ML20136H916
| ML20136H916 | |
| Person / Time | |
|---|---|
| Site: | Seabrook, 05000000 |
| Issue date: | 09/30/1984 |
| From: | Bayless P, Chambers R EG&G IDAHO, INC. |
| To: | NRC |
| Shared Package | |
| ML20136H886 | List: |
| References | |
| CON-FIN-A-6354, FOIA-85-503, REF-GTECI-A-44, REF-GTECI-EL, TASK-A-44, TASK-OR EGG-NTP-6700, NUDOCS 8508200546 | |
| Download: ML20136H916 (78) | |
Text
. . _ _ _ . ._ _ . _ . - - . ._ - .-- . _ _ _ _ .- . -- . - --_. - -- ._.-
s 0 Y l?
l
, EGG-NTP-6700 l September 1984 l ANALYSIS OF A STATION BLACK 0UT TRANSIENT AT THE i SEABROOK NUCLEAR POWER PLANT j ,
/
4'L N
\ Paul D. Bayless Rosanna Chambers l
Idaho National Engineering Laboratory Operated by the U.S. Department of Energy
' 'N .
_ .. _ gj T 'E F ~ ~ '
j'
- Nst ,
N N ====a ummma imme, .,
.ly & -
' 4 m summmmamme e fM 2* ' l i ,% '_ ', E
~
m
. 7,-
- .s g =
. _ c:.s~yu,_
2 usumuurunsur gagg--- --
. gy,y I s 1
j n - WW,3* - -
~~ ~
(.;; s, . '
^
r This is an informal report intended for use as a preliminary or working document 8508200546 850727 hatt -503 PDR Q
l l
Pref.aredfortheNUCLEAR u.
Under DOE Contract No. DE-AC07-761001570 REGULATORY COMMISSION M EEEOldaho FIN No. A6354
_.__.__..__...__.._.4
, , e EGG-NTP-6700 ANALYSIS OF A STATION BLACK 0UT TRANSIENT AT THE SEABROOK NUCLEAR POWER PLANT Paul 0. Bayless Rosanna Chambers i
Published September 1984 EG&G Idaho, Inc.
l l
Prepared for the i U.S. Nuclear Regulatory Commission Washington, D.C. 20555 )
i Under DOE Contract No. DE-AC07-761001570 '
FIN No. A6354 i
d I
- - - - . - . . .- - - , - -n,,-
1 I
l ABSTRACT 1
A postulated station blackout transient at the Seabrook nuclear power l plant was analyzed in support of the Nuclear Regulatory Commission's Severe 1 Accident Sequence Analysis Program. The RELAP5/M002 and SCDAP/ MODI computer codes were used to calculate the transient from initiation through severe core damage. The base transient, the TMLB' sequence, assumed no offsite power, onsite power, emergency feedwater, or operator actions. -
Additional analyses investigated the sensitivity to the core modeling and a potential mitigating action.
s i
FIN No. A6354--Severe Accident Sequence Analysis (SASA) Program.
11
i *
SUMMARY
A postulated station blackout transient at the Seabrook nuclear power plant was analyzed. The analysis was performed at the Idaho National Engineering Laboratory as part of the Nuclear Regulatory Commission's Severe Accident Sequence Analysis (SASA) Program. The objectives of the SASA Program are to evaluate nuclear plant response for accident sequences that could lead to partial or total core melt and to evaluate potential mitigating actions.
With the plant operating at full power, a loss of offsite power occurs. This is followed by the failure to provide onsite power or emergency feedwater. No operator actions are taken. This scenario, which was the base transient for the analyses, has been designated the TMLB' sequence, and has been identified by the Accident Sequence Evaluation Program as a dominant sequence leading to core melt for pressurized water reactors.
The RELAP5/M002 computer code was used to calculate the system thermal-hydraulic response until fuel cladding temperatures exceeded 1000 K (1340*F), the onset of calculated cladding oxidation. The SC0AP/M001 computer code was used to calculate the core thermal, mechanical, and chemical behavior from cladding oxidation until fuel melting began.
The TMLB' sequence led to core heatup starting near 8300 s, and cladding oxidation beginning by 9100 s. Oxidation of 4% of the cladding
^ occurred before being stopped by steam starvation. Ballooning occurred over 80% of the fuel rod length. Fuel melting began by 14000 s. By the end of the calculation (15000 s), 84% of the cladding had relocated below the bottom of the fuel rods.
A calculation was performed to investigate the sensitivity of the calculations to the core modeling. Using a three channel core yielded little difference in the results, iii l
l l
Opening the power-operated relief valves (PORVs) to depressurize the system and allow accumulator injection prior to cladding oxidation was investigated as a mitigating action. Whether the PORVs were opened just after the steam generators could no longer remove all of the decay heat or after the core heatup began, the depressurization resulted in earlier oxidation than when they were not held open. The accumulator pressure was reached before fuel and cladding relocation began. The results also showed that the PORVs can be used to control the reactor coolant system pressure ,
during the core damage phase of the transient.
The analyses also demonstrated the need for a single computer code containing the thermal-hydraulic capabilities of RELAPS and the fuel damage capabilities of SCDAP.
The analyses presented are interim results. When the investigation is completed, a final report on the station blackout transient for the Seabrook plant will be issued.
)
4 0
iv
i <
CONTENTS ABSTRACT .............................................................. 11
SUMMARY
............................................................... iii
- 1. INTRODUCTION ..................................................... 1
- 2. INITIAL AND BOUNDARY CONDITIONS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3 Initial Conditions .........................................
2.1 3 2.2 Boundary Conditions ........................................ 3
- 3. ANALTIICAL RESULTS ............................................... 6 3.1 Base Transient ............................................. 6 3.2 Three Channel Care Case .................................... 26 3.3 Late Depressurization Case ................................. 36 3.4 Early Depre s suri zation Ca se . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 46
- 4. CONCLUSIONS AND RECOMMENDATIONS .................................. 50
- 5. REFERENCES ....................................................... 53 APPENDIX A--INPUT MODEL DESCRIPTIONS .................................. 54 APPENDIX B--CDMPUTER CDDE DESCRIPTIONS ................................ 64 FIGURES
- 1. Pressurizer pressure from the base case . . . . . . . . . . . . . . . . . . . . . . . . . . 9
- 2. Steam generator pressures from the base case ..................... 11
- 3. Steam generator downcomer liquid levels from the base case ....... 11
- 4. Pressurizer liquid level from the base case ...................... 12
- 5. Hot and cold leg liquid temperatures from the base case .......... 14
- 6. Hot leg mass flow rates from the base case ....................... 14
- 7. Core bypass inlet mass flow rate from the base case . . . . . . . . . . . . . . 15 y
- . - - . . . m_ _
I ,
l 1
- 8. Core void fractions from the base case ........................... 17
- 9. RELAp5-calculated fuel cladding surface temperatures from the base case ............................. .... 17
- 10. Steam temperatures in the core, upper plenum, and hot l egs f rom the ba se ca se . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18
- 11. Single loop hot leg steam temperatures from the base case ........ 18
- 12. RELAPS-calculated metal surface temperatures for the fuel .
I cladding, upper plenum, and single loop hot leg nozzle -
and steam generator tubes from the base case ..................... 20
- 13. SCDAP-calculated fuel cladding surface temperatures '
from the base case ............................................... 20
- 14. Fuel assembly flow areas from the base case ...................... 22
- 15. Fuel assembly hydrogen generation rate from the base case ........ 22
- 16. Fuel assembly soluble fission product release rate from the base case ............................................... 25
- 17. Ratio of the cladding oxidation heat release to the core decay heat from the base case ............................... 25
- 18. Core exit mass flow rates from the three channel core case ....... 28 ',
i
- 19. Mass flow rates out the top of three outer channel volumes from the three channel core case ......................... 28
- 20. Top core volume void fractions from the three channel core case and the base case ..................................~.... 29
- 21. RELAPS-calculated top core volume fuel cladding surface temperatures from the three channel core case and the base case ........................................... 29
- 22. SCDAP-calculated fuel cladding surface temperatures at node 4 from the three channel core case and the base case ............... 31 ,
a
- 23. Total core cumulative hydrogen generation from the three channel core case and the base case .............................. 33 .
- 24. Ratio of the cladding oxidation heat release to the core decay heat from the three channel core case ................. 35
- 25. Total core soluble fission product release rate from the three channel core case and the base case ............... 35 vi
- 26. Pressurizer pressure from the late de case and the base case .............. pressurization
............................. 37
- 27. Core void fractions from the late depressurization case and the base case ........................................... 37
- 28. RELAP5 calculated fuel cladding surface temperatures from the late depressurization case and the base case ............ 38
- 29. SCDAP-calculated fuel cladding surface temperature at node 1 from the late depressurization case and the base case ............ 40
- 30. SCDAP-calculated fuel cladding surface temperature at node 4 from the late depressurization case and the base case ............ 40
- 31. SCDAP-calculated fuel cladding surface temperature at node 9 from the late depressurization case and the base case ............ 41
- 32. Ratio of the cladding oxidation heat release to the core decay heat from the late depressurization case ................... 43
- 33. Total core cumulative hydrogen generation from the late depressurization case and the base case .......................... 43
- 34. Total core cumulative soluble fission product release from the late depressurization case and the base case ................. 44
- 35. Pressurizer pressure from the early depressurization case and the base case ........................................... 47
- 36. Core void fractions from the early depressurization case and the base case ..............:............................ 47
- 37. RELAPS-calculated fuel cladding surface temperatures from the early depressurization case and the base case ........... 48 A-1. RELAP5 nodalization of the Seabrook plant for the station blackout transient ....................................... 56 A-2. RELAP5 nodalization of the three channel core . . . . . . . . . . . . . . . . . . . . 57 A-3. Cross-sectional view of the three channel core ................... 58 A-4. SCDAP nodalization for the Seabrook station i blackout transient ...............................................
! 61 i i -
A-5. Cross-sectional view of a typical Seabrook fuel assembly ......... 62 l
t 1
vii l
. i i
s TABLES 1.
- 1. Comparison of computed and desired initial conditions . . . . . . . . . . . . 4 I 2. Sequence of events for the TMLB' sequence . . . . . . . . . . . . . . . . . . . . . . . . 7
,. B-1. RELAP5/M002 updates used in the station blackout analyses ........ 66 -
i E
e 0
1 I
l 1
4 1
i I
l t
i J
i l
}
h 1
i j .
i l
1 i i
i I
l- .
L 1
I t
s Vili
~
i 1 . - _ , _ - . _ _ _ _ _ _ , _ - . . . . . - . . _ . . - - _ _ _ _ _ - . . _ . - _ .
i 1
ANALYSIS OF A STATION BLACK 0UT TRANSIENT AT THE SEABROOK NUCLEAR POWER PLANT
- 1. INTRODUCTION A postulated station blackout transient at the Seabrook nuclear power plant was analyzed. The analysis was performed at the Idaho National Engineering Laboratory (INEL) as part of the Severe Accident Sequence Analysis (SASA) Program, which was formulated by the United States Nuclear ;
, Regulatory Commission (NRC) to evaluate postulated reactor accidents over a broad spectrum of accident sequences. These sequences may extend beyond the current design basis in terms of system failures, core damage, and release of fission products to the environment. The objectives of the SASA Program are to evaluate nuclear plant response for accident sequences that could lead to partial or total core melt, to determine the timing of significant events, to determine the magnitude and timing of fission product release from the fuel rods and the hydrogen generation rate, and to evaluate the effect of operator actions on accident mitigation.
The base transient is initiated by a loss of offsite power with the plant operating at full power. The diesel generators then fail to provide onsite power. It is further assumed that the steam-driven emergency feedwater pump fails to provide cooling water to the steam generators, and that operator intervention does not occur. This sequence has been designated the TMLB' sequence, and, although it is a very unlikely event, has been identified as a dominant core melt sequence by the NRC-sponsored Accident Sequence Evaluation Program.I An analysis of the TMLB' sequence for the Bellefonte nuclear plant, a Babcock and Wilcox reactor, has also been performed as part of the INEL SASA program.2 The transient was analyzed from initiation through severe core damage 3
using the RELAP5/M002 and SC0AP/M0014 computer codes. RELAP5/M002 was used to calculate the system thermal-hydraulic behavior until fuel cladding 1
l oxidation temperatures were exceeded. The SCDAP/ MODI code was then used to calculate the detailed core and reactor vessel upper plenum behavior until the fuel started to melt. The transient calculation was terminated shortly thereafter.
Three calculations were performed in addition to the base case transient. In the first, the modeled core was divided into three parallel flow paths to investigate the effect of the core modeling on the transient /
results. In two additional calculations, it was assumed that the operators opened both power-operated relief valves (PORVs) sometime during the -
transient. Opening the PORVs was investigated as a potential mitigating action in which the reactor coolant system (RCS) would be depressurized so the accumulators would inject er ding water, thereby delaying core damage.
The analyses presented in this report are interim results of a continuing investigation of the station blackout transient at the Seabrook plant. Plans for the analysis include using a merged RELAPS/SCDAP/ TRAP-MELT computer code to calculata the core damage portion of the transient, further investigation of potential mitigating actions, and consideration of multi-dimensional flows in the reactor vessel. When the analysis is completed, a final report on the station blackout transient at the Seabrook plant will be issued.
Seabrook is a Westinghouse-designed, four loop pressurized water reactor with U-tube steam generators and a rated thermal power of 3411 MW.
The core contains 193 fuel assemblies. The RELAPS model consisted of two loops. The loop with the pressurizer was modeled as a single loop, with the other three loops lumped together and modeled as one loop. The SCDAP model was a representation of one fuel assembly and the associated portion of the upper plenum.
- The following section describes the initial and boundary conditions used for the calculations. Subsequent sections present the results of the analyses, conclusions and recommendations based on the analyses, and references. Appendix A provides a more detailed description of the computer input models. The RELAPS/M002 and SCDAP/M001 computer codes are described in Appendix 8.
2 J
I t
- 2. INITIAL AND BOUNDARY CONDITIONS The initial and boundary conditions for the RELAPS and SCDAP calculations are described below.
2.1 Initial Conditions A RELAP5 calculation of steady state, full power operation was performed to provide the initial conditions for the transient. The initial
. conditions are summarized in Table 1. In general, the calculated values were in excellent agreement with the desired values. The low steam generator pressure (5%) and pressurizer level (10%) did not have a significant effect on the transient events or timing.
The initial radial temperature profile in the fuel was calculated 5
using the FRAPCON-2 computer code. The axial power profile was obtained from the Seabrook Final Safety Analysis Report.6 Both of these profiles were based on 275 effective full power days of operation, which corresponds to the end of the first cycle. The fuel radial temperature profiles for each region of the three channel core model were also calculated using FRAPCON-2.
The initial fuel rod temperatures, core void fraction, and core inlet fluid enthalpy for the SCDAP calculations were obtained from the RELAPS transient calculations. The SCDAP calculations were started shortly before the RELAPS-calculated maximum fuel cladding surface temperature reached 1000 K (1340*F). This is the temperature at which SCDAP begins to calculatecladdingoxidation. The initial fission gas inventory was calculated by SCDAP. The internal gas pressure at transient initiation was
- . calculated by FRAPCON-2. The perfect gas law was then used to calculate the pressure at the beginning of the SCDAP calculations.
2.2 Boundary Conditions At transient initiation, the reactor scrammed, the reactor coolant pumps tripped, the turbine stop valves started to close, and the main feedwater started a five-second flow coastdown. There was no charging I
3 I
TABLE 1. COMPARISON OF COMPUTED AND DESIRED INITIAL CONDITIONS
, Parameter RELAPS Desired i
Core thermal power (MW) 3411 3411 Pressurizer pressure (MPa) 15.5 15.5 (psia) 2250 2250 Pressurizer level (m) 7.96 8.82 (in.) 313 347
~
l Hot leg temperature (K) 598.5 598.6
(*F) 617.6 617.8 Cold leg temperature (K) 565.2 565.2
(*F) 557.6 557.6 Total loop flow (kg/s) 17741 17741 (lbm/s) 39111 39111 Steam generator pressure (MPa) 6.53 6.89 (psia) 948 1000 Steam flow (kg/s)" 474.2 476.7 (lbm/s) 1046 1051 Feedwater flow (kg/s)* 474.2 476.7 (ibm /s) 1046 1051
, Feedwater temperature (K) 500 500
(*F) 440 440 Steam generator fluid mass (kg)a,b 49216 49216 (1bm) 108500 108500
- a. Per steam generator.
! b. Shell side.
e l
5 I
i 9
4
i l
letdown, or emergency feedwater flow during the transient. The steam generator pressure was controlled by the main steam line safety relief valves; the steam dump to the condenser was unavailable, and the atmospheric dump valves were operable only by operator action. RCS j pressure control was provided by the safety relief valves and PORVs. The l PORVs are air-operated valves that operate automatically following a loss of offsite and onsite power. It was assumed that sufficient air pressure and battery power were available to keep the valves operating throughout the transient. The pressure was assumed to be constant during the
. SCDAP-calculated portion of the base transient, and the core inlet flow was adjusted so that the core dryout time was as close as possible to that calculated by RELAP5. The decay heat for the entire transient was calculated by RELAPS using the 1979 American National Standard.7 No operator actions were considered in the base transient.
The boundary conditions for the three channel core calculations were l the same as for the base transient RELAPS and SCDAP calculations, except j that the fuel assembly relative power in each region of the core was different from the core-average power used in the base calculation. The 0 enter and middle channels had above-average powers, while the outer channel had a below-average power. Appendix A contains more details on the three channel core model.
Two calculations were performed in which the PORVs were held open at ,
some point in the transient. In the first, they were opened when the upper plenum steam temperature was 5 K (9*F) superheated. The depressurization rate from the RELAPS calculation was extrapolated and input to the accompanying SCDAP calculation. In the second, the PORVs were opened shortly after the steam generators were no longer an effective heat sink.
l
, No SCOAP calculation was performed for this transient. The remaining boundary conditions were identical to the base and three channel core calculations, l
i 5
l
I F
- 3. ANALYTICAL RESULTS The results of the various station blackout transjent calculations are presented below. The base transient will be discussed first, followed by the three additional calculations. In the context of this analysis, core ,
damage is defined to begin when fuel cladding oxidation is first calculated, at a temperature of 1000 K (1340*F).
3.1 Base Transient The base transient was the TML8' sequence. A brief, general description of the plant response during the station blackout transient will be presented. This will be followed by a detailed discussion of the system behavior up to core damage, as calculated by RELAPS. The core damage phase of the transient, as calculated by SCDAP, will then be examined. .
Following the loss of power, the steam generator liquid was boiled off by the core decay heat. As the liquid was depleted, the heat transfer across the steam generator tubes decreased until it could'no longer remove all the decay heat, and the RCS began to heat up. The coolant expansion caused by the temperature increase raised the system pressure until the PORVs opened. The PORVs were large enough to relieve the RCS pressure, and -
the pressure did not reach the safety relief valve opening pressure of 17.24 MPa (2500 psia). The coolant temperature continued to increase until saturation was reached. Steam filled the higher portions of the system, with the liquid draining to the reactor vessel and the loop seals. As the boiloff continued, the core uncovered and began to heat up. The fuel rod cladding oxidized, ballooned, and ruptured. The heat generated by the cladding oxidation, together with the decay heat, led to melting of the ,
cladding and fuel. The molten material relocated below its original location.
The calculated sequence of significant events for the transient is 1 presented in Table 2. The RELAP5 calculation was terminated near 10000 s, about 900 s after the maximum cladding temperature reached 1000 K (1340*F).
i 6
l
, o !
l I
4 TABLE 2. SEQUENCE OF EVENTS.FOR THE TMLB' SEQUENCE Time Event (s)
Reactor scram 0 Reactor coolant pump trip 0
, Main feedwater coastdown begins 0 Turbine stop valves start closing 0 Loss of effective heat sink 4903 PORV i'.itial opening 5010 Hot legs reach saturation temperature 6514 Natural circulation ends 6797 Fuel cladding heatup begins 8289 Fuel cladding oxidation begins 9060 t
i Fuel cladding ballooning begins 9720 Fuel cladding fails 9820 Steam starvation in upper part of core 9840 Dryout at bottom of fuel 11000 l Zr0 rupture 11400 2
Zr-U-0 relocation 11400 i
Fuel melting begins 13860 Transient terminated 15000 4 e 4
i 4
7
The calculated RCS pressure is shown in Figure 1. The pressure decreased in the first 25 s because the heat transferred to the secondary coolant system during the reactur coolant pump coastdown was greater than the decay heat. This cooled the liquid in the loops, reducing its specific volume, and caused the steam in the pressurizer to expand and reduce the pressure. As the pump coastdown ended and natural circulation was being established, the heat transfer decreased, causing the loop average temperature to increase, resulting in a pressure increase. Natural .
circulation was fully established at about 200s. The gradual increase and decrease in the RCS pressure from 200 to 4000 s reflected the balance .
between the decay heat and the heat transfer to the steam generators. When the decay heat was larger, the pressure increased; when it was lower, the pressure decreased. The pressure increased after 4000 s as the steam generator liquid levels continued to decrease and the heat transfer decreased more rapidly. The steam generators were no longer an effective heat sink shortly after 4900 s, and the RCS pressure'then rapidly increased until the PORVs opened. The pressure cycled between the PORV open and 4
j close setpoints of 16.20 MPa (2350 psia) and 16.06 MPa (2330 psia),
respectively, until about 6600 s. The pressurizer filled with liquid near 5900 s, and the saturation temperature was reached in the hot legs shortly after 6500 s. Between 6600 and 7600 s, the flow out the PORVs was .
8 insufficient to reduce the pressure. The increase in the reactor coolant specific volume caused by the boiling in the core could not be relieved by the volumetric flow out the PORVs. Steam formed in the core displaced saturated liquid from higher parts of the RCS, which then drained to the core where it could be boiled. This continued until enough steam was drawn into the pressurizer that the volumetric flow out the PORVs was again sufficient to relieve the pressure. This occurred at about 7600 s, when i the void fraction upstream of the PORVs was about 0.75. The pressure was
! maintained between the PORV setpoints for the rest of the transient. .
The oscillations in the pressure during the first 4900 s of the transient were reflections of the steam generator pressure, which was
- cycling between the steam line relief valve setpoints. When the steam
! generator pressure decreased, the cold leg temperature decreased (following the steam generator saturation temperature), reducing the average temperature and the pressure. Conversely, when the steam generator 8
, e l
17 , , , ,
16.5 '- CleMing 2400 PCRV cycling hootup starts 2
o.
3
\ v.-
M C-v 16 - -
v
- 00 ,
Inifici value g,,
n so tur a tion 3 e 15.5 -
Clodding - N g % oxidation
- g Q.
2200 15 -
i)g Decay hoot greater then steam generator heat removal 14.5 .
0 2000 4000 6000 8000 10000 Time (s)
Figure 1. Pressurizer pressure from the base case.
1 9
1
O pressure increased, the cold leg and average temperatures increased, raising the pressure. The magnitude of the oscillations decreased when the pressure oscillations in the two steam generators no longer coincided.
Figure 2 presents the pressure in the two steam generato s. The pressures cycled between the main steam line safety relief valve setpoints of 8.27 MPa (1200 psia) and 7.58 MPa (1100 psia) throughout the transient.
The time it took to pressurize the steam generators from the closing setpoint to the opening pressure increased significantly after about 4900 s, when the steam generators dried out, because the heat transfer rate from the tubes to steam was much lower than that from the tubes to liquid.
The steam generator downcomer collapsed liquid levels are shown in Figure 3. The large decrease in level as the transient started was caused by the void collapse in the riser after the scram occurred, which tended to equalize the riser and downcomer levels. The levels in both steam generators then decreased fairly steadily to zero near 4900 s. The oscillations in level reflected the pressure oscillations. When the pressure decreased, there was more steam in the riser, and the downcomer .
level increased. Conversely, when the pressure increased the downcomer level decreased.
Figure 4 presents the pressurizer collapsed liquid level. The level response was coupled with the RCS pressure response for the first 5000 s.
When the loop average temperature decreased, liquid was drawn out of the pressurizer, causing the level and the pressure to drop. Similarly, a level increase resulted in a pressure increase. After the steam generators were lost as an effective heat sink near 4900 s, the coolant expansion associated with the heatup caused the pressurizer level to increase until the pressurizer was full at about 5900 s. The level was re-established -
near 6700 s, when steam was drawn in from the hot leg. The level decreased as rore steam was drawn in from the hot leg to replace liquid flowing out ,
the PORVs. The level increase after 7100 s was caused by the pressure increase described previously, which caused some of the vapor in the pressurizer to condense. The level decreased slowly between 7800 and 8400 s because there was sufficient water in the core to cause a relatively 10
i
- I I
S.5 , ,
y'.
e '"
l9 /
8 i ', f, -
9
' q q/
' 115 0
% h jj U w
e s
llis C
[ y/$ $- i v
u 7.5 - 110 0 ,
3 -
u a s E
a 1-Loop
- w O 3-Loop e C. -1050 6
7 - c.
1000 6.s " ' ' ' ' '
0 2000 4000 6000 ~
8000 10000 Figure 2. Time (s)
Steam generator pressures from the base cose.
is , , , ,
O A 1-Loop O 3-Loop m 40 E n v
to :
v i> -
30 3 >
.c. e 4
t-
- t 20 g I s -
a -
E a
10 0 ' ' ' * * * ' ~
O 2000 4000
'O 6000 8000 10000 Figue) 3. Time (s)
Sloom generator downcomer liquid levels from the base case.
11
- e i 3 6 6 Pressurizer fun S t oom fr om 50 Not leg 14 - * }_ -
a ^ .
E -
v 12 -- - -'O v e e e e
- 10 -
D
- 3g T-3 3 F g . Reactor coolant heetus .
7 a _J i
(:"uv.w: _: : .
4"- --20
- i . .
4 i
.I O 2000 4000 6000 8000 10000
~
Time (s) -
Figure 4. Pressurizer liquid level from the base cose.
12
, a rapid repressurization of the RCS after the PORVs closed. The PORVs were not closed long enough to allow liquid in the pressurizer to drain into the hot leg. As the core liquid inventory was depleted, the RCS did not pressurize as quickly, and the ' pressurizer level decreased more rapidly.
The pressurizer remained more than half filled with liquid until after the core heatup had begun. As was the case in the Three Mile Island accident, the pressurizer level was not indicative of what was happening in the Core.
The hot and cold leg temperatures are shown in Figure 5. The temperatures in the two loops agreed very well. The slight offset in the cold leg temperatures was the result of a difference in the steam generator pressures. The hot leg liquid vas subcooled until shortly after 6500 s, then remained at the saturation temperature for the rest of the transient.
The cold legs did not reach the saturation temperature until about 6750 s.
Figure 6 presents the hot leg mass flow rates. The setle on the figure is reduced to show the flow behavior after the reactor coolant pump coastdown. Following the pump coastdown, a steady natural circulation flow was established. The flow rate decreased when the steam generators dried out because the temperature difference between the hot and cold legs, which provides the driving force for the flow, decreased. Natural circulation flow through the loops ended shortly after the cold leg reached saturation at about 6750 s.
Figure 7 shows the mass flow rate from the lower plenum to the core bypass. Again, the scale is reduced to highlight the behavior after the pump coastdown. The flow was positive or near zero for the first 4900 s of the transient. When the natural circulation flow decreased near 4900 s, the flow reversed in the core bypass. The liquid in the bypass channel was cooler than that in the core, and the density gradient pulled water from the upper plenum into the top of the bypass channel while the cooler water flowed into the lower plenum. The density difference became more pronounced and the flow more negative when boiling started in the core at 13
o ,
1 l
1 640 , , .
A 1-Loop ho t leg O 3-Loop hot leg C 1-Loop cold leg
^
P ** '*9 7 = "'
^ 620 -
650 D e
c 6 '
6 3 3
o o 600 @ 6 6 e e c. .
- c. 600 E E
580 0 -
l
. s:n. . . . za. . . . n - _
-550 ~
560 10000 2000 4000 6000 8000 0
Time (s)
Figure 5. Hot one cold leg liquid temperatures from the base case.
i 2000 - ,
A 1-Loop - 4000 0 3-Loop m
1500 f "
3000
'7 -
E N
oi .
- 0 6 1000 -
2000 V
3 R o
500 (; 7 1000 M
0 O 3
i O -h; ;, 2 m 'x
- ,_ .n m, 0 0__
e e i i -1000
-500 - 8000 10000 0 2000 4000 6000 Time (s)
Figure 6. Hot leg mass flow roles from the base cose.
I I
l 14
i 10 0 . . i i e 4
t
- 200 I
m Q 50 , -
. goo (
N cn E 3
v 0 Core bomng begins
" ~
0 +- . - N - 0
)
2 e
_- -10 0 3 i
Bypots drains
-200
- 10 0 O 200C 4000 6000 8000 10000 Time (s)
Figure 7. Core bypass inlet mass flow rate from the base case.
i 15
L about 6300 s. The magnitude of the flow decreased after 8000 s as the bypass started to drain. The bypass flow remained slightly negative after the bypass had drained because colder steam still provided a density difference between tne core bypass and the core.
The void fractions in the six core volumes from 5000 to 10000 s are presented in Figure 8. Boiling in the core started near 6300 s, and the void fractions increased for the next 500 s. The void fractions then -
leveled off for about 1300 s as the liquid that was boiled in the core was replaced by water draining from the reactor vessel upper head and upper .
plenum and the loop piping. The void fractions resumed their increase at about 8100 s, proceeding from the top of the core down. The decreases in the void fractions just prior to their final increases was caused by liquid draining from the downcomer and the core bypass.
Figure 9 shows the fuel cladding surface temperatures from the RELAPS calculation from 5000 to 10000 s. The temperatures were very close to the core fluid temperatures until the core heatup began shortly before 8300 s.
As a volume dried out, the heat transfer coefficient decreased, and the fuel rod began to heat up. As was seen in the void fractions, the core heatup progressed from the top of the core to the bottom.
Figure 10 presents a comparison of the steam temperatures in the top of the core, the upper plenum, and the hot legs from 5000 to 10000 s. The hot steam leaving the core flowed through the upper plenum into the hot legs. The single loop hot leg temperature was higher than the three loop hot leg temperature because the open PORVs drew more steam into that hot leg. The steam temperature distribution in the single loop hot leg from 5000 to 10000 s is shown in Figure 11. As expected, the hotter steam was located between the pressurizer surge line and the reactor vessel. The steam temperature in the first section of the steam generator tubes was only slightly above the saturation temperature. The reactor coolant pump -
loop seals still contained water, so hot steam could not flow through the loops.
16
I l
t > >
g j ;;; ri E < :
A 0.3 m above core bottom O 1.0 m above core bottom e O 1.6 m above core buttom o X 2.2 m above core bottom ~
- 0.75 -
0 2.9 m above core bottom 8 7 3.5 m above core bottom b
o -
0.50 -
o L
o ;
O. 1
$ 0.25 -
_m - '. j cx- -
9 L -a u 0__ -
E&- ._ ." ~ , .
~
5000 6000 7000 8000 9000 10000 Time (s)
Figure 8. Core void fractions from the base case.
1600 , , , ,
6 0.3 m above core bottom O 1.0 m above core bottom 1400_- 0 1.6 m above core bottom --2000 X 2.2 m above core bottom m x 0 2.9 m chove core bottom b v
- V 3.5 m above core bottom .
1200 -
e c 5 -1500 $ -
- c .
O 1000 - c 6 6 e e O- c.
-1000 h 800'- h
- ~ ~
2 E = E E = = 0 -
600- - - -
- 500 400 5000 6000 7000 8000 9000 10000 Time (s)
Figure 9. RELAPS-col eul a t ed f uel clodding surf ace t empe r atur es from the base cose.
. 17
a ,
1600 , , ,
a Core. 3.5 m above bo tt om ,
O Upper plenum 0 1-Loop hot leg 1400 - _
X 3-Loop hot le9 2000 m M
v b
v 0 1200 - -
E
0 -1500 0
- 6 6 e e
- o. 1000 - -
- c. .
E E
- e 800-- - 1000 600 "" ~
5000 6000 7000 8000 9000 10000 Time (s)
Figu r e 10. Sloom t empera t ures in the cor e, upper plenum, and hot legs from the base cose.
900 , , , ,
A Near vessel O Noor surgo line q 1100 C Necr stoom generator X in steem generator tubes m 6 800 -
L1000 b e , ,
' L 3 3
]6 r 900 g L
e . i e Q-c.
h700-- -800 h
-700 c, c 600 5000 6000 7000 8000 9000 10000 Time (s)
Figure 11. Single loop hot leg steam terrporatures from the base case.
18
i Metal surface temperatures from the fuel cladding, upper plenum, and I single loop hot leg nozzle and steam generator tubes from 5000 to 10000 s are compared in Figure 12. The structural temperatures are important because of the potential for mechanical failure prior to the core melting through the reactor vessel. The location of sucn a failers is important l because a rupture of the steam generator tubes would provide a direct path ,
to the atmosphere for the fission products. The figure indicates that, based on this calculation, a steam generator tube rupture is unlikely, i However, any recirculation of superheated steam in the tubes or hot leg )
that might occur cannot be accounted for with RELAP5. i The SCDAP calculation was started at 9000 s, 60 s before the maximum i fuel clhdding surface temperature reached 1000 K (1340*F) in the RELAPS calculation and before the onset of calculated cladding oxidation. A constant pressure of 16.21 MPa (2351 psia) was used for the calculation.
The SCDAP calculation was terminated at 15000 s because fuel and cladding temperatures exceeded 3000 K (4900'F) and significant melting and !
relocation of fuel and cladding had occurred. l The core damage will be discussed in four areas: cladding ballooning and rupture, cladding oxidation and the resultant hydrogen generation, ,
rupture of the cladding oxide shell and relocation of material, and fission product release. Total core values were obtained by multiplying the single bundle SCDAP calculation by the number of fuel assemblies in the core (193).
The cladding surface temperatures at three of the ten axial nodes, node 1 [0.2 m (0.7 ft) above the bottom of the fuel), node 4 [1.3 m (4.3 ft)],- and node 9 [3.2 m (10.5 ft)], are presented in Figure 13. In the following discussion, reference will be made to two specific times, 12000 and 13860 s. At 12000 s, cladding surface temperatures first reached 2400 K (3860*F). The existing data base 8,9,10 haspermittedSCdAPcode 4
assessment 'II only up to this temperature. At 13860 s, the temperature of the fuel and cladding had reached the fuel melting temperature in node 1 and all of the fuel and cladding was gone from this node. The temperatures were then set to a constant value equal to the temperature previously attained, as seen in Figure 13, and the rod decay power was reduced by the 19
1600 , ,
a Clodding, 3.5 m above core bottom O Upper plenum 1400 _- O Hot leg nozzle -2000 m X Steem generator tubes ,
d b 1200 - -
- e
-1500 $ .
0 1000 - -
a
' 6
- e E 1 h 500 "- --1000
- h 600 ar--- - O C E E E E E E " U
-500 400 .
5000 6000 7000 8000 9000 10000 Time (s)
F igu r e 12. RELAPS-colculated metal surf ace t empe r a t ur es for the fuel clodding, upper plenum, and single loop hot leg nozzle end steam generator tubes from the base cose.
3500 , , , , i Fuel Aelting /
- h. "
y5000 3000"-
0x ida ti,3 Zr0 rupture 1 M l'-
2500 -- Ballooning
--4000
- e boTotion ~
0 u 2000 -
1
-3000 g
- 1 e
- o. I L o.
1500 ~
[ Zr-U-0 relocotion
-2000 A 0.2 m above f uel bo ttom 1000 0 1.3 m above f uel bottom 0 3.2 m above f uel bottom 1000 500 9000 10000 11000 12000 13000 14000 15000 Time (s)
Figure 13. SCDAP-colculated fuel cladding surf ace t empe r a t ur es from the base case.
20
amount generated in that node. When the temperature at other rod elevations reached the fuel melting temperature, the same calculations were made in the SCDAP code. With no fuel or cladding in node 1, the existence of unmelted fuel and cladding at higher nodes is questionable. Future model development should address this aspect of fuel and cladding relocation.
Cladding ballooning was calculated over 80% of the rod length.
Ballooning that spreads over a major portion of the fuel rod, as was calculated nere, is referred to as sausage-type ballooning. Determination of the location, timing, and extent of cladding ballooning is important for several reasons. Ballooning often provides space for axial fuel relocation, results in reduced fuel rod cooling, and leads to cladding failure and release of fission products. In particular, sausage-type ballooning can lead to coplanar flow blockage. Rod ballooning was calculated at 9720 s fr.om node 2 [0.6 m (2.0 ft) above the bottom of the fuel] through node 9. The effect of the ballooning on the cladding surface temperature can be seen in Figure 13 when comparing the curves corresponding to nodes 4 and 9, where ballooning was calculated, to the curve corresponding to node 1, where ballooning was not calculated.
Because the fuel-cladding gap size increased, heat transfer across the gap was reduced and the cladding surface temperature rise decreased in nodes 4 and 9. A corresponding decrease in flow 4.rea can be seen in Figure 14, in which the bundle cross-sectional flow area is plotted at the same axial eleva.tions used in Figure 13. The fuel rod ballooned to the maximum extent i allowed by SCDAP (approximately 90% flow area reduction) in two of the nodes and to nearly that extent in six others. The assumption of a maximum
{
balloon size is based on results of ballooning experiments.12 The '
cladding failed shortly after ballooning at 9820 s at node 8 [2.9 m (9.5 ft)] when the cladding temperature was 1710 K (2620*F).
Figure 15 presents the fuel assembly hydrogen generation rate.
Zircaloy oxidation and the accompanying generation of hydrogen is calculated by SCDAP when temperatures exceed 1000 K (1340*F). Oxidation first began at 9060 s at nodes 8 and 9, and the oxidation and hydrogen generation rates increased until 9500 s. Then, because the core water level had decreased, the mass flow rate of steam in the bundle slowly l
l 21
O 0.03 . , , ,
- 0.3
=
^
i a a - a a a a a a a 1
0.02 - '
^
n . A 0.2 m above tuet bottom . o,; n[
E o 1.3 m above tuoi bottom "
O 3.2 m above fuel bottom O O e e w 6 4 4
~
- - C C C C C C C C C C U 0.1 O O O O O O O O O O O Basconing 0.00 O.0 9000 10000 11000 12000 13000 14000 15000 Time (s)
Figur e 14. Fuel assembly flow areas from the base cose.
'o 4.0 7 , , , , ,
, - 0.0008 ,
o 0 6
3.0 -
hl -
w C C O - 0.0006 0 I' =
0 l O^
w I wM e
C eN 2.0 - - C
- 0.0004 $60 C C e e G G b t.0 . Steam starvation f n .
8
'O
~
) upper half of core 0.0002 y I I -
J t
0.0 O.0000 9000 10000 11000 12000 13000 14000 15000 Tim. (3)
Figure 15. Fuel ossombly hydrogen generation role from the base cose.
I 22 l
decreased, causing a decreased oxidation rate. A lack of steam, or steam starvation, implies that little or no oxygen is available at the cladding surface for oxidation. At 9860 s no steam was flowing past the cladding in the upper half of the bundle. Oxidation ceased in those nodes (6 through
- 10) as a result of steam starvation and no heat was generated by the l
, oxidation reaction. If steam had been present, accelerated oxidation would have been calculated by SCDAP beginning at 10180 s when cladding surface i
temperatures reached 1850 K (2870*F) and cladding temperatures would have increased rapidly. However, with no steam present, the rate of the temperature rise decreased and the fuel rods did not form a thick oxide shell. When dryout at the bottom of the fuel was calculated at about 11000 s, oxidation ceased completely. Because of the steam-starved environment, only about 30 kg (66 lba) of hydrogen was generated as a result of the oxidation of approximately 4% of the cladding in the core.
When the cladding temperature increased to 2125 K (3365*F) at 11200 s at node 9, the zircaloy inside the oxide shell began to liquefy and dissolve the outer portion of the fuel pellets. At 11400 s, the relatively thin Zr0 2shell ruptured at node 9 because of increased stress and decreased strength at a temperature of about 2200 K (3500*F). A hot mixture of liquefied fuel and cladding relocated downward. As a result, a step increase in the cladding surface temperature can be seen in Figure 13 at 11400 s at node 4, where the liquefied mixture resolidified. That rapid temperature rise was calculated at nodes 2 through 4. At 12000 s, the Zr02 shell had ruptured at four of the ten axial nodes, and 25% of the original Zr and 1% of the original UO had relocated below the bottom of 2
the fuel rods. However, the amount of UO relocated is undercalculated 2
by a factor of 6 to 7.13 This had a small effect on the fission product release rate. By 13860 s, the cladding oxide layer had been breached at eight of the ten axial elevations and downward flow of fuel and cladding l L
had resulted in the relocation of 84% of the original zircaloy below the bottom of the fuel rods. Of the original U0 , 4.6% had relocated below 2
i the bottom of the fuel rods. Interactions between the molten core material and the liquid in the lower plenum were not considered in this analysis.
Molten material interactions would have boiled the lower plenum liquid, l
providing steam to the core and allowing further cladding oxidation.
l 23 1
The calculated release rate of soluble fission products (Csl and Cs0H) from a 17 x 17 fuel assembly to the coolant is shown in Figure 16. SCDAP models the release of cesium and iodine from a failed fuei rod with a burst component and a diffusion component. The burst component is a model of the initial release when the cladding fails. The diffusion component models the release of the remaining cesium and iodine. Figure 16 shows that cladding failure and the accompanying burst release occurred at 9820 s.
Release of fission products from the fuel to the fuel-cladding gap and from the gap to the coolant is highly temperature dependent. As the temperature increased, more fission products were released from the fuel and from the cladding. The step increases seen at 11460 s, 13860 s, and 14900 s were a result of fuel liquefaction (melting or dissolution). At liquefaction, all fission products are calculated to be released instantaneously from the fuel to the fuel-cladding gap. The increase at 11460 s corresponded to the Zr02 shell rupturing at nodes 1 and 7. Fuel had been dissolved by molten zircaloy shortly before the rupture. At 13860 s, the fuel in node 1 reached the melting temperature. At 14900 s, the fuel melted in four additional nodes. At 13860 s, 3.1 kg (6.8 lbm) of soluble fission products hrJ been released from the fuel rods in the core. Most of this release was calculated at temperatures abeve 2400 K (3860*F). At 12000 s, when temperatures reached 2400 K (3860*F), only 0.3 kg (0.7 lbm) of soluble fission products had been released to the coolant in the core.
Several deficiencies in the calculations affected the results. The RELAPS calculation did not account for the. cladding oxidation energy or the effect of replacing steam with hydrogen. Figure 17 shows the ratio of the heat released by the oxidation reaction to the core decay heat from the SCDAP calculation. Inclusion of this energy would result in higher steam and structure temperatures than those shown in Figures 10 through 12. The RCS pressure would have been affected, but only in the rate at which it cycled between the PORV opening and clo:ing pressures. The core dryout rate would not have been directly affected since the oxidation energy would ,
be added to the cladding in steam-filled regions above the liquid in the core.
i I
24
o
- o
'- 4 I I I
- e 0 _
$ = =
E 10 E ' "'
3 f 'N fuel melting e
"q - -
u%
n3 10"
=
=
mg o
- L u - -
Q. . -
10-' F"*'d'***'"*"
E - --
m : :
e . -
6 . .
10-*
S000 10000 11000 12000 13000 14000 15000 Tim. (s)
F igu r e 16. Fuel ossembly soluble fission product release rate from the base case.
0.3 , , , , ,
0.2 - -
o 0
0.1 - -
t b
L J
0.0 9000 10000 11000 12000 13000 14000 15000 Tim. (3)
Figure 17. Rotto of the cladding oxidation hoot release to the core decay heat from the base cose.
25 l
l I
The core dryout times were different in the two calculations. RELAPS predicted core dryout at about 9600 s, while in SCDAP it was about 11000 s. This difference could have a large effect on the cladding oxidation and hydrogen generation. The more accurate dryout time is felt to be that calculated by RELAP5, because RELAPS was designed to calculate fluid behavior, whereas the SCDAP calculation used a simple -
thermal-hydraulic model which assumed a constant pressure and no interaction between the core and the rest of the reactor vessel. -
In light of these problems, a more accurate calculation could be performed with a computer code that linked the system thermal-hydraulic behavior with the core damage phenomena. Such a code should also be able to account for the geometry and flow changes associated with ballooning.
3.2 Three Channel Core Case A second analysis of the TMLB' sequence was performed in which the core was modeled as three parallel channels rather than as a single channel. Each channel represented a radial region of the core, and therefore had a different average fuel assembly power. The center and middle channels had approximately the same power, while the outer channel power was about 20% lower. The channels were connected to each other at six axial locations to allow radial flow. Appendix A contains a more detailed de'scription of the RELAP5 model. The SCDAP analysis modeled one fuel assembly from each of the three regions. The purposes of this analysis were to investigate two-dimensional flow effects in the core region and to determine if a more detailed model of the core would yield significant differences in the calculated behavior.
i The single channel core was replaced by the three channel core at 6000 s in the base calculation, while the system was liquid-filled and subcooled. The RELAPS calculation was stopped at 10000 s, about 750 s .
after the maximum cladding temperature reached 1000 K (1340*F). The RELAPS )
[ calculation will be presented first, followed by a discussion of the SCDAP 1
l calculations.
l l-
{
26
The overall plant response calculated by RELAPS was nearly the same as in the single channel core calculation. The main interests in the calculation were in the flows in and between the three core channels, and in providing initial conditions for the SCDAP calculations.
The mass flow rates at the exit of each of the three core channels are shown in Figure 18. When the loop natural circulation flow stopped near 6800 s, the flows in the center and middle channels decreased and the flow in the outer, low power channel reversed. This established a recirculating flow from the high power channels to the upper plenum, down into the outer channel, then back into the middle channel. Figure 19 presents the flow out the top of three of the six axial nodes in the outer channel. The flow recirculating back from the upper plenum went about 2/3 of the way back down the outer channel when the flow first reversed, and about half way down at the end of the calculation. The behavior in the outer channel and the core bypass was similar to that of the core bypass in the base transient.
Figure 20 presents the void fractions in the top core volume in each of the three channels and in the base transient. The good agreement between the three channels indicates there was good mixing between the channels. The void fractions in the two calculations were nearly identical, except for the final increase. The earlier dryout (about 100 s) in the base transient was caused by a slight difference in the PORV flow.
In the base transient, the flow out the PORVs contained a little more liquid after 7800 s than the three channel core case, causing the PORVs to stay open longer, resulting in a higher mass loss. The lower mass loss in the three channel core calculation caused the later dryout. This timing difference is not significant, and is within the uncertainty of the calculation.
The fuel cladding surface temperatures in the top core volume of the l three channel core and base cases are compared in Figure 21. The heatup l started slightly later in the three channel core case, consistent with the dryout difference. The outer channel heated up more slowly because it was a lower power region. The two inner channels heated up at a slower rate than the average core (base case) because of the cooler steam that was l 27
l l
400 , , ,
3 A Conier channeI - 800 3 O Widdle channel 300C -
m Q -
- 600 (
\
E E .c v 200g A
,, O 3 Lam .itw (3
? . V 1 },p 2 W j
- u -
4 1 -
,, 10 0 _- l l -- 200
$ N 2 L - ---- .k_ ,
O
- m g_ a 2
0--
, l }[
y,,, -
- _p 30
- -200
- 10 0 ~
6000 7000 8000 9000 10000 Time (s)
Figure 18. Core exit mass flow rates from the three channel core case.
300 , , ,
A 1.3 m cbove core bottom - 600 0 1.9 m above co r e bo tt om
. O 2.6 m above core bottom h - 400 6 O v
k l C.
10 0 ,-
-200 -
! l 0 *-
l0
, i , - -200
-10 0 -
6000 7000 8000 9000 10000 Time (s) -
Figur e 19. Moss flow rates out the top of three outer channel volumes from the three channel core case.
28
l l
1 i i s u u wo w .C ,
3.5 m above core bottom A Center channel (g O viddle channel i
g I 0.75 -
O Outer channel l g X Bose ecse -
o I
o O
E 0.50 - -
m I
o 0.25 - -
_ m -
N OG ,
6000 7000 8000 9000 10000 Time (s)
Figure 20. Top core volume vold fractions from the three channel core cose and the base case.
2000 , , ,
3.5 m above core bottom 3000 A Conier chonne1 O Middle chonnel
^ 1600
~
O Outer channeI _-2500 m v
x X Bose case p v
- - e 6 $2000 g ,
3 3 O 1200 -
4 C 6 6
- . -1500 e G- C.
E E
= c N
>=
800-- --1000 e
- O X CX C ?C X OX Os' w 500 400 6000 7000 8000 9000 10000 Time (s)
Figure 21. RELAPS-cal cul a t ed top core volume f uel clodding surf ace temperatures from the three channel core cose and the base cose.
i l
t ,
l l 29
being recirculated back into the core from the upper plenum via the outer channel. This resulted in the temperature reaching 1000 K (1340*F) nearly 200 s later in the three channel core case than in the base case.
The SCDAP calculations were started at 9200 s, about 50 s befora the maximum fuel cladding temperature reached 1000 K (1340 F) in the RELAPS calculation. The calculations were terminated at 15000 s, when significant fuel melting and relocation of fuel and cladding had occurred. The fuel -
rod cladding surface temperatures will be discussed first, followed by discussions of the cladding oxidation and hydrogen generation and the -
fission product release. Total core values were obtained by multiplying the single bundle SCDAP results for each channel by the number of fuel assemblies in the channel, then adding these three values.
The calculated cladding surface temperatures in the three channels and the base case at node 4 [1.3 m (4.3 ft)] are shown in Figure 22. The four calculations show similar trends. When comparing the calculated temperature for the three regions, the higher the power, the higher the temperature. That is, the calculation representing the 92 bundles in the middle channel had the highest power and, therefore, had the highest calculated temperature of the three. The calculation representing the 37 bundles in the center channel had the second highest power and temperatares. The lowest temperatures were calculated for the 64 bundles in the cuter core channel. The same trend was not noted, however, when comparing the calculations for the three regions with the base case calculation. Although the powers used for the middle and center channels (1.0813 and 1.0359 power factor, respectively) were larger than the power used in the base case (average rods, power factor 1.0), the middle channel .
temperature stayed near the base case temperature and the center channel temperature remained below it. This apparent inconsistency occurred only because the three channel core calculations were started 200 s later than the base calculation.
The effect of cladding ballooning can be seen in Figure 22 by the slope change at 9845 s, 9865 s, and 10165 s in the middle, center, and outer channels, respectively. Cladding ballooning was calculated at 9720 s 30
3000 i , , , .
3 2500,- -
m 4000
^
x v l'-
Relocated Ir-U-O v e 8 w 2000 ~
] -3000 $
o E
$ ScBooning L
- c. 1500 -
f, -
[
h l' o 2000 Ee 1.3 m above f uel bottom 1000 -
Bosooning a Center channel _
O Middle channel O outer channel -1000 X Base case 500 9000 10000 11000 12000 13000 14000 15000 Time (s)
Figure 22. SCDAP-colculated fuel clodding surf ace temperatures of node 4 from the three channel core cose and the base case.
1 31
in the base case. Ballooning over 80% of the rod length was calculated in the three channel core calculations and in the base calculation. When cladding ballooning occurred in the middle channel, a 90% flow area reduction was calculated. An increased coolant flow in the center and outer channels would have resulted, causing increased oxidation and temperatures in those regions. Interaction between bundles, however, cannot be considered in SCDAP calculations. Cladding failure was calculated within 100 s after ballooning at node 8 [2.9 m (9.5 ft)] when .
the cladding temperature was about 1700 K (2600*F) in all cases. Steam starvation at node 4 occurred between 10000 s and 10200 s in all four calculations. As a result, the temperature rise rate showed no increase from oxidation heat after ballooning. At 11365 s, 11425 s, and 12065 s, cladding temperatures at node 8 reached about 2200 K (3500*F) in the middle, center, and outer channels, respectively, and the oxide shell ruptured. In the base case calculation, the oxide shell ruptured at 11400 s. The step changes in the temperatures seen at those times in Figure 22 were caused by the molten mixture of cladding and fuel which had relocated from above node 4. At about 13500 s, the temperature at node 1
[0.2 m (0.7 ft)] in the middle channel calculation exceeded the fuel melting temperature and all of the fuel and cladding in that node melted -
and relocated below the bottom of the core. All calculations performed after 13500 s are considered highly uncertain.
The cumulative hydrogen generated throughout the core as calculated by SCDAP is shown in Figure 23 for the three channel core and base cases. The three channel core calculations began generating hydrogen later and generated less hydrogen than the base case. Oxidation began when cladding temperatures reached 1000 K (1340*F) at 9050 s in the base case and at 9245 s in the center and middle channels in the three channel core case.
This difference in time was discussed previously. The difference in the '
final amount generated was caused by two factors. First, the initial coolant inventory (calculated by RELAPS) was smaller in the three channel core calculations than in the base case calculation. Since most of the water was used for oxidation in both cases, less water means that less hydrogen was generated. Second, cladding temperatures in the outer channel 32
30 , g g g y a y g g y a a a a a a a a a e g -
60 E O
2 0 3 *"*""*: cose O
e - 3o c o O Base cose
- 20 -
,0 o 40 0' w
e #
C C o '*
3 30 m C C 0
m 10 -
o -20
- w '
2 V Z -10 I 4
4 0' a O 9000 10000 11000 12000 13000 14000 15000 Time (s) i Figure 23. Total core cumulative hydrogen generation from the ihree chonneI core cose orld ihe base cose.
I r
i, 8
33
in the three channel core case were well below 1000 K (1340*F) for nearly 100 s. Since no oxidation was occurring in these bundles, the little amount of steam that was generated was not used for producing hydrogen but was flowing out of the core. When dryout at the bottom of the core was calculated (about 11000 s in all calculations) 27.5 kg (60.6 lbm) of hydrogen had been generated in the three channel calculations and 29.6 kg (65.3 lbm) in the base case. The difference, about 7% of the base case amount, is not felt to be significant. In severe fuel damage experiments, '
9 an instrument error of 5% ,10 in measuring hydrogen is reported.
Figure 24 presents the ratio of the oxidation heat to the decay heat for each of the three core channels. As in the base case, the energy released during the cladding oxidation was significant, and would have affected the thermal-hydraulic behavior if it had been accounted for.
As in the base case, a significant difference in the core dryout times calculated by the two codes existed. This, together with the inability of SCDAP to model the flow between channels and reverse flow in the top of the outer channel calculated by RELAPS, points out the need to analyze this transient using a linked thermal-hydraulics / fuel damage computer code.
Recirculation from the uppe plenum may enhance oxidation by providing steam, or it may lead to earlier steam starvation by recirculating hydrogen.
Figure 25 shows the soluble fission product release rate from all of the fuel rods in the core as calculated by SCDAP. Both the base and the three channel core calculations are presented. The burst release accompanying cladding failure is seen at 9820 s for the base case calculation and at 9945 s, 9985 s, and 10260 s when the middle, center, and outer channel rods, respectively, failed in the three channel core case.
The magnitude of the initial burst release was smaller in the three channel ~
core calculations since it represented only 92 of the 193 fuel rods.
However, after 10260 s, when cladding failure had been calculated for all -
three channels, the cumulative fission product release was nearly identical to the base case. The increased rates at 11365 s, 11425 s, and 12065 s were the result of fuel dissolution.
34
o
- 0.3 . . , ,
A Center chonnel O Widdle channel O outer chonnel 0.2 -
b 2
o '
= .
O a:
3 -
0.1 -
J 0.0 C -- 3 'O 'O 3 C C 0 9000 10000 11000 12000 13000 14000 15000 Time (s)
Figure 24. Ratio of the cladding oxidation heat release to the core doccy heat from the three chonnel core cose.
16' s . . . . i g
A 3. chonnel cose e : o go e case 3
- 'G'
2
- rr$
-4
= _ _
' ciodding ourst u 10-' r e :: :
0 % 5 5 3 G * ~
Vfo 10-. r -
- c. :
Fuel dissolu tion 8
a 10-' r E
/ -
e t : :
10 * -
9000 10000 11000 12000 13000 14000 15000 Time (s)
Figure 25. Total core soluble fission product release rote from the three chonnel core cose and the bcss cose.
35
3.3 Late Depressurization Case A scenario was calculated that was the same as the base transient, except that when 5 K (9'F) of superheat was achieved in the upper plenum both PORVs were latched open. This was investigated as a potential mitigating action for the transient, in that core damage may be delayed if the accumulator liquid can be injected into the system. The upper plenum superheat criterion was chosen as a point at which the operator would know (by the core exit thermocouple readings) that the core was heating up and damage would eventually result. The normal operating pressure of the accumulators is 4.59 MPa (665 psia).
The results of the RELAPS calculation are presented first, followed by those of the SCDAp calculation. A discussion of the possible effects of accumulator injection is also presented.
A comparison of the RCS pressures from this calculation and the base case is presented in Figure 26. The pressure decreased continuously after the PORVs were opened at 8392 s. The depressurization was not fast enough, however, to achieve accumulator injection before core damage began. .
Figure 27 shows a comparison of the core tcp and bottom volume void fractions from the late depressurization case and the base case from 5000 to 10000 s. The void fractions in the top volume were identical since the PORVs were not latched open until after this volume dried out. The bottom of the core dried out about 500 s earlier with the PORVs open because the depressurization allowed the hot liquid in the RCS to flash.
The fuel cladding surface temperatures for the same core volumes and calculations from 5000 to 10000 s are presented in Figure 28. As expected,
- the lower volume heatup started earlier than in the base case. The temperature increase at the top of the core was more rapid in the late depressurization case because of the earlier core dryout. The steam that was available to cool the top of the core was hotter, so less heat could be transferred out of the fuel rod.
36
-2500 m
s m
- v
=~
m/ <
_~ -.
y 16 g %+= w y Q.
2 14 - -
2000 Ec.
. v v e e 6 12 - -
3 3 E
=
. .iS00 m b - 6 10 .
Q. a.
8 -
A Late depressurtzotion -
O Base case A 1000 6
0 2000 4000 6000 8000 10000 f
Time (s)
Figure 26. Pressurizcr pressure from the late depressurization cose l cnd the base case. I i
1 . .
O?s i O )' ROC I
Elevation above core bottom A Late dooressur;zation 0.0 m 0 Lat e depressurization 3.5 m l
i 0.75 -
O Base case 0.3 m -
e X Base case 3.5 m o
=
u O
D 0.50 - -
t t 0 {
0.25 - -
A
- o. _ _ m g, _s &. , ,
5000 6000 7000 8000 9000 10000 Time (s)
Figure 27. Core void fractions from the late depressurization case and the base case.
37
, l l * !
l I
l 1600 , . . .
Elevation above core bottom a Late depressurization 0.3 m .
1400 _- O Late dooressurization 3.5 m -2000 m O Base case 0.3 m ,
- w: X Base case 3.5 m E
" 1200 -
- e
$ -1500 $
3 1000 - -
3s O e Q.
h 800 -
1000
{
600;'r C = = = = ===
-500 400 '
5000 6000 7000 8000 9000 10000 Time (s)
Figure 28. RELAP5-calculated fuel cladding surf ace temperatures from the late dooressurization cose and the base case.
i
+
l t
l ,
l l
38 i
l
The RELAP5 calculation was terminated near 9800 s, about 800 s after the maximum fuel cladding temperature reached 1000 K (1340 F). The SCDAP calculation was initiated at 9000 s. The depressurization rate from the RELAPS calculation was used for the SCDAP pressure input. The SCDAP calculation was terminated at 10480 s, when the extrapolated pressure reached 4.14 MPa (600 psia) and the accumulators would have been injecting.
The cladding surface temperatures calculated by SCDAP for the late depressurization and base cases from 9000 to 10500 s are shown in Figures 29, 30, and 31 for nodes 1 [0.2 m (0.7 ft)], 4 [1.3 m (4.3 ft)],
and 9 [3.2 m (10.5 ft)], respectively.
Figure 29 shows that in the late depressurization calculation heatup began in the bottom of the core 240 s sooner than in the base transient calculation. The initial liquid level in the SCDAP late depressurization calculation was 0.06 m (0.2 ft), well below the midpoint of node 1 [0.2 m (0.7 ft)]. The initial liquid level in the base case calculation was 0.5 m (1.6 ft), above the top of node 1 [0.38 m (1.25 ft)]. The heatup rate, i
however, was similar for the two calculations after the node was uncovered.
I Figure 30 shows that the node 4 heatup rate was slightly lower in the late depressurization calulation than in the base case calculation until 9720 s. At that time, cladding ballooning was calculated in the base case
{
and a decrease in the heatup rate was seen. However, in the late depressurization case, ballooning was not calculated at this elevation and {
the temperature rise rate remained relatively constant. At the end of the calculation, temperatures at node 4 were higher than in the base case calculation. At 10480 s, higher temperatures were calculated for nodes 1 through 4 and lower temperatures for nodes 5 through 10 in the late depressurization case than in the base case.
The cladding surface temperatures at node 9 are presented in z
Figure 31. The cladding temperature at node 9 rose at a much lower rate in
{
the late depressurization calculation than in the ba'se case calculation for l two reasons. First, the cladding ballooned earlier in the late depressurization case than in the base case (9000 s and 9720 s, respectively) because of the large pressure difference across the i l
l 39 l
2000 , ,
-3000 0.2 m above f uel bottom A Late depressurization O Base cose 00 m ,
^ 1600 '- b M v v
I - -2000 E ,
3 .s.
0 1200 -
o
' w
- - -1500 e O- CL E E
- e 0=
800 -- --1000 9
C ~
-500 400 9000 9500 10000 10500 Time (s)
Figure 29. SCDAP-calculated f uel cladding surf ace temperature at node 1 from the late depressurization cose end the base case.
2000 , , '
' @ 00 l.3 m above fuel bo tt om A Late depressurization O Base cose 2500 m
^ 1600 '- b D ~
v e 2000
' E 3
? -
0 1200 -
' w
- -1500 e ,
- Q- c. '
E E
- l 800: --1000
(
-500 400 9000 9500 10000 10500 Time (s)
Figure 30. SCDAP-cal cula t ed f uel cladding surf ace temperature at node 4 from the late copressurization cose and the base case.
l
O O
2000 , i 3.2 m above f uel bo tt om -3000 a Late dep.ressurization O Base case
^ 1700 -
- ^
3 Bellooning 2500 b 6
6 3 3
. 0 1400 ,-
. 2000 8 CL .
Q.
E e E.
>=
>=
110 0 .- --1500
(
800 - ' '
2 00 9000 9500 10000 10500 Time (s)
- Figure 31. SCDAP-colculated f uel cladding surf ace temperature at nose 9 from the late depressurization case end the base Cose.
e 9
6
,. 41 l
! i
cladding. The pressure difference was larger in the late depressurization case than in the base case because the system pressure was lower. In both calculations, heat transfer across the fuel-cladding gap decreased at the time of ballooning, resulting in a decreased heatup rate. Second, since the liquid level was lower at the beginning of the late depressurization calculation than in the base case calculation, less steam was being generated in the late depressurization calculation. Less steam reached node 9 and steam starvation at this node occurred earlier (9780 s as opposed to 9840 s) in the late depressurization case than in the base case. Less steam resulted in less oxidation which in turn resulted in less heat generation from oxidation and lower temperatures. I Figure 32 shows the ratio of the heat released by oxidation to the core decay heat. Since the oxidation energy release rate was less than 10%
of the decay heat, accounting for this energy in the RELAP5 calculation would have had only a small effect on the depressurization.
The cumulative hydrogen generation for the entire core is shown in Figure 33 for both the late depressurization case and the base case.
Hydrogen generation began at about the same time in both calculations, when cladding temperatures reached 1000 K (1340*F) and oxidation started.
Initially the late depressurization calculation showed a larger hydrogen generation rate than the base calculation because the lower liquid level -
exposed more zircaloy to oxidation. After about 9240 s, however, more hydrogen was generated in the base case than in the late depressurization case because the temperatures were higher, and more steam was available for oxidation. At the termination of the late depressurization calculation, .
10.5 kg (23.1 lbm) of hydrogen had been generated as a result of the oxidation of about 1.4% of the cladding.
Figure 34 presents the cumulative soluble fission product release to ,
the coolant from all of the fuel rods in the core for the late !
depressurization and base cases. The burst release is shown at 9660 s when the cladding failed in the late depressurization calculation and at 9820 s in the base calculation. Cladding failure occurred at node 8 in both calculations. The earlier failure resulted from earlier cladding deformation caused by lower system pressure. The mass of fission products 1
, 42
. =
i i
0.05 , ,
l 0.06 -
o 0.04 -
l -
ck: )
0.02 -
0.00 '
9000 9500 10000 10500 Tim. (s)
Figure 32. Ratio of the clodding oxidation heat release to the core decoy heat from the late depressization cose.
30 , ,
A Late depressurization g O Base case 60 g
- 2. 2 v
g -
-50 g
- 20 -
- O e -
g
-40 0 C @
c C
-30 $
c 10 -
a ^+ i 5 Ch 6 -20 0 o
>. V" I -
-10 $
o- m 0
9000 9500 10000 10500 Time (s)
Figure 33. Total core cumulative hydrogen generoflon from the late depressurization cose and the base cose.
43
=
1 10" : i
- 5 l g !A Late depressurization 5
~
O Base case - )
1 l
@ 10 ' E
]
e u
e m
8 10" r - ,
1 o .
7- :
u . -
a _ -
E 10 = 1 e
n :
6 _ .
10-'
9000 9500 10000 10500 Time (s)
Figure 34. Teetal core cumulative soluble fission product release f' om the late depressurization case and the base cose.
t O
l 44
l l
released in the burst release was smaller in the late depressurization case than in the base case because lower temperatures resulted in a smaller fission product inventory in the fuel-cladding gap prior to failure. After failure, lower average core temperatures led to a smaller release rate of fission products from the fuel to the fuel-cladding gap and from the gap to the coolant in the late depressurization calculation than in the base calculation. At tne termination of the late depressurization case, i
_ 0.002 kg (0.004 lbm) of soluble fission products had been released from the cladding, compared to 0.01 kg (0.02 lbm) at this time in the base case.
Rupture of the cladding oxide shell and relocation of material had not occurred at the end of the calculation. This means that accumulator injection may be able to delay the co e melt.
If the accumulators were permitted to inject ifquid, the initial effect would be a depressurization as the cold water condensed steam in the cold legs, which would draw more liquid from the accumulators. As the liquid entered the core it would boil, and the steam would oxidize more ci:dding. If the coolant expansion from the boiling and the heat input from the oxidation were sufficient to stop the depressurization, a slow draining of the accumulators would occur during which extended oxidation and a relatively more rapid heatup would take place. This would generate more hydrogen and may lead to an earlier core melt. If the boiling did not stop the depressurization, the accumulator liquid may be able to cool or quench part or all of the core because there is enough water in the dCCumulators to Completely cover the core. The core melt would then be delayed. Again, though, more hydrogen would be generated by the cladding oxidation. Because of the complex interaction between the core and system responses, a computer code that can account for both the fuel damage and the system thermal-hydraulics simultaneously is necessary to further investigate this scenario.
It should be noted that accumulator injection will eventually occur unless the accumulators are isolated. The RCS will be depressurized as a result of operator action, failure of the RCS boundary caused by high temperatures, cr the core melting through the reactor vessel.
45
- 3.4 Early Depressurization Case Since the late depressurization of the RCS was unable to delay core damage, the earliest reasonable time to start the depressurization was
- selected for investigation. It was assumed that an operator would not depressurize while the steam generators were still able to remove the decay
~
heat, so the PORVs were opened at 5000 s, approximately 100 s after the steam generators were no longer an effective heat sink. This was about 3300 s prior to the calculated core uncovery in the base case.
, A comparison of the RCS pressures from this case and the base case is shown in Figure 35. The pressure decreased rapidly after the PORVs were opened. The depressurization stopped when the saturation pressure of the liquid in the RCS was reached. The RCS then repressurized as liquid was boiled in the core. This behavior was similar to that seen in the base case after the system reached saturation. The pressure increased until the j volumetric flow out the PORVs exceeded che volume generation rate in the RCS. The pressure then decreased continuously until the calculation was -x terminated. )
The void fractions at the top and bottom of the core are presented in Figure 36, together with these parameters from the base transient. Since the RCS reached saturation sooner in the early depressurization case than in the base case, void formation in the core occurred earlier. The void fractions during the respective boiloffs were nearly the same. The top core volume dried out, and the core heatup began, about 1000 s sooner in the early depressurization case than in the base transient.
The fuel cladding surface temperatures for the top and bottom core -
volumes from the early depressurization and base cases are shown in
, Figure 37. The early depressurization calculation was terminated at 8000 s, after the temperature in the top volume had reached 1000 K j (1340*F), and when core damage had begun. Only the bottom core volume had not yet started to heat up at the end of the transient. Thus, j depressurization results in earlier core damage.
i 46
_ _ . . .. - - . . -= -. . .. - . ._ -
1
)
l l
18 i i i
- A Early depressurization -2500 O Base case -
- u v -
16 n .. 1n . . ^
o g u - m w.vn .3 gc.-g --- 0 Q- M 2
v 14 c.
--2000 V e e 6 L s s M M
$ 12 - -
y
' 6 0- n.
-1500 10 -
8
+ 0 2000 4000 6000 8000 10000 Tim. (s)
Figure 35 Pressurizer pressure from the early depressurizollon ecse end the base case.
1 i i 4 OX 1 Il Elevation obove core bottom A Early depressurization 0.3 m O Early depressurization 3.5 m O Base case 0.3 m ~
e X Base cose 3.5 m o {
u O O
A 0.50 - -
.v_
o 0.25 - -
1
" * " " "'~~- '
0" -* -
O 2000 4000 6000 8000 10000 Time (s)
Figure 36. Core vold fractions from the early depressurization case and the base case.
1 i
l 1
l i
1 47 l
t
1600 , , . .
Elevation above core bottom A Early depressurization 0.3 m 1400 -
O Early depressurization 3.5 m --2000 0 Base cose 0.3 m ,
x X Bose case 3.5 m p.
" 1200 - -
e e
-1500 $
0 1000 - -
o
' s '
e e E Q.
h 500 -- l1000 h
_ o o -
600 E-~-a - Ly" -
-500 400 O 2000 4000 6000 8000 10000 Time (s) -
Figure 37. REl.AP5-calcula t ed fuel cladding surf ace t emper a t ures from the early depressurization ecse and the base case, l
l l
48 l 1
No SCOAP calculation was performed for this case since the plant conditions were similar to those at the end of the late depressurization case RELAP5 calculation.
The ability of the PORVs to depressurize the RCS during the core heatup could be useful. If the PORVs remain operable, the operators could use them to keep the RCS pressure within a desired band. For instance, it may be determined that accumulator injection is not desired. If a steam generator tube rupture occurred, the PORVs could be used to keep the RCS pressure below the main steam line relief valve opening pressure to prevent the re-lease of fission products directly outside containment, while keeping the pressure high enough (assuming the tube rupture is small enough) to prevent accumulator injection.
e 49
i l
l
- 4. CONCLUSIONS AND RECOMMENDATIONS '
Several conclusions and recommendations based on the analyses performed are presented below.
- 1. The Seabrook station blackout base transient (TMLB' sequence) resulted in core damage beginning at about 9060 s and cladding rupture and fuel relocation at about 11400 s. -
With no offsite or onsite power and no emergency feedwater, -
the plant remained stable as the steam generator liquid inventory was boiled. When the steam generator liquid was gone, the RCS heated up and pressurized. The PORVs were able to keep the pressure beluw the safety relief valve opening pressure. When sufficient mass was lost through the PORVs, the core uncovered and heated up. The heatup continued until fuel melting began.
- 2. Use of a more detailed core model did not significantly affect ,
)
the results or chronology of the transient.
The core dryout, heatup, and final condition was nearly the same in the three channel core case as in the base case.
However, the limitations of the computer codes were evident. Recalculation using a linked thermal-hydraulic and l fuel damage code is required.
1
. I
- 3. Holding the PORVs open before core damage occurs reduced the time
)
to the onset of core damage. .
The PORVs were opened in an attempt to mitigate the transient by depressurizing the RCS to allow accumulator injection before core damage occurred. However, the open PORVs accelerated the liquid depletion; more mass was l removed because the PORVs were open constantly rather than intermittently, and the decreasing pressure caused the hot 50 l
. +
liquid in the system to flash to steam. Whether the PORVs were opened after the core heatup began or shortly after the steam generators dried out, core damage began before the DCS could be depressurized to the accumulator pressure.
- 4. Accumulator injection will occur before the core melts if the PORVs are opened. A net delay in the time to core melt may result.
The accumulator pressure was reached before cladding oxide shell rupture and molten cladding and fuel relocation occurred in the late depressurization case. Depending on the rate of injection, the core may heat up more rapidly and generate a significant amount of hydrogen, or it may be cooled or quenched, delaying the core melt. Further investigation is needed.
- 5. The PORVs can be used to control the pressure at which core damage occurs.
It was demonstrated that the PORVs can depressurize the system. It follows, then, that by using the PORVs the plarit could be controlled to maintain a desired pressure. For example, if a steam generator tube rupture occurred, the RCS could be depressurized below the steam line relief valve opening pressure to minimize the fission product release outside the containment.
- 6. A computer code that links the system thermal-hydraulics with the
. fuel damage is needed to better calculate the transient after core damage begins.
RELAPS cannot account for the heat generated by cladding oxidation or for flow area changes caused by cladding deformation and molten material relocation. SCDAP does not l
l 51 I
allow reverse flow in the top of the core or flow between parallel channels. These and other important events such as the time to core dryout and complex interactions such as those that may be encountered with accumulator injection I need a linked code to be calculated accurately.
4 I h i
?
e D
52
l S. REFERENCES l
- 1. A. M. Kolaczkowski et al., Interim Report on Accident Sequence Likelihood Reassessment (Accident Secuence Evaluation Program, February 1983.
- 2. C. A. Dobbe and R. Chambers, Analysis of a Station Blackout Transient for the Bellefonte Pressurized Water Reactor, EGG-NTP-6704, September 1984.
I
- 3. V. H. Ransom et al., RELAPS/M002 Code Manual, Volumes I and 2, EGG-SAAM-6377, April 1984.
- 4. G. A. Berna, C. M. Allison, and L. J. Stefken, SCDAP/M001/V0: A Computer Code for the Analysis of LWR Vessel Behavior During Severe Accident Transients, IS-SAAM-84-002, Rev.1, July 1984.
, 5. G. A. Berna et al., FRAPCON-2: A Computer Code for the Calculation of Steady State Thermal-Mechanical Benavior of Oxide Fuel Rods, NUREG/CR-1845, January 1981.
- 6. Public Service Company of New Hampshire, Seabrook Station Final Safety Analysis Report, Docket 50-443, revised through Amendment 48, January 1983.
- 7. R. J. Wagner, Addition of ANS 1979 Standard Decay Heat to RELAP5, WR-NSMD-83-0003, January 1983.
l 8. R. K. McCardell et al., Severe Fuel Damage Test Series, Severe Fuel Damage Scoping Test, Quick Look Report, (Draft), EG&G Idaho, Inc.,
l December 1982.
- 9. R. K. McCardell et al., Severe Fuel Damage Test 1-1 Quick look Report, (Draft), EG&G Idaho, Inc., October 1983.
- 10. S. Hagen and S. O. Peck, " Temperature Escalation of Zircaloy-Clad Fuel Rods and Bundles Under Severe Fuel Damage Conditions", Paper No. TS-1.9, International Meeting on Light Water Reactor Severe Accident Evaluation, Cambridge, Massachusetts, August 28-September 1, 1983.
i 11. W. D. Driskell et al., Developmental Assessment of the Severe Core l Damage Analysis Package: SCDAP/M000, EGG-NTAP-6212, March 1983.
- 12. S. Kawasaki et al., "Recent MRBT Tests in JAERI," Sixth i
American-Japanese-German-French Fuel Behavior Workshco, Tokai-Mura, Japan, May 20, 1981.
i
- 13. C. M. Allison, D. L. Hagrman, and G. A. Berna, "The Influence of Zircaloy Oxidation and Melting Behavior on Core Behavior During a l Severe Accident," Fifth International Meeting on Thermal Nuclear l Reactor Safety, Karlsruhe, Federal Republic of Germany, September 9-13, 1984.
53 l
l
e D l
l APPENDIX A INPUT MODEL DESCRIPTIONS e
54
APPENDIX A l
INPUT MODEL DESCRIPTIONS l The input models used to represent the Seabrook plant in the RELAP5 and SCDAP calculations are described.
RELAp5 Input Model The model of the Seabrook plant used for the RELAP5 calculations included all of the major systems and components necessary to characterize the station blackout transient response. The model was based on a RESAR-35 model developed previously,A~1 with changes made to reflect plant specific information, to remove unneeded components, and to add new components required. Modeled components included the reactor vessel; two coolant loops, each of which contained a U-tube steam generator, a reactor coolant pump, and associated piping; the pressurizer; and the secondary sides of the steam generators. A nodalization diagram of the model is presented in Figure A-1.
The reactor vessel modeled the downcomer, lower plenum, core, core bypass, upper plenum, and upper head regions. The path from the upper plenum to the upper head region through the guide tube assemblies was modeled, as were the leakage path from the vessel inlet nozzles to the ,
upper plenum and the upper head spray nozzle flow path. It should be noted that the upper head region was an " inverted top hat", so there was no path for some of the liquid in the upper head to drain to the upper plenum.
The nodalization of the three channel core is shown in Figure A-2.
Three parallel flow paths were modeled in the core, with the core bypass remaining unchanged. Volumes at the same elevation were connected using cross-flow junctions. Figure A-3 defines the regions of the core the channels represent. The core average linear heat generation rate was 17.8 kW/m (5.44 kW/ft). The average linear heat generation was 18.5 kW/m 55 i
. S oss are ses 47s d e4s
[a47l-pa-R jasa[4>4-es7 naa
--] M a.
) aan
}. . .
7-hlM see D '-
[4 ase M1
( see E > OE 8 '
li 34o )~#
sa M IN l~
sai
-a- see -- - , - -
- t. ***
y 1:4 l ',
__e j
- =- 4
- c; tea _ , % - -
~[ [I p...........\......;
l***
k' [_la a_.
{ 'f' F1 I .. . . . . . . .R dia Jan__L J-mia
._ I ... ...daaJ- I
-vts.-
g - -
g 7,p
,s . -
- .5 p1 ne 1 ..
_L '"
{ '
^
fr*f*rH-Mf k-L4**ck- ,l,
~
- tr M *, H ,*A#i H ,422 ,
ise Ii_iis -,- 334
- T y
f,;--,/, q,
,m r/
w r; d _2 I sin +
l t'd ses l t
iso Figure A-1. RELAPS nodalization of the Seabrook plant for the station blackout transient.
6
l 11e l l l 1 1
/ / /
/ / /
/ ise / tes / ire
/ / /
/ / / . /
i i i . . . . . .
/ / /
/ / /
/ tes - ies - srs
/ / /
/ _
/ / ,_ j i i i ......
/ / /
/ / /
II'
/ 154 / 104 / 1T4
/ / /
/ / / . . /
i i ) . . . . . .
/, / /
/ / /
/ tes / 1e3 / tys
/ p/ /
/
V / . )
i l l . . . . . .
. < / ,
l / /
! / 152 / to2 / 172
/ / /
/ / / /
1 i i ......
/ / /
/ / /
/ 151 / 101 / 171
/ l / /
/ _
! / / /
1 1 I l
- sta Figure A-2. RELAPS nodalization of the three channel core.
l 1
57 i
i l
l l
l l
l I
l Region 3 Region 2 s l l l l Region 1 bl i' I I I q j N '
I I I N I II N I I N g l ,
) I ,
I I E
! l I I I I I I I I I i
5 Figure A-3. Cross-sectional view of the three channel core.
58
o - o
, (5.64 kW/ft) for the 37 fuel assemblies in regior, I (center channel),
19.3 kW/m (5.88 kW/ft) for the 92 fuel assemblies in region 2 (middle channel), and 15.4 kW/m (4.69 kW/ft) for the 64 fuel assemblies in region 3 (outer channel).
Three of the four coolant loops were lumped together and modeled as a single loop. The flow areas and fluid volumes were three times those for a single loop to preserve the fluid velocities. The remaining loop was modeled explicitly.
- The pressurizer was connected to the hot leg of the single loop by a surge line. The piping to the relief valves was also modeled. Since the two power-operated reitef valves operate identically, they were modeled as l a single valve. Similarly, the three safety relief valves were modeled as a single valve.
The steam generator primary and secondary sides were modeled.
Included in the secondary coolant system were the main feedwater system, main steam line, safety reitef valves, main steam isolation valves, and turbine stop valves. The steam dump system, atmospheric dump valves, and emergency feedwater system were not modeled because they were not available for use during the transient. Only the relief valve with the lowest pressure setpoint was modeled for each steam generator since it had the capability to control the pressure by itself.
Heat structures were used to model the stored energy and heat transfer surfaces of the reactor coolant system loop piping; the pressurizer wall; the steam generator walls, internals, and tubes; and reactor vessel walls, internals, and fuel rods. Ambient heat losses were not modeled.
SCDAp Input Model The core of the Seabrook plant contains 193 fuel assemblies arranged l on a square lattice to approximate the shape of a cylinder. For the SCDAP t calculations, a single fuel assembly was modeled along with a portion of 59
the vessel upper plenum. Information for the model was obtained primarily from the Seabrook Final Safety Analysis Report.A-2 Figure A-4 shows the nodalization used for the model.
A single fuel assembly consists of 464 fuel rods, 24 control rod guide tubes,1 instrument tube, spacer grids, and end fittings. A 17 x 17 array of fuel rods and tubes is formed within a structure of spacer grids, end fittings, and guide tubes. A cross section of a typical fuel assembly is shown in Figure A-5. Of the 193 fuel ass?mblies in the Seabrook core, 57 will contain control ruds. The remaining 136 assemblies will contain axial
- power shaping rods, burnable poison rods, or empty plugged guide tubes in the control rod positions. The modeled fuel assembly contained no control rods, since preliminary SCDAP calculations of the station blackout transient showed very little difference between the calculated parameters for assemblies containing control rods and those containing none. Two components were used to model the fuel assembly; one fuel rod component representing the ,264 fuel rods, and a second component representing the 24 control rod guide tubes and the instrument tube. The two components .
were divided into ten equal axial segments called nodes. For the fuel rod ,
)
component, six radial nodes were used. Four radial nodes were used for the guide tube component.
For the upper plenum, only the region between the top of the core and the hot leg' centerline above a single fuel assembly was modeled. This region was divided into two axial zones of equal length, as shown in Figure A-4.
60
O
- O Upper plenum sone 4 Hot les centerline 2
Core node Elev bottom (ft) : Top of core
, t0 3.6 (ii .e) s a
3.2 (i o.5) h 2.s (o.5)
@ ruelrods 7 85(8'8) youtoe tubes 6 2.1 (6.9)
S 1.7 (5.6) 4 1.3 (4.3) 3 1.0 (3.3) 2 0.6 (2.0) 1 0.2 (0.7)
Inle$
Figure A-4. SCDAP nodalization for the Seabrook station blackout transient.
1 l
l 61 l
I l
I C Il () Control rod
[] l C4 guide tube O C l. C O I
C '[ () (3 O
+ )
l \
[: (' [1 (3 [' '
\ \\
[] I \ (;
Fuel rod
[] L'\ C assembly I l \
l l \\
\
Instrumentation tube Figure A-5. Cross-sectional view of a typical Seabrook fuel assembly.
62
1 1
REFERENCES l A-1. J. E. Blakeley and J. M. Cozzuol, Best Estimate Analysis of a Small Break LOCA in a RESAR-35 Pressurized Water Reactor, EGG-NTAP-6032, September 1982. 1 A-2. Public Service Company of New Hampshire, Seabrook Station Final Safety Analysis Report, Docket 50-443, revised through Amendment 48, January 1983.
1 i
(
1 I y t
63
, 8 >
APPENDIX 8 COMPUTER CODE DESCRIPTIONS e
64 n
APPENDIX B COMPUTER CODE DESCRIPTIONS The two primary computer codes used in the station blackout transient analyses were RELAP5/M002B-1 and SCDAP/M001.8-2 A brief general description of each of the codes, together with information on the specific versions used, is presented below.
RELAP5/M002 The RELAP5/M002 computer code was developed for best estimate transient simulation of pressurized water reactors and associated systems.
It is a one-dimensional, two fluid, nonequilibrium thermal-hydraulic code utilizing a six equation hydrodynamic model. This model provides continuity, momentum, and energy equations for both the liquid and the vapor phases within a control volume. The energy equation contains source terms which couple the hydrodynamic model to the heat structure conduction model by a convective heat transfer formulation. The code contains special process models for choking, abrupt area changes, branching, cross-flow junctions, pumps, accumulators, valves, core neutronics, and control systems.
The cycles of the code used in the analysis were 17 and 22, which are stored under Code Configuration Control Number F01581 and tape number A33840, respectively, at the Computer Science Laboratory at the Idaho National Engineering Laboratory (INEL). In addition, three updates to the code were used. These are listed in Table B-1. Cycle 17 was used until just before the core heatup began, and cycle 22 was used to calculate the rest of the transient.
, SCDAp7 MODI The SCDAP/M001 computer code calculates the behavior in a light water reactor vessel during extenced periods of severe overheating. SCDAP simulates core and vessel plena disruption by modeling heatup; geometry changes; material relocation; and debris formation, heatup, and melting.
SCDAP allows detailed modeling of cylindrical and slab heat structure 65 ~
TABLE B-1. RELAP5/M002 UPDATES USED IN THE STATION BLACK 0UT ANALYSES Cycle 17 Updates
- IDENT RJWX018
- COMPILE DEFINE,SEGOIR -
- D PUMP.149 IF (SHIFT (IPURVI(I),3) .GE. 0) PMPOLD(I) = 0.0 ,
- INSERT,PSTDNB.136 l
IF(VOIDF(IV) .LE. 0.0)GO TO 460 Cycle 22 Updates
- COMPILE DEFINE,SEGDIR
- DELETE RRESTF.164,RRESTF.172 l
l 66
geometries. Thus, fuel rods, control rods, instrument tubes, and flow shrouds can be represented. All structures of the same type, geometry, and power are grouped together and one set of input parameters is used for each of these groupings or components. Code input identifies the number of rods or tubes in each component and their relative positions for the purpose of radiation heat transfer calculations. Models in SCDAP calculate fuel and cladding temperatures, zircaloy and stainless steel oxidation, hydrogen generation, cladding ballooning and rupture, fuel and cladding liquefaction, flow and freezing of the liquefied materials, release of fission products, fragmentation during reflood, and subsequent debris behavior.
SCDAP also contains two models for calculating the thermal-hydraulic conditions in the vessel. One of these models is the CHAN component from 8-3 the TRAC-BD1 computer code. The second is a coolant boiloff model which, when used, replaces the TRAC portion of SCDAP with a simplified calculation of heat transfer from component to coolant and from component to component. The simplified model was used for the calculations described in this r'eport.
The SCDAP code is being used to help plan,and analyze the results of severe core damage experiments, to aid in the determination of probabilities and uncertainties in risk assessment analyses, to identify the major contributors to vessel behavior during core uncovery accidents, and to help determine the order and timing of the events observed by a plant operator during core uncovery. A complete description of the code is contained in References B-2, B-4, and B-5.
The version of SCDAP-M001 used to perform these calculations <as version 10, which is stored on tape number A35563 at the Computer Science Laboratory at the INEL.
i 67
REFERENCES B-1. V. H. Ransom et al., RELAP5/MDD2 Code Manual, Volumes 1 and 2, EGG-SAAM-6377, April 1984. .
B-2. G. A. Berna, C. M. Allison, and L. J. Stefken, SCDAP/M001/V0: A Computer Code for the Analysis of LWR Vessel Behavior During Severe Accident Transients, IS-SAAM-84-002, Rev. 1. July 1984.
- B-3. J. W. Spore et al., TRAC-BD1: An Advanced Best Estimate Computer Program for Boiling Water Reactor Loss-of-Coolant Accident Analysis, -
NUREG/CR-2178, EGG-2109, October 1981.
B-4. L. J. Stefken, Coolant Boil-Off and Component-to-Comoonent Radiation .
Models for SCDAP, IS-NSMD-83-023, June 1983.
B-5. G. A. Berna et al., Vessel Models for SCDAP, Final Design Report, IS-SAAM-84-03, January 1984.
)
d 4
4 68
, .e Po.es 35 g g eeuCLE.A mioWL.704T Commmaoes ' at'oa t *wassem vaasw e, reoc , ye, e. , ., ,
, I"',"$' 8IBLIOGRAPHIC DATA SHEET EGG-NTP-6700 ggg on esTau CYeohs Om Test AGvtalt 3 aT LE .'so Sve f efit J Lt.WESL***
ANALYSIS OF A STATION BLACK 0UT TRANSIENT AT THE SEABROOK NUCLEAR POWER PLANT
. o.f t espont Co.*LE Tio 4Ao4, te eG.R S .wTaomisi $p n t pmhp F* IQnd e O.,4 etPon? 155wtD Paul D. Bayless l Rosanna Chambers tentomh.c tona
,....o. o.u a...o .. ....... .oo..., _ ,.c , . ,.o,. c , , .w.o.. . ..
EG&G Idaho, Inc. nr-arn7_7sinntq7n P. O. Box 1625 Idaho Falls, ID 83415 A6354
.. o .o.. o.a. a..o.......o..6. . o..., - <.c . . r, i oa aiaoar U.S. Nuclear Regulatory Comission Technical Washington, D.C. 20555 . .. .. co ..to c .,
,, ou,.t. . . .. . a r. .
i3 .asTe.cf 400 morse er wee A postulated station blackout transient at the Seabrook nuclear power plant was analyzed in support of the Nuclear Regulatory Commission's Severe Accident Sequence Analysis Program. The RELAPS/ MOD 2 and SCDAP/ MODI computer codes were used to calculate the transient from initiation through severe core damage. The base transient, the TMLB' sequence, assumed no offsite power, onsite power, emergency feedwater, or operator actions.
Additional analyses investigated the sensitivity to the core modeling and a potential mitigating action.
. . ooco.. .. . . ... . . . . .o-o... .c . .. . o-. .. .,,.gg, .
'O $E Cumif v CL.sseseC.f'04
,rev ears,
)
e *0f 4f'8st*Se0Pth 69049 ttaess '
t r.e ,ayere, l
6, %v.04 4 0. p.(,0 8 bS i
.....c.
l
~ l