ML20116K286
| ML20116K286 | |
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| Site: | 05200003 |
| Issue date: | 04/26/1996 |
| From: | Knudson D IDAHO NATIONAL ENGINEERING & ENVIRONMENTAL LABORATORY |
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Text
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4 A SCDAP/RELAPS ANALYSIS OF AN AP600 3BE TRANSIENT WITH EX-VESSEL FLOODING 1
i 4
1 l
t D.L.Knudson April 26,1996 Idaho National Engineering Laboratory (daho Falls, ID 83415 9608150007 960522 gDR ADOCK0520g3
e-1 1
CONTENTS l
l
- 1. INTRODUCTION........
.. I
- 2. SCDAP/RELAPS ANALYSIS.............
.2 2.1 SCDAP Input...
.2 2.2 RELAP5 Input
.3 2.3 Lower Head Input
... 3 2.4 Description of Transient Simulation.
.5
- 3. RES ULTS...............................
.7
- 4. REFERENCES
.8 FIGURES 1.
AP600 core cross section showing the radial division used in the subject analysis 10 2.
AP600 core radial power profile.
11 3.
AP600 core cross section showing RCC and GRC assembly locations 12 1
4 Cross section of an AP600 fuel assembly with a RCC assembly
.. 13
............ 13
- 5. Cross section of an AP600 fuel assembly with a GRC assembly...................
6.
Existing AP600 RELAP5 model nodalization....................
........... 14 7.
AP600 core cross section showing axial nodalization (without the thimble bypass and the core cross flow junctions).............
15
- 8. AP600 core axial power profile
. 16
- 9. Cross section of the AP600 reactor vessel lower wall and lower head.......................... 17 1
- 10. Finite element mesh representing the AP600 reactor vessellower wall and lower head.......... 18
- 11. Nusselt number ratio as a function of the angle from molten pool centerline...........
. 19 i
- 12. Azimuthal position of major AP600 vessel penetrations.........
............... 19
- 13. MAAP containment pressure used as a SCDAP/RELAP5 boundary condition.......
. 20 s
8
- 14. MAAP containment vapor temperature used as a SCDAP/RELAP5 boundary condition..
. 20 8
- 15. MAAP containment water level used as a SCDAP/RELAP5 boundary condition.
. 21 ii
- 16. Orientation for calculation of boiling heat transfer from a hemispherical surface 21
- 17. Predicted heat flux to a subcooled pool from lower head heat transfer correlations as a function of position and temperature difference.7 22 ' '
6 22
- 18. RCS and containment pressures..
23
- 19. DVI2 break flows.
23
- 20. DVI2 integrated break nows.
- 21. Accumulator liquid volumes.
24
- 22. CMT collapsed liquid volumes 24
- 23. Core collapsed liquid level relative to the bottom of the lower core plate 25 1
M
- 24. Calculated core collapsed liquid level compared to ROS A/AP600 experimental data l
25 relative to the bottom of the lower core plate...
- 25. Reactor vessel liquid mass 26
- 26. Maximum core surface temperature...
26 27 l
- 27. Total hydrogen generated.
- 28. ADSi now 27
- 29. ADSl integrated flow 28 28
- 30. ADS 2 flow
. 29
- 31. ADS 2 integrated flow
- 32. ADS 3 flow
. 29 30
- 33. ADS 3 integrated flow
- 34. ADS 4 flows.
.. 30 31
- 35. ADS 4 integrated flows.
31
- 36. Average core outlet vaporirmperature...
32
- 37. Fraction of fission products released from the fuel.
32
- 38. Decay power distdbution iii
TABLES 1.
SCDAP components used to represent the AP600 core.
. 33
- 2. Material masses in the active core region represented by SCDAP components.
34
- 3. Coeffic.ients used in a subcooled nucleate boiling correlation as a function of position on the exterior surface of the reactor vessel lower hemispherical head.
34 4.
Subcooled boiling correlations and CHF for nodes on the exterior surface of the COUPLE mesh.
35 5.
Sequence of transient events 36
- 6. Summary of relocation events..
. 38 i
Ne iv
l A SCDAP/RELAP5 ANALYSIS OF AN AP600 3BE TRANSIENT WITH EX-VESSEL FLOODING i
- 1. INTRODUCTION l
Accidents involving full reactor coolant system (RCS) depressurization and failure of gravity injection 1
into the reactor vessel from the in-containment refueling water storage tank (IRWST) are the largest con-tributors to AP600 core damage frequency. A SCDAP/RELAPS analysis of such accidents, which are I
l identified as 3BE transients, was recently completed to evaluate lower head integrity.2That analysis con-sidered accidents initiated by a double-ended off-set break in one of two direct vessel injection (DVI) lines.
l Three variations of the transient were considered including a case with a dry reactor vessel cavity, a case with a sustained cavity water level at the top of the reactor vessel lower hemispherical head, and a case j
with a sustained reactor vessel cavity water level above the break (which led to vessel reflooding).
l Results from the recently completed analysis indicated that the lower head would fail by melting fol-1 lowing the relocation of core materials with or without ex-vessel Gooding. The importance of lower head integrity prompted a thorough review of those results. The review indicated a need for a revised analysis with recommended SCDAP/RELAP5 code and AP600 input model refinements to be used in the calcula-tion of lower head response. The revised analysis is the subject of this report.
Code refinements that were recommended and incorporated into this analysis included a modification l
for the appropriate decay of power in the low M debris, extension of molten pool natural convection l
heat transfer calculations after complete abla.. r citing of any frozen oxide materials on inner vessel sur-faces, and adjustment of molten pool natural cc.wection heat transfer correlations consistent with current l
recommendations. Those refinements were incorporated into Version 8dl of the code. Version 8di, which 3
was the most current version and a pre-MOD 3.2 release, offered the advantage of including other code corrections and refinements that were developed since completion of the earlier analysis. (It should be noted that recommendations were also made for the addition of dead load and buoyancy {as a result of the ex-vessel flooding] forces in lower head integrity calculations and solution algorithm modifications to l
improve time step performance, particularly during the low pressure core oxidation phase. However, a l
dead load / buoyancy evaluation indicated that SCDAP/RELAPS refinement was not justified because the l
specified forces are insignificant compared to the temperature-dependent ultimate strength of the vessel wall.4 Work was also initiated on solution algorithm modifications to improve time step performance, but those modifications were not completed in time to be used.)
AP600 input model refinements that were recommended and incorporated into this analysis included s
reactor kinetics input for a best-estimate calculation of decay heat, extension of the lower head finite ele-ment mesh to avoid any potential for thermal isolation (at the top of the debris bed) imposed by the current version of the code if the mesh is filbd, a 3.3 K (6 F) increase in steady state loop operating temperatures consistent with current design specifications, adoption of subcooled boiling cortelations for ex-vessel heat transfer (based on experimental dataO), use of an ex-vessel flooding history from a previous MAAP cal-culation, and addition of input to model stainless steel slumping if intemal vessel structures are predicted 8
to melt. "Itose refinements were incorporated into Version 3.0 of the simplified AP600 model.9 Version 3.0, which was the most currentansion, offered the advantage of including other AP600 design changes that have been announced since completion of the earlier analysis.
r 1
O
)
Like the earlier analysis, lower head failure is assumed if heat loads exceed the CHF or creep mpture is calculated. In addition, a double-ended off-set break in one of two DVI lines was assumed to be the tran-sient initiator. In this revised analysis, however, the reactor vessel cavity water level was assumed to increase as a function of time until reaching a sustained level above the break (consistent with the previous 8
MAAP calculation ). Although ex-vessel flooding resulted in a sustained cavity water level above the break, vessel reflooding was not considered. That modeling decision was made to allow direct comparison with a current MAAP calculation at a future date.M The SCDAP/RELAPS model and details associated with the transient simulation are described in Section 2, calculated results through the time of core relocation are outlined in Section 3, and references are listed in Section 4. Calculations to evaluate lower head response to the core relocation are in progress.
A revision of this report is anticipated after the lower head calculations and a comparison with the current MAAP calculation are complete.
- 2. SCDAP/RELAP5 ANALYSIS SCDAP/RELAP5 is an integrated computer code package designed for nuclear reactor accident analy-sis. Modules for simulation of severe core damage, thermal-hydraulics, and heat transfer are included The l
user may develop input for those modules needed to simulate the problem of interest. In this analysis, an appropriate model required SCDAP input to simu! ate AP600 core components (including fuel rods, control l
rods, gray rods, instrumentation thimbles, empty guide thimbles, grid spacers, intermediate flow mixing
[IFM] spacers, and the stainless steel reflector); RELAPS input to represent thermal hydraulics in the core, throughout the remainder of RCS, and in selected portions of the secondary systems; and finite element input to represent the lower head during any thermal attack associated with relocated core materials. Corre-sponding input descriptions for those modules are provided in Sections 2.1,2.2, and 2.3, respectively.
Details specifically associated with the transient simulation are described in Section 2.4.
2.1. SCDAP Input The radial division of the AP600 core, shown in Figure 1, was the basis for.tll SCDAP input devel-oped for this analysis. Specifically, the division was established to reflect the AP600 radial power profile (shown in Figure 2) and define parallel paths for axial flow through the core. (It should be noted that the j
assembly relative power was normalized for an unrodded core at hot full power with equilibrium xenon i
near the beginning of life, which appeared to be the best data available.) Once the radial nodalization was establisheo, the number of fuel assemblies, rod cluster control (RCC) assemblies, and gray rou cluster (GRC) assemblies within each radial division / flow channel could be determined as indicated in Figure 3.
That information and the RCC and GRC assembly cross sections shown in Figures 4 and 5 respectively, were then used to develop input for each of the components summarized in Table 1.
The stainless steel reflector, which is not illustrated, was modeled as two components connected to the outer radial division / flow channel as indicated in Table 1. That approach was used because the reflector consists of a ielatively thin flat section and a thicker section that was fabricated to fit the core barrel arc.
The use of two components allowed a more accurate representation of the thermal mass and resistance of each section, which should result in a beuer approximation of reflector heating. An appropriate area was provided within each reflector component to represent embedded flow holes.
2
o l
There are four IFM spacers, two Inconel grid spacers, and seven Zircaloy grid spacers in the AP600 l
core. Input was developed to represent all four IFM spacers and the seven Zircaloy grid spacers in the l
SCDAP model. The two Inconel spacers were not explicitly simulated because they lie outside (above and j
below) the active core region that can be modeled with the current version of SCDAP, However, loss coef-l Ocients were appropriately included to provide hydraulic simulation. A summary of all material masses l
specified through SCDAP input is provided in Table 2.
1 l
l 2.2. RELAP5 Input j
Input used to simulate AP600 thermal-hydraulics was based on an existing RELAP5 model shown in Figure 6. The model includes the reacter vessel; both primary coolant loops; all four primary coolant pumps; the pressurizer; passive safety systems (including accumulators, an automatic depressurization sys-tem [ ADS], core makeup tanks [CMTs], the IRWST, and a passive residual heat removal [PRHR] system);
selected portions of the secondary systems as necessary to complete the subject analysis, and associated heat stmetures. Additional details regarding the RELAP5 model are available in existing documentation.9 It should be noted that PRHR piping downstream of the ADS 4 tee (numbered 832,833,834,835,836, 838, and 839) and IRWST drain piping (numbered 811 through 817 and 821 through 827) were deleted because those systems were assumed to fail in this analysis.
In order to simulate core damage, existing RELAP5 How paths and heat structures representing the active core region (numbered i14,115, i16,121, and 122) had to be replaced with the SCDAP compo-l nents described in Section 2.1. The region affected by that substitution is shown in Figure 7. Five radial l
divisions defining axial flow paths through the core are shown, consistent with Figures 1 and 3. The axial flow paths, with the axial power profile shown in Figure 8, were cross flow connected at each elevation. As indicated, cavity bypass and embedded reflector flow holes were connected to the outer core flow channel.
2.3. Lower Head input A detailed representation of the reactor vessel lower hemispherical head was needed to evaluate lower hesd response following the relocation of core materials. In addition, a representation of some portion of the reactor vessel cylindrical wall had to be considered. A portion of the cylindrical wall had to be included because the current version of the code will set heat transfer from the top of the debris bed to zero if the total volume considered is too small to contain all relocated materials. In this case, a relatively large por-tion of the cylindrical wall was included along with the lower head (as indicated in Figure 9) so that the corresponding volume cannot possibly fill. De portion of wall considered should also be adequate for sim-ulation of any axial conduction effects.
The specified lower head and cylindrical wall region was represented by the finite element mesh shown in Figure 10, which replaced corresponding RELAP5 heat structures. The axisymmetric mesh included a total of 572 nodes with 525 elements. Three elements were used to represent the combined thickness of the carbon steel wall and the stainless steel liner. Those wall elements, which were modeled as a homogeneous slab of carbon steel since the liner is relatively thin, are defined by Nodes 1 through 88 as
{
shown in the figure. A layer of zero-width gap elements was aligned with the inner surface of the liner j
(Nodes 67 through 110) to provide a wa,yjg represent the contact resistance between the debris and the 2
vessel. A constant coefficient of 500 W/m -K was used to represent that resistance. The remaining ele-3 l
l
ments were initially filled with primary coo' ant, which can boil off and/or be displaced during relecation of molten core materials.
After molten materials collect in the lower head, SCDAP/RELAP5 calculates natural convection heat transfer at interfaces between the molten pool and adjacent solid materials. That heat transfer is based on average coef5cients from steady state correlations. Specifically, the mean heat transfer coefficient to the ll upper crust covering the molten pool was developed by Steinberner and Reineke and is given by h, =
0.345 Ra""
(1) while the mean downward heat transfer coefficient applicable to all other surfaces was developed by May-i2 inger and is given by h, =
0.54Ra" (2) where k = thermal conductivity of the melt in the boundary layer adjacerit to the interface, R = effective radius of the molten region (based on the volume of all molten elements in an assumed hemispherical geometry), and Ra = Rayleigh number associated with the molten pool.
The Rayleigh number is defined as Ra = g QR' (3) avk where g = gravitationalconstant,
= coefficient of volumetric expansion, Q = volumetric heat generation rate, a = thermaldiffusivity and
= kine' atic viscosity of the molten materials.
v m
Experiments conducted by Jahn and Reineke indicate that the local downward heat transfer coefficient is a function of the angular position from the centerline of the molten pool.13 Specifically, the ratio of the local Nusselt number to the mean Nusselt number (Nu/Nu ) varies from -0.15 at 0 = 0 to -1.5 at 0 = 90 m
in a hemispherical geometry. That result is approximated in SCDAP/RELAP5 as shown in Figure 11.
Accordingly, calculation of natural convection heat (Tansfer from the lower head molten pool to the vessel wall involves identincation of molten finite elements adjacent to elements containing solid materials, use of Equation (1) or (2) to determine the appropnate heat transfer coefficient, and application of the analyti-cal representation of Figure 11 to modify the downward coefficients as a function of angular position.
4
I.
Heat transfer from the molten pool through the vessel wall was then rejected into containment through the exterior surface of the lower head (Nodes 1 through 22 shown in Figure 10). A discussion of that l
boundary condition is provided in the following section since it is transient specific.
2.4. Description of Transient Simulation The transient-initiating break was assumed to occur at time zero in the DVI2 line, which is shown rel-ative to other major AP600 vessel penetrations in Figure 12. (Note that the position of the DVI lines rela-tive to hot leg one [HLl] is of interest because the pressurizer is connected to HL1 and because the l
pressurizer represents the only potentially significant asymmetry of interest. As indicated, however, the DVIlines are positioned symmetrically with rupect to HL1 and the pressurizer. Consequently, the break
)
could have been modeled in either DVI line without a significant impact on the results.) The break was opened between the (0.102 m) DVI2 flow venturi and the flow restrictors downstream of the accumulator, CMT2, and the IRWST. Since the DVI2 line has an inside diameter of 0.173 m, each ride of the double-ended off-set break was assigned a flow area of 0.0235 m. However, break flow was actually controlled 2
by the DVI2 venturi and the other flow restrictors. (A flow coefficient of 0.8 was applied to both DVI ven-turis, which is consistent with the current practice in ongoing RELAPS analyses of AP600.) All break flow l
was appropriately directed into containment.
AP600 containment was represented by an adiabatic volume in the existing RELAPS input model.
However, that approach is not adequate for simulation of containment response in long term transients l
because condensation is not represented. Although existing RELAPS input could have been modified to 8
account for condensation, a decision was made to use the containment pressure from a MAAP calculation j
l as a time-dependent boddary condition for break and ADS flows. That pressure is shown in Figure 13.
Consistent with the definition of 3BE transients, gravity injection from the IRWST was assumed to fail. In addition, the PRHR system and one train of the ADS (one of two ADS 1 valves, one of two ADS 2 valves, one of two ADS 3 valves, and two of four ADS 4 valves) were assumed to fail. Both accumulators and CMTs were assumed to be available. However, the inventories of one accumulator and one CMT were lost to containment through the passive safety system end of the DV12 line break.
1 Initially, heat was transferred from the exterior surface of the lower head (Nodes 1 through 22 shown 2
in Figure 10) to a dry containment environment using a constant coefficient of ~15 W/m -K and a MAAP containment air temperature. The coefficient was assumed to be reasonable for simulating convective heat 8
losses to containment. The MAAP temperature, shown in Figure 14, was used as the sink temperature in lieu of modifying RELAP5 containment input (as previously discussed).'
'Ihe dry containment environment was altered relatively early in the subject transient because a portion of RCS, accumulator, ahd CMT inventories collect in the reactor vessel cavity after break initiation. In addition, it was assumed that operators initiate cavity flooding (using IRWST drain lines) at the time core 8
outlet vapor temperatures reach 1366 K. The MAAP prediction of the resulting water level, which is shown in Figure 15, was used in this analysis because the required containment input was not included in the existing RELAPS model. Accordingly, heat transfer from the exterior surface of the lower head shifted from an air convective mode to a water convective / boiling mode as each surface node was submerged as discussed below.
1 5
l
i 6
A set of subcooled nucleate boiling correlations were developed for calculation of heat flux (in 2
W/m ) from a hemispherical surface in the form q = a AT + BAT * + c AT' (4) where a, b, and c = position-dependent coefficients from Table 3 for the orientation shown in Figure 16 and AT = the difference between the wall surface temperature and the pool saturation temperature (in K).
Nucleate boiling curves derived from Equation (4) are valid from a AT of ~4 K to the AT associated 2
7 with the critical heat flux (CHF) for subcooled boiling (in MW/m ) given by q,,, = 0.4 (1 + 0.036/1
- (l + 0.0210 - (0.0070)')
(5) where ATsub = 10 K (assumed to be a constant in this analysis) and 0 = the surface contact angle in degrees for the orientation shown in Figure 16.
Experimental data embodied in Equations (4) and (5) were then applied to each lower head surface node (numbered I through 22 in Figure 10) to simulate natural convection, nucleate boiling, and transition boiling regimes as follows.
+ Equation (4) was used to calculate a position-dependent heat flux for AT = 4 K. Linear interpolation between zero and the resulting flux (at AT = 4 K) was performed to estimate the flux for any AT from 0 to 4 K which should adequately simulate natural convection to a subcooled pool.
l
. Equations (4) and (5) were solved simultaneously to determine the position-dependent AT associ-ated with the CHF. (Resulting CHF values are summarized for all surface nodes in Table 4.)
Equation (4) was then applied to determine the subcooled nucleate boiling flux for any AT between 4 K and the AT at the CHF.
1
- The transition boiling heat flux was linearly extrapolated from 19 position-dependent CHF to the 2
minimum flux that could occur assuming a heat transfer coefficient of 375 W/m -K. The extrapola-tion was performed for any AT greater than the AT at the CHF based on an estimate of the slope associated with transition regime experimental data.
l Application of the data is depicted graphically in Figure 17 for several positions on the exterior surface of the lower head. It should be noted that experimental data for the film boiling regime were not provided; I
although the lack of that data is not expected to be important because lower head failure will be assumed if heat fluxes from core materials in the lower head are high enough to drive the exterior surface beyond the CHF. (The potential effects associated with creep damage an: also considered in evaluation of lower head i
integrity.) As previously indicated, however, calculation of lower head response (based on the foregoing) is in progress. A revision of this report innticipated when the lower head calculations and a comparison i
with the current MAAP calculation are complete.
l l
6 i
l
l
- 3. RESULTS A double-ended off-settreak in the DV12 line was opened (between the DVI2 flow venturi and the flow restrictors downstream of the accumulator, CMT2, and the IRWST) to initiate the transient at time zero. A sequence of key events that developed thereafter is summarized in Table 5, which may be helpful in conjunction with the following discussion.
He break flows led to a rapid RCS pressure reduction (from an initial pressure of 15.51 MPa) as indi-cared in Figure 18. That pressure reduction triggered a reactor scram at 10.89 s. Continued flow through the breaks (see Figures 19 and 20) resulted in actuation of both CMTs by 12.86 s. The accumulator and CMT connected to the DVI2 line drained directly into containment through the passive safety system side I
of the break and, consequently, emptied relatively early as indicated in Figures 21 and 22. Flows from the accumulator and CMT connected to the intact DVI line (DVil) were directed into the vessel, but they did not prevent an early uncovery of the top of the active fuel as shown in Figure 23.a However, a relctively cool two-phase mixture in the upper half of the core was maintained by inventories from that accumulator and CMT, which was sufficient to delay core heatup as shown in Figure 26. When the accumulator and CMT connected to the DVII line emptied, the two-phase mixture in the upper half of the core was gradu-ally replaced by superheated vapor as core heatup progressed. Oxidation followed (see Figure 27) as the core temperatures approached the appropriate reaction temperature range. He energy associated with the exothermic oxidation reaction led to rapid core heating and, ultimately, the first fuel melting at 6595 s.
He ADS was triggered by decreasing CMT levels. Specifically, ADSl was actuated at 176.2 s because the CMT2 level dropped to 67.5%. He resulting flow and integrated flow from the ADS 1 valve are plotted in Figures 28 and 29, respectively. (De actuation timing for remaining ADS valves is given in Table 5. Corresponding flow information is provided in Figures 30 through 35 for reference.) ADS actua-tion is an important phase of AP600 accident mitigation since it is normally needed to reduce RCS pres-sures so that gravity injection of water from the IRWST can reflood the core and arrest any heatup. In this transient, however, IRWST gravity injection was assumed to fail. Given that assumption, the IRWST can be used to flood the reactor vessel cavity in an alternate strategy to try to preserve vessel integrity.
A core outlet vapor temperature of 1366 K will be tised as the signal for operator initiation of AP600 reactor vessel cavity flooding through IRWST drain lines. In this analysis, a temperature of 1366 K (based on an average of temperatures at the top of Flow Channels 113 and 114) was reached by 5666 s as indi-cated in Figure 36. As previously discussed, however, cavity flooding was not explicitly calculated in this 8
analysis. Instead, the flooding history from a MAAP calculation shown in Figure 15 was used as a bound-ary condition. From that figure, it appears that MAAP reflooding did not begin until ~6450 s. (ne water level of -2.4 m shown prior to the reflood is presumably consistent with the portion of RCS, accumulator, and CMT inventories that would collect in the cavity as a result of the DVIline break.) The discrepancy in the initiation of cavity flooding is not particularly critical in this analysis because flooding is well ahead of any core relocation into the lower head. As previously discussed, heat transfer from the exterior surface of
~
the reactor vessel lower head (and lower wall) shifted from a vapor convection mode to a convection /sub-cooled nucleate boiling mode as the cavity flooded.
- a. It is important to note that the SCDAP/RELAP5 core liquid level for first 1000 s compares closely with the level M as indicated in measured during a DVI line break experiment conducted in the ROSA /AP600 test facility Figure 24. After -1000 s, the ROSA /AP600 level began to increase (and the ROSA /AP600 level began to deviate from the calculated level) as a result of IRWST injection, which was assumed to fail in this analysis. The calcu-lated reactor vessel liquid mass is provided in Figure 25 as an additional reference.
7
a Core uncovery was completed by 10 350 s with core melt spreading radially and axially to the renector boundary shortly thereaf*er (by 10 780 s). The heat flux from the in-core molten pool to the reflector was much larger than the heat that could be removed by steam in the core barrel / reflector cavity. Consequently, the reflector was predicted to melt relatively quickly (by 10 970 s). An assumption of local reflector melt-ing is embodied in the current version of SCDAP/RELAP5. That assumption appears to be reasonable given the asymmetries that would be expected in the core region due to the location of the break and the potential temperature variations that could develop in the vicinity of major vessel penetrations. Given that assumption, the code allowed a side-way relocation of all molten materials, at or above the location of reflector failure, into the lower head without addition of reflector stainless steel. The timing, composition, and temperature of the associated relocation (and all other predicted relocations) are summarized in Table 6.
Four subsequent side-way relocations followed as molten materials spread to intact reflector bound-aries at progressively lower core levels (see Tables 5 and 6). A final relocation through the bottom of the core followed. Spcifically, molten materials reached the bottom of the fueled region by 12 840 s. At that time, the code allowed relocation into the lower head based on the assumption that the melt could flow through holes in the lower core support plate and/or melt through the reflector in order to complete the relocation. In total, ~64 000 kg of molten fuel, or ~85% of the AP600 active core, was relocated to the lower head during the transient. Other lower head debris constituents are identified in Tables 5 and 6. Cal-culations are in progress to evaluate lower head response to the relocated materials. A revision of this report is anticipated when those calculations are complete.
Betwcen 60 and 70% of the volatile fission products were released from the fuel during the core degra-dation process (see Figure 37). The corresponding effect of fission product release on core decay power is illustrated in Figure 38. Specifically, the total decay power was reduced by ~14% (calculated at the time of the first relocation) as a result of the fission product release that began at ~4500 s. Subsequent reductions in the core power correspond with relocations into the lower head that were discussed above. It should be noted that the single increase in (total and core) power shown in Figure 38 at 10000 s is consistent with the American Nuclear Society Standard for the calculation of decay heat, which was appropriately considered in development of the best-estimate reactor kinetics input used in this analysis!
- 4. REFERENCES 1.
C. M. Allison et al., SCDAP/RELAP5/ MOD 3.1 Code Manuals, Volumes )-5, NUREGICR-6150, INEL-95/0609, (formerly EGG-2720), December 1995.
2.
D. L. Knudson (INEL) letter to Y. Chen (NRC), " Transmittal of Revision 1 of Task 14 Letter Report Under JCN L2230',', DLK-12-95, July 7,1995.
3.
C. M. Allison et al., Design Report on SCDAP/RELAPS Model improvements - Debris Bed and Mol-ten Pool Behavior, INEL-94/0174. November 1994.
4.
S. A.Chavez (INEL) letter to Y.Chen (NRC), " Corrections Made per Telephone Conversation of July 31,1995", SAC-8-95, October 30,1995.
5.
J. L. Judd (INEL) letter to W. Jensen_(NRC), "AP600 Best-Estimate Decay Heat Input for RELAP5 based on ORIGEN2 Calculations (L2419-T10)", JLJ-09-95, October 19,1995.
8
6.
W. Cheung (Pennsylvania State University) fax to L Rempe (INEL), October 25,1994
- 7. W. Cheung (Pennsylvania State University) fax to E. Coryell(INEL), February 16,1995.
- 8. R. Palla (NRC) fax to D. Knudson (INEL), transmitting a summary of a MAAP calculation entitled
" Description of DVI Line Break Cases" February 6,1995.
- 9. G. E. Wilson (INEL) letter to T. Lee (NRC), " Benchmark of Updated AP600 Simplified input Deck, 16008", GEW-61-95, October 27,1995.
- 10. R. Palla (NRC) fax to D. Knudson (INEL), transmitting a summary of a MAAP calculation entitled
" Description of DVI Line Break Cases", April 8,1996.
I1. U. Steinberner and H. H. Reineke, " Turbulent Buoyancy Convection Heat Transfer with Internal Heat Sources", Proceedings of the 6th International Heat Transfer Conference, August 1978.
- 12. F. Mayinger et al., Examination of Thermal-Hydraulic Processes and Heat Transfer in a Core Melt.
BMFT RS 48/1, Institute for Verfahrenstechnik der T. U. Hanover,1976.
- 13. M. Jahn and H. H. Reineke, " Free Convection Heat Transfer with Intemal Heat Source, Calculations and Measntements", Proceedings of the International Meeting on Thermal Nuclear Reactor Safety, NUREG/CR-0027. February 1983.
- 14. R. Shaw, T. Yonomoto, and Y. Kukita, Quick Look Report for ROSA /AP600 Experiment AP-DV-01, JAERI-memo 07-189 September 1995.
MM 9
....... -............~ -. -.-...-.
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270*
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I h
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I 180" 1.048 0'
O.968 1.073 8
1.112 1.01 1.1 1
i 1.054 1.287 1.
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l 1.14
.02 1.081 1.15' 1.018 1.1 1.
895 0.712 h.93 1.07
.95 Anembly relauve power (where core has octant symmetry) l I
Radia! Division Number of l
Relative Power Normalized Relative Power
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(numbered from center to periphery)
Fuel Assembh.es f
1 25 1.040 1.040 2
16 1.208 1.209 i
3 32 1.039 1.039 4
40 1.014 1.013 f
5 32 0.809 0.809 i
i Figure 1. AP600 core cross section showing the radial division used in the subject analysis.
i i
4 1
10
.i b
- i
1.30,
c
,Q 4
o m-1.20 5
3 s
oC.
.$ 1.10
_E E
ue# 1.00 e
E OC 1o 0.90 -
Es
~O>
0.80 O.0 0.5 1.0 1.5 Equivalent core radius (m)
Figure 2. AP600 core radial power profile.
11
i 1
i 270*
l E
l l
l l
l R
P i
I il i
I O
180" RCC O'
I
[
"CC i
I GRC GRC 1
l RCC RCC i
1 l
GRC RCC and GRC locauons 1
(where core has actant symmery)
Radial Division Numberof Fuel Assemblies Number of RCCs Number of GRCs i
(numbered from center to periphery) l t
l 1
25 5
4 l
2 16 12 4
l 3
32 12 0
)
4 40 16 8
5 32 0
0 l
j Figure 3. AP600 Core Cross section showing RCC and GRC assembly locations.
i 1
)
a 4
12 i
A
l O
l
~.
r3 r 3 r3 L J L )
k )
rm r,
t r m r m r,
r3
<m d ContTOI rod t ;
t s
t >
t }
t ;
r3 r,
r3
( m L }
k)' -N Q J N)
Fuel rod htruments location r,
r, c,
r, thimble e>
e>
e>
w >
v>
r, r,
k J LJ rm r,
o e J Figure 4. Cross section of an AP600 fuel assembly with a RCC assembly.
B 3
A A
I I
[]
E E'.
- Gray rod
+
N Control rod -
o L
J
,N L J
'I Fuel rod Instrument-location thimble g
]
g r
l l
I Figure 5. Cross section of an AP600 fuel assembly with a GRC assembly.
l I
l I
13 I
l 4
l E
i em I
a I
3 l
3
,\\
j r
13 ia L
a L
M w__=__i___ ___m.........
a
- /EEE q
kE
% Duunad3Eu f:
a r
c
=.
a e
. :.: :. L ic 5 :
\\
r a
- "j
= mu u g
D,,,
V:;
l
-1l i
p-g l-E E
~
,8 i
/
-u d
d O,
3
{
'/
Q, "JL
- 1=
If i
h l
x-i I.E r.
s i -
E
.............. ;a..J,L.5./. : ::
. i.
i..
Q T
-g ll
- p,.
ms +].s 73-*-
2, 3; i.
l l.
.r u us.2
.u.
- ~
t J
i s m.s.
I
.I
[
g d h
i s
/ $.
anf 3 pli I
k.
E li i
sa l
w
_ :~ ;t l 1
..:~:
e
- as s
s
____ D _B.__l, l:
~
.a
_i:
C,i "L
4 i
g e
a
. a
- ; a :
_j.
.J.__ j
%1 j
]
__c=o______sil.i.__i m.
n u
.r a
T
.._.... n::
i aA l
..........g..........".....,
2...........
g g
e s
j s
I a
l a
nA a
e e
me a n y i
t l l
- ,l J
/
L l
.W:
. a -._f C
a
- : Y g
E C=
szzzu sxxxx::
~
" " ~
m d
[
____,....a y
r i
- har,
- /p,.
h ass a
1;;
h
- k o
l,V V
W I*
k 8E ui l
%....s
=*
l tW l
c a
E !
g x
k _._..
p6 3.,,
I
.u.
14
103/107 upper plenum l
- L a
n
- L JL
)
{
upper core plate l
l l
A,.
17 17 17 17 17 12.05" j
j 8
8 j
5 16 16 16 16 16 17.95" vessel wall-e-j j
15 15 15 15 15 9.7375" s
s 7
7 core banel j
j 14 14 14 14 14 10.8125" 1
1 cavity bypass
=
13 13 13 13 13 9.7875" (123/124) 3 j
6 6
j j
12 12 12 12 12 10.7625"
(
s,,
j l
11 11 11 11 11 9.7375" 5
5 l
10 10 10 10 10 10.8125" l
[
active fuel region reflector j
j 9
9 9
9 9
9.7875" d
s
?
4 4
reflector flow f
f 8
8 8
8 8
10.7625" holes (121/122) 7 $
E 7
7 7
7 7
9.7375" j
j 3
3 s
6 6
6 6
6 10.8125" e
r m
1
?
y 5
5 5
5 5
10.275" 3
2 2
i j
j 4
4 4
4 4
10.275" i
8 h
3 3
3 3
3 8.117" 1
I f
i 2
2 2
2 2
12.883" h
f h
f
\\
f f 14"-lower core plate I
l I
I I
downcomer 117 116 115 114 113 core flow channels Q
Figure 7. AP600 core cross section showing axial nodalization (without the thimble bypass and the core cross flow jt:nctions).
4 15 i
1 I
1.4 e
C
.9 U
M h 1.2 e
NOQ.
.$ 1.0 E
l E
u DO af 0.8
~
m O
?5 0.6
~
aC 0.4 0,000 n,73g Figure 8. AP600 core axial power profile.
l 16
- i f
9.
i 4.0173 m 2.2027 m (86.72 in)
-Top of lower head mesh (158.2 in)
(-tmddle of active core) l i
i i
f 2.1669 m (85.31 in) ~~
l
)
J 1.9220 m 4
(75.67 in) :g 1.9939 m (78.5 in)*
2.1530 m (84.76 in)*
\\
1 l
1.5624 m
- Bottom of lower core plate (61.51 in)
\\
1.9158 m (75.42 in),
2.0808 m (81.92 in) i 2.0089 m (79.09 in) j 0.15799 m (6.22 in)-
0~
j 0.15799 m (= 6 in carbon steel + 0.22 in stainless steel clad) 9, i
Figure 9. Cross section of the AP600 reactor vessellower wall and lower head.
in'dicates that the number was derived using engineering judgement 17
,e i
- I H
2.2027 m
=,
j y,,,,,,,,,,
T 1
Node 22 h
?
2 l
m MW/M M\\\\\\\\WlllL I
\\
'\\\\\\\\\\\\V///llll!Ih)iG t
l
(\\\\\\\\\\\\W///////////////
l
' \\\\\\\\\\\\////////////////
t t \\ \\\\\\\\\\///////////////
l
\\ \\ \\ \\ \\\\//////////////
\\ \\ \\ \\ \\/////////////
i
\\ \\ \\ \\ \\////////////
l
< t \\ \\ \\ \\/////////
t\\\\\\\\///////
\\
t\\\\ W E S "
\\ \\ \\///
l Nodes 67,89-I
\\'#
1 jNode 1 i
Figure 10. Finite element mesh representing the AP600 reactor vessel lower and lower head.
i j
l i
18 1
o 1
1.5 S
j 1.0
.2t C 0.5 sZ 1
a 0.0 0.0 10.0 20.0 30 0. 40.0 50.0 60.0 70.0 80.0 S'3.0 Angular position relative to molten pool centerline (degrees)
Figure 11. Nusselt number ratio as a function of the angle from molten pool centerline.
HL2 n
CL2B CL2A Rea DV12 --.
[
.-- DV11 CL1B CL1A y
HL1 Figure 12. Azimuthal position of major AP600 vessel penetrations.
l 19
.a
.0 0.25 i
- 0.20 E
2L e
a E._
' O.15 0.10 O.0 5000.0 10000.0 15000.0 Time (s) 8 Figure 13. MAAP containment pressure used as a SCDAP/RELAP5 boundary condition.
450.0 i
i l
tempg-7500100001 e
j 400.0 B
E
\\
2 o
E e 350.0 E
- :s
-o i
i i
300.0 O.0 5000.0 10000.0 15000.0 Time (s) 8 Figure 14. MAAP containment vapor temperature used as a SCDAP/RELAP5 boundary condition.
)
20
o 8.0
[
6.0 l
cntrfvar-9750 l 7
i i
'E 3
4.0 j
b 1
5; i
8 j
2.0 ci k
0.0 g
f Bottom exterior surface of reactor vessellower hemispherical head i
-2.0 0.0 5000.0 10000.0 15000.0 Time (s) 8 Figure 15. MAAP containment water level used as a SCDAP/RELAP5 boundary condition.
9.
h 0
l I
Location of interest D
1 Figure 16. Orientation for calculation of boiling heat transfer from a hemispherical surface.
21
1.5 i
i C
O UD=0.00 0--e UD=0.20 L
h UD=0.35 k--* UD=0.50
^ 1.0 -
C D UD=0.75
=
h1 5
a::
i
$ 0.5 1
0.0 :.
=
0.0 100.0 200.0 300.0 Temperature difference (K)
Figure 17. Predicted heat flux to a subcooled pool from lower head heat transfer correlations as a func-tion of position and temperature differenceh7 l
1.0 i
i I
C ORCS O--O Containment 7
CL2 w
2 0.5
?n E
ct 4
=
~
=
0 0.0 O.0 5000.0 10000.0 15000.0 Time (s)
Figure 18. RCS and containment pressures.
22
l i
1000.0
]
C O Vessel side 800.0 1
e-4 PSS side 7h 600.0 6
2
'E 400.0 E
200.0 -
M. M.
.b c.AA.A _mr. _ A._..
e -
0.0 o, 0.0 500.0 1000.0 1500.0 2000.0 Time (s)
Figure 19. DV12 break flows.
200000.0 i
150000.0 -
0 0
0 0
0 EE it 8
n n
n n
n
^
~
v 100000.0 2
8 E
.c 50000.0 C
o Vessel side O-4 PSS side A
0'0 O.0 5000.0 10000.0 15000.0 Time (s)
Figure 20. DVI2 integrated break flows.
23
d 4
l 50.0,
i 1
l a7 40.0 C
ointact side j
{
e--e Broken side e-1 E
o
~
} 30.0 v
l
$ 20.0
_m i
)
o E
8 l
M - 10.0 i
s
)
0.0 5
0
'O O
0.0 500.0 1000.0 1500.0 2000.0 i
Time (s)
Figure 21. Accumulator liquid volumes.
100.0 0 C
80.0 -
C Olntact side g
e--e Broken side E20 60.0 s2
.eQ1 40.0 a
i
~O l
o p
i 2
20.0' I
o A
0.0 0
0 0
t 0
'O
'O 0.0 1000.0 2000.0 3000.0 4000.0 Time (s)
Figure 22. CMTliquid volumes.
24
5.0 Top of active fuel y 4.0 h
n O
3.0 l
lll f
f I
entrivar-18 l E
Y u>
S 2.0 -
=
8 e
~
~
Bottom of active fuel s
i t
0.0 O.0 5000.0 10000.0 15000.0 Time (s)
Figure 23. Core collapsed liquid level relative to the bottom of the lower core plate.
10.0 SCDAP/RELAP5 C
MOSNAP600 AP-DV-01
^
8.0 -
E I
y s.O z
k Ton of active fuel 4.0 J
% fY
- 4 0
2.0 Rntinm of nr4ive foal 0.0 O.0 1000.0 2000.0 3000.0 4000.0 Time (s)
I Figure 24. Calculated core collapsed liquid level compared to ROSA /AP600 experimental data # rela-tive to the bottom of the lower core plate.
25
.. ~.
1 5.
80000.0 i
i
~.
cm 6 60000.0 eem-E 32 I
o
.7 40000.0 R.
8 m2 W
20000.0 e
e 0.0 O.0 5000.0 10000.0 15000.0 Time (s)
Figure 25. Reactor vesselliquid mass.
1 4000.0 i
_E e
3000.0 eo.
E S
e U
q 2000.0 o=
l Eo i
E'se 1000.0 I
bgmet-0 l E
e B
A o
t 9
0.0 0.0 5000.0-10000.0 15000.0 i
Time (s)
I Figure 26. Maximum core surface temperature.
26 I
400.0 1
l cntrfvar-9800 l
_cn6 300.0 2
E J
E 200.0 e.
ba
.c
-.g 100.0 H
0.0 O.0 5000.0 10000.0 15000.0 Time (s)
Figure 27. Total hydrogen generated.
80.0 i
i
~
l mflowi-454000000 I
~
l b
40.0 6
20.0 c
0.0
._20,0 O.0 1000.0 2000.0 3000.0 4000.0 5000.0 Time (s)
Figure 28. ADS 1 flow.
27
1 a
l i
8000.0 i
t i
l i
i
[
6000.0 - /
l 3
.g I
w 3
7 4000.0 l
E en i
M S
)
i 2000.0
)
i l
i l
l l
O.0
~
O.0-5000.0 10000.0 15000.0 Time (s)
Figure 29. ADS 1 integrated flow.
-i 40.0 2: ";
l l
l mflowi-456000000 l 7h 20.0 6
.2 1
E 10.0 E
N A
k 0;
l l
-10.0 0.0 1000.0 2000.0 3000.0 4000.0 5000.0 Time (s) 2,.
Figure 30. ADS 2 flow.
j j
i i
28 p.
4 1
t 1
l 15000.0 l
22 10000.0
[
k h
ilii 4
n-e 1
i s
c-5000.0 i
4 i
i 0.0 i
0.0 5000.0 10000.0 15000.0 Time (s)
Figure 31. ADS 2 integrated flow.
1 i
l 20.0 i
i
.i i
15.0 -
I mflowi-458000000 l i
10.0
,{
7b6 P
5.0 g
0.0
-5.0 ^
-10.0 O.0 1000.0 2000.0 3000.0 4000.0 5000.0 Time (s)
Figure 32. ADS 3 flow.
29
l O
l l
6000.0 l
[o 4000.0 i
V l
3 8
eo E
cn o
\\
l
.E_
2000.0 l
1 0.0 0.0 5000.0 10000.0 15000.0 Time (s)
Figure 33. ADS 3 integrated flow.
i 80.0 l
60.0 -
C O Broken side 0--e intact side l
l 7
i b
40.0 6
5 20.0 I
{
E I1l W&l b
A 0.0,,
g__--
i i
l
-20.0 O.0 1000.0 2000.0 3000.0 4000.0 5000.0 Time (s)
Figu 9 34. ADS 4 flows.
6 30
i 1
40000.0 ---
0 8
i J
g-l l
30000.0 -
c-O 0
0
~
l 5
6 3
8 e
20000.0 O
5 S
C O Broken side E
9--e intact side
~
10000.0 0.0 0 O.0 5000.0 10000.0 15000.0 Time (s)
Figure 35. ADS 4 integrated flows.
3000.0 i
i entrivar-9996 l l
k j
2000.0 -
7 W#w/*{
1366 K (RV cavity flooding temperature)
H 1000.0 i
1
~
l i
L O.0 0.0 2500.0 5000.0 7500.0 Time (s)
Figure 36. Average core outlet vapor temperature.
31
i l
la 80.0 i
i l
C O Xe/Kr L
a Cs/l l
.] 60.0 c
E.
I
@3
$ 40.0
'8 15.
c
.9e
[ 20.0 l
I 0.00 l
0.0 5000.0 10000.0 15000.0 l
Time (s) l Figure 37. Fraction of fission products released from the fuel.
30.0 i
i i
c 20.0 8
~
f E
E-l lii i
e
.C 10.0 C
OTotal Em e--e in core I
8z i
i 0.0 0.0 5000.0 10000.0 15000.0 i
Time (s) j Figure 38. Decay power distnbution.
i 3
i a
32
=
1 j
o Table 1. SCDAP components used to represent the AP600 core.
Component Radial Division / Flow Channel l Descnption 1
1/113 Fuel rods 2
I/I13 Control rods, instrument thimbles, empty guide thimbles 3
1/ 113 Gray rods 4
2 / 114 Fuel rodt 5
2 / 114 Control rods, instrument thimbles, empty guide thimbles 6
2 / 114 Gray rods 7
3 / 115 Fuel rods 8
3 / 115 Control rods, instrument thimbles, empty guide thimbles 9
4/116 Fuel rods 10 4/116 Control rods, instrument thimbles, empty guide thimbles i1 4 / 116 Gray rods i
12 5 / 117 Fuel rods 13 5 / 117 Instrument thimbles, empty guide thimbles 14 5 / 117 Stainless steel reflector, flat sections 15 5 / 117 Stainless steel reflector, corner sections j
l l
d 1
33
)
Table 2. Material masses in the active core region represented by SCDAP components.
Mass Material (kg)
UO (fuel) load 75 376 2
Stainless steel (reflectors, gray rods, and cladding; the reflector mass of ~6487 kg lying above and 46 562 below the active core region was included in separate RELAPS heat structures)
Zr (cladding, guide tubes, and spacers) 16 562 Ag-In-Cd (control rod absorber) 2771.1 Table 3. Coefficients used in a subcooled nucleate boiling correlation as a function of position on the exterior surface of the reactor vessel lower hemispherical head.6 CorTelation Coefficients Data Correlation Applied Subcooled Boiling Correlation Reported At At a
b c
1 L/D = 0 0 < L/D $ 0.1 0
319
-2.83 j
2 L/D = 0.20 0.1 < Le D s 0.275 4016 430
-4.13 3
L/D = 0.35 0.275 < L/D 5 0.425 0
337 2.61 4
L/D = 0.50 0.425 < L/D s 0.625 0
891
-9.04 5
L/D = 0.75 L/D > 0.625 0
529 0.08 i
34
i P,'
Table 4. Subcooled boiling correlations and CHF for nodes on the exterior surface of the COUPLE mesh.
2 Surface Node Number 0
UD l Subcooled Boiling Correlation CHF(MW/m )
1 0
0 1
0.544 i
2 9.97*
0.111 2
0.655 j
3 19.0 0.211 2
0.751 I
4 24.7*
0.275 2
0.811 5
30.6 0.340 3
0.869 6
35.4*
0.394 3
0.915 1
l 7
40.4*
0.448 4
0.962 8
45.9*
0.510 4
1.01 l
9 51.1" 0.568 4
1.06 l
l 10 55.0 0.612 4
1.09 11 59.l*
0.657 5
1.13 12 63.5 0.705 5
1.16 13 68.l*
0.756 5
1.20 i'
14 72.9 0.810 5
-1.24 t
15 77.9*
0.866 5
1.27 16 83.5*
0.928 5
1.31 17
> 90 5
1.36 18
> 90
>1 5
1.36 19
> 90*
>1 5
1.36 20
> 90*
>1 5
1.36 21
> 90*
>1 5
1.36 1.36 22
> 90
>1 5
j l
l i
3 i
l, i
l i
35 i
l' i
a
,o l
~
Table 5. Sequence of transient events.
Event Time (s)
Double-ended break in the DV12 pipe (transient initiation) 0 Low-l pressurizer pressure (< 13.20 MPa) 10.185 Low-l pressurizer pressure signal (low-l pressurizer pressure plus 0.7 s delay) 10.885-I Reactor scrams on low-l pressurizer pressure signal 10.885 Low-2 pressurizer pressure (< 12.86 MPa)
I1.655 Low-2 pressurizer pressure signal (low.2 pressurizer pressure plus 1.2 s delay) 12.855 S-signal on low 2 pressurizer pressure signal 12.855 CMT actuation (both sides) on S-signal 12.855 RCPs trip and begin coastdown on CMT actuation plus 15 s delay 27.865 Core collapsed liquid level falls below the top of the active fuel 37.868 j
Reactor vessel cavity water level reaches the bottom of the lower hemispherical head 132.65 CMT2 level < 67.5%
176.21 j
~ ADS 1 actuation signal (after CMT actuation and either CMT level < 67.5%)
176.22 ADSl (1 train /l valve) begins to open on actuation plus 20 s delay 196.23 j
ADS 2 actuation signal (60 s after ADSl actuation signal) 236.22 i
CMT2 level < 20%
246.80
)
ADS 2 (1 train /l valve) begins to open on actuation plus 30 s delay 266.22 1
i CMT2 empties 325.76 i
ADS 3 actuation signal (120 s after ADS 2 actuation signal) 356.22 j
ADS 3 (1 train /l valve) begins to open on actuation plus 30 s delay 386.23
- }
Accumulator-2 empties 460.00 4
j ADS 4-1 actuation signal (120 s after ADS 3 actuation and either CMT level < 20%)
476.23 ADS 4-2 actuation signal (30 s after ADS 4-1 actuation) 506.23 ADS 4-1 (1 train /l valve) begins to open on actuation plus 30 s delay 5 %.23 ADS 4-2 (1 train /l valve) begins to open on actuation plus 30 s delay 536.23 Accumulator-1 empties 640.00 CMTl level < 67.5%
908.60 CMTl level < 20%
1573.9 36
- s. i ' '
g Table S. Sequence of transient events. (continued) m Event Time (s)
CMTl empties 1971.2 First fuel clad failure (Component 4, Fuel Level 12) 4455.0 Core outlet tempemture reaches 1366 K 5666.3 First fuel (ceramic) melting occurs (Component 4, Fuel Level 8) 6594.9 Reactor vessel cavity water level reaches elevation of the bottom of the lower core plate 6777.0 Core collapsed liquid level falls below the bottom of the active fuel 10 346 First fuel melting adjacent to the reflector (Fuel Level 7) 10 782 First reflector failure (melt through, Component 14, Fuel Level 7) 10 970 First relocation into the lower head (via reflector failure, Component 14, Fue! Level 7) 10 970 Reflector failure (melt through, Component 14, Fuel Level 6) 11126 Relocation into the lower head (via reflector failure, Component 14, Fuel Level 6) 11 126 Relocation into the lower head (thru failed reflector, Component 14, Fuel Level 6) 1I232 Reflector failure (melt through, Component 14 Fuel Level 5) 11513 Relocation into the lower head (via reflector failure, Component 14, Fuel Level 5) 11513 First relocation of control rod materials (Ag-In-Cd) into the lower head 11891 Relocation of control rod materials into the lower head 11987 Reflector failure (melt through, Component 14, Fuel Level 4) 12 047 Relocation into the lower head (via reflector failure, Component 14, Fuel Level 4) 12 (M7 Relocation of control rod materials into the lower head 12 189 Relocation of control rod materials into the lower head 12 263 Relocation of control rod matenals into the lower head 12 431 First molten penetration through the bottom of the active fuel (Component 4) 12 844 First relocationinto the lower head through the lower core plate 12 844 End of calculation with lower head debris including 12 844
-63 995 kg of UO.
2
-7586.4 kg of ZrO.
2 I
-5572.6 kg of Zr,
-2771.1 kg of Ag-In-Cd, and
-3350.9 kg of stainless steel l
I 37 i
i
i I
1 Table 6. Summary of relocation esents.
Relocated Mass (kg)
Temperature Relocated Time (s) of Relocated Material Decay j
2 ZrO:
Zr l S. Steel Ag-In-Cd Material (K)
Power (MW) l'0 l
i 10 970 7 145.4 806.72 ! 822.48 556.24' l l
3076 1.730 I
11 126 8 428.0 964.24 863.46 1341.7b 3095 2.116 i
11232 4 559.3 524 77 I 463.93 157.25C 3115 0.982 d
11 513 10 478 1226.1 1046.3 342.76 3145 2.295 11 891 956.41 1073 11987 627.79 1073 s
12G47 6 596.3 792.43 638.07 205.13' 3182 1.416 12 189 954.20 1073 12 263 174.20 1073 l
12 431 i
! 58.500 1073 f
12 844 26 788 3277.6 1732.9 j 747.82 l
3199 5.779 l
l 63 995 l 7591.9 l 5567.1 l 3350.9 l 2771.1 l Relocated Mass Totals
- a. Includes 453 06 kg of lower plenum structural stainless steel at a temperature of 393 K that was submerged dunng relocation.
b includes 1220.0 kg of lower plenum structural stainless steel at a temperature of 416 K that was submerged dunng relocation.
- c. Includes 91410 kg of lower plenum structural stainless steel at a temperature of 438 K that was submerged during relocation.
1 Includes 191.45 kg of lower plenum structural stamless steel at a temperature of 427 K that was submerged dunng relocation.
- c. Includes 109 88 kg oflower plenum structural stainless steel at a temperature of 417 K that was submerged during relocation.
- f. Includes 36100 kg of lower plenum structural stainless steel at a temperature of 410 K that was submerged dunng relocation.
l 38