BDO410004

INTERN,.,IL REPORT INST-90IRPED-22, MA Y 2003

Analysis of the DNB ratio and the Loss-of-Flow Accident (LOFA) of the 3 MW TRIGA

MARK 11 Research Reactor

M. Q. Huda, M. S. Mahmood, T. K. Chakrobortty,M. Rahman and M. M. Sarker

REACTOR PHYSICS AND ENGINEERING DIVISION (RPED)

INSTITUTE OF NUCLEAR SCIENCE TECHNOLOGYATOMIC ENERGY RESEARCH ESTABLISHMENT

GANAKBARI, SAVAR, GPO BOX 3787, DHAKA-1000 BANGLADESH

,q! 101- 9m�Wfw

BANGLADESH ATOMIC ENERGY COMMISSION

INST-90/RPED-22

CONTENTS

Page

1. Introduction 1

2. Analysis of DNB Ratio 1

2.1. Effect of Operating Power 3

2.2. Effect of the Flow Rate 3

2.3. Effect of the Hot-Rod Factor 3

2.4. Effect of the Inlet Temperature 4

3. Loss-of-Flow Accident (LOFA) Analysis 4

3.1. Relative Power and Flow during LOFA 4

3.2. Fuel Temperature Distribution during LOFA 4

3.3. Occurrence of Flow Reversal 5

4. Conclusions 5

5. Acknowledgments 6

6. References 6

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ABSTRACT

The PARET code was used to analyze two most important thermal hydraulic designparameters of the 3 MW TRIGA MARK research reactor. The first design parameter is theDNB (departure from nucleate boiling) ratio, which is defined as the ratio of the critical heatflux to the beat flux achieved in the core and was computed by means of a suitable correlationas defined in PARET code. The reactor core should be designed so as to prevent the DNBRfrom dropping below a chosen value under a high heat flux transient condition for the mostadverse set of mechanical and coolant conditions. Over the length 0381 in of the hottestchannel the DNB ratio vaiies, starting from 38951 to 5403 1, with a minimum of .785 1. Thepeak heat flux occurs at the axial center of the fue e therefore the DNB ratio isM1111mum at this location.

The second design parameter is the loss-of-flow accident scenario of te TRIGAreactor. The Bergles-Rohsenow correlation was selected for detecting onset of nucleateboiling, te transition model with the McAdams correlation was included for fully developedtwo-phase flow, and the Seider-Tate correlation was used for the single-phase forcedconvection regime. The lss-of-flow transient after a trip time of 408 sec at 85% of loss ofnormal flow for the TRIGA core shows a peak temperature of 709.220C in the fuel centerlineand 131.94'C in the clad and 46.63'C in the coolant exit of the hottest channel. The transientwas terminated at 15% of nominal flow after about 48.0 sec. The time at which the reversalof coolant flow starts is about 67.0 sec.

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1. Introduction

The 3MW TRIGA MARK II research reactor was commissioned in late 1986 at Savarnear Dhaka, Bangladesh. The principal objective of this study is to provide thorough analysesof two most important thermal hydraulic design parameters of the reactor, namely the departurefrom nucleate boiling ratio (DN-BR) and the loss-of-flow accident (LOFA). These analyses arean extension of some other studies reported recently 131. The core neutronic parametersneede& f6i the analyses were obtained from the work of Bhuiymi et al 4]. This investigationwill provide a basis for updating the Final Safety Analysis Report (FSAR) of the TRIGAreactor and also for upgrading the current core configuration through bericbmarking theavailable TRIGA experimental and operational data both for steady-state and transient mode ofreactor operations.

The PARET [5] computer code that has already been used successfully elsewhere inthe world for te analysis of TRIGA research reactor 6] was applied for this analysis. Theimportant behaviors that have been studied relating DNB analysis are: (i) effect of operatingpower, (ii) effect of the flow rate, (iii) effect of the hot-rod factor, and (iv) the effect of theinlet temperature. For LOFA analysis the parameters those have been studied area (i) relativepower and flow for the exponential 16ss-of-flow transients, (ii) the resulting fuel centerfine,clad surface and coolant outlet temperatures after the loss-of-flow transients, and (iii) theoccurring of coolant flow reversal.

2. Analysis of DNB Ratio

'flee power output of the reactor is limited by three dependent thermal andhydrodynamic variables: the DNBR, the maximum fuel temperature (Tma,), and the corepressure drop (A ). The variations of these variables are discussed in the following sections.The first design limitation is related to the DNB flux, which can be computed by means of asuitable correlation. When the heat flux become sufficiently large, the small bubble forined innucleate boiling coalesce into a vapor film that covers the surface, since then the heat4ransferefficiency drops dramatically, and the clad surface temperature will rise by several hundreddegrees. This situation is known as the DNBR. The DNBR is defined as the ratio of the cticalheat flux to the heat flux achieved the core. The reactor core should be designed so as toprevent the DNBR from dropping below a chosen value under a high heat flux transientcondition for the most adverse set of mechanical and coolant conditions.

The PARET code was used to analyze the thermal and hydraulic design for the 3 WTRIGA MARK 11 research reactor. Two-channel model of PARET code was used where onechannel represents the hottest rod and associated flow channel, and other average channel torepresent the remaining 99 fuel rods a volume weighted sense. The axial source distributionwas represented by 15 axial regions. The correlations from Dittus-Boelter 71 for single-phase,McAdams [8] for two-phase and Bernath 9] for ctical heat flux analysis were used in PARETcalculation and the mode of transition model was selected. The Nusselt number used in therevised single-p ase heat transfer coefficient subroutine was 140 for laminar flow and theenhance effects of laminar flow were also included. Detailed input preparations were reportedby Huda et al [ 1 3].

in forced convection, the heat flux and the wall and bulk temperatures are related byq j -_ h (T, - T)

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where, q = forced convection heat fluxh forced convection heat transfer coefficientT, = wall temperatureTb coolant bulk temperature

The forced-convection heat transfer coefficient h was determined with the Dittus-Boeltercorrelation as recom ended by Tong and Weisman 7], namely,

Nu = .023 Re 0.1 r 0.4

where, Nu, Re, and Pr are the Nusselt, Reynolds, ad Prandtl numbers, respectively, based onthe bulk water properties,

At higher ,eat fluxes, the surface is uniformly covered by a dynamic bubble layer, andthe heat is removed by fully developed, subcooled ucleate boiling. The correlation used forthis mode of heat transfer is from McAdams eal. [8], namely,

q, 0074 A Ts'-"where, qfd is the heat flux for fully developed subcooled nucleate boiling, and AT, is thedifference between te surface and saturation temperature.

In the fully developed nucleate boiling regime it is possible to increase the heat fluxfurther without an appreciable change the surface temperature, unti te bubble motion o theSurface becomes so violent that a hydrodynamic crisis occurs with the formation of acontinuous vapor Film in the surface. This is termed DN13, and the heat flux is the critical heatfILIX (CHF). The ratio of the CHF to the actual heat flux is the DNBR. In subcooled boiling theCHF is a function of the coolant velocity, the degree of subcooling, and the pressure. Thecorrelation used to predict the CHF is from Bernath 9] and is given by

q = k (Tw, TO

where, h, = 10,890 [D,/(D, + Di)] + sVs = 48/DO-' for D < 003048 n

= 90 I De) for D > 003048 nTw, I S 57 In P 54[P/(P + 15) - V/41 32, (critical wall temperature)D, hydraulic diameter of coolant channelDi diameter of heated surfaceV coolant velocityP pressureTB bulk coolant temperature

The design flow Tate has a lower limit determined by the value of the CHF at that flowrate- the larger the flow rate the larger will be the CHF and hence the safety margin. he flowrate also has an upper limit, which is determined by the maximum allowable pressure dropthrough the core to avoid cavitation in the flow system. The design conditions used for thePARET calculation are tabulated in Table .

Since the enthalpy of te coolant increases as i flows through the core, the computedDNB flux decreases correspondingly, as depicted in Fig. 1, The heat flux along the hot channeland the DNBR are also sown in the figure qualitatively. Over the length 03 8 1 m: of the hottestchannel the DNBR varies, starting from 38951 to 5403 1, with a minimum of 2785 1 . The peakheat flux occurs at the axial center of the fuel elements-, therefore the DNB ratio is minimum atthis location.

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Table 1. The design conditions used for the PARET calculation for DNB analysis

Operating Power 3MWFlow rate 3208.9 kg/mz-s 3500 gpm)Inlet temperature 40.5556 C (105 OF)Pressure L60654XI0`kPaHot Rod Factor (BFR) 1.854 (C4 element)Inlet pressure loss coefficient 1.81Outlet pressure loss coefficient 2.12

2-1. Effect of Operating Power

The mm'imum DNBR and the corresponding maximum fuel temperature as a functionof operating power are shown in Fig. 2 The correlations [5] used for calculating the criticalheat flux/DNBR in the PARET code were Tong, Bemath and Mirshak among which Bemathgives the minimum value, and according to the suggestions made by the General Atomics (GA)the Bemath correlation was selected [5]. It can be seen that the DNBR is 27851 at 3 MW andis 23 03 at 120% over power. Even at 4 MW it is 20162, but the maximum fuel temperaturewas computed to be 1024.76C, that is well above the maximum operating temperature limit forTRIGA LEU (low enriched uranium) fuel.

2.2. Effect of the Flow Rate

The variation of the maximum fuel temperature. the DNBR, and the core pressure dropwith the reactor flow rate is shown in Fig. 3 According to the Berriath DNB correlation a flowrate as low as 1833.7 kg/m2.s 2000 gpm) provides a DNBR of 22233 with a pressure drop of4.89 kPa. At a flow rate of 4584.2 kg/m2.s (5000 gpm) the DNBR is 33063 and thecorresponding pressure drop is 40.12 kPa which is acceptable. As suggested by GA, Ap mustnot be larger than about 103.42 kPa, so that with the inlet pressure of 160.65 kPa, cavitationwi e avoided at all points in the flow system.

A design flow rate of 3208.9 kg/m2.sl 3500 gpm) was selected under forced convectionmode. At this flow rate the DNBR is 2.785I (average Reynold number Re was found to be95983.66 over the total length of the hot channel) and the pressure drop is 102 kPa (pressuredrop due to friction was 842, for elevation 738, for spatial acceleration 003, and no drop dueto transient acceleration). There is a sufficient margin of safety with regard to cavitation andmaximum fuel temperature. Snce the design flow rate has a conservative safety margin, theeffect of flow maldistribution will be minimal. As can be seen from the figure, the maximumfuel temperature is relatively insensitive to variations in flow rate.

2.3. Effect of the Hot-Rod Factor

The hot-rod factor (HRF) accounts for the maximum variation in element powergeneration in the cores and the design value is 1854. This factor is applied to the core averageheat generation per fuel element in order to determine the heat generation at the hottestlocations in the core. These locations occur at the center of the tore.

Figure 4 shows the variation of Ta, and DNBR with HRF (the Ap curve is not shownbecause HRF has no effect on Ap). Below an HRF of 1.25, no boiling is predicted and Tna,,decreases rapidly. Above 1 .25, boiling is predicted and the element surface temperature

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remains nearly constant; however, Tnax increases with BRF because the temperature dropthrough the fuel ad fuel-cladding interface increases with increasing heat generation. TheDNB ratio at the design value is 27851 and decreases with HRF, since the calculated heat fluxin the denominator of the DNBR increases with HRF and the CBF in the numerator remainsiiearly constant.

2.4. Effect of the Inlet Temperature

The design inlet temperature of 40,55560C (1050F) corresponds to the heat exchangeroperating on the hottest day anticipated at the reactor site. Because surface boiling is predictedfor the hottest elements, the inlet temperature has virtually no effect on the maximum fueltemperature, as shown Fig. The DNBR is affected because a higher bulk watertemperature reduces the degree of subcooling. At the design value the DNBR is 27863.

3. Loss-of-Flow Accident (LOFA) Analysis

When the reactor is shut down from power under normal operation, the main coolantpumps will continue to be operated for a short time until the fuel temperature declines to a near-ambient value. Should the pumps fail or be shut off because of an emergency during full-poweroperation, the reactor would scram on a loss-of-flow signal. Experiments conducted on otherforce-flow- cooled TRIGA reactors show that the flow reverses direction to the naturalconvection mode very quickly and smoothly, with essentially no interruption in the fueltemperature decay rate. Thus, te afterheat from te shut-down reactor will be removed bynatural convection following pump failure or emergency shutdown.

'Me PARET code was used to analyze the loss of flow accident scenario for the 3 MWTRIGA MARK 11 reactor. Two-channel model of PARET code was used where one channelrepresents the hottest rod and associated flow channel, and other "average" channel to representthe remaining 99 fuel rods in a volume weighted sense. The axial source distribution wasrepresented by 15 axial regions. The Bergles-Robsenow correlation 7 was selected fordetecting onset of nucleate boiling, the transition model with the McAdams correlation [81 wasincluded for fully developed two-phase flow, and the Seider-Tate correlation 7] was used forthe single-phase forced convection regime. Detailed input preparations were reported [ 1 -3

3.1. Relative Power and Flow durin2 LOFA

Figure 6 shows the relative power and flow for the exponential loss-of-flow transientswith a time constant of 25.0 sec in the TRIGA core. This loss-of-flow transient ischaracterized uantitatively in Table 2 The flow coast-down was initiated after 1.0 sec at apower of 3 MW. Thus, a 33 MW overpower factor was included.

3.2. Fuel Temperature Distribution during LOFA

The resulting temperatures at the fuel centerline, clad surface, and coolant outletduring LOFA are shown in Fig. 7 The loss-of-flow transient after a trip time of 408 se at85% of loss of normal flow for the TRIGA core shows a peak temperature of 709.29T in thefuel centerline and 131.940C in the clad and 46,630C in the coolant exit of the hottest channel.

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Table 2 Tabulated results for the transient response of the TRIGA core to a loss-of-coolant flow with a decay time of 25.0 sec a scram trip at 85% flow and a .015 sec delay

Initial Power MW) 3.0Trip Time at 85% Low Flow sec) 4.08Time for Nominal Flow of 15% (see) 48.0Time after Flow Reversal sec) 67.0

---HOT AVERAGESteady State Temperatures Cc):Fuel Centerline 791.60 460.30Clad Surface 231.44 89.56Coolant 44.84 41.69Peak Temperatures OQ after Loss-of-Flow:Fuel Centerline 709.22 396.8Clad Surface 131.94 74.77Coolant 46.63 40.77Peak Temperatures Q at 15% Nominal Flow:Fuel Centerline 225.91 I3&34Clad Surface 111.79 78.32Coolant 46.57 42.44Peak Temperatures after Flow Inversion (0c):Fuel Centerline 207.95 135' 27Clad Surface 128.29 95�74Coolant 55.12 49.54

3.3. Occurrence of Flow Reversal

The transient was terminated at 15% of nominal flow after about 48.0 sec. Realistically,the flow would be expected to reverse direction and establish a natural convection flow ratewhich should be adequate to cool the core. The time at which the reversal of coolant flow startsis about 67.0 see. It should also be noted that at low flow rates the peak temperature in thecoolant occur upstream from the outlet (the heated slug has not yet reached the outlet).

0Even if a reactor scram did not occur due to the loss of flow, experience indicates that

the reactor power would decrease to the range of 6 to 2 MW because of changes in the corewater temperature and density. This power range is near the limit of natural convection cooling,but it is expected that the heat would be removed by the natural convection flow through thecore.

4. Conclusions

It is found that the PARET code can be ud effectively for analyzing importantthermal hydraulic design parameters of the 3 MW TRIGA MARK II research reactor. Thetesting of the PARET model calculations through benchmarking the available TRIGAexperimental and operational data showed that PARET can successfully be used to analyzethese type of design conditions of the reactor to predict the safety margins. The maximumpower density remains, by a substantial margin below the level at which the DNB couldoccur. Snce the enthalpy of the coolant creases as it flows through the core, the computedDNB flux decreases correspondingly. The lss-of-flow transient after a trip time of 4.08 sec at

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85% of loss of normal flow for the TRIGA core shows a peak temperature of 709.22'C n thefuel centcrline and 131.94'C in the clad and 46.63'C in the coolant exit of the hottest channel.Even if a reactor scram did not occur during LOFA, experience indicates that the reactorpower would decrease to the range of 16 to 2 MW because of changes in the core watertemperature and density.

5. Acknowledgements

The authors wish to thank Professor T. A. Parish of Texas A&M University, USA forhis valuable suggestions: and assistance in modeling the various portions of the TRIGAreactor. Thanks are due to the members of the TRIGA Reactor Operation MaintenanceUnit of AERE for their cooperation throughout this work. The authors also wish to thankProf Dr. Abdul Jalil, Director, INST for his encouragement and administrative supportthroughout the work.The support of the Ministry of Science Technology, Government ofthe People's Republic of Bangladesh is acknowledged gratefully.

6. References

1. M. Q. Huda, S. I. Bhuiyan, T. K. Chakrobortty, M. M. Sarker and M. A. W. Mondal,"Thermal Hydraulic Analysis of the 3 MW TRIGA MARK 11 Research Reactor",Nuclear Technology, 135(l), 51-66 2001).

M. Q. Huda, S. . Bhuiyan, T, K. Chakrobortty and M. A. W. Mondal, "ThermalHydraulic Analysis for the Transients of the 3MW TRIGA MARK II ResearchReactor", INST-70/TRPSPD-2, AERE, Savar, Dhaka, December 1999.

3. M. Hda, S. I. Bhuiyan, T. K. Chakrobortty and M. A. W. Mondal, "Steady StateThermal Hydraulic Analysis of the 3MW TRIGA MARK 11 Research Reactor",INST-7 I TRPSPD-3, AERE, Savar, Dhaka, December 1999.

4. S. 1. Bhuiyan, M. A. W. Mondal, M. Rahman, M. M. Sarker, M. S. Shahdatullah M.Q Huda, T. K. Chakroborty and M. J. H. Khan, Criticality and &fety ParameterStudies of 3MW TRIGA MARK II Research Reactor and Validation of GeneratedCross Section Library and Computational Method", Nuclear Technology, 130(2), 1 1 -131 2000).

5. C. F. Obenchain, PARET - A program for the analysis of reactor transients, IDO-17282, January 1969,

64 & T. Rearden, Engineering Analysis of a Power Upgrade for the Texas AM NuclearScience Center Reactor, A&M University, Texas, USA, August 1995,

7. L S. Tong and Weisman, "Thermal Analysis of Pressurized Water Reactors,"American Nuclear Society, Hinsdale, Illinois, 1970.

8. W. H. McAdams et al, "Heat Transfer at High Rates to Water with Surface ilin inIndustrial and Engineering Chemistry", Reinhold, New York, p. 1 945, 1949.

9. L. Bemath, "A Theory of Local Boiling Burnout and its Application to ExistingData," Heat Transfer- Chemical Engineering Progress Symposium Series 56, 95,

(1960).

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DNBRBERNATH

0

z

0 �- IVIINIIVIUM-2.7851X

CRITICALLL HEAT FLUX

Ujr

HOT CHANNELHEAT FLUX

COOLANT DISTANCE UP CHANNEL COOLANTIN OUT

Fig. 1. Qualitative representation of heat flux and related conditions alongthe hot channel in the 3 MW TRIGA MARK 11 reserach reactor.

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9- 1000

8- 900

7 DNBRBERNATH TMAX 800 2

6 :3

- 700 ID5 0.z E

r') a)- 6004

3 - - 500

U2 - 400 7j-D

=3U-

300

2 3 4Power MW)

Fig. 2 Fuel performance at different power levels of the TRIGA reactor.

50 6...... .....

800 1.35e+5

40 5

1. 1 5e+&OE30 4 - RAv.'-' 775 z

CL

-j z E0 91.50e+41

n'- 20 - 3 - (D<

I/ DNBRBERNATH 4)750C

(.5 7'.50.e+410 2 - 7a-AP TOTAL U-

0 1 725 - 5.50e+4

2000 2500 3000 3500 4000 4500 5000Flow Rate (kg/M2S)

Fig. 3 Fuel performance at various flow rates.

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6 900

5 800BERNATH

:3

4 700 TCDCL

z ECD

3 600

4-C:02 500 2

CDI

U-

1 400

1.00 1.25 1.50 1.75 2.00

Hot Rod FactorFig. 4. Fuel performance at different hot rod factors.

4.0 - 825

TMAX3.5 800

CD

z 3.0 775 Er) IDNBR BERNATH

CD

2.5 750

2.0 - 725

30 35 40 45 50

Inlet CooInat Temperature C)Fig. 5. Fuel performance at different inlet temperature.

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Scram1.0

0.9

- Trip0.8

0 0.7 -

Cu 0.6 -

0n 0.5 -Flow

4-10.4 -

0.3Power

0.2

0.1

0.0

0 10 20 30 40 50 60

Time (sec)

Fig. 6 Flow coast down transient response of TRIGA core to a loss-of-Coolant flowwith a decay time of 25.0 sec a scram trip at 85% and a .01 5 sec d lay.

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800

700 - HOT CHANNEL

600 -Fuel Centerline

0 500 - CladCoolant

400CL

300 -

200 -

100 -- - - - - - - - - - - - - ---

0 1 1 1 1 1 1 1 1 1 - I I0 10 20 30 40 50 60 70 80 90 100

Time (sec)

500

AVERAGE CHANNEL

400

300Fuel CenterlineCladCoolant

Ecl- 200

100

------------------------

0 - I I I I I I I I -- F-

0 10 20 30 40 50 60 70 80 90 100

Time (sec)

Fig. 7 Flow coast down temperatures at the fuel cnterline, clad and coolantexit after the transient response of the TRIGA core to loss of flowaccidents for the hot and average channel.