ML20206F071

Flow-Pattern Results for Tmlb' Accident Sequence in Surry Plant Using Melprog

ML20206F071
Person / Time
Site:	Surry, 05000000
Issue date:	04/08/1987
From:	Dearing J LOS ALAMOS NATIONAL LABORATORY
To:	Han J NRC OFFICE OF NUCLEAR REGULATORY RESEARCH (RES)
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RTR-NUREG-1150 LA-UR-85-3668, NUDOCS 8704140103
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CONTENTS b .

ABSTRACT ................................................................... 1

.l. I NTRODUCT I ON . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . I

11. VESSEL MODEL .................................................... 2 111. INITIAL AND TRANSIENT BOUNDARY CONDITIONS ....................... 3 IV. BASE CASE ....................................................... 3 V. ONE-DIMENSIONAL CASE ............................................ 5 VI. NO RADIATION CASE ............................................... 5 VII. COMPARISON WITH COBRA NC CALCULATION ............................ 6 A. Modeling differtRCes ....................................... 6 B. Comparison of results ...................................... 6 Vllt. NONPHYSICAL EXTENSION OF BASE CASE .............................. 7 IX. CONCLUSIONS ..................................................... 7 X. REFERENCES ...................................................... 8 l-e

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FLOW-PATTERN RESULTS FOR A 1MLB' AOCIDENT SEQUENCE IN THE SURRY PLANT USING MELPROG* 1

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by l James F. Dearing

. ABSTRACT A preliminary version of MELPROG/ MOD 1 was used to calculate in-vessel flow patterns for a station blackout sequence at the Surry Nuclear Power Plant. The calculation, which was terminated with control-rod failure, showed significant mining of mass and energy

• between the core and upper plenum. This mining was caused by natural convection, driven by density differences between the core center and periphery. These results differ significantly from previous work in which one-dimensional flow was assumed.

1. INTRODUCTION The new two-dimensional, four-fluid FLUIDS module' of the MELPROG eode' allows simulation of multidimensloral transport processes within a pressurised water reactor (PWR) vessel during a core disruptive accident. MELPROG is designed to give an integrated, mechanistic treatment of PWR core meltdown accidents from accident initiation to vessel melt-through,. The accident sequence analyzed in this report is a station blackout (1MLR') sequence for the Surry plant. This analysis terminates with control-rod failure. The analysis will be continued to vessel melt-through in the near future, when the '

, fourth fluid (molten corium) has been fully integrated with the other modules in MELPROG.

In addition to the base calculation, two parametric cases are presented.

A case in which flow in the core is limited to the antal direction was run to U show the effects of natural circulation of steam and hydrogen between the 1

'This work was funded by the US PatC Office of Nuclear Regulatory Research, '

l Division of Accident Evaluation.

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l 2-i j. . upper plenum and core. A case in which the radiation heat-transfer module was j

deactivated was run to show the effects of radiation heat transfer between the structures and betwen the structures and fluids. Finally, the base case was estended using nonphysical assumptions (no loss of geometry) for comparison with a calculation by another code.

i II. VESSEL MODEL 1

! The model used for this analysis is shewn in Fig.1. Five radial and thirteen asial nodes are used in a cylindrical grid representing the reactor vessel. The calculation is bounded on the lower surface by the lower head, on f.' the upper surface by the upper head, and on the outer surface by the vessel well. The first three radial nodes are used to model the core, the fourth node re pre se'nt s the core-bypass region. and the ,fif th node represents the downconer. All of the asjor vessel structures are modeled, as shown la Fig. 1. llent conduction occurs between hydraulic cells through walls and plates (for example, betwen the core bypass and the downcwer through the core barre!). A mass flow boundary condition is used at the cold-leg

, junction, while a pressure boundary condition is used at the hot-leg junction.

A realistic simulation of the upper head requires more than a two dimensional model. Flow paths esist between the downconer and the upper head (via cooling spray nostles) and between the upper plenum and the upper l bead (via the control-rod drive covers). These flow paths need to be ,

represented in addition to the flow paths simulated by the two dimensional ,

mesh. The FLUIDS module has been modified to allow the user to specify a I Ilmited number of additional interfaces between cells. Three of these

" embedded" interfaces are used in this analysis to represent the flow pathe Inside the control-rod drive covers, and one is used to represent the head ,

! spray cooling nostles (Fig. 1).

Steady state values of pressure drops and flows were used to calibrate the model. Specifically, it of the total vessel flow cools the upper head via l

., the spray cooling nogales. The amount of flow penetrating the core barrel and flowing dowe the, core bypass was adjusted to 0.52%. Flow resistances wre .

l adjusted to achieve the required pressure drops between the inlet nogales and  !

diffuser plate, and betwen the diffuser plate and outlet nogales. Aslal

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3 lengths of cells representing the upper and lower heads were adjusted to

'. achieve the proper fluid volume of these hemispherical regions. Finally, the Final Safety Analysis Report (FSAR) values of total vessel flow and power produced the required temperature rise and core average velocity.

III. 1NITIAL APO TRANSIENT BOUNDARY CXWOITIONS A TRAC (Ref. 3) calculation

• for Zion-1 PWR was used to provide initial conditions for MELPROG 6400 e into the transient, or 80 s befc,re boiling began in the core. The cold-leg flow was scaled to represent Surry (Zion is a four-loop, while Surry is a three-loop Westinghouse plant) and was equal to 3.5% cf nominal at 6400 s. This flow is due to natural convection within the primary system, and was set to zero at the initiation of boiling in the core.

(The TRAC-calculated flow approached zero soon after boiling initiation). The cold-leg water temperature at 6400,s was 615 K and slowly increasing. The pressure boundary condition representing the hot legs was set to the power-operated relief valve (PORV) set point of 16.3 lWa.

I Boundary conditions for the convective terms la the event of inflow at

' the hot legs were determined assuming a zero first spatial derivative (i.e..

what comes in is identical to what borders the pressure boundary conditlos).

I i These boundary conditions n're necessary for this type of transient (in which the system is heating up and expanding, causing outflow at the hot legs) only during bolling oscillations.

l IV. BASE CASE The sequence of events for the base case is given in Table 1. The calculation was terminated at 9804 s. when the control rods in cell (1.7) '

(radial, antal node. as shown in Fig.1) melted. Figure 2 graphically shows the state of the vessel at this time.

The top number in each cell represents the temperature (K) of the 'luid with the largest volume fraction, while the bottom number represents the surface temperature of the structure or fuel pins (in cells that contain both fuel pins and structure, the pin surf ace temperature is shown). Dotted line l

I density is propoYtional to , liquid water volume fraction, wblie the angle of .

I the dotted lines from the vertical is proportional to hydrogen partial

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4 pressure. The fuel-pin volume fraction is represented by the vertical solid

"(. lines in the core regions, while impenetrable (in the two-dimensional sense) walls are represented by closely spaced parallel lines. The velocity vectors show a four-point average of the axial and radial cell interface velocities for the fluid with the largest volume fraction in that cell. Additional embedded flow paths that represent the control-rod drive covers and the upper-head spray nozzles are not shown on Fig. 2 for clarity. i A strong recirculating flow exists in the upper plenum at 9304 s. This recirculation extends to a smaller entent to the bottom of the core, and is

- removing approximately 40% of the core decay heat. The recirculation is l driven by density differences, which are caused by both tempera ture differences and differences in the relative fractions of steam and hydrogen.

Temperatures are botter in the central regions of the core, causing the Zircalloy to oxidize faster, creating more hydrogen. The larger source of hydrogen near the center increases circulation above the level that would be ,

given by thermal natural convection alone. flydrogen is well mixed in the core (with a partial pressure of 40%), with only a slightly lower partial pressure in the upper plenum. ilydrogen has penetrated the upper head (through the control rod drive covers) to a lesser entent, and is virtually absent from the downconer. The total amount of hydrogen produced to this time is 45 kg.

I The upper plenum is acting as a heat sink for the energy transported out  ;

of the core by the recirculating flow, so that structural surface. temperatures t are lagging cladding surface temperatures by only about 300-400 K, on the (

average. Although the upper head is also being heated by natural convection i through the control-rod drive covers, this flow path is much more restrictive, '

and the surface temperatures in the upper head are at least 150 K cooler than l the corresponding cell directly below in the upper plenum. l The liquid water level has fallen below the botta of the lower core ,

support plate, and is evaporating slowly. The water level in the dowmeer is ,

significantly lower, because of the difference in static head between the cold i steam in the downcomer and the hot steam / hydrogen misture in the core. The l

' I temperature of the water in the lower head is actually slowly decreasing because of thermal conduction through the lower head.

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V. ONE-DIMENSIONAL. CASE To show clearly the effect of the two-dimensional recirculati,on into the core, the vessel model was modified to limit flow in the core to the axial direction. This was done by collapsing the central three radial nodes into one, as shown in Fig. 3. This figure shows conditions in the vessel at 9020 s, when the control rods melted in node (1,8). The time of control-rod melting is 13 min earlier in this case than in the base case, because there is no recirculation pattern to remove energy from the core. Because of the lack of thermal mixing, thermal gradients are much steeper. The bottom of the core is 400 K colder than in the base case (so that the Zircalloy is not even oxidizing), while the upper plenum is also approximately 400 K colder. The lack of mixing produces a high concentration of hydrogen in the hottest cell (1,8), but very little elsewhere. The total amount of hydrogen produced to this time is 29 kg. Two-dimensional recirculation obviously plays a crucial role in transporting energy and mass within the vessel, and making them available for possible transport in the hot legs.

VI. NO RADIATION CASE To show clearly the effect of MCLPROG radiation heat-transfer module on the calculation, the calculation was run with this module disconnected (using an input patameter). The results, again at the time of initial control rod melting, are shown in Fig. 4. This time, 9773 s, is 31 s earlier than in the base case. More leportantly, the cladding and structural surface temperatures show steeper antal and radial temperature gradients because of the lack of temperature radiation heat transfer between the surfaces. The local dif ferences between the surfaces and the fluid are larger because of the lack The maximum

. of radiation heat transfer between the surfaces and the fluid.

cladding surface temperature is 90 K at (2,7). These

, difference in differences would be even larger were it not for the strong mining effect of the flow recirculation. They can be espected to be significantly larger later

' ' in the calculation, when thermal gradients are also larger. The total amount of hydrogen produced is 49.kg, more than in the base case, because of the higher cladding temperatures.

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• 6-S V11. COMPARISON WITH COBRA-NC CALCULATION The following comments are based on my review of the draft document. i A. Modeline Differences COBRA is a fixed-geometry code, and as such has been pushed to and beyon MELPROG, on the other the limit of its capabilities in this calculation.

hand, is a variable-geometry code and has exercised only a small fraction o its capabilities.

The 00 BRA model uses a finer mesh (10x24) than the MELPRO Use of a fine mesh makes possible the definition of small-scale (5x13).

. effects such as stagnant regions in the circulation pattern, which also turbulence modeling (for transport of energy and mass) for requires correct proper s imul'a t i on. Our lack of data makes accurate simulation of such small-scale effects questionable.

The COBRA calculation used initial conditions obtained fro 3 case, while the HELPROG calculation used initial conditions obtained from a The COBRA calculation began with core uncovery at 5730 s, while TlyC case. The difference may be in the crude MELPROG predicted core uncovery at 7100 s.

MAROI calculation of mass lost through the PORVi that is, the TRAC calculation less of the primary system water was lost before the may have predicted that onset of boiling. The main effect of the different initial conditions is that the MELPROG decay power is 14% lower.

The COBRA calculation does not include radiation heat transfer, and it is unclear whether cell-to-cell structural beat conduction is included.

B. Comparison of Results Qualitatively, the results are very similar. given CX) BRA-calculated flow patterns are given, but velocities are not quantitatively.

The flow patterns appear similar to those calculated by The MELPROG mesh is too large to resolve the MELPROG on a larger mesh size.

i downflow in the central nine fuel assemblies predicted by COBRA, but in any case such a downflow would require a significant dip in the power distribution

, l in the central subassemblies.

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7 Calculations show similar timing between core uncovery and the time that f '

the liquid level falls below the fuel rods, but the subsequent time to Zircalloy oxidation is much smaller in the COBRA case. Possible reasons for this are that .the decay power level is higher in the COBRA case (earlier la time), hot channels are better defined in the finer noding, and wall heat-transfer coefficients may be slightly lower. After the onset of oxidation in the 00 BRA case, comparison of results is no longer meaningful because of the oxidation reaction energy.

Vill. NONP)lVSICAL EXTENSION OF BASE CASE The NRC requested that the base-case calculation be extended from control-rod failure, with the assumption that no geometry changes occur, so that the results could be compared with COBRA. Fig,ure 5 shows the results of this extension at 9986 s, when the hottest node (1,7) is almost completely full of hydrogen. Subsequent to 9986 s, the rest of the core fills up with hydrogen, shutting down the Zircalloy oxidation reaction. After this occurs, comparison with COBRA is no longer meaningful. The COBRA calculation showed a f, maximum partial pressure of hydrogen of 70% when the calculation was terminated. The difference. In results is probably due to the differences in timing (power) and hot-channel factors discussed above.

It was necessary to assume nonphysical behavior in order to create Fig. 5. In particular, the melting control rods would have created a significant source of steam when they fell into the lower he'ad, and the steam starvation depicted in Fig. 5 may not have occurred.

i IX. CONCLUSIONS These calculations show the need to include multidimensional effects in I

realistic models of degraded-core accidents. Natural convection between the core and upper plenum provides sufficient thermal mining to change event timing significantly from that of a one-dimensional model. Material components are also well mixed, in contrast to distributions predicted with

/ low "once through" velocities. This mining of mass and energy will be very important to the* transport and deposition of fission products later in the accident.

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. , ,. REFERENCES L

b 1. J. F.

Dearing,

"A Four-Fluid Model of PWR Degraded Cores," Third International Topical Meeting on Reactor Thermal Hydraulics, to be held October 15-18, 1985, Newport, R.I.

. 2. ' W. J. Camp. M. F. Young, J. L. Tomkins, J. E. Kelly, P. J. Maudlin, R. J. Henninger, "MELPROG-PWR/ MODO

• A Mechanistic Code for Analysis of Reactor Core Melt Progression and Vessel Attack Under Severe Accident Conditions," Sandia National Laboratories report SAND E5-0237 (February 1985).

3. Safety Code Development Group, " TRAC-PF1, An Advanced Best-Estimate Computer Program for Pressurized Water Reactor Analysis," Los Alamos National Laboratory report LA-9944-MS (NUREG/CR-3567) (February

. 1984).

4. Brent E. Boyack, " Loss of Offsite Power Transient for the Zion-1 PWR " Los Alamos National Laboratory document LA-UR-83-1714 (June 1983).

5. M. J. Thurgood T. E. Guidotti, C. L. Wheeler, " COBRA-NC Analysis of a Station Blackout Transient (TMLB') For the Surry Plant (draft),"

Battelle Pacific Northwest Laboratory report FATE-85-103 (February J , 1985).

6. J. A. Gieseke et al., "Radionuclide Release Under Specific LVR Accident Conditions'" Battelle Columbus Laboratories report BMI-2104, Vol. 5 (July 1984).

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