ML20206F086

Draft COBRA-NC Analysis of Station Blackout Transient (Tmlb') for Surry Plant. Inel Viewgraphs Entitled Structural Failure Studies of RCS Also Encl

ML20206F086
Person / Time
Site:	Surry, 05000000
Issue date:	02/28/1985
From:	Guidotti T, Thurgood M, Trent D Battelle Memorial Institute, PACIFIC NORTHWEST NATION
To:	NRC
Shared Package
ML20204G644	List: ML20206C066 ML20206E142 ML20206F071 ML20206F086 ML20206F182 ML20206F313 ML20206G821 ML20206G928 ML20206H081 ML20206H113 ML20206H144 ML20206J252 ML20206J270 ML20206J368 ML20206J403 ML20206J600
References
RTR-NUREG-1150 FATE-85-103, FATE-85-103-DRFT, NUDOCS 8704140111
Download: ML20206F086 (52)

Text

?~g'o

, .T. 3.1S Mk l FATE-85-103

. .e.;/ma ;

i l ..u.:

ar.fi d NJ hi . [

COBRA-NC ANALYSIS OF A STATION BLACKOUT TRANSIENT (TMLB') FOR THE SURRY PLANT r

e

.k .

M. J. Thurgood .

T. E. Guidotti

. C. L. Wheeler February 1985 i

APPROVED d .

D. 5. Trant, Manager /

Fluid and. Thermal Enginedrin(g/ Section f, Engineering Physics Department *

. BATTELLE PACIFIC NORTHWEST LABORATORY RICHLAND, WASHINGTON 99352  ;

1 l

Prepared by Numerical Applications, Inc. l Richland, Washington 99352 I I

This report is a working paper intended f6r the sponsor and other contributors l to the program. Do not reference in open literature at this time.

8704140111 870408 PDR NUREG 1150 C PDR .

=

This report was prepared as'~ an account of work sponsored by the United States Nuclear Regulatory Commission. The work has.been performed by Numerical Applications, Inc., under subcontract to Battelle Northwest Laboratories, Richland, Washington. Numerical Applications, Inc., nor any agent or officer thereof nor any of their employees, makes any warranty, expressed or implied.

C. or assumes any legal liability of responsibility for any carty's use, nor the ,

results of such use, of any information disclosed in this report.

i.

i l

(m L. . ,

iii

? .

, 6. . .

l .

W., - .

d i

o .

c ABSTRACT

~'

An analysis of convection loops between the reactor core and the upper plenum during a TMLB' accident sequence has been performed using the COBRA.NC computer code. The SURRY plant has been used as the reference plant for the calculation. The calculation has shown tnat significant convection loops do develop between the core, upper plenum, and upper head. This convective flow does provide extra cooling to the core. However, it also delivers additional steam from the upper plenum and upper head to the core. This steam was  !-

effective in maintaining tne metal / water reaction until zircaloy melting

" temperatures were reached . Whether upper plenum or loop components will fail before significant core slumping occurs must be determined by further analysis.

• 1

.i, 9

e b

f o

5

/-

l h .

> 1v 6

9 8

l

CONTENTS

~

ABSTRACT iv EXECUTIVE

SUMMARY

viii

1.0 INTRODUCTION

1

2.0 DESCRIPTION

OF THE SURRY REACTOR VESSEL 3 i 3.0 MODEL DESCRIPTION 4 3.1 Input Model 4 3.2 Computational Model 7

- 4.0 RESULTS 9

5.0 CONCLUSION

S AND RECOMMENDATIONS 25 i 30 .ig REFERENCES

.* Y j e

FIGURES l

1 Reactor vessel internals 3 2 Upper plenum internals 4 3 Computational mesh 5 4 Core liquid fraction vs. time 10 5 Clad surface temperature for rod number 1 10 6 Clad surface temoerature for rod number 2 11 7 Clad surface temperature for rod number 3 11

- 8 Clad surface temperature for rod number 4 12 9 Clad surface temoerature for rod number 5 12 10 Clad surface temperature for rod number 8 13

• 11 Support column surface temperature above rod 4 at ,

not leg elevation 14 ,.-

12 Upper core plate temperature above rod 5 14 13 Guide tube temperature above rod 8 at not leg elevation 15 14 Support column temperature above bundle 8 at not leg elevation 15 15 Core barrel temperature at hot leg elevation 16 16 Steam temperature at hot leg inlet 16 17 Comparison of temperatures for several comoonents 17 18 Hydrogen concentration in bottom 5 cells of radial ring 4 18 19 Hydrogen concentration in too 5 cells of radial ring 4 18 20 Vapor / gas velocities at 401 seconds 20 21 Vapor / gas velocities at 813 seconds 20

, 22 Vapor / gas velocities at 1203 seconds 20 23 Vapor / gas velocities at 1607 seconds 20 k .~ . , Vi 9

~ . . *

o. -

w 24 Vapor / gas velocities at 1668 seconds 21

25. Vapor / gas velocities at 1768 seconds 21 26 Vapor / gas velocities at 1809 seconds 21 27 Vapor / gas velocities at 1869 seconds 21 28 Vapor / gas velocities at 1909 seconds 22 29 Vapor / gas velocities at 1949 seconds 22 30 Vapor / gas velocities at 1990 seconds 22 31 Gas temperature at 1600 seconds 22 32 . Gas temperature at 1668 seconds -

23 33 Gas temperature at 1708 seconds 23 34 Gas temperature at 1768 seconds 23 ,

! 35 Gas temperature at 1809 seconds 23

• 36 Gas temperature at 1869 seconds 24 *)

37 Gas temperature at 1900 seconds 24 38 Gas temoerature at 1949 seconds 24 39 Gas temperature at 1990 seconds 24 40 Hydrogen concentratio.1 at 1607 seconds 26 41 Hydrogen concentration at 1668 seconds 26 42 Hydrogen concentration at 1700 seconds 26 l

43 Hydrogen concentration at 1768 seconds 26 44 Hydrogen concentration at 1909 seconds 27 ,

j

! 45 Hydrogen concentration at 1869 seconds 27 l 46 Hydrogen concentration at 1909 seconds 27 47 Hydrogen concentration at 1940 seconds 27 48 Hydrogen concentration at 1990 seconds 28 l

.,' i vii Q...'

. O EXECUTIVE

SUMMARY

An analysis of the effect of natural circulation between the reactor core and -

the upper plenum during a TMLB' accident sequence was performed using the COBRA-NC computer code.

The Surry nuclear power plant was selected as the reference plant for the calculation. Components of the entire reactor vessel (downcomer, lower plenum, core, core-by-pass, upper plenum, and upper head) were included in the .

calculational model. The model used an axisymmetric mesh which incorporated the following noding distributions:

Core region 8 radial cells,10 vertical cells Core bypass 1 radial cell, 10 vertical cells Downcomer 1 radial cell,17 vertical cells Upper plenum Igradialcells,7verticcicells Guide tubes 8 radial cells,1Q, vertical cells Upper head 17 radial cells, 6 vertical cells The fuel rods in each bundle were modeled using a single average fuel rod of tne overall rod geometry for that bundle. The heat flux from the average rpd

- is multiplied by the number of rods in a region to give the total energy * *,

transcort to the fluid. All of the structures within the vessel were modeled using the unheated conduction ootion. The hot and cold leg connections were' represented by constant pressure boundaries of 2371.2 osi. The calculation'was initiated at the start of core uncovery, with all structure temoeratures set to the saturation temperature of 660.3*F. The effect of metal / water reaction of airconium is modeled using the Cathcart reaction rate coefficients.

The presented results show: ,

e core liquid fract. ion versus time e surface temperature versus time for selected rods and structures e hydrogen generatipn 'versus time at selected locations e vapor velocity vector plots at selected times e temperature contour maps at selected times e hydrogen concentr'ation contour map at selected times.

4

• Part of the 16 radial nodes in the UDoer plenum, and, also part of the 17 nodes in the upper head

• • Down to three vertical cells on the periphery to model dcwncomer curvature. ,

h.,,

viii 4

4 1

. o The results of the simulations which are subject to several modeling

. constraints as discussed in the report show:

. . e Convection loops between the core, upper plenum, and upper head do develop.

dI e The convection loops cause the clad temoer3ture to lag those predicted using one-dimensional models by about 10 min.

• The convection loops bring steam down into the core from the uocer plenum and upper head. This steam maintains the metal / water reaction in the Core.

e The heat generated from the metal / water reaction dominates the temperature rise in the clad.

e The cladding reaches melting temperatures while the upper plenum structure is relatively cool (2000'F).

e The vessel average steam volume fraction is about 50% at the end of the calculation so the metal / water reaction will be maintained.

• The modeling constraints (no zircaloy heat of fusion, no radiation neat ,'

transport, no convecti,on from loop components) applied to this analysis are f such that the results are inconclusive with respect to answering the question as to whether or not upper plenum and loop components will fail before 4

significant core slumping occurs. The question can only be answered by further, more detailed, analysis.

(

9 0

4 m *

(9 ..

ix I

e

. o COBRA-NC ANALYSIS OF A STATION BLACKCUT TRANSIENT (TMLB') FOR THE SURRY PLANT C-

1.0 INTRODUCTION

Several accident sequences have been postulated that may result in core meltdown, reactor vessel failure, and ultimately failure of the nuclear containment and the release of radioactive material to the environment. Among these, the station blackout transient (designated TMLB') is among the most U dominant. The TMLB' transient has been analyzed by several investigators (Ref. 1 and 2). These analyses have been performed using one-dimensional models for the tnetsnal hydraulics of the reactor core and upper-plenum and have, therefore, neglected the possibility for convection looos between the core, the upper-plenum, and the upper head. Such convection flows could play an important role in providing additional cooling to the core following core uncovery. More importantly, however, is the possibility that these flows will

- transfer heat to upper plenum structures and to the reactor vessel wall in the vicinity of the hot leg nozzles. If the heat transfer rate is sufficiently high and the heatin.g rate in the core is significantly delayed, the failure .of' the primary system may occur in the upper elevations of the system (i.e., hot leg nozzle welds, instrument penetrations, oumo seals, steam generator tubesg etc.), changing the scenario for the accident from a molten fuel ejection'at, high pressure due to bottom head failure to a low-oressure ejection at a later time in the accident. The consequences and probabilities of radioactive (

release to the environment would then be significantly different than are now predicted for the TMLB' accident.

k The purpose of this analysis is to determine if, indeed, convection flow does occur between the core and upper plenum and, if it doet, how the heatuo of the core and structures in the upper olenum are affected. This analysis does not investigate the possibility of convection flow between the reactor vessel and  ;

> the remainder of the primary loop, altnough it is recognized that such flows I are a real possibility.

The SURRY nuclear power plant has been selected as the reference plant for the calculation. The entire reactor vessel is included in the model (downco:ner, lower plenum, core, core-by-pass, upper-plenum, and upper head), altnough tne lower-plenum has been represented with a ratner coarse mesh.

Tne COBRA-NC computer code has been used for the analysis. This code has i appropriate fluid dynamic and convective heat transfer models to calculate the l thermonydraulics within the reactor vessel during the transient. The code does not have models for fuel rod ballooning and deformation and, therefore, the results of the analysis must be viewed witn some skepticism once clad reaches

~

temperatures at which these %ffects cause significant alteration in the flow geometry.

A simple surface-to-steam radiation model is available in the code; however, it was not applied in this calculation. A more sophisticated radiation model would have to be applied to correctly credict the radiative heat transfer between the steam / hydrogen mixture and the heat transfer surfaces.

l h 9

I g e

A brief description of the SURRY reactor vessel is given in the following section to farniliarize the reader with the geometry which is being analyzed.

[ Th'is is followed by a description of the COBRA-NC model for the vessel that is

\' used in performing the analysis. Following the model description, the results of the analysis are presented.

l

)

. k,

- 2

(? .

I kw -

2.0 DESCRIPTION

OFTHESURRYREACTORVkSSEL i 48.u

(- '

The SURRY nuclear power plant is a fotfr-loop. 2441 MW thermal PWR of Westinghouse design. The core consists of 15715x15 fuel assemblies, each containing 104 fuel rods. Each rod has a 144 in, heated length. The ohysical arrangement of the reactor vessel internals is shown in Figure 1. The upper plenum contains 65 guide tubes and 36 suoport columns. The ohysical arrangement of the upper plenum internals is shown in Figure 2. The location of the upper plenum internals is important in the current analysis because they determine the distribution of heat sinks within the upper plenum which, in l

turn, influence the flow patterns between the upper plenum and the core.

' I The upper support plate is a deep beam design and provides additional heat

. transfer surfaces in the upper plenum. The guide tubes provide flow paths between the upper plenum and upper head. The reader should refer to the plant j .

4 FSAR for a more detailed description of the vessel geometry. l l

l ,

.P" '% .

l

- sg ,,,,

~

(. _-c

O ,

- c..

/ "

t t

c  !

i i

i ,

)

l Figure 1. Reactor vessel internals i

(

e A.., 3 i

1 I

- ..v._---..~... -

- - . - - - - _ - _ - - - . - - . - - _-. . _ . . -. -- . = - - - - _ = _ . __

~. -

I to the upper plenum and upper head. The mesh in the upper head is nine cells radial by six cells vertical, and it also contains the local cells representing t

the guide tubes. The number of vertical mesh cells is reduced towards the -

I(- outside diameter of the upper head to represent the curvature of the upper head. The lower plenum and core inlet regions were modeled with a one cell high mesh with connections to the downcomer and core to allow water to enter

the core from the downcomer as boiloff occurs in the core. -

The fuel rods in each bundle were modeled using a single average fuel rod having the same radial and axial dimensions as the actual fuel rods. The heat transfer and conduction calculation is performed on this single rod. The heat flux from the single rod is then multiplied by the number of rods in the region to give the total heat flux to the fluid. Two radial heat transfer nodes were used in the clad, and six radgal nodes were used in the fuel. A constant gao conductance of 2000 8tu/hr-ft *F was used in the calculation. The dynamic gao conductance was not used. ,

l All of the structures within the vessel were modeled using the unheated j

conductor model. The vessel wall,' core barrel, guide tube and support columns were modeled using the tube conductor geometry with aooropriate inside and outside diameters of the actual structures specified as input. The baffle .

plates, core plates, . top fuel nozzles, deep beam weldments, and uoper supp~ers plate were modeled with the slab conductor geometry. The fuel too nozzle and.'

core plate were lumped into a single slab for each radial ring tnat had the g-same surface area and volume as the actual components. The heat flux to a i single guide tube or support column is multiplied by the total number of guide

, tubes or support columns in tne radial zone to get the total heat flux from :ne i

fluid.

The outside and, where appropriate, inside surfaces of each conductor have been connected to the fluid cell containing the structure so that the correct radial distriaution of heat sinks witnin the reactor vessel was maintained. Since

! thermal connections to slabs cannot be made in the vertical direction of the mesh; the upper-support plate was divided at' the mid-thickness into two conduction slabs. One slab was connected to the fluid cells in the upper plenum and the other to fluid cells in the upper head.

Pressure boundary conditions were specified on the third vertical cell of the

, outer radial rings of tne upper plenum and downcomer to represent the hot and cold leg connections. The pressure at these boundaries was set equal to the

• pressurizer relief valve set point, 2371.2 osia. A loss coefficient representative of the 10o0 resistance was also specified at the boundaries.

1 The decay heat as a function of time was specified for the heat flux to the

fuel rods. The decay heat was calculated based on a core uncovery time of 5730.0 sec.

The calculation was initialized with all of the structural surfaces and the fuel rods at a temperature of 660.3*F, the saturation temperature for the j i system pressure. The core and downcomer void fractions were set so that the l i

two-phase level was at the top of the heated length. The upper plenum, upper

, head, and the top of the downcomer were initialized with saturated steam.

(L +

6 1

--- -- _ , _ . . , _ . . . - . _ -._,.m _ _ _ _ _ m. __ ~. _ ,__,,_.,._,-_,.m_...~_ , _-m.-..m . .. .

- =-

L -

. < - , .x

{ '

},{

C .

\

Figure 2. Upoer plenum internals i ,

3.0 MODEL DESCRIPTION 3.1 Inout Model i The reactor vessel was mocialed using a two-dimensional mesh. A two-dimensional mesh was used to maximize the number of radial mesh cells while maintaining reasonable computational costs. The mesh was developed by arranging the fuel assemblies in rectangular rings. This was done to simplify the distribution of fuel assemblies and upper plenum structures among the radial nodes. This mesh will create some distortion of actual flow patterns since not all fuel I assemblies in a radial ring are equidistant from the center of the core and,

! also, since. the hot leg is assumed to be distributed around the perimeter of the outer ring. These limitations are acknowledged but are believed to be

acceptable for the purposes and limitations of the overall analysis.

A vertical section of the mesh is shown in Figure 3. The core region is l divided into 10 cells vertically and 8 cells radially. Each cell r2 presents ,

one row of rod bundles. Bundle 1 is at the core center line and bundle 8 is at l

the outer parameter of the core. The core bypass and downcomer are modeled '

with one radial cell each and have the same length as the corresponding mesh cells in the core. The upper plenum is modeled using eight cells radially and 1

seven cells vertically. In addition, eight columns of local mesh cells are l i placed within each radial ring to model the flow path through the guide tubes 6:..

t k

l

( L II I I I I~

T~ 'l

• pperU Support

]

~ '

Plate 3 . ..

I i l " l- -

l l  : Guide Tube Cells i t

. Ig .

__ P= I

~ ~

Cold Leg

~

and Hot Leg l ,--

Boundary y

J ~-

g Conditions - L

'h i!][ I f l T i . . .pp., Co,. P,...

1 2 ,4l5,6 7 8 rL

(~ I I

Li~ I~ l I l l l Downcomer ,

! j_ l ~

I -I -~ ~ ~^

I __ ~ C_

_y . _

j

~

l l ,

Lower Core Plate l II l.  ;

I l l 1 l l i

, Figure 3. Computational mesh s .

\.-

l i

to the upper plenum and upper head. The mesh in the upper head is nine cells radial by six cells vertical, and it also contains the local cells representing

(e

, the guide tubes. The number of vertical mesh cells is reduced towards the ,

~

outside diameter of the upper head to represent the curvature of the upper

head. The lower plenum and core inlet regions were modeled with a one cell

! high mesh with connections to the downcomer and core to allow water to enter .

l 1 the core from the downcomer as boiloff occurs in the core.

4 The fuel rods in each bundle were modeled using a single average fuel rod having the same radial and axial dimensions as the actual fuel rods. The heat transfer and conduction calculation is performed on this single rod. The heat flux from the single rod is then multiplied by the number of rods in the region to give the total heat flux to the fluid. Two radial heat transfer nodes were used in the clad, and six radgal nodes were used in the fuel. A constant gap conductance of 2000 Btu /hr-ft 'F was used in the calculation. The dynamic gao conductance was not used.

I All of the structures within the vessel were modeled using the unheated conductor model. The vessel wall, core barrel, guide tube and support columns l l

were modeled using the tube conductor geometry with appropriate inside and  ;

outside diameters of the actual structures specified as inout. The baffle . - 1 S.

plates, core plates, too fuel nozzles, deep beam weldments, and upper support .

plate were modeled with the slab conductor geometry. The fuel too nozzle and .

core plate were lumped into a single slab for each radial ring that had the i; same surf ace area and volume as the actual components. The heat flux to a single guide tube or support column is multiplied by the total number of guide tubes or support columns in the radial zone to get the total heat flux from the

! fluid.

The outside and, where appropriate, inside surfaces of each conductor have been  !

connected to the fluid cell containing the structure so that the correct radial 1 distribution of heat sinks within the reactor vessel was maintained. Since thermal connections to slabs cannot be made in the vertical direction of the

! mesh; the upper-support plate was divided at the mid-thickness into two 1 conduction slabs. One slab was connected to the fluid cells in the upper ,

j plenum and the other to fluid cells in the upper head. '

\

Pressure boundary conditions were specified on the third vertical cell of the outer radial rings of the upper plenum and downcomer to represent the hot and

cold leg connections. The pressure at these boundaries was set equal to the

!- pressurizer relief valve set point, 2371.2 psia. A loss coefficient

! representative of the 1000 resistance was also specified at the boundaries.

The decay heat as a function of time was specified for the heat flux to the fuel rods. The decay heat was calculated based on a core uncovery time of 5730.0 sec.

l The calculation was initialized with all of the structural surfaces and the fuel rods at a temperature of 660.3*F, the saturation temperature for the I i system pressure. The core and downcomer void fractions were set so that the l 1

two-phase level was at the too of the heated length. The upper plenum, uoper head,' and the top of the downcomer were initialized with saturated steam. ~

l 0 '

i -

6 I .

i

. _ __. . _ _ __ .____,__ __-_.... __ _ ____-_.w_m._ -

_ . . - . _ _ . ,_ __I. . ... _ - _ _ _ _ _ . . _ _ _

3.2 Comoutational Model The details of the COBRA-NC code are provided in the code documentation

( - (Ref. 3). The reader should refer to this documentation for a comolete description of the equations, numerical solution methods, and emoirical correlations used in the code. Models that play significant roles in this particular simulation will be discussed here briefly. COBRA-NC uses a two-

• l fluid, three-field, two-component representation of two-phase flow. The two-  !

fluids are a vapor / gas mixture and liquid water. The three velocity fields are l the vapor / gas mixture, the continuous liquid, and the dispersed liquid )

fields. The dispersed liouid field is of no importance in this apolication.

The three temperature fields are the vapor / gas mixture, the combined liquid, and the solid structures. The two components are water and a noncondensable gas mixture that may be composed of any number of gas species. The concentration of each component of the gas mixture is comouted using separate mass transport equations. Only the concentrations of steam and hydrogen are of interest here.

The metal / water reaction of zirconium is modeled using the chemical reaction Zr + 2H2O + Zr 02 + 2H2 + heat ., i The Cathcart reaction rate equation .

l w $? . A,-B/T

, dt is assumed to be valid provided steam and zircaloy are available to sustain the reaction, wnere w is the total mass of oxygen consumed oer unit area of cladding oxidized. The coefficients for the reaction equation are A = 16.8 kg2/m#s for T < 1853.0 K 2

= 5.426 kg 7,4s for T 1 1853.0 K B = 20070 K for T < 1853.0 K

= 16610 K for T 1 1853.0 K The rate equation given above is converted into a rate equation for the reacting surface radius

-C(R,-r)h=Ae(-B/T) l 7

l g ..

This equation is integrated between the old and new time reacting surface radius to yield an equation for the reacting surface radius:

r"+1 = R, - t(R, -n e ) +ye -8/T at)

The heat released into the clad is given as q = 6.45 x 106 ," zr Ifn

• I +1) A*/at The hydrogen mass rate generated from the reactor is In  !

H 2 0'0442

• x ZI !(rn * ~ I n+1) ax/at The steam consumed by the reaction is *

.(

l I

b0=8.92*k 2

H 2

4 1 -

The metal / water reaction is assumed to continue until the partial pressure of steam is less than 1.0 osia or until the clad is completely oxidized. The heat transfer from the fuel rods to the fluid and from the fluid to structural j

surfaces is calculated by the ORNL (Ref. 4) natural convection heat transfer coef ficient for Reynolds numbers below 10,000. This model is as follows:

Nu = maximum of (

4.0 Pr"0.4 0.021 Re O Pr0.4 i

where GD Re mw " p N = modified wall Reynolds number 9w Pr, = vapor / gas mixture Prandtl number evaluated at the wall y temperature Gg = vapor / gas mass flux 8

h-

' "

• e+ , - - - . - , , - - - - , , , _ , - g - - - - n-

.-,,..,.---en. --- n-

t e

t pI" = vapor / gas mixture viscosity evaluated at the wall  !

temoerature

( -

Dh = hydraulic diameter.

l 1

1

}

The Dittus-80elter equation is used for Reynolds numbers greater than 10,000.

Radiation from the surface to the vapor / gas mixture is not included in this l I

calculation.

4.0 RESULTS I l

The calculation was initiated at the beginning of core uncovery and continued until the peak clad temperature exceeded the temperature limit for the equation

- of state for the fluid. Since COBRA-NC does not contain models for zirconium melting, the clad temperature will continue to rise above the melting point of i

, zirconium.

Since no emergency water is available to replace the water boiled away in the l core, the liquid level continuously droos, uncovering the core as it does so.

The liquid' level reaches the bottom of the core at about 6930 sec after the initiation of the transient, just 1200 sec after the beginning of core uncovery.

, j .

J ,

i The core liquid fraction as a function of time is shown in Figure 4 As the',"

l core uncovers, the bare rods receive poor cooling from the steam, and they heat i

up as a result of the decay-heat generated in the fuel rod. When the clad j temperature reaches a sufficiently high valu. the zircaloy cladding begins to react with the steam. The heat from the chem; cal reaction causes the cladding 1 -

heatuo rate to increase. The clad surface temperatures for three coints on i

several rods are shown in Figures 5 through 10. The rod numbers run from 1 at  ;

the core center to 8 at core periphery. The temperatures predicted on rods 1, 5, and 8 are compared with temperatures calculated with the MARCH code i (Ref. 1). As can be seen from these curves, the rod temperatures calculated by l COBRA-NC lag those calcula,ted by MARCH by about 10 min. This delay is caused by the additional cooling provided by convective flow between the upper plenum i and core. Since the MARCH calculation is one-dimensional, no convective l

currents could be calculated. In the two-dimensional COBRA-NC model, such j

convection currents are possible and they were calculated. This will be discussed in more detail later. .

i The sudden increase in the rate of rod temperature rise at about 7800 see is '

i caused by the increasing heat generated from the metal / water reaction as the i cladding temperature increases. Rods 2, 3, and 4 become the hottest. The temperature-versus-time profile is nearly vertical for these rods at l 7800 sec. The upper part of rod 2 heats up rapidly during this period, because i

a flow stagnation occurs in bundle ring number 2. The stagnation provides poor i

cooling to the rod, causing it to heat up very rapidly because of cladding

! oxidation. Rods 3 and 4 also heat up, since they are the higher-powered rods j and also because they receive the hottest steam / hydrogen mixture in the core.

l' i

! e. 9

+

1 l l f

,. e

. . , - , - . - - - - - _ , - - - . _ , . _ . . .g- %- ----,.-_----,c...,--y., ,.,-q--y...,,__._m,_ +,-_..-_.,,y.,.,m- _.-_. - , , . - ~ , - . . , , , - , _ , ,, ,w v2.,-...., .a**-+~%mw,!

i l

{ to .

. ...... .. ,- ,- . n- s- - .

: i <

0.8 5-i : .

0.6 :- i 0.4 i.-

4, 0.2 E  :

0.0 5400 5900 6400 6900 :7400 7900 8400 nut (stcoNos)

• J:

Figure 4 Core liquid fraction vs. time

• t j

Sooo , ' ' ' *

- l'- sortou or soo 2 - e t.temew tLtvanow 3 - rop or noo 4000 - . --- uancM c4Lcutanon e

, l

, ,e  !

. I e

C w 3000 -

e .

o h l

._ i e

< e er /

• g 2000 -

i 2 00 - .

t 0

+

0 t200 2400 3600 4soo sooo 7200 8400 9600 TlWC (SECON05)

Figure 5. Clad surface temoerature for rod number 1 Y; * ' 10

\

l

[ ._

P" **\ . - . , , . _ _ . . _ .

)

. ..g.v iyy avyy , .y i g

.. ......... .g.... . .- v v v ,, , ,,, ,

2 - er.it*4cw CLtvarion  :

~ [

3 - not spot ou noo -ise.re n j 4000

-  : I

: 1 000 F .i l

: l 000 i- e 1000 i-

- .s.n 0' " -

O ?200 2400 3600 4800 6000 7200 8400 9600.

ME (SECONOS) ' .

e e

Figure 6. Clad surface temperature for rod number 2 ,

5000 '

. ... . .. .... ... .. ......... . .. ........ , .... ... n; ,6 3 g g g .. . . . ... ..

[ 2 - er.iie.cw n.sv Arion i

s - not spot on noa - mes m.  : i 4000 !- -

. .i

[:

. i.  :

l-k000  !- .  !

$ i i

i 2000 E- -i 1000 F -

b .t. . t. .t. .t. .t. .1. .t.

0 1200 2400 3600 4800 6000 7200, 8400 9600 TIME (SECONOS)

Figure 7. Clad surface temperature for rod number 3 ,

h; . . 11 l

I w

1 5000 ,,. .

1,' ,

.g .,.ggg .g . ,,. .. .

i - er. men ruvanou  :

(-  : s - not spot ou noo - ix : =ca !

i l

l y >%

4000 .

--"^"* *""*" j -

000 E-

-i 2000 F 1000 F - * -i 0 ' '

O 1200 2400 3600 4800 6000 7200 '

8400 9600 TIME (sEcorros) .

J.

. g Figure 8. Clad surface temperature for rod number 4 .

( 1000 e-eittou W 800 2 = #7taasCM (L(Wafl0es

• 3 = T08 Of 800

. 4000 - *--

"'*C" c,ucuta n0= -

e I

I t::. 3000 -

i -

W y e e

5 /

y 2000 -

j -

• /

o s000 =

S_'n 1

o , , , , . ,

l t 0 200 2400 3600 4e00 6000 7200 e400 9600 f

. T1WC (5CCON05)

Figure 9. Clad surface temperature for rod number 5 (t. .

12 e *

, o o

3 .

• ' '~ ' '

p'. 00ff0M o/ A00 2 = 47 inNCH (LCv4n04

\ ,

3=f070fR00 gggg ,

~~= Wa#CM CECULAMON .

I l b 3000 f -

. g  !  : .

a 2000 i -

\

= -

1000 -

.<3wy , e i

0 '- '- "~

0 I200 2400 3600 4000 6000 7200 8400 9600 Twt (stCoNos) -

Figure 10. Clad surface temoerature for rod number 8 .'

i

[' The steam / hydrogen mixture leaving the core transhrs beat to the structures in

\, the upper plenum and upper head. The surface temperature for seve.ral structures within the upper plenum are shown in Figures 11 througn 15.

• The structural temperatures above the hot bundles are higner tnan those above peripheral bundles. This is expected since the flow exits the core above the hot bundles transferring heat to the structure immediately above first. As the i

flow circulates through the upper plenum, it cools. It reaches its lowest i

temperature before re-entering the core at the outer bundles (Figure 16).

Since the vapor is coolest at the outside perimeter of the upper plenum, it i

transfers less heat to the structures in that region.

The relationship between the temperatures of the various components of the system is illustrated in Figure 17. The peak clad temperature, steam / hydrogen exit temperature, support column temperature, and deep beam temperature for radial ring number 4 are shown. The guide tube temperature for radial ring 3 is also shown. The steam exit temperature is only about 100*F less than the peak clad temperature. Since radiant heat transfer between the rods and steam is not included in the calculation, it must be concluded that the convective heat transfer, combined with the large surface-area-to-fluid-volume ratio of the fuel bundle, is adequate to heat the steam / hydrogen mixture up to near the clad surface temocrature. The guide tube and support column temperatures at the hot leg elevation are from 100*F to 500*F lower than the core exit steam temperature. The margin increases with time. The larger temoerature ,

4 13

]

S000 , ,

- C0 sea /mc /8COICnO4 4000 -

• t l

C' w 3000 -

1 2000 -

W t000 -

_ ,t 0 200 2400 ss00 400 6000 7200 a400 se00 TNC (SCcoNDs) *

, 3,j Figure 11.

Supportcolumnsurfacetemperatureaboverod4athotlegelevatfon 5000 , , , '

- Cossa/mc PeC0icnog '

l 4000 -

C

• w 3000 -

t.a ,

E

~

E -

2000 -  !

- I

=

1000 -

0 ' '- '- ' ' ' '

O 1200 2400 3s00 400 s000 7200 8400 9600 3

Twc (stcoNoS)

Figure 12. Upper core plate temperature above rod 5 1

14

0. . .

l l

.g .  !

3

(.

3000 ,, ,,,,,, , ,,,,,, - ... m - ~ ~i m " " i g g gcei; w " " i " " ;

i OUTSDc SURFACE  :

4000 F i .

- i l 000 F 2000 F _

- i 1000 F -  :

a p -  :

I .

E E

• 3

. , ... ... .i. .i. .i. '- -

0 e,

0 1200 2400 3600 4800 6000 7200 8400 9600 ,

TIME: (SECONDS) .-

i i '

i Figure 13. Guide tube temperature above rod 8 at hot leg elevation 5000 .

outsor svaract  :

4000 F i  !

!  :  : 1 l

' c 9000 3-l g  :  : l 2 i 2000 F 2 i - .

i .

1000 F  :

e

.i

^^ '- '- '- '- '- '-

i O O 1200 2400 3600 4800 6000 7200 8400 9600

' TIME (SECONOS)

Figure 14 Support column temperature above bundle 8 at hot leg elevation o

4 (A. ..> 15 i

4

^

.t. . . . - - :n J. - a

O O

f, 5000 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ; ;4;g .ggg . . . . . . . . .

[ OUT$eDE SueFACC 3

4000 E i k000 5- 5 i

E 2000 i- 7 1000 t

i .

~

O

• 0 1200 2400 3600 4800 6000 7200 84,00 9600 .i Ttut (SECONOS)

Figure 15. Core barrel temperature at hot leg elevation

('\ ,

j 5000 ....................................... .g .........;

I

2- TOP OF SUNDE 4 .

l 4000 E- .-

e  :  :

9000 E- 2

. 2000 :- i j 1000 E- .:

gege *  :

O ,

5400 5900 6400 6900 7400 7900 8400 I TIME (SECONOS) )

Figure 16. Steam temoerature at hot leg inlet l

.. i

b. -

16 l

\

l l

l l

._-. ++,i- ,, - li

a , .

, /

o

(

5000

.. . ....... .. ..... ... .... .. . ,... .. ...,._. w , g g .... .

h 2 -VAPOR AT TOP 0F SUNDE 4 "

3 - SUPPORT COLUM SUNO2 4  :

• 4000 {: , * - " " ' " " " ' 3

,, , 5 - DEEP BEAW SUNDG 4  :

i i k000 5-

$ i i i i i

2000 i- -i

. 3 1000 gesme-r _ m * , , u.

' ~

.. 1 0 '

' Y. 1 5400 ' 5900 6400 6900 7400 7900 8400 .

nME (SECONDS) *

,, Figure 17 Comparison of temoeratures for several components 1

i difference for these structures can be attributed to the lower surface-area-to-fluid-volume ratio of the uocer plenum and also to some precooling of the steam i due to mixing caused by convection currents in the upper clenum and heat i 1

transfer to structures in the bottom of the upper plenum. '

l l

Radiant heat transfer could become a significant factor at the higher temperatures. Structures near the top cf the upper plenum are cooler, probably because of lower fluid velocities that provide poorer heat transfer. Structure l

temperatures in the upper head remain extremely low because of the very weak l convection currents there. Radiant heat transfer may be the dominant mode of heat transfer in the upper head. The fluid temperature at the hot leg nozzle, Figure 16, is lower than the fluid temperature exiting the core because of the heat loss to upper plenum structures.

Hydrogen is generated as a result of the metal / water reaction in the core. The oxidation rate increases as the clad temperature increases. The increasing oxidation rate is tilustrated by plots of hydrogen concentration as a function of time for radial ring number 4 (Figures 18 and 19). The oxidation reaction

)

will continue unt'il eitner all of the steam is consumed or all of the zircaloy '

clad is reacted. The convection currents between the core and the upper plenum l l

y'; . . 17

\',

I

^

9

_h # e-

• a o

40 , , , , , , , ,

50 -

I q

4 40 -

E  !

I l8 30 -

• 20 .

I K. N -

, __- - - - - = s ' , ,

5500 $750 6000 6250 6500 6750 7000 7250 7500 7750 DML SCCONoS . j gure 18.

• Hydrogen concentration in bottom 5 cells of radial ring 4 10 s s s .

40 -

]

a 50 -

N 5 40 - -

d 30 -

5 -

@ 20 -

E E

10 -

0 """ " " " " " " " " "'#'t" " " " " '

' )

, 5300 5750 6000 6250 6500 4750 7000 7250 7300 7750 l

, )

TlWE.SCCONoS Figure 19. Hydrogen concentration in top 5 cells of radial ring 4 ,

h.. . 18 t

l l

. , _ _ _ _ - . ._. . . + . . . ~ . - -

, 'J replenish the steam in the core as it is consumed by the chemical reaction.

Thus, while the convection currents provide more effective cooling during the

(-

' early stages or core heatup, they are detrimental later in the transient since they supply the core with additional steam from the upper head and upper plenum to sustain the metal / water reaction. The hottest ooints on the rod have already reached the melting temperatures of the cladding, and the peak hydrogen concentration is less than 70%. The deformation, ballooning, and rupture of the -fuel rods are all factors that affect the convection flows and the metal / water reaction, and these factors are not included in this calculation.

The convection of steam into the core from the upper plenum is significant,

~

however, and it may completely overwhelm all other heat transfer benefits in determining whether the fuel will melt before failure of the primary loop components occurs.

4 The overall thermal-hydraulic behavior of the fluid within the reactor vessel is illustrated by a series of vapor velocity vector plots, gas temperature contour plots, and hydrogen concentration con.our plots. The vector plots are shown in Figures 20 through 30. The transient, time given for these plots is the time since the beginning of too-of-the-core uncovery. Two principal convection currents persist through out the transient. The first consists of:.

the flow going down the periphery of the ccre and up in the third, fourth,* anti fifth bundles. This current connects to one in the upper plenum where flow is up in the third, fourth, and fifth radial rings and down in the three outer '

ri ngs. Some of this flow recirculates above the core plate mixing cooler uoo'er plenum steam with hot steam coming from the core. A second convection current carries flow down the center two bundles and up the third and fourth bundles.

This convection current is caused by two factors. First, the radial power (n '

profiles for the first and second bundles is lower than that of the fourth and fifth bundles, so less heat is transferred to the steam in the center two bundles. Second, the amount of structure in the upper plenum is larger for the first and second radial rings, so more heat is extracted from the steam in this region. Each bundle in this region has either a guide tube or a sucoort column above it. Radial rings three, four, and five only have structure above every other or every two out of three bundles (Figure 2). Therefore, the fluid in the center two rings is cooler (i.e., more dense) causing down flow in the center two bundles. Later in the transient, increased core temperatures and hydrogen generation cause the flow to reverse in bundle 2 resulting in a flow stagnation near the top of bundle 2. This results in a very rapid increase in the clad temperature of rod 2 (Figure 28). A third convection current flows through the guide tubes into the upper head in the tnird, fourth, and fifth radial rings and down through the center and peripheral guide tubes. Weak ,

convection currents also exist in the upper head.

Contour plots of vapor / gas temper'tures a are given in Figures 31 through 39.

The contour lines are plotted at 240*F intervals. Temoerature in the upper plenum are fairly uniform (witnin the 240*F range). A temperature stratification occurs in the upper head caused by hot steam exiting the too of the guides tubes which extend part way up into the upper head. Cooler, denser steam remains below the top of the guide tubes. The steam near the bottom of ee y- ,-vr -- -, ---mn,v-s- - - g -e , e-- , - v-.. --

,-y, --. _r-g e,,,- - -,..,~--,4 --,,-n,- , vY--~m--- w.,.,--c.-rw, w w w-og,.p-,-

( . . . w-

. .i

........I

........ c g;. 1,'* . ,

go( ,,..*

j *,

• g l 1"~%.

/.

f. i e.gg

') i.._ gs

,. 1 3. ,.' .. , ,

t

.s; y e \ s , . J.

, l 4*pi3'1

\'

' ,4

'I t i ', J ' 4 I I e

4

.* .* ., ., .' ., ,i l

... ..... -, 4, t t te i s , . ;;5 4

,....=.., g

. .....I.

tefel t ........

ffettIi! '

t.t. tt.. .

1. ..}III..ft

/  !

itittie*

l =

}

- 1 Figure 20. Vapor / gas velocities at Figure 21. Vapor / gas velocities at .

401 seconds 813 seconds -

s. .

. .... ..... i

,,.....l

,, t ..,

I \ '

I t . ,' i l *t \

t .' ' i in,i.,!

i.,1 1.t.'l t

s n . ,' !

i .

d

'jie*i' h

i

, '\ t l 't '* 4'I 4

• ttt ' 6n  !

i 6

,g,*is 6 *ff ,f i, e n 6:

4 3

899 .,g g.ftt .' g%gg Figure 22. Vapor / gas velocities at Figure 23. Vapor / gas velocities at 1203 seconds 1607 seconds G.. . . 20 Q=

- 1 a . .

1 I

. i

( ~

• • • -~- -

i * '. '. *. * *. *. *.

e- *-~. ........

lA/,'lI IlI/*i j/*'f/'L I ',

6 y,

.f., '

f

'*o f.'I1 I ,.  :

\ <'*<h { ,1,1.f qt(s ,

, ' t ' t 43

\

l' p 'g oas ttt ' *\ }

8

• ' fs't f i *' I i i s.Iit.*5g i8 .

s ,gg i

l. <. 8.f1i*tI s'.; '

l l .,., .. ,,.3I

,j

-o

.'.....j j ,

. . .t . ..., 1 s.

Figure 24 Vapor / gas velocities at Figure 25. Vaoor/ gas velocities at 1668 seconds .k 1768 seconds *

.......m oa ..**.. ,, .....

J i < ' ', i l j f , ' '. . . .

$ 'e l. \ $*. / '\

< f 1

f,,

t

, t, I , ,' ,

. l t ,

i t\' ,,a <

e ) ~ . .' ;

,\

ft*33 #

II ,k / f t '6 I68i .

8 .'Ifff I

, /

3 it*ii 8.tti'IL 5L ,

I e ,t t i * \ i e

!.et$,....lI

!..ii...ll Figure 26. Vapor / gas velocities at Figure 27. Vapor / gas velocities at 1809 seconds 1869 seconds

^a

\l'.s .

A l

[ '

. . .... i .......l

........ . , ......i It'st. ,,..e..,

f '} z / ,If'*. I

• f f ," ,' \!

s.

g r,i8 s.,

e.*I ..

.l, '.a o s,4*./ ) ,t

• i 4 i

} ,. Ife ..

is d 8i

! , ,f g,,.64 I'ttt,

ni

~ ' '

[.'tt' ... ..;j [ .,ll'.,,

. . . ...,,l ......,,

.... .... ........ . 7 s.

Figure 28. Vapor / gas velocities at Figure 29. Vapor / gas velocities at

• 1909 seconds 1949 seconds

• F.. .. .....

r t im ,

, IMO v g i1.,. . . . - - .. ,. ,,,, ,

l, { < 't . . s, teoo r

,g{ ....

.e N

• I)t.

, I t t . .' .' .'

. '.. t...,

. 9 ...., .

isoo r A

........ i* r\_

V ixa r\ x

• Figure 30. Vapor / gas velocities at Figure 31. Gas temperature at 1990 seconds at 1600 seconds 1

4 e

4

. . . _ , , . , . . - - . - . - - . , ,, , , - - , - . -.- _ , , . , _ . _ . , _ , . , , . , , , , _ , .. _ _ , , , _ , - - - - - --....,.-r--

. I

-e .

v3 i -

i

( '

l l

1 I

l T>t360 F , 1360 r

? >t360 r 1

^#  ; 1560 r% _ g a320 r ,

e I

1000 F==

,"'" T >1000 F , 4000 r * '

4 Y >t900 r '

i

. I is00 r-

! l i 1000 rg 3,,, , .

+

T >1000 r l 't 2060 t j

1s00 r g 1j60 r gea, , j 1960 r 13:0rg ~ ,  !

w -

, y

. l 4

j, Figure 32. Gas temperature at Figure 33. Gas temperature at '-

1668 seconds .

! 1708 seconds 4

f T >t$60 r T >tJ60 r 7 1540 r '

t 540 r ,,,,

t320 r-

• I  !

i

• j l

T >t000 r l 1,

20h0 r a

i (40 r ,,

asso r 4 l 20n0 r-20m0 r-

teco r s 1e00 r i 1360 r s i360 r-Figure 34 Gas temperature at Figure 35. Gas temperature at

1768 seconds 1809 seconds .

23 1

- - - .e.,s -n,-,n,n , - - . - - - - - , . ----,---.n-,-,--n vn-,-- ,- . - - , . - , _ ---n-. - , - - . - - - - - . - , ~ - - - . , - ,v- - - - -- - ~ ,-..,---~-n-

e .

I seco r - \ l rxsoo r I '

tsoo r -

136o r - 1 g g ,_ -

I 2oso r- 22eo r- V 20"o !

\ \ .

22so r- _ 228o r-l 2k6o F N

2040 'N -

20so r c~ r~ ..~ r2

~

.y Figure 36. Gas temoerature at Figure 37 Gas temperature at

• 1869 seconds 1900 seconds 1

TM8co r 2040 r *

-r- ,,00 ,_

I#

E 156o r-w l+

2760 r- . p 2760 r 2200 r " l

Wrr

) I. '

2200 r- Mo r. .

276o r " V i

233o r- 2100 Y'  ;

25to rs 228o r y

., 2ano rs J 228C \

20"o (-- -

teco r - ~

Figure 38. Gas temocrature at Figure 39. Gas temocrature at 1949 seconds 1990 seconds

.g 24 ,L 6

v -- - . - . - - - - - - - - - - - - - , .

f steam near the bottom of the core as well as the fuel rods are actually heated by steam convected down from the upper plenum. The center two bundles are cooler due untilflow to the laterstagnation.

in the transfer, where the second bundle rapidly heats up Hyroden concentration contours are given in Figures 40 through 48. The 1

hydrogen concentration in the upper plenum is fairly uniform throughout the transient.

i plenum. The uniformity is caused by the convection currents in the upper The upper head has a stratified distribution of hydrogen because of the elevation of the too of the guide tubes. The core also has a large t

gradient in hydrogen concentration, since most of the hydrogen is being generated in the upper elevations of the bundle. Hydrogen in the lower elevations is transported there by the convection currents. It appears that there is still plenty of steam available to sustain the metal / water reaction unti.)

(Figurethe48).

majority of the core has reached the melting temperature The oxidation reaction will stop when the steam is depleted; whether this will occur before most or all of the core has melted cannot be determined from thij simulation. The code would have to be modified to account for the heat of ' .

fusion of the zircaloy, so that the clad temperatures cannot exceed the melting temperatures before the calculation can be run out further. '

5.0 CONCLUSION

S AND RECOMMENDATIONS k The results of a COBRA-NC simulation of convection currents between the upper

\ plenum and core during the TMLB' accident sequence have been presented. The results of the simulation hav'e shown that:

o Convection loops between the core, upper plenum, and uoper head do develop.

e The convection causes the clad temperature to lag those predicted using one-dimensional models by about 10 min.

e The convection loops bring steam down into the core from the upper plenum and upper head. This steam maintains the metal / water reaction in the core.

• 1 l

e The heat generated from the metal / water reaction dominates the temperature rise in the clad.

e The cladding reaches melting temoeratures while uoper plenum structure is i

relatively cool (2000'F).

e The vessel-average steam volume fraction is about 50% at the end of the calculation, so the metal / water reaction will be maintained.

l 25

• e

e .

\

e .

(

C*S ,

ces I

C.n 49 %

~. . 5 ,

C)$8 l

r .

-s  ;

l i

3 .>

t Figure 40. Hydrogen concentration at Figure 41. Hydrogen concentration at.'.

1607 seconds 1668 seconds *

( ,

m** L x-  :: -

C 95 CCA 5%

Ct%

Cat 2n i

C>ss im ~

"~

m i

5- Q Figure 42. Hydrogen concentration at Figure 43. Hydrogen concentration at 1700 seconds 1768 seconds i

h.--

t

(.* '

~ ~

em  !

s- tm 5 -

cats 15 ~

- so~ 7 y 15 (

f 205 -

g

• $_ T  !DD '

, j s

Figure 44 Hydrogen 1809 seconds concentration at Figure 45. Hydrogenconcentratinnatl-1869 seconds -

c.nn I i M-in-nous 35 ,

is-

, c.en I

N- 25 ss

{ 39 ~

f-1 35-an-am- '

M8N /

a$3 Q3 Figure 46. Hydrogen concentration at Figure 47. Hydrogen concentration at 1909 seconds 1940 seconds ,

27 e

e O

I

. N g !E

(- -: -

o N

n' ..

= .

I M~ 1

)

65 -

J T- J

~

45% N 1

• w-35 7 AT

"/

as

, Figure 48. Hydrogen concentration at 1990 seconds .;

The calculation:

the above statements must be considered in view of the several constraints on-e A two-dimensional noding was assumed.

e No fuel clad deformation or ballooning has been considered in the calculation.

e Radiation heat transfer between the structure and the steam or between structures has not been included in the calculation.

e The heat of fusion for melting zircaloy is not in the heat transfer model; thus the clad temperature will exceed the clad melting temperature. This will cause the fluid temperature to ultimately exceed the limits on the code's equation of state causing the code to fail. For this reason, the calculation was not run out further in time.

. The loops were not included in the model. The loops may contribute to core cooling due to convection between the looos and the core. They will also sucoly more steam to sustain the metal / water reaction.

• A relatively coarse mesh has been used. This may have resulted in some underprediction of the convective heat transfer in the deep beam regions of the upper plenum.

e k; - 28

.7.._,,,_....._.,.-.,.m..,-.,..,_..,,_,_#..~. . . _ . - . . . _ . , _ . , , . ~ _ . _ . _ _ . - - ~ , , . . _ _ - , , . _ - . . . . . _ _ . , , _ . _ , . . . . -

• ~ ~

t i .

With these limitations in mind it is reconsnended that:

e An improved radiative heat transfer model be isolemented into C08RA-NC.

! [\

e A clad melt model that will account for the heat of fusion of zircaloy be

isolemented to allow the calculation to be carried out until all of the .

steam is consumed or until all of the clad has reacted or melted. This i will allow the temperature of the uocer alenum structures to be calculated j as a function of fraction of core melted. Since only a small fraction of the core is now at the melting point, significant structural heating may occur before much of the core slumos to the lower plenum.

e i

$1mplified clad deformation models should be added to allow the code to estimate the degree of flow blockage occurring.

1

! e i *

• The loop components can now be and should be included in future calculations.

' i 4

e i

The code should be assessed against the Westinghous's natural convection data, i

• 5 To summarize, from the current results it appears that if the upper structur(

is to fail first, it will be a tight horse race between the metal / water -

reaction system.

in the core and heat transfer to the uoper components in the reactor

{ which horse Unfortunately, wins. additional work probably is required before we can say i

i i

)

1 i

i i

I i

29 1

i

(<. . . .

i I

._v -,,_,,.,_..w, -

,m__, -,,,,,,.y.._. .,_m.._.,_-r_r. ,,_,,.,gm.g_, ._,W . . , _ . ..,,..,.-.,_._..._,._,,,,,,,,_r-.

\ .

i o

REFERENCES t 1.

J. A. Gieseke, et al., "Radionuclide Release Under Specific LWR Accident 0' . Conditions", BMI-2104, Volume 5, Battelle Columbus Laboratories, July 1984.

2. .

Rosanna Chambers, " Analysis of Core Behavior During a Station Blackout Transient (TMLB')

NUREG/CR-3979, Augustfor the Bellefonte Pressurized Water Reactor".

1984

3. et al. " COBRA-TF:

M. J. Thuracod, Analysis of Nuc lear Rea,ctor Components"A Thermal-Hydraulic Code for Tr

, NUREG/CR-3262, March 1985.

be published). (To 4

T. M. Anklam, et al., " Experimental Investigations of Uncovered-Bondle

- ' Heat Transfer and Two-Phase Mixture-Level Swell Under High-Pressure Low

' Heat-Flux Conditions", NUREG/CR-2456, March 1982.

. q 9

l t

e 30 k': "'

1.

r a s. m aa y- ,

~

Structural Failure l Studies of RCS

/

i i

W e /DA HO F NAT/ONAL l- EN6/NEER/No Presented by:

l LABORATORY , V. N. Shah i

i t

i 1

ldeh0, bMll.

I M A006,42 l

l

- ^ ~ ~ ~ - - -- -

p; .

I Outline l

• .0bjective 1

• Hot leg nozzle / piping materials and design i

l.

• Master creep rupture curves Estimate of f allure time during

\

l TMLB' sequence

• Conclusion I

I'

Objective Determine failure modes and timings for the press'ure boundary components

~

during TMLB' sequence.

ll

• Alters accident sequence a

1 '. ,

~

^

i

• Changes severity of accident .

i-

• Affects containment loadings .

l E A0064,4

~

i

• f e

j- . .

Failure Modes During i

L i

! Severe Accident

? -

Failure by melting if primary l

4 pressure is low -

i .

Failure by creep rupture if primary pressure is high l

i l E A006,45

- ^- --

- ^ - -

.., I Schematic of Hot Leg Nozzle 1

~

Fieldweld Nozzle Forging Reactor Coolant Piping (316 S. S.)

(A-508, Class 2)

M(

l i

N w upg;; .

a

- .[ n

\ S. S. Clad 29 in. )

(1/ s" Thick) l i

' Outlet 1r

't

j m . . .

.  :-u 2.5 in.

l

~ f .

T M Safe-end

~

\ .

9eactor Vessel Wall .'

(A-508, Class 2 and A-533, Grade B, Class 1)

L IEA W T

~

-~ ' '

~~ -

Creep Rupture

]

l h (61 < U2 '< 63 e 64) i I

creep G3 rupture

.E G4 (t3 Accelerating D

l' cn creep (Stage III) l U2 a

l O O

G t O l a1_

Steady-state creep (Stage II) i i Transient creep (Stage I)

I Time h IEA00639 9

k

[

Mastor Rupturo Curvo f@r l

A-533, Gr. B, CL1 Steel i

i i -

10 0-- - T-

~

o ~

x ~

o_

O Q ,

N1 o -

G OA e ^

l CD oee -

d ~

A l ll ^e#

0 - 728 K (850 *F)

'55 x \(e A - 783 K (950 *F) l -

- e - 811 K (1000 *F)

, c + - 867 K (1100 F) _

j X - Min TS*

O - Max TS j.

, e m

G) . b-Jr l

u)

s y' ,

l 10 .

25 30 y) g "35 40 15 20 _

Larson-Miller parameter T(20+1ogt ? r x 10-3__,

l

~

k F. R. Larson and J. Miller. A Time-Temperature Relationship for Rupture an .

i

Reference:

Transaction of the ASME July 1952, pp. 765-775.G. V. Smith, " Evaluatio C-Mo. Mn-Mo and Mn-Mo-Ni Steels". ASTM Publication. DS 47, 1971.

l l -- - _ _ _ _ _ - _ - __ _ . _ _ _

! Master Rupture Curve for 316 Stainless Steel i

i r i e i 10 0 _ i

- ~ _

m _

-- .. -9#

" " ~ -

D409 AJ

[

.. -- ASME Code Case N-47

} _

e Larson-Miller dato

$, for 18Cr-8Ni S.S. '

f -

I ll ~. ' .8 N

u) g- 'g' 7

• ~

O - 700 K (800 *F) .

' X - 755 K (900 "F) '.3 0 - 811 K (1000 F)

~

' .^- - ~

j O - 922 K (1200 F) l

' " A - 1033 K (1400 F)

~

I O V - 1089 K (1500 "F) -

4 m

0) -

! u) ' ' ' '

l 1

35 40 45 50 15 20 25 30 Larson-Miller parameter [T(20+1ogtr) x 10-3[ 16

Reference:

F. R. Larson and J. Miller. A Time-Temperature Relationship f or Rupture and Creep Str .

Transaction of the ASME. July 1952 pp. 765-775.

l

Rupture Times vs. Temperatures f@r

!, Hot Leg Nozzle (ID = 29", t = 2.5")

r r - i , i, i i ' , , r,- r cr i ' , , ,,,

l 110 0 ,

1 i, e l M .

4 -

.~ ... ,- N o

s_

1

]

g NN . .

l

! 8 1000- x- N~N s ..~~~~ . .. s

~

l (D N-x ...,' N Q, N N N ..'

'N l g 'N  %~% ... . ~

j y N, N. ....,'

j ___ N, '-  % ... ,

l

~

to System Pressure N~

x, ~ ~ . _

3 -

N r 900- 4 MPa t

O g .............8MPa N-  %

i (0 -

12 MPs t_

(D -

16 MPa

> ~

i <C

, ,..i i i ' ' ' ' '

i , , ,

i

' , , , . .ii J 800 ~ 10 10 0 1

0.1 i Rupture time (hours? vmsuma-i:'

, s 4 s

l

~

Preliminary Failure Studies are .

l Assumptions Following l Based on the i

1 ID = 29", t = 2.5" l

• Nozzle dimensions:

f considered.

Average wall temperatures are 1 .;

loss of ductility in the

• Residual stresses.and not accounted for.

heat-affected zone are J

due to aging are not j

• Material degradations accounted for.

l(

. . i, x

- n n <

f f;

'\

-f '

R i & '~

l ,

z '

-. g -

j

~

':?

~S 3 =

g -

gg o '

hl - ,

c,g

-l

., m

, -g l

p l- @ o;np;adu;ej, pg

(,. .

~

Creep Rupture Failure .

of Surry Components-Preliminary Results l

for pressure and temperatures

• MELPROG results ~

during TMLB' sequence are considered. Pressure l

l 16.2 MPa {2350 psi? and remains constant.

l is equal to l

10,710 s. The

• Hot leg nozzle will rupture at of rupture

[- average wall temperature at the time i i

! is.1015 K.

• Hot leg pipe will rupture at 11,620 s. The of average wall temperature at the time l

E rupture is 1135 K. DEA 40642

?:

i

~

1

.l '

Conclusion L.

i l Creep rupture is a most liekly structural fallure mode during

! TMLB' sequence.

e

l 4

m aoes,rs j

i N

A