Text

Draft Revision 1 to Analysis of Nominal Heat Removal Capacity of Crbr in Natural Circulation Mode
	ML20065A370
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Site:	Clinch River
Issue date:	07/16/1982
From:	; FAUSKE & ASSOCIATES, INC.
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- _ -

l FAI/82-17 REVISION 1 - 7/16/82 DRAFT l

Fauske & Associates, Inc.

l 16WO70 West 83rd Street Burr Ridge, Illinois 60521 (312) 323-8750 lI ANALYSIS OF NOMINAL llEAT REMOVAL CAPACITY OF THE CRBRP IN Tile NATURAL CIRCULATION MODE I

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. 1982 8209130211 820813 PDR ADOCK 05000537 A

PDR

_ DRAFT TABLE OF CONTENTS 1

1.0 INTRODUCTION

2.0 SODIUM RE-ENTRY AND DRY-0UT CRITERIA UNDER 3

DECAY HEAT POWER CONDITIONS 2.1 Simple Re-Entry Criterion for Sub-4 Assembly Geometry..

3.0 NATURAL CIRCULATION HEAT REMOVAL ANALYSIS 7

4.0 NATURAL DRAFT CAPACITY FOR Tile PROTECTED 19 AIR COOLED CONDENSERS (PACCs) 19 4.1 General Configuration......

s 26 4.2 Pressure Difference For Natural Draft

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26 4.3 System Flow Capacity 4.4 Air Side lleat Transfer Coefficient 26 4.5 overall Heat Transfer Coefficient 30 4.6 Energy Removal from the PACCc.

32 4.7 Heat Removal Rates by Ultimate

{'

Heat Sinks 34 41 5.0

SUMMARY

43

[

6.0 REFERENCES

l I

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l Em I eikruii iBg'i

- DRAFT LIST OF TABLES r

L

~

iable No.

Page, 2.1 Comparison of Ileat Fluxes from Empirical Correlation [Eq. (2.4] and the Simple Re-Entry Criterion [Ineq. (2.2)]......

6 4.1 PACC Air Flow Rate Vs Pressure Drop

.............28 I

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- iii -

DRAFT l

LIST OF FIGURES l

l Figure No.

Pg 3.1 Ilydraulic Ilead for Natural Circulation.

8 3.2 Ilydraulic Pressure Drop Equation for the CRBRP Primary System.

. 10 3.3 Comparison of Best Estimate and DEMO Re-sults at 1000s.

. 11 3.4 Hot and Cold Leg Temperature-Time IIistories for One Loop Natural Circulation.

. 12 3.5 liot and Cold Leg Temperature-Time IIistories for Two Loop Natural Circulation.

. 13 3.6 Ilot and Cold Leg Temperature-Time Histories for Three Loop Natural Circulation.

. 14 3.7 Decay Power and IIcat Removal Rates vs. Time for One Loop Natural Circulation.

. 15 3.8 Decay Power and lleat Removal Rates vs. Time for Two Loop Natural Circulation.

. 16 I

3.9 Decay Power and IIcat Removal Rates vs. Tice for Three Loop Natural Circulation

. 17 4.1 Protected Air Cooled Condenser.

. 20 4.2 Protected Air Cooled Condenser.

. z1 4.3 Fin Tube Coils.

. 22 4.4 PACC Closed Loop Schematic (Shown During Normal Plant Operation - PACC liot Standby)

. 23 4.5 The Piping and Instrumentation Diagram for the PACC Portion of the Steam Generator I

Auxiliary lleat Removal System

. 24 4.6 Dimensions and Calculated Surface Area of the Finned Tube.

. 25 4.7 Schecatic of PACCs for Natural Draft Operation

. 27 4.8 PACCs Flow Rate vs. Pressure Drop

. 29 lk_

- iv -

DRAIT I

LIST OF FIGURES (Continued)

Figure No.

Page 4.9 Average Friction and IIcat Transfer Data for Flow Over Five Different Arrangements of 3/8 in. Diameter Tube Bundles in the Laminar and Transition Regime

. 31 4.10 Steam and Air Temperature Variation in the PACCs

. 33 4.11 Heat Removal Rate by Venting vs. Time for One Loop Natural Circulation.

. 35 4.12 lleat Removal Rate by Venting vs. Time for Two Loop Natural Circulation.

. 36 4.13 Heat Removal Rate by Venting vs. Time fnr Three Loop Natural Circulation.

. 37 4.14 Cumulative Mass of Steam Vented During a One Loop Natural Circulation Event

. 38 4.15 Cumulative Mass of Steam Vented During a Two Loop Natural Circulation Event

. 39 4.16 Cumulative Mass of Steam Vented During a Three Loop Natural Circulation Event

. 40 I

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_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - - D_RA_FT

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1.0 INTRODUCTION

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As part of the probabilistic risk assensment (PRA) of the CRBRP, a variety of loss of heat sink (LOHS) accidents are being examined as events which might lead to core damage. Current evaluations of whether an event vill lead to core damage have been based on design criteria which allow for little

(

structural damage to a component and also include the normal design conserva-tiEms to account for uncertainties in material properties, operating condi-

[

tions and loads, etc.

Because of these conservatisms many branches of the LOHS event trees have been designated as leading to core damage, although some heat sinks of limited capacity are available during the accident sequence.

[

Two examples are a one-loop natural circulation event with a forced draft PACC and SGAHRS vent as heat sinks and a one-loop natural circulation event with a

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natural draft PACC and SGAHRS vent as heat sinks.

The PRAs are based on best estimates and permit in some instances more extensive damage before failbre than do design analyses.

Three areas of conservatism which may have a sig-(

nificant, unfavorable impact on the FRA have been identified and are discussed in this report.

The replacement of the conservative evaluations with "best estimate" ones may result in climinating a significant fraction (N 507.) of the

(

core damage states currently being predicted.

The three areas of conservatism discussed in the following three sections of the report have a direct effect on the natural circulation heat

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removal capacity of the CRERP systems.

The areas deal with (1) localized boiling in the core under decay power with flow coastdown, (2) natural cir-(

culation heat removal under " quasi-steady state" conditions and (3) PACC heat removal capacity under natural air draft conditions. Accepting boiling in the core on a "best estimate" basis results in a large extension of the time

DRAFT l I

before core damage would occur and permits one to ignore the short term (< 1/2 hour) ef fects in the core during a natural circulation event.

Then in the longer term (> 1/2 hour) the heat removal from the primary system by natural circulation can be estimated by assuming that quasi-steady conditions exist through the heat transport systems.

Finally a heat sink credit is taken for the heat removal capacity of the PACCs under natural draft conditions by using "best estimate" analysis.

Combining the results from three areas gives the best estimate of the CRBRP response to the loss of all AC power, i.e., natural circulation with turbine-driven feedwater and a natural draf t PACC unit. The results of the analyses in the three areas of conservatism are presented in the next three sections; in the last section a summary of the current work is presented and further work is discussed which will extend the best estimates of the CRBRP heat removal capacity during natural circulation events.

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I

- UnAFT 2.0 SODIUM RE-ENTRY AND DRY-0UT CRITERIA UNDER DECAY HEAT POWER CONDITIONS In the natural circulation analysis for CRBRP one criterion used for design purposes is that the sodium temperature must remain below boiling

point, i.e., 938'C.

For a best estimate analysis local boiling in the core is permitted because it does not prevent, in itself, the reactor to be brought

{

eventually to a cooled-down condition.

This section provides the technical basis for permitting local boiling and establishes sodium re-entry and dry-out criteria.

Based upon previous studies [1] of the accident progression for a hypothetical loss-of-heat-sink with scram in an LMFBR, incipient sodium

{

boiling is expected to occur as the primary system sodium temperature ap-proaches the saturation temperature.

At this time (3 to 4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months into the accident) the decay heat has fallen to approximately 1% of full power.*

At this power level the critical assumption is made that sustained dryout is

[

taken to coincide with incipient boiling [1,2]

i.e.,

no sodium re-entry is

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allowed after boiling initiation with the result that fuel meltdown is pre-dicted to occur with sodium still present in the upper plenum.

However, in this report we will illustrate through the use of first principle analysis of hydrodynamic requirements related to the sodium re-entry

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phenomenon which clearly suggests that the above assumption is overly conser-vative and that in fact the core must be uncovered of sodium before' sustained

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dryout and fuel overheating are possible. On this basis, fuel disruption, if

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If not natural circulation is assumed following pump coastdown, incipient boiling would initiate at about 3 to 4% decay power.

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f

DRAFT

_4_

1 it can occur at all*, will not take place before about 100 hours0.00116 days 0.0278 hours 1.653439e-4 weeks 3.805e-5 months into the postulated accident as compared to earlier estimates of 3 to 4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months .

The simple re-entry criterion developed specifically for the subassembly geometry is also compared with a low flow convection, high quality critical heat flux criterion based upon experimental data including liquid metals.

Heat flux values predicted by these two criteria are shown to be in excellent agreement.

2.1 Simple Re-Entry Criterion for Subassembly Geometry In considering the potential for sodium re-entry the effects of the rod bundle geometry and associated two-dimensional phenomena must be taken into account [4].

Since the subassembly wall is not a heat sink and the flow to power ratio is considerably higher in the wall subchannels than in the central subchannels, the liquid flow next to the hexcan wall has no limitation in terms of the critical heat flux phenomenon, i.e., the wall channels can be running full of liquid while the central subchannels are experiencing dryout.

In order to sustain dryout within the pin bundle portion of the subassembly, the vapor velocity and corresponding frictional pressure drop gradient must at least equal or exceed the gravitational pressure drop gradient corresponding to liquid sodium in the unheated hexcan wall channels so as to avoid refilling I

or re-entry of coolant.

For the interconnected channel system typified by the subassembly geometry this physical requirement can be simply stated as fol-lows:

2p D E gh u >

(2.1) gp 88 1

1

The effect of natural insulation heat losses at elevated temperatures may prevent such an event altogether [3).

I h

DRArt where p and p are the density of liquid and vapor respectively, D is the h

g hydraulic diameter,.f is the vapor friction factor and g is the gravitational constant. The heat flux that corresponds to Ineq. (2.1) is given by

[

2p p EU gg h (A/wDi N)(Ah q>

g h

g+hg)

(2.2)

{

g g

where A is the flow cross-sectional area of the subassembly, D is the fuel pin diameter, i is the heated length, N is the number of fuel pins in the assem-h

[

bly, Ah is the liquid subcooling enthalpy and h is the latent heat of g

g vaporization. Thus for prototypic core conditions such as CRBRP one obtains a value for the minimum required heat flux to prevent re-entry at zero sub-2 cooling of approximately 8.4 w/cm. This heat flux corresponds approximately

[

to 6% of the average nominal operating power.

Only a modest increase of

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approximately 20% is seen by accounting for maximum subcooling (Ah = 582 g

J/g), i.e., q = 10 w/cm.

A potential re-entry limitation provided by the single-channel geometry portion of the subassembly, i.e., downstream of the fuel pin section, also need to be examined.

For the channel dimension involved, the flooding

{

condition can be determined by the well known Kutateladze criterion [5].

u

>Kp

{(og(p - p )] !

(2.3) g g

Where K is the Kutateladze stability parameter, and o is the liquid surface

(

tension.

Using a value of K of 3.0 which corresponds to flooding of liquid films or transition from slug to annular flow, one obtains u

= 2400 cm/s.

The corresponding velocity in the fuel pin section is (2400)(101.4/43.4) =

56-7 cm/s.

Since this velocity is slightly larger than the value given by Ineq. (2.1), we conclude that Ineq. (2.2) provides a reasonable estimate of

(

the fuel pin heat flux that must be exceeded in order to prevent refilling or re-entry of liquid sodium.

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T

L DRAFT F

The equivalent critical heat flux value provided by Ineq. (2.2) is compared below to the following modification of the original Katto [6] cor-relation

by considering a typical LMFBR subassembly geometry [7].

[opg h0.00 - 1+y D

P t

h h

l i

(2.4) g l(G t )

q = 0.25 Gh l

'h k Pf fg c

g h

h where G is the mass flow rate.

The above correlation is compared to values predicted by Ineq. (2.2) in Table 2.1; the agreement is excellent.

Table 2.1 COMPARISON OF HEAT FLUXES FROM EMPIRICAL CORRELATION

[EQ. (2.4)] AND THE SIMPLE RE-ENTRY CRITERION [INEQ. (2.2)]

CHF -

Re-Entry T

Eq. (2.4)

Ineq. (2.2) 1 2

420'C s 10 w/cm s 10 w/cm 800*C m 8.5 w/cm N 8.7 w/cm No Subcooling 4

8 w/cm N 8.4 w/cm I

I I

Katto's original correlation was based upon critical heat flux data (liquid metals) relevant to the low flow convection (natural circulation) high quality nearly saturated inlet conditions regime of interest.

_ _ _.. _ _ _ _ _ DMR 3.0 NATURAL CIRCULATION HEAT REMOVAL ANALYSIS

(

Natural circulation analysis of the CRBRP shows that the system performance (e.g., temperature histories, heat removal rates) is strongly influenced in the short term by several factors which causes some temperatures and flows to change markedly within the first half-hour. Detailed calcula-tions, by the DEMO code for example, are needed to predict the system vari-(

ables with accuracy.

The large heat capacity of the primary system would limit fn the early stage of the event the average temperature increase to N

[

$5' *c/ hr-t for the conservative case of an adiabatic primary system. This fact

' ^ * " ' - - -

and the acceptability of local boiling in the core means that the evaluation

(

of core damage during a natural circulation event can focus on the longer term effects.

In this section the assumptions used in long term heat removal

{

analysis of natural circulation are explained and the results are presented and discussed for circulation in one, two, and three loops.

The scope of this analysis is to determine the heat removal rate from the primary syscem to the steam generator system; the ultimate heat sink for the cases without any

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feedwater is analyzed in the following section.

The best estimate natural circulation calculations were based on f

basic principles and CRBRP performance characteristics known from detailed DEMO analyses [8]. The hydraulic head equation given in Fig. 3.1 describes the system adequately for the quasi-static conditions which exist af ter N 1/2 hour of natural circulation.

In this time regime the cold leg temperature,

{

T, has reached equilibrium with the saturation temperature in the steam drum

(

because the IHX and evaporator are very ef ficient at low flow rates.

Decay power changes relatively slowly and inbalances between Q and Q are h

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dampened by the large heat capacity of the primary system hot leg.

Thus the

[

DRAFT

_g.

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HYDRAUUC HEAD I

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TH

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O OUT

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g

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AL IN U

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b TC

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6PH =ffp AT AL where AT = TH-TC

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Fig. 3.1 Hydraulic Head for Natural Circulation.

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s DRAFT

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hot leg can also be characterized by a single value of T at any given time H

after natural circulation has been in effect for over N 1/2. hour. The hot leg

(

temperature is obtained by an iterative process of allocating the decay heat input between natural circulation heat removal and heat capacity storage.

The hydraulic head is balanced by the frictional pressure drop, which is a function of the flow rate through the sections of CRBRP primary system.

The pressure drop equation is given in Fig. 3.2.

Plots of the

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frictional coefficients (K ) vs. normalized flow rates in the most recent CRBRP natural circulation study [8] were used in the best estimate calcula-tions. The flow rate is calculated from the hydraulic balance equation, and then the heat removal rate is obtained from the product of the flow rate, sodium heat capacity and temperature drop along the IHX.

The best estimate

{

values of the flow rate and system temperatures compare quite favorably with DEMO results at a 1000 seconds into natural circulation, as shown in Fig. 3.3.

The heat removal rate for natural circulation was calculated for one, two, and three loop operation using the model described above. The decay

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power values used in the CRBRP design calculations were reduced by 17% to obtain best estimate values. The final cold Icg temperatures were taken as

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315*C, which is the steam drum saturation temperature at SCAHRS operating

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conditions. The hot and cold leg primary system temperatures are given in Figs. 3.4, 3.5, and 3.6 for one, two, and three loop operation.

In all three cases the hot leg temperatures peak early in the transient a few degrees above

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the full power temperature and then steadily drop as the primary system cools down. The cold leg temperature is programmed to approach 315'C early in the transient. The plots of the decay power and the heat removal rates by natural circulation in one, two, and three loops in Figs. 3.7, 3.8, and 3.9, respec-tively, show that natural circulation heat removal overtakes decay power

[

_a

l DRAFT [

r l

l l

l HYDRAULIC PRESSURE DROP' 2

2+KW

+KW

+KW APf=KW3 2

2 3

3 4

4 3

I where APf = frictional pressure drop K; = flow-dependant constant for section i l

W; = flow velocity through section i i=1 core i=2 pump with locked rotor i=3 check valve i=4 piping, IHX, etc.

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,.- DRAFT r

b l

b APPROX.

METHOD DEMO T ('C) 511 493 H

T ('c) 268 257 c

AT('C) 243 236 W/W (%)

2.9

% 3.1 w

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1

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__ DRAFT F

production in about 1000 seconds, even with only one loop.

In less than a half-hour the natural circulation heat removal rates pea'ks and then steadily decreases as the AT driving force also diminishes. The results show that the CRBRP has very good natural circulation capabilities and that primary system temperatures correspond to those experienced in a mild transient, as long as a heat sink exists.

1 b

- DRAIT l

4.0 NATURAL DRAFT CAPACITY FOR THE PROTECTED l

AIR COOLED CONDENSERS (PACCs) l In assessing the response of the CRBRP system to a transient such as a total loss of AC power or an inability to supply auxiliary feedwater to the steam generators, the PACCs would play an essential role in both establishing the ultimate heat sink and the response time of the steam side.

Since these units are part of the heat transfer circuit during normal operation, they would be available for immediate heat renoval in such a postulated sequence.

At the beginning of the sequence the in1 it and outlet louvers to these units are assumed to open. The resulting natural draf t flow and the corresponding heat removal capacity is evaluated below.

4.1 General Configuration I

The PACCs units are illustrated in Fig. 4.1 with the flow paths being given in Fig. 4.2.

Each unit consists of 50 finned tubes arranged in a spiral as illustrated in Fig. 4.3.

These tubes are connected to an inlet and outlet header and are supplied steam from the top of the steam drum with the condensate draining back to the recirc pump header as illustrated in Figs. 4.4 I

and 4.5.

The major resistances to heat removal are through the wall of these tubes and the convective heat removal of f the outer surfaces of the tubes and the fins.

Both the tubes and fins are constructed of carbon steel with the details of the fins being illustrated in Fig. 4.6.

With this configuration, the air flow passes through the bank of spiral tubes and into the central region of the coil with the exhaust flow traveling through the outlet louvers and up through the exhaust stack.

For these evaluations the exhaust stack is assuced to be 10 m high.

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AND C ALC UL. ATlo N ) 4 0,'15 NUHbER OF SECNENTS PER RCW e 4

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DRAFT 4.2 Pressure Difference For Natural Draft Considering the schematic illustration for the PACCs shown in Fig.

4.7, the driving pressure for natural draft through the units can be expressed by AP = (p,9 -pd)Eh (4*1) and considering the air to be a perfect gas this can be expressed in terms of the inlet and outlet temperatures as P

1 1

(4.2) j gh AP = y I

~

T (T

ai) ao A typical air inlet temperature would be 300*K (80*F) with the outlet tempera-tures under these conditions ranging from 500*K to 550*K (440*F - 530*F). For these two ranges, the pressure difference av.allable for air flow is between 45.6 and 51.8 Pa.

4.3 System Flow Capacity Evaluations of the forced convection flow through the PACCs unit have been made and are reported in Ref. [9]. Considering all the resistances in the system, the resulting values are. tabulated in Table 4.1 in both English and SI units.

In addition, these are graphically represented in Fig. 4.8 and compared with an expression relating the pressure drop and the square of the flow rate.

2 AP = 0.43 W (4.3)

For an average driving pressure of 48 MPa, the air flow rate would be 10.6 kg/sec.

4.4 Air Side Heat Transfer Coef ficient As discussed in Ref. [10], the heat transfer coefficient over tube bundles can be represented as a function of the dimensionless grouping

. DRAFT

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n

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l a

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Fig. 4.7 Schematic of PACCe for Natural Draft Operation.

(

l i

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_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ DRAFI Table 4.1 PACC AIR FLOW RATE VS PRESSURE DROP Mass Flow Pressure Drop lbm/hr kg/sec Inches of Water Pa 124,259 15.7 0.4281 106.5 165,679 20.9 0.7810

'194.4 207,099 26.1 1.1939 297.2 248,519 31.4 1.7025 423.8 289,939 36.6 2.3077 574.4 331,358 41.8 2.9713 739.6 372,778 47.1 3.7085 923.1 I

I

__ DRAFT l

1000 -

800 -

as o.

h 600 x

O w

E 3

400 ew&n.

2 AP = 0.43 W 200 I

I I

0 O

10 20 30 40 50 MASS FLOW, kg/sec.

Fig. 4.8 PACCc Flou Rate vs. Pressure Drop.

s DMN D, l

(4.4)

{

h D

Pr /3 2

a max a

4 (E )l kN j

J

/

b a

P max where the term G is the maximum flux density between the minimum tube spacing.

Ignoring the presence of fins and the viscosity variation and assuming that the air flows uniformly radially inward toward the center of the

{

cooling unit, the maximum flux density is N 0.41 kg/m -sec.

The above dimen-sionless number is then N 1000 and the corresponding value for j, as shown in Fig.

4.9, is s 0.025.

Evaluating the gas side heat transfer coefficient yieids a value of 0.013 kw/m

K, or in English units 2.4 btu /hr-ft /*F.

This value is well within the range of those values considered for natural draf t

{

behavior.

4.5 overall lleat Transfer coefficient For this analysis the thermal resistance on the steam side is neglected and the principle thermal resistances are conduction through the tubes and convective heat removal off the tube and the fins by the air.

Since these are series resistances the overall heat transfer effectiveness can be i

expressed by (4.5)

^ " In(r /r )

9 y

2n k L + il I

where r is the uter radius of the tube and the fin base (2.68 cm) and r is 0

1 the inner diameter of these tubes (2.09 cm).

For the values given above, the product of UA would be 13 kw/*C.

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Fig. 4.9 Average Friction and Heat Transfer Data for Flou Over Five Different Arrangemente of 3/8 in. Diameter Tube Bundles in the Laminar and Transition Regime.

- _ _ _ _ DRAFT 4.6 Energy Removal from the PACCs As an air cooled condenser, the temperature variation for the steam side and air side within the PACCs can be represented as shown in Fig. 4.10.

The heat transfer rate can be expressed as Q = UA E = Wcp (T,9 - T,1)

GO where I

-O a

b H.

(4,7)

AT,

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,UA (4,g)

AT N#P(T,9 - T,1)

We b

p or UA ai,,We T

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s As illustrated in a previous section, the gas flow rate varies as the square root of the pressure drop and the pressure difference developed by the natural draft is relatively insensitive to the outlet air temperature. As an average value sufficient for these evaluations, we will consider the pressure differ-ence to be 48 MPa which provides for a gas flow rate of 10.6 kg/sec.

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AIR TEMPERATURE 2m I-U 300 DISTANCE Fig. 4.10 Steam and Air Temperature Variation in the PACCc.

DRAFT 4.7 Heat Removal Rates by Ultimate Heat Sinks The heat which is removed from the primary system by natural circu-lation is, assumed to be transferred through the IHX and evaporator in a very efficient manner.

The amount of heat removed from the primary system is assumed to equal the amount of heat used to generate the steam which flows to the steam drum. The heat is removed from the steam drum by condensing steam in the PACCs and by venting the excess steam to the atmosphere. The amount of the heat removed by venting at a given time is obtained by subtracting the constant PACC heat removal rate of 4.24 Mw/ loop from the current natural circulation heat removal rate.

The venting rates in the are plotted for one, two, and three loop natural circulation cases in Figs. 4.11, 4.12, and 4.13.

The venting rate is converted into the mass of the steam which is integrated to give the cumulative mass of steam vented from the system. These results are given in Figs. 4.14, 4.15, and 4.16.

The heat removal rate by venting peaks within a half hour and then decreases steadily until it reaches zero.

The three cases show significant differences in the time to the end of venting, about 9600 seconds (2 2/3 hrs) for three loop operation and over 100,000 sec (N 27 hours3.125e-4 days 0.0075 hours 4.464286e-5 weeks 1.02735e-5 months ) for one loop opera-tion.

The total mass of the steam vented also dif fers significantly, between 50,000 kg and 177,000 kg for three and one loop operation, respectively. Note that the total water inventory in the drums, pipes and evaporators is N 50,000 kg and in the protected water storage tank is % 250,000 kg.

The analysis shows that most of the steam is vented as a result of cooling down the primary system rather than as a result of removing decay power.

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_ _ _ _ _ _ _ _ _ - - _____ DRAFT 5.0

SUMMARY

l The report has analyzed three areas of conservatism which is suit-able for design purposes but is unneeded and undesirable for PRA purposes.

Replacing the conservative quantities with best estimate one results in a number of benefits to the PRA study of a LOHS due to loss of all AC power and/or feedwater.

Establishing sodium re-entry and dryout criteria results in

.at least a postponement of core damage for several hours.

In less than 1/2 l

hour natural circulation heat removal under one loop operation is greater than decay power so the reactor could be cooled down with one loop if an adequate heat source is available.

The natural draft heat removal capability of the PACCs is substan-tial (* 25-30% of the forced draft capacity).

Using best estimate values in place cf the conservative design ones results in predicting the cooldown of the system for about two hours of completely natural circulation operation without feedwater before the dry-out of the steam system.

The study also shows that it would take several more hours (> 10 hours1.157407e-4 days 0.00278 hours 1.653439e-5 weeks 3.805e-6 months ) to heat the system to a temperature at which core damage may occur.

The availability of this additional time may prove to be very beneficial in a PRA study.

The dif ficulty in achieving cooldown without AC power and feedwater pumps is with the short term heat removal.

The calculations predict that excessive steam would most likely be vented before the PACCs could assume the full load.

However, some conservative assumptions were still used in ob-taining these results and the removal of the conservatism should be investi-gated.

One remaining area of conservatism is that of assuming the heat exchangers (IHX and evaporator) to be highly efficient, even as the water

DRAFT L inventory is decreasing. Analysis to obtain the best estimate values for heat p

L exchanger efficiency is now being conducted.

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6.0 REFERENCES

l.

R. A. Bari, et al., " Accident Progression for a Loss-of-Heat-Sink with

(

Scram in a Liquid-Metal Fast Breeder Reactor," Nuclear Technology, Vol.

44, August 1979.

2.

K. R. Perkins, R. A. Bari, and W. T. Pratt, "In-Vessel Natural Circula-tion During a Hypothetical Losli-of-Heat-Sink Accident in the Fast Flux Test Facility," 79-WA/HT-66, Paper presented at the ASME Winter Annual r

l Meeting, New York, New York, December 2-7, 1979.

3.

H.

Vossebrecker and A. Kellner, " Inherent Safety Characteristics of f

Loop-Type LMFBRs," Proc. Intl. Mtg. on Fast Reactor Safety Technology, Seattle, Washington, August 19-23, 1979.

4.

T. C. Chawla and H. K. Fauske, "On the Incoherency in Subassembly Voiding

{

in FTR and Its Possible Effect on the Loss-of-Flow Accident Sequence,"

Trans. Am. Nucl. Soc., Vol. 17, p. 285, 1973.

5.

Kutateladze, " Elements of the Hydrodynamics of Cas-Liquid Systems," Fluid Mechanics - Soviet Res., Vol. 17, pp. 29-46, 1972.

6.

Y. Katto, "A Generalized Correlation of Critical Heat Flux for the Forced f

Convection Boiling in Vertical Uniformly Heated Round Tubes," Intl. J.

Heat Mass Transfer, Vol. 21, p. 1527, 1978.

7.

M. Ishii and H. K. Fauske, " Boiling and Dryout Behavior in an LMFBR Subassembly Bundle Under Low Heat Flux and Low Flow Conditions," to be published, 1982.

6.

CRBRP-ARD-0308, " Summary Report on the Current Assessment of the Natural Circulation Capability with the Heterogeneous Core," February 1982.

9.

K.

Chen, S.Wong, and D.

Drendel, " Thermal Evaluation of the CRBR Protected Air Cooled Condenser," Draft Report, March 1, 1982.

10.

F. Kreith, " Principles of Heat Tranrfer," International Textbook Company, 1958.

_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _

ML20065A370

Contents

Text

1.0 INTRODUCTION

SUMMARY

6.0 REFERENCES

1.0 INTRODUCTION

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4 (E )l kN j

SUMMARY

6.0 REFERENCES

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ML20065A370

Text

1.0 INTRODUCTION

SUMMARY

6.0 REFERENCES

1.0 INTRODUCTION

=

4 (E )l kN j

SUMMARY

6.0 REFERENCES

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