located outside of the containment building (e.g. spent fuel pool, radwaste systems) insof ar as the magnitudes are large i enough to impact upon the issue under consideration here.
4 r
RESPONSE 48 We would like to note that while it is true that WASH-1400 assumed a high probability of containment failure, it is not true that bypass sequences were l
relatively less important. Most of the containment f ailure probability in WASH-1400 for PWRS was assigned to the base mat penetration and leakage failure modes and these modes did not result in high contributions to early fatality l risk. The V-sequence in fact dominated category PWR-2 which, in turn, dominated j early fatality risk. Hence, the V-sequence type of bypass was very important in
- the WASH-1400 results.
i a. Modes 2-6 i
The Seabrook Station technical specifications generally require an equivalent f level of safety equipment to be available in modes 1, 2, and 3, consistent wi t h l
i the actual mode of operation, for the reactivity control systems, instrumentation, j coolant circulation, and ECCS subsystems. The technical specifications require that
! only one train of ECCS subsystems be available for mode 4 and that the accumulators be isolated. The saf ety injection pumps are not required to be operable in Mode 4. Technical specifications for containment integrity, containment cooling
' and support systems such as of f-site power, on-site power, component cooling and
.! service water pumps are generally the same for modes 1, 2, 3, and 4. Considering i
f
that the plant is in modes 2, 3, and 4 for only a small portion of time and the technical specifications generally require an equivalent level of safety equipment to be available, the SSPSA and RMEPS analyses account for the mode 1, 2 and 3 events. The ef fect of reduced safety equipment and reduced containment integrity requirements permitted by the technical specifications while shutdown or ref ueling are addressed in the response to request for additional information #21.
- b. Direct Paths For completeness, all sequences in the SSPSA were considered with regard to their potential for early release. As shown in Figure 1, there are four ways of having an early large release (Release Categories S1, S6 and S7). Each is described below:
Containment Bypass There are two types of containment bypass initiating events in the SSPSA.
The interf acing LOCA (V) is quantified in SSPSA Section 6.6 and an enhanced new model is provided in the RMEPS. As shown in Table 1, the V sequences contribute approximately 12% to early release frequency (1.3% of which is in release category S1 and 10.8% of which is in release category S7).
A steam generator tube rupture (SGTR) initiating event can result in an early bypass release if the secondary side is not isolated and the ability to cool the core is lost. The SGTR model has not changed since the SSPSA (SSPSA Sections 5.3.11 and 5.4.4). SGTR is an insignificant contributor to early release f requency for two reasons. First, the frequency of core melt is low (Aprox. 10-6 / ye a r) . This low frequency stems from a high degree of plant equi pment availability for core cooling and a very long time available for the operator to terminate break flow and maintain basic safety functions.
Second, the conditional frequency that the secondary side is open and the operators don't isolate paths to the environment is extremely low. Therefore, SGTR core melt sequences with bypass are insignificant contributors to early release frequency.
External Events i
External initiating events have the potential of f ailing containment and starting an initiating event. Two such events (aircraf t crash and turbine i missiles) are in the SSPSA model but neither provide a significant contribution j to early release frequency or early health risk. It should be noted that other external events were considered. For example, the seismic capacity of the Cont ai nment (SSPSA, Section 9.2) was found to be greater than 2g and excluded i from the detailed event tree quantification.
i
Aircraf t crashes into the containment that can potentially penetrate the containment are assumed to cause a large LOCA initiating event (SSPSA, Section 9.3). The mean frequency of core melt and early release is 1.03 x 10-10/ year (SSPSA, Section 13.2). This analysis has not changed since the SSPSA. As shown in Table 1 air crashes ( APC) make an insignificant contribution to early release frequency.
Turbine Missiles that penetrate containment are also assumed to cause a large LOCA (SSPSA, Section 9.9). Core melt frequency is assessed in the plant model (SSPSA, Section 5.3.7 and 5.4.3) with a mean core melt frequency equal to 5.23 x 10-10/ year (SSPSA, Section 13.2)
In conclusion, external events impacting containment are insignificant contributors to early release f requency as shown in Table 1 and are minor contributors to early health risk in RMEPS. To the extent that WASH-1400 (and NUREG-0396) risk curves do not explicitly contain risk f rom external events, the RMEPS results are conserv-ative in comparison.
Containment Isolation Failure Containment isolation system failure and failure to recover during a core melt would result in a potential early release. The systems analysis without recovery is documented in SSPSA Section D.13. The systems analysis model was updated to include operator recovery that was included at the sequence level in the SSPSA.
This update is provided in the Risk-Based Evaluation of Technical Specifications For Seabrook Station, PLG-0431 (Section 4.5) . Presently, the update and the original SSPSA analysis are being integrated. The top contributor to early release f requency and risk is f ailure of containment isolation caused by seismic initiating events that fail support systems such as the SSPS. As shown in Table 1 containment isolation (CIS) failure dominates early release frequency (89%, all of which is S6). Again, RMEPS risk is more complete WASH-1400 and (NUREG-0396) since risk from earthquakes were not explicitly included in WASH-1400.
Containment Structural Failure Containment structural failures due to core melt phenomena is assessed in SSPSA Section 11. This assessment includes an analysis of the ultimate strength of the containment, failure modes, and failure times as well as effects of hydrogen burns, and steam explosions. This analysis has not changed since the SSPSA. BNL reviewed the SSPSA severe accident phenomena, containment response, and radiological source terms and generally concurred (NUREG/CR-4540) that early f ailure of Seabrook containment is very unlikely. As shown in Table 1, steam explosion makes an insignificant contribution to early release frequency.
- c. Other Sources As discussed in Section 5.2.2 of the SSPSA (PLG-0300), Initiating Event Identification, if a significant radiological release occurs at a nuclear power plant , "...it must originate either in a damaged core or in a non-core source of radioactivity such as the spent fuel storage pool or the gaseous, liquid, and soild waste f acilities. Past releases from the experience and analysis (Reference 5.2-1) have clearly shown that core are by far the only significant source of risk at a nuclear power plant
...It is generally recognized that sources of radioactivity at the plant (other than the reactor core) and the possible mechanisms for their release are such as to provide a negligible risk of public health impact." We also note that contributions from class 3-8 accidents, modes 2-6 events, external events and many common cause f ailures were not included in RASH-1400 or NUREG-0396.
Ref erence 5.2-1 is NUREG/CR-0603, "A Risk Assessment of a Pressurized Water Reactors for class 3-8 Accidents," R.E. Hall, et al, October 1979.
l Table 1 Distribution of Large Early Release Frequency Release Mean Fraction of Sequence Mean % of Fraction of Category Frequency + Total Type
- Frequency + RC Total S1 5.8 x 10-9 .016 V (pipe break) 4.6 x 10-9 79 .013 Stm Expl 5.9 x 10-10 10 .002 TMLL 5.2 x 10-10 9 .001 APC 1.0 x 10-10 2 .000 S6 3.2 x 10-7 .889 Earthquakes /CIS 3.2 x 10-7 100 .889 S7 3.9 x 10-8 .108 V 3.9 x 10-8 100 .108 (RHR pump seal) 3.6 x 10-7
+ events per reactor year
- V- interf acing LOCA Stm. Expl- Reactor vessel steam explosion TMLL- Turbine missile resulting in large LOCA APC- aircraft crash into primary containment (resulting in large LOCA)
- CIS- containment isolation system failure RC- release category
-- .. .,--.m.-,.-- ..-.. - . --m,.--..._ _ . - , _ _ _ _m.. .--e_.,- y,,,,_-.m ,, , , _ , . . _ . _ _ - , , , _ . .
PRINCIPAL CONTRIBUTORS TO EARLY RELEASE FREQUENCY INITIATING EVENTS j
WITH CONTAINMENT l BYPASS 1
l
- INTERFACING LOCAs ,
- STEAM GENERATOR TUBE RUPTURE l
i EXTERNAL EVENTS WITH -
POTENTIAL CONTAINMENT DAMAGE FOR I
- AIRCRAFT CRASH > EARLY
- TURBINE MISSLE REMASE I
i .
' LOSS OF CONTAINMENT . ,
STRUCTURAL INTEGRITY ALL OTHER INITIATING EVENTS g CONTAINMENT ISOLATION FAILURE 3 l-
RAI 52
Page 3-7 contains a discussion of vault behavior in response to RHR system breaks. The emphasis is upon loss of equipnant due to flooding. What i consideration has been given to breaks which are small enough that the vault is not flooded, but there is a significant thermal energy release that 4
may impact equipment operation? Please include consideration that enough energy may be released to activate the fusible links in the ventilation system, thereby terminating ventilation and indirectly causing f ailure to pumps due to overheating of pump motors, and that this could occur at a time earlier than might occur due to flooding.
RESPONSE 52 on subsequent pages in this chapter (e.g. pp 3-22 thru 3-28 and referenced tables) it described how environmental damage to pumps in the RHR vault due to causes other than sube.ergence was assessed. Calculations are presented that show that the leak area of the RHR pump seals must be less than .09 in2 to enable the RHR vault sump pumps to keep up with the rate of floodi . In the event trees of Figure 3-4 and 3-5, leak areas of l 0 to .9 in are covered by sequences in which the top event L1 is successf ul (sequences No. 4-42 in the VI tree and 4-34 in the VS tree, respectively).
1 In addition to submergence by RHR vault flooding, consideration was given
. - to other f ailure modes to RHR vault pumps such as the thermal and moisture I
ef fects of the steam environment. The f ault tree in Figure 3-9 together with the quantifications in Tables 3-11, 3-12 and 3-13 of PLG-1432 were used to assess the probabilities of environmental failure of the RHR, CBS, and SI pump, respectively. The result of these assessment is summarized in Table 1. As seen in this table, there is a high probability assigned to environmental f ailure probability of each type of pump even for the range of 0 .09 in2 in which flooding of motors may not occur. While
- these assessments are highly subjective, they demonstrate adequate consid-1 eration of these failure modes and are believed to be conservative. Note that for larger leak rates there is some chance of non-submergence in the injection path (VI) event tree because of the possibility of operator action to ISOLATE the leaking check valves.
In spite of the conservative assessment of environmental damage of RHR vault pumps, it is clear that the conclusions of the sensitivity study i
are insensitive to these assessment. The total f requency of non-core melt
- sequences resulting f rom success of event L1 is on the order of 2 to 3 x 10-8/per reactor-year. Even if no credit were taken for any unflooded pumps in this vault, the results would be unaffected because the charging pumps located outside the vault would still be available for core cooling. Even if it is further assumed that the charging pumps would also fail during these sequences, there would be no impact on the conclusions of the sensitivity study because of the low frequency currently assigned to the non-melt sequences in which the pumps in the RHR vault I do not flood.
1
The conservative assessments of environmental damage in Table 1 accounts for the direct ef fects of water jets, steam, humidity and thermal damage. These effects were assessed to overshadow the rather indirect mechanism of fusible link actuation and overheating due to lack of ventilation. It is not clear whether the fusable links would actuate or not. Even if they did, the incremental thermal stresses acting on the pumps would be small in relation to the direct environmental stresses. Note that sensitivities were assessed on the impact of fusable links on source terms in Section 4 of RMEPS.
J TABLE 1. Assessment of Environmental Damage to Pumps in RHR Pump Vault RHR Pump Seal Environmental Failure Proability of RHR Vault Pumps Leak Area RHR Pumps CBS Pumps SI Pumps (in 2)
VI 0 .09 .55 .1 .1
.09-1.05 .85 .44 .33 1.05-2.6 1.0 .75 .64
>2.6 1.0 1.0 1.0
.V.S_
0 .09 .56 .11 .11
.09-1.05 1.0 1.0 1.0 1.05-2.6 1.0 1.0 1.0
>2.6 1.0 1.0 1.0 i
1
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. , , .- . - . - -