A Study on Critical Heat Flux in Gap

KAERI/TR-1321/99

KR9900259

A Study on Critical Heat Flux in Gap

1998

1999. 4.

°f

II.

TMI-2

l-71 4]5>^ ^ ^ s > 3 1 $lfe SONATA-IV(Simulation Of Naturally Arrested Thermal Attack In-Vessel) W ^ I ^ ^ r - ^ - S . ^ CHFG(Critical Heat Fux in Gap) -M^^r ^ ^ 5 > a l

#$171

III.

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71^51 %& tt

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71^21

CCFL

CHFG

CHFG

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5514.

SUMMARY

I. Report Title

A Study on Critical Heat Flux in Gap

II. Objective and Importance of the Study

As a part of the SONATA-IV(Simulation Of Naturally Arrested Thermal Attack In-Vessel) program, performing to verify the cooling mechanism of the corium in the lower plenum during severe accident in the TMI-2 nuclear power plant, an experimental study of the CHFG(Critical Heat Flux in Gap) have been performed to measure the critical power and to investigate the inherent cooling mechanism, because there has been no experimental data on CHF in hemispherical narrow gaps

III. Scope and Content of Study

The scope and content of this technical report is to perform the test on critical heat flux in hemispherical narrow gaps using distilled water and Freon R-113 as experimental parameters, such as system pressure from 1 to 10 atm and gap thickness of 0.5, 1.0, 2.0, and 5.0 mm. The experimental results on critical power were compared with the existing correlation, developed in flat plate and annuli gaps. The distilled water data on critical power were also compared with the Freon R-113 data. The conduction heat transfer in the copper shell has been performed to verify

the experimental results using CFX-F3D computer code. The CCFL(Counter Counter Flow Limit) test is being performed to evaluate the CHFG test results on critical power in hemispherical narrow gaps.

IV. Results of the Study

The test results have shown that even if local dryout occurs, there exists a quasi-steady state and the temperature of the dryout region is limited within a certain value. When the heater power is large enough, however, there is no quasi-steady state. The dryout region expands by itself without an increase in heater power, finally leads to a global dryout, and the temperature of the heater surface monotonically increases. The heat flux bringing about that situation is defined as the critical power in the present experiments.

The CHFG test results have shown that the measured values of critical power are much lower than the predictions made by empirical CHF correlations applicable to flat plate gaps and annuli. The pressure effect on the critical power was found to be much milder than predictions by those CHF correlations. The values and the pressure trend of the critical powers measured in the present experiments are close to the values converted from the CCFL data. This confirms the claim that a CCFL brings about local dryout and finally, global dryout in hemispherical narrow gaps. Increases in the gap thickness lead to increase in critical power. The measured critical power using R-113 in hemispherical narrow gaps are 60 % lower than that using water due to the lower boiling point, which is different from the pool boiling condition.

The CFX-F3D results on conduction heat transfer have shown that the copper shell contacts with heater very well during CHFG tests. The heat conduction in the copper shell is effective on copper shell temperature in CHF(Critical Heat Flux) condition. The CCFL(Counter Counter Flow Limit) test facility was constructed and the test is being performed to estimate the CCFL

phenomena and to evaluate the CHFG test results on critical power in hemispherical narrow gaps.

V. Proposal for Application

The present experimental data on critical power in hemispherical narrow gaps, which is only in the world, will be used computer code development on in-vessel corium coolability issue. Further studies are needed to investigate the high-pressure effect on critical power because the electrical power is limited in the present test.

CONTENTS

Chapter 1. Introduction 1 Chapter 2. Experimental Contents and Method of CHFG 5 Section 1. Experimental Facility 5 Section 2. Experimental Methodology 12 Chapter 3. Experimental Results and Discussion of CHFG 15

Section 1. Experimental Results and Discussion using Distilled Water 15 Section 2. Experimental Results and Discussion of using Freon R-113 24 Section 3. Estimation ofThermal Conductivity Effect in the Copper Shell 29 Section 4. Comparison and Discussion of Experimental Data 37 Chapter 4. Experimental Contents and Method of CCFL 45

1. JQAUClilllClltCll

Section 3. Experimental Methodology 50 Chapter 5. Conclusion and Recommendation 53 Chapter 6. References 55 Appendix 1. Detailed Description of the CHFG Test Facility 59 Appendix 2. CFX-F3D Input 73 Appendix 3. Published Paper on CHFG Test in International Journal 81

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1.1 T. G. Theofanous, C. Liu, S. Addition, S. Angelini, O. Kymalainen, and T. Salmassi, "In-Vessel Coolability and Retention of a Core Melt," DOE/ID-10460, Vol. 1&2, 1995 1.2 O. Kymalainen, H. Tuomisto, and T. G. Theofanous, "In-vessel Retention of Corium at

the Loviisa Plant," Nuclear Engineering and Design, Vol. 169, pp. 109-130, 1997

1.3 J. L. Rempe, J. R. Wolf, S. A. Chavez, K. G. Condie, D. L. Hagrman, W. J. Carmack, "Investigation of the coolability of a continuous mass of relocated debris to a water-filled lower plenum," EGG-RAAM-11145, 1994

1.4 J. R. Wolf et. a l , "TMI - 2 Vessel Investigation Project Integration Report," NUREG/CR-6197 (EGG - 2734), 1994

1.5 S. B. Kim et.al., "Recent Progress in SONATA-IV Project," OECD/NEA CSNI PWG-2, The Third Mtg. Of TG-DCC, Rockville, MD, USA, May 9-10, 1997

1.6 K. Y. Suh et al., "SONATA-IV Simulation Of Naturally Arrested Thermal Attack In-Vessel," Proc. Int. Conf. on PSA Methodology and Applications, pp. 453-460, Seoul, November 26-30, 1995

1.7 K. H. Kang et al., "Experimental Investigations on In-Vessel Debris Coolability through Inhererent Cooling Mechanisms, OECD/CSNI Workshop on In-Vessel Core Debris Retention and Coolability," Garching, Germany, March 3-6, 1998

1.8 J. H. Jeong et al., "Experimental Study on CHF in a Hemispherical Narrow Gap," OECD/CSNI Workshop on In-Vessel Core Debris Retention and Coolability, Garching, Germany, March 3-6, 1998

1.9 W. Kohler, H. Schmidt, O. Herbst, W. Kratzer, "Experiments on heat removal in a gap between debris crust and RPV wall," OECD/CSNI Workshop on In-Vessel Core Debris

Retention and Coolability, Garching, Gennany, 1998

1.10 K. H. Bang et al., "Boiling Heat Transfer in Narrow Spaces and Its Implications for Lower Head Integrity during a Severe Accident," Proc. Int. Top. Meet, on Probabilistic Safety Assessment, pp. 1206-1211, Park City, Utah, 1996

1.11 t W , ^ 1 ^ , ?M3L, o l « , ^%Tt, "?% ^± *H^HS} ^ t ^

7}X\$- ^^,"cfl^7l^l^-sl^7il^-#tHs] fe^ B, pp. 105-109, 1996

1.12 3*1$ S], "tiV^tg ^ o f l ^ o ] «1^7M^- ^ , " ^ € * r 3 * N '97 ±A] ^ - # ^ ^ 5 1 fe-g-^1, *fl 1 ^ , pp. 575, a^tfl^SL, 1997 Vl 5 ^ 30 ^ -31 «g 1.13 J. H. Jeong, R. J. Park, and S. B. Kim, "Visualization Experiments of the Two-Phase

Flow Inside a Hemispherical Gap," Int. Commu. Heat & Mass Transfer, Vol. 25, No. 5, pp. 693-700, 1998

-^,"KAERI/TR-1027/98, ^ ^ f ^ T 1 ^ , 1998 Vl 4 ^

3.1 P. B. Whalley, Boiling, "Condensation and Gas-Liquid Flow," Oxford, U.K., Oxford University Press, 1987.

3.2 W. Peyayopanakul and J. W. Westwater, "Evaluation of the Unsteady-state Quenching Method for Determining Boiling Curves," Int. J. Heat Mass Transfer 21, pp. 1437-1445,

1978

3.3 M. S. El-Genk and A. G. Glebov, "Transient Pool Boiling from Downward-facing Curved Surfaces," Int. J. Heat Mass Transfer 38, pp. 2209-2224, 1995

3.4 "CFX-F3D Code User Manual," AEA Technology, 1995

3.5 "MATLAB New Features Guide Version 4.0, " The MATH WORKS Inc., March 1003 3.6 Y. Chang and S. Yao, "Critical Heat Flux of Narrow Vertical Annuli with Closed

3.7 M. Monde., H. Kusuda, and H. Uehara, "Critical Heat Transfer during Natural Convective Boiling in Vertical Rectangular Channels Submerged in Saturated Liquid," J. Heat Transfer 104, pp. 300-303, 1982

3.8 Y. Katto and Y. Kosho, "Critical Heat Flux of Saturated Natural Convection Boiling in a Space Bounded by Two Horizontal Co-axial Disks and Heated from Below," Int. J. Multiphase Flow 5, pp. 219-224, 1979

3.9 R. E. Henry and R. J. Hammersley, "An Experimental Investigation of Possible In-Vessel Cooling Mechanisms," CSARP Meeting, Bethesda, Maryland, 1997

3.10 "Modular Accident Analysis Program User's Manual," EPRI, 1994

3.11 Y. Koizumi, H. Nishida, H. Ohtake, and T. Miyashita, "Gravitational Water Penetration into Narrow-gap Annular Flow Passages with Upward Gas Flow," Procs. Of NURETH-8

1, pp. 48-52, 1997

3.12 N Zuber, "Hydrodynamic Aspects of Boiling Heat Transfer," AECU-4439, 1959

4.1 F. Mayinger, P. Weiss, and K. Wolfert, "Two-Phase Flow Phenomena in Full-scale Reactor Geometry," Nuclear Engineering and Design 145, pp. 47-61, 1993

4.2 S. C. Lee, C. Mo, S. C. Nam, and J. Y. Lee, "Thermal-hydraulic Behaviours and Flooding of ECC in DVI Systems," KAERI Report, KAERI/CM-045/95, 1995

4.3 L. Y. Cheng, "Counter-Current Flow Limitation in Thin Rectangular Channels," BNL Report, BNL-44836, 1990

4.4 K. Mishima and H. Nishihara, "The Effect of Flow Direction and Magnitude on CHF for Low Pressure Water in Thin Rectangular Channels," Nuclear Engineering and Design 86, pp. 165-181, 1985

4.5 Y. Sudo and M. Karhinaga, "A CHF Characteristic for Downward Flow in a Narrow Vertical Rectangular Channel Heated from Both Sides," Int. J. Multiphase Flow 15, pp. 755-766, 1989

4.6 G. B. Wallis and S. Makkenchery, "The Hanging Film Phenomenon in Vertical Annular Two-Phase Flow," J. Fluids Engineering 96(3), pp.297-298, 1974

4.7 S. S. Kutateladze, "Heat Transfer in Condensation and Boiling," USAEC Rep-tr 3770, 1952

4.8 M. Osakabe and Y. Kawasaki, 'Top Flooding in Thin Rectangular and Annular Passages," Int. J. Multiphase Flow 15, pp. 747-754, 1989

4.9 H. J. Richter, "Flooding in Tubes and Annuli," Int. J. Multiphase Flow 7, pp. 647-658, 1981

bd

I NOZZLE TABLE SERVICE DEMI WATER IN VAPOR OUTLET CONDENSATE INLET PT CONN. PSV CONN. UT CONN LT CONN.

TYPE MAT'L RATING BW/C FLO 8WC BW/C BW/C BWt BW/C SUS3O4 SUS304 16KO 16KO BUTT WELD CONNECTOR

(H0KE-8CBW8) DESIGN CONDITIONS PRESSURE IS kq/cma TEMPERATURE: 300 OPERATION CONDITIONS PRESSURE 10 TEMPERATURE: 200

CODE CODE STAMP

MAWP

SHELL SPEC.. SUS 304 B0WID x 320H xt7t JOINT EFF OISH S P E C . 85S SUS 304 600(10 X 300H x 201 JOINT EFF.

CLADDING OR LININO SPEC

85X

OASKET SPEC. SUPPORT SPEC. EXTERNAL PIPE S P E C . INTERNAL PIPE SPEC._ FIANGF 5PEC. METAL 4-3UPP0RT LUO SUS 304 RAT1N0L. FACING-JIS 18 kg/em* SO/RF SUS PART SANOINO GLASS WOOL lOOt

_OPERATINO WT_ TFST WT

PAINT INSULATION FtREPROOFIN0_ SHIPPING WT._

ELEVATION MEASURED FROM BASELINE.

ORIENTATION MEASURED CLOCKWISE FROM 0* SUREO PROJECTION MEASURED FROM C OF VESSEL TO EXTREME FACE OF FLANGE. PROOF

LUO NOZZLE STANDARD TITLE STD-1020 STD-1007 DRAWING NO. REFERENCE DRAWING

SHINSUNG CAM PLANT CO., LTD.

_{Sheet No. EDS-4002}

HEAT EXCHANGER DATA SHEET

Revision Date Checked by Plont SONATA-rV / CHFG Client KAERI Item No. E—111 P r o j e c t N o . 61217 Service CONDENSER

Locotion TAEJON. KOREA No. Required ONE

Heot Duty 3 4 . 0 0 0 Kcol/Hr Shells/Unit

Ht. Trans. Area per Unit 0.5 ml Ht. Trans. Areo per Shell 1.0

Code TEMA

PERFORMANCE OF ONE UNIT CONSTRUCTION

Shell Side Tube S i d e No. of Tubes per Shell

Fluid VAPOR C.W Tube S i z e , m m 14.85 'D X 19.05 0 0 . (BWG 14 )

Tube Length ₆₀₀

Tube Loyout Pitch _{25.4 mm}

Fluid Vop'd or Conds'd WATER Shell Oio. 8B

Flow Rote Kg/Hr 64 6800 Boffle Spocing 100

Density Kg/m 1000 Boffle Cut Hori. D Vert. H •40

Viscosity CP 1.0 Insulot'n, Shell/Chonn.. Hot/Cold

Sp. Heot Kcol/KgX 1.0 Pointing SANDING

T h e m . Cond. Kcol/mHrX

Latent Heot Kcol/Kg MATERIAL

Temp. In _•c 180 32 Shell SUS 304 Channel _{SUS 3 0 4}

Temp. Out 37 Chonn. Cov. SUS 3 0 4 Flating Heod SUS 3 0 4

Op. Press. K g / c m3 10 Backing Device Poss Port. SUS 304

No. ot Posses per Shell Tub* SUS 304 Tube Sh't SUS 304

Velocity m/sec Boffle SUS 304 Impingem't

Press. Drop. Kg/cm Tie Rod SUS 304 Bofl Spocer SUS 3 0 4

Design Temp. •c 300 100 Saddle SUS 304 Gosket

Kg/cm'-G

Design Press. 16 10 No22le Shell Chann. METAL

Press. Test. Hydro K g / c m ' - G ₁₀ 10 Shell side SUS 3 0 4 Flooting METAL Pneu. K g / c m2 - G 10 10 Chonn. side SUS 3 0 4 Shell nozz. METAL

Rodiogroph Flonge Chonn. nozz. PTFE

Con*. Allow. Shell nozz. SUS 3 0 4 Bolting

LMTD Corrdcted 145 X Chonn. side SUS 3 0 4 Shell Chonn. FC25C

Uo. Colc./Service 500 Kcol/W HrT Chann. end SUS 3 0 4 Flooting FC25C

CONNECTIONS Chonn. nozz. SUS 3 0 4 Shell nozz. FC25C

Mork Size Roting/Focing S e r v i c e Chonn. nozz. N - 1 11/2B 16K/RF VAPOR INLET

N - 2 1/2B 1 6 K / B W C COND. OUTLET

Tie Rood SUS 3 0 4 N - 3 IB 5K/SOC C.W INLET NOTE: BAFFLE WITH V-NOTCHING

N - 4 IB SK/SOC C.W OUTLET

N - 5 1/2B 16K/BWC VENT

N - 6 1/2B 1 6 K / B W C DRAIN

N - 7 1/2B 5K/SOC DRAIN ENG. DW5. NO. EQD-4102

SKFI.TON DRAWING

SHINSUNG CAM PLANT CO., LTD.

_{Sheet No. EDS-4001}

TANK & VESSEL DATA SHEET

Revision Dote Cheeked by Plant SONATA-IV / C H F C Client KAERI Item No. R-101

Project No. 61217 Service REACTOR Location TAEJON. KOREA No. Required ONE

Regulation Code. Type/Norm. Cop. VIR. / 906.

OPERATING CONDITIONS

Process fluid WATER

Specific gravity •c Viscosity C.P. @

•c

Heat Transfer N o D Heating, Cooling.D Ht. Trans. Device

Agitation Yes O Not Agitofn Device M e c h . p Uq. Circ. Q

Op. Temp.. Trim 180 'C Op. Press.. Trim 10 Kg/cm' —G. mmHg—A

Jkt/Coil "C Jkt/Coil Kg/em1 —G. mmHg—A

DESIGN CONDITIONS Design Temp. 300 *C

Design Press.. Trim 16 Kg/cm' - C Jkt/Coil Kg/cm' - G Press. Test.Hydro./Pneum. Radiograph Wind Load 10/10 Kg/cm' - G Stress Relief Seismic Coeff. Corr. Allow. CONSTRUCTION Shell ID/Length * 600 mm/ 620

Head type 10%-DishedD 2 : 1 Elip. D Support Saddle D LugQ Leg D

Insulation: (fig). Cold 100

Ladder Yes D No Plotform Yes D No fa* Insulation Ring Yes D No (2 Pointing SANDING

Weight: Empty Kg, Op Kg

MATERIAL Shell SUS 304 Head Support SUS 304 Flange SUS 304 SUS 304 Nozzle SUS 304 Gasket METAL Lining Spec. CONNECTIONS Mark N - 1 N-2 N-3 N - 4 N-5 N-6 N-7 Size 1/2B 11/2B 1/2B 1/2B 1/2B 1/2B 1/2B Roting/Facing 16K/BWC 16K/RF 16K/BWC 16K/BWC 16K/BWC 16K/BWC 16K/BWC Service

DEMI WATER INLET VAPOR OUTLET COND. INLET FT CONN. PSV CONN. LT CONN. LT CONN. SKELTON DRAWING iwmt ENG. DWG. NO. NOTE: EQO-4101

SWNSUNO CAM PLANT CO..LTD. EOO-410IB

R-1O1 REACTOR / SHELL DETAIL

soHAU-iv / oro

Preset H»

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SHINSUNG CAM PLANT CO..LTD. SmlK*. EQ0-4101D

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R-101 REACTOR / MOLD HEATER DETAIL

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20t COPPER (SEE OET.)

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EQUIPMENT LAYOUT

S-t/50 SCHATA-IV/CHfO Cllwt KAEBI *50 350 I 500 1500 200 CL+1580 TOP VIEW S-1/ftO • M i l ,

SHINSUNG CAM PLANT CO..LTD. LOT-30021 /

SUPPORT FRAME

S-1/20

Wont SOWU-IV / CWO HtvWcri

IOO1I Pro»ct H t 81217 1500 4-CHWNEL / (100x50x5) c-100x50x5 TOP VIEW C-75x40x5 OL+1280 4.5t CHECKED PLATE a+soo OH-0.00 500 C-100x50x5 200 600 . 1 . 400 ,

•"T~1

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SHINSUNO CAM PLANT CO..LTD. EQO-4102A

E-111 CONDENSER / T U B E SHEET DETAIL

SOWAU-IV / O f 0KAOtl

81217

8B FLANGE (16KO)

35tf»x XS^x 261

34-3/4B TUBE A P1TCH-25.4 50 400 (53) 100 too 100 100 JL uA. T SO 50

4

¥ ¥*

Spj.70 100 500 5050 50 100 ASSEMBLY NOTE

• TUBE CONNECTION TO TUBE SHEET TO BE SEAL WELDEO WITH EXPENDING • BAFFLE CUT 40X (VERTICAL) • BAFFEL WITH V-NOTCHINO • BHA: : BUTT WELD CONNECTOR

(HOKE : 8C8W8)

HEAT EXCHANGER DATA SHEET

NOZZLE TABLE M.K. N-1 N-2 N-3 N-5 N-6 N-7 SIZE I V i B 1/2B IB IB 1/2B V2B SERVICE VAPOR INLET COND. OUTLET C.W INLET C.W OUTLET VENT DRAIN DRAIN TYPE FLO BVI/C SOC. SOC. 8W/J BW/C SOC MAT'L RATING SUS304 SUS304 SGP SOP SUS304 SUSJ04 SOP 16kg 16kg 5kg 5kg 16kg 16kg 5kg NO77I F ORIFNTAT1ON COOE : KS B623O FLUIO ALLOCATION SHELL SIDE TUBE SIDE FLUID ALLOCATION SHEU SIDE TUBE SIDE

RADIOGRAPHED D FULL D SPOT [ NO JOINT EFFICIENCY INSULATION SIZE LENGTH 85X 500 NO. 34 METAL DESIGN CONDITIONS <T U BE ) PRESSURE 16 kq/cm3 TEMPERATURE: 300 OPERATION CONDITIONS (TUBE) PRESSURE 10 kg/cm TEMPERATURE: 200

COPE COPE STAMP

SHELL SUS304 8B (200») x 5C0L JOINT EFF TUBE S P E C . 85% SUS304, SEAMLESS 3/48 x BWS14 (2TU) JOINT EFF. CLAOO1N0 OR LINING S P E t _ ssx GASKET SPEC. SUPPORT SPEC. EXTERNAL PIPE S P E C . . INTERNAL PIPE SPEC._ R A N G E SPEC SUS METAL 2-SAODLES SUS304 RATINO- FACINO-16ko/fcm» SO/ttF PAINT INSULATION FIREPROOFINO. SHIPPING WT._ SUS SANDINO _OPERAT1NO TFST WT

ELEVATION MEASURED FROM BASELINE.

ORIENTATION MEASURED CLOCKWISE FROM 0 ' SUREO PROJECTION MEASURED FROM t OF VESSEL TO EXTREME FACE OF FLANGE. PROOF

STEEL SADDLE NOZZLE STANDARD TITLE STO-1021 STO-1003 0RAWINO NO. REFERENCE DRAWING

bd

/* CHFG Heater Configuration Simulation */ /* CASE 2 : Heater is in contact with 2/3 of the inner surface /* PURE CONDUCTION PROBLEM V

»CFXF3D »OPTIONS TWO DIMENSIONS HEAT TRANSFER USE DATABASE END » M O D E L TOPOLOGY »CREATE PATCH

PATCH NAME 'SHELL'

PATCH TYPE 'CONDUCTING SOLID' BLOCK NAME 'BLOCK-NUMBER-1' PATCH LOCATION 130 110 11 PATCH GROUP NUMBER 1 END

»CREATE PATCH PATCH NAME'Centerline'

PATCH TYPE 'SYMMETRY PLANE' BLOCK NAME 'BLOCK-NUMBER-1' PATCH LOCATION 30 30 1 10 1 1 HIGH I

END

»CREATE PATCH

PATCH NAME 'Outer Wall'

PATCH TYPE 'CONDUCTING BOUNDARY1

HIGHJ END

»CREATE PATCH

PATCH NAME 'Heated Inner Wall1

PATCH TYPE 'CONDUCTING BOUNDARY'

BLOCK NAME 'BLOCK-NUMBER-11

PATCH LOCATION 1130 11 11 LOWJ

END

»CREATE PATCH

PATCH NAME 'Naked Inner Wall'

PATCH TYPE 'CONDUCTING BOUNDARY' BLOCK NAME 'BLOCK-NUMBER-1' PATCH LOCATION 110 11 11 LOWJ END »MODEL DATA »MATERIALS DATABASE »SOURCE OF DATA PCP END »FLUID DATA FLUID 'AIR' MATERIAL TEMPERATURE 293.0 MATERIAL PHASE 'GAS'

END »TITLE

PROBLEM TITLE 'CHFG COPPER SHELL CONDUCTION PROBLEM' END

»PHYSICAL PROPERTIES

»SOLID HEAT TRANSFER PARAMETERS PATCH GROUP NUMBER 1

SCALAR CONDUCTIVITY 379.9 END

»SOLVER DATA

»PROGRAM CONTROL

MAXIMUM NUMBER OF ITERATIONS 25 OUTPUT MONITOR POINT 25 5 1

MASS SOURCE TOLERANCE 1.0E-6

ITERATIONS OF HYDRODYNAMIC EQUATIONS 0 END

» M O D E L BOUNDARY CONDITIONS

»CONDUCTING BOUNDARY CONDITIONS PATCH NAME 'Outer Wall1

TEMPERATURE 373.4 END

»CONDUCTING BOUNDARY CONDITIONS PATCH NAME 'Heated Inner Wall1

HEAT FLUX 105000. END

»CONDUCTING BOUNDARY CONDITIONS PATCH NAME 'Naked Inner Wall'

HEAT FLUX 0. END

»OUTPUT OPTIONS » D U M P FILE FORMAT

FORMATTED

NUMBER OF SIGNIFICANT FIGURES 8 END

1/3°1 dryout

/* CHFG Effect of Dryout Patch Simulation */ /* CASE 2:1/3 Outer Surface is Dried Off */ /* PURE CONDUCTION PROBLEM */

»CFXF3D »OPTIONS TWO DIMENSIONS HEAT TRANSFER USE DATABASE END »MODEL TOPOLOGY »CREATE PATCH

PATCH NAME 'SHELL'

PATCH TYPE 'CONDUCTING SOLID' BLOCK NAME 'BLOCK-NUMBER-1' PATCH LOCATION 1 30 1 10 1 1 PATCH GROUP NUMBER 1 END

»CREATE PATCH PATCH NAME 'Centerline'

PATCH TYPE 'SYMMETRY PLANE'

BLOCK NAME 'BLOCK-NUMBER-11

PATCH LOCATION 30 30 1 10 1 1 HIGH I

END

»CREATE PATCH

PATCH NAME 'Heated Inner Wall'

PATCH TYPE 'CONDUCTING BOUNDARY' BLOCK NAME 'BLOCK-NUMBER-1'

LOWJ END

»CREATE PATCH

PATCH NAME 'Dry Outer Wall'

PATCH TYPE 'CONDUCTING BOUNDARY' BLOCK NAME 'BLOCK-NUMBER-1'

PATCH LOCATION 1 10 10 10 11 HIGHJ

END

»CREATE PATCH

PATCH NAME Wet Outer Wall1

PATCH TYPE 'CONDUCTING BOUNDARY' BLOCK NAME 'BLOCK-NUMBER-1' PATCH LOCATION 11 30 10 10 11 HIGHJ END » M O D E L DATA »MATERIALS DATABASE »SOURCE OF DATA PCP END »FLUID DATA FLUID 'AIR' MATERIAL TEMPERATURE 293.0 MATERIAL PHASE 'GAS'

END » T I T L E

PROBLEM TITLE 'CHFG COPPER SHELL CONDUCTION PROBLEM' END

»PHYSICAL PROPERTIES

»SOLID HEAT TRANSFER PARAMETERS PATCH GROUP NUMBER 1

SCALAR CONDUCTIVITY 379.9 END

»SOLVER DATA

»PROGRAM CONTROL

MAXIMUM NUMBER OF ITERATIONS 25 OUTPUT MONITOR POINT 25 5 1

MASS SOURCE TOLERANCE 1.0E-6

ITERATIONS OF HYDRODYNAMIC EQUATIONS 0 END

»MODEL BOUNDARY CONDITIONS

»CONDUCTING BOUNDARY CONDITIONS PATCH NAME 'Heated Inner Wall'

HEAT FLUX 70000. END

»CONDUCTING BOUNDARY CONDITIONS PATCH NAME 'Dry Outer Wall'

HEAT FLUX 0. END

»CONDUCTING BOUNDARY CONDITIONS PATCH NAME 'Wet Outer Wall'

TEMPERATURE 373.4 END

»OUTPUT OPTIONS » D U M P FILE FORMAT

FORMATTED

NUMBER OF SIGNIFICANT FIGURES 8 END

Heat and Mass Transfer 34 (1998) 321-328 © Springer-Verlag 1998

Thermal-hydraulic Phenomena Relevant to Global Dryout

in a Hemispherical Narrow Gap

J. H. Jeong, R. J. Park, S. B. Kim

Abstract A series of experimental investigations on the cooling mechanism in hemispherical narrow gaps has been carried out A visualization experiment, VISU-II, was done as the first step to get a visual observation of the flow behaviour inside a hemispherical gap and to understand the mechanism inducing global dryout. It was observed that the counter-current flow limitation (CCFL) phenom-enon prevented water from wetting the heater surface and induced dryout The CHFG test was performed to measure the critical power corresponding to global dryout and to investigate the inherent cooling mechanism in hemi-spherical narrow gaps. Temperature measurements over the heater surface show that the two-phase flow behaviour inside the gaps could be quite different from the other usual CHF experiments. The measured values of critical power are lower than the predictions by existing empirical CHF correlations based on the data measured with small-scale horizontal plates and vertical annulus.

1

Introduction

During the TMI-2 accident, the reactor pressure vessel (RPV) received no damage and molten corium was kept inside the pressure vessel and cooled down, despite the fact that all severe accident analysis codes predicted it would fail. It means that there might be inherent cooling mechanisms that are not known. In order to explain the safe cool-down of the relocated corium, three cooling mechanisms have been suggested and a gap-cooling mechanism is considered to be the most plausible one that played the major role in corium cooling [1, 2]. Such a mechanism works like this: When molten corium relocates to the bottom of a pressure vessel filled with water, it shapes like a hemispherical pool and its surface forms crust As the corium is solidified it shrinks and the pres-sure vessel experiences creep due to high temperature and pressure, causing a gap to develop between the crust and the pressure vessel. The water penetrates the gap and the corium is cooled. In order for gap cooling to be effective,

Received on 14 April 1998

J. H. Jeong, R. J. Park, S. B. Kim Severe Accident Research Lab

Korea Atomic Energy Research Institute

the gap size should be large enough and water should continue to be supplied through the gap so that the boiling heat transfer is maintained. The maximum power of a heat source removable through boiling is the critical power. At the critical power, the whole of heated surface will be dried off. Therefore, a study on a critical power in hemispherical narrow gaps is needed to assess the gap cooling mecha-nism.

If it is realistically possible to retain the molten corium inside the lower plenum of the RPV, it is very beneficial to the nuclear power plant designer. It is because they do not need to consider the problems that may occur if the molten corium penetrates the RPV, like a direct contain-ment heating, an explosive interaction between molten fuel and water and a concrete-melt interaction in the cavity. Such a situation is also desirable in terms of public ac-ceptance. Therefore, the gap-cooling concept became a hot issue in the field of severe accident research to achieve invessel retention of molten corium. As the gap-cooling concept was considered to be of importance recently, there has not been many researches related to the concept. Some of them have focused on gap formation [3, 4, 5] and some upon the heat transfer through the gap [6, 7].

Some CHF studies have been carried out in various gap geometries. Sudo and Kaminaga [8] and Katto and Kosho

[9] performed CHF experiments in gaps of rectangular channels and horizontal plates, respectively. Chang & Yao [10] carried out CHF experiments with test sections of vertical annulus at atmospheric pressure and developed the following correlations:

0.38

(1)

where, <JCHF> g> D> Pi» Pg» hg> L and s are the critical heat

flux, gravitational acceleration, diameter, liquid density, gas density, latent heat of evaporation, heated length and gap size, respectively. Monde et aL [11] carried out an experimental study of the CHF at atmospheric pressure in vertical rectangular channels. Three different heating surfaces (20, 35, 50 mm in length and 10 mm in width) were used and the gap size varied from 0.45 to 7.0 mm. They reported the following empirical correlation:

ICHF

Pg?s

0.16

capacity through gaps and this equation is currently used in the MAAP4 code [12].

Up to now, however, there has been no experiment concerning critical power and few studies on the two-phase flow behaviour in hemispherical gaps. An experi-mental investigation of which geometry is similar with a hemisphere was done by Kohler et al. [7]. The test section of their experimental facility is a scaled RPV and solidified corium of the TMI-2 plant That is, the curvature of the lower plenum surface is shaped such that the angle of inclination of the heating surface at the outer edge is the same as the angle of inclination exhibited by the RPV lower head of TMI-2 at the edge of the solidified debris crust. With the measurements, they reported that a gap size of 1 mm is capable of transferring the decay heat produced during the TMI-2 accident. In the TMI-2 acci-dent, around 19 tons of material relocated to the lower head, which is around 20% of the core materials (the core contained 93.1 tons of fuel). So, in order to assess the cooling capacity through the gap for a more general situ-ation, it is necessary to do more experimental and ana-lytical investigations in addition to Kohler et al.'s [7] work which is case-specific to the TMI-2 accident.

A series of experimental investigations of the cooling mechanism in hemispherical narrow gaps, focusing on a critical power, have been carried out. The visualization experiment, VISU-II, and the first stage of the CHFG tests have been completed. The later one has been conducted with water. The VISU-II test aims to get a visual obser-vation of the two-phase flow behaviour inside a hemi-spherical gap and to understand the mechanism inducing global dryout. The purpose of the CHFG experiments is to investigate the inherent cooling mechanism in a hemi-spherical narrow gap and find out critical power. The name of the tests, CHFG, implies that the phenomena we are interested in are related to the critical heat flux. However, the term of critical power will be used instead of CHF in the present work and the reason will be given later in this paper.

Visualization Experiment: VISU-II

Hydrodynamic instability was proposed by Kutateladze [13] as a CHF mechanism and his theory is generally ac-cepted. Therefore, understanding the hydrodynamic phe-nomena of boiling water in hemispherical narrow gaps is important in the study of critical power. Concerning the physical phenomena associated with two-phase flow inside a gap, Kohler et al. [7] thought that an oscillating water/ stream flow performs the necessary heat removal function between the debris crust and the RPV wall. However the two-phase flow behaviour is not still well understood. We carried out experiments aiming to visualize boiling water inside a hemispherical gap and see the hydrodynamic phenomena triggering dryout. In order to provide visual observations, water and a hemispherical heater were placed in a transparent pyrex-glass vessel.

video camera to visualize the flow inside the gap. We in-tended to make a 1 mm gap between the heater and pyrex-glass vessel itself. However, there exists some non-uni-formity due to the difficulty in machining the pyrex-glass vessel. The top of the pyrex-glass vessel is open to the atmosphere. Figure 1 shows a cross-section detailing the test section, including the heater. An electric heater wire is located inside a hemispherical copper shell and the shell is filled with Wood's metal of melting point 70 °C. The thickness and outer diameter of the copper shell are 20 mm and 238 mm, respectively. The maximum heater power is 6 kW. Below the test section, a mirror slanting at 45° is installed to provide a visual observation of the test section bottom area.

The steam bubbles generated go upward alongside the hemispherical heater walL At the same time, water goes down in the counter direction with the bubbles. The two phases flow violently in the gap and this flow pattern prevails from the bottom to the top end of the gap. In the vicinity of the top end, steam tries to penetrate into the water pool above the heater while water flows into the gap. They flow in counter directions through separated flow paths. These flow paths are randomly established and disappear quickly. Figure 2a shows the multiple flow paths, the steam flow path is about 2 cm wide. At a certain elevated heater power, the mass velocity of steam reaches a critical value corresponding to the counter-current flow limitation (CCFL). Jeong & No [14] named this type of CCFL as entrance flooding because the CCFL is initiated due to flow instability at the liquid entrance. That type of CCFL occurs when the liquid entrance geometry is sharp and the liquid flow rate is large enough. In this case, hydrodynamic and geometric conditions around the top end of the gap influence the CCFL significantly but those conditions of the gap below the top end do not. The CCFL is also affected by the gap size. In our facility, CCFL first occurs at some part of the top end where the gap size is small compared with the other part. According to obser-vations, steam and water still flow actively through mul-tiple paths in regions of larger gap sizes while CCFL occurs in regions of smaller gap sizes. The heater surface just

Pyrex glass Power line / ^ f / C lead wire Water Heater-

Copper-Top area

b Bottom area

Fig. 2a,b. Images of flow in the gap. a Top area; b bottom area

below the region where CCFL occurs is locally dried out because CCFL prevents water from penetrating the gap. With the heater power that initiated CCFL, the local dryout region was small and often re-wetted by the water coming up from the bottom. Since the test section was not so big and the boiling two-phase flow fluctuated dynamically, water was able to reach the dryout region. With the further increase in heater power, however, the dryout region was enlarged and water could not reach it any more.

Figure 2b shows the dryout region when the heater power is 5.5 kW. The left-hand-side is filled with water while the right-hand-side is dried out.

CHFG Experimental Facility

The VISU-II test gave us some phenomenological under-standings. However, it was not possible to perform boiling experiments with the facility at elevated pressures since the vessel is made of fragile glass and the heater does not provide uniform heat flux. So we designed a CHFG facility

1500 100 200 _ T _ 20 250 + Gap Drain j - X — ;—T Level guage jlnsulatiorj P : pressure T : temperature Unit: mm copper shell gap pressure vessel

Fig. 3. Schematic Diagram of CHFG facility

an electric heater, a pressure vessel, a heat exchanger and a coolant control system. An electric heater is put inside a hemispherical copper shell, which provides the maximum average heat flux of 90 kW/m2 at the outer surface. The

thickness and outer diameter of the copper shell are 25 and 498 mm, respectively. Four units of stainless steel pressure vessel were manufactured to provide gap sizes of 0.5, 1.0, 2.0 and 5.0 mm between the copper shell and the pressure vessel itself. The experiments were performed using de-mineralized water. The measurements of critical power were made in the range of 1 to 10 atm. The heat generated by the electric heater is removed in a heat exchanger installed 150 cm above the top of the pressure vessel to maintain a near-saturated condition of the working fluid. The heat exchanger takes a role in system pressure regu-lation as well. A level gauge is installed in the pressure vessel to confirm that the heater is always covered with water during the experiments. The occurrence of dryout is noticed by 66 K-type thermocouple readings. The ther-mocouples are embedded in the copper shell, as shown in Fig. 4. From each pair of T/Cs, local heat flux is calculated and found to be within a ±20% variation of average in a nucleate boiling regime. The temperatures and mass flow rates are processed by a Hewlett Packard data acquisition system and HP-VEE program.

As the experimental facility constitutes a closed loop, the first step necessary to carry out experiments is to purge the air accumulated in the loop. If the air remains in the loop, it obstructs the heat transfer in the heat exchanger so that the working fluid might not circulate. Initially the heater power is maintained at a low level and the set-value of the pressure control system is set at a pre-determined

45

outside (5-8,21-24) outside _{Fig. 4. Thermocouple locations}

temperature readings are believed to reach a quasi-steady state, the heater power is increased step-wisely. When all the temperature readings increase monotonically without a limit, the heater power is cut off. Usually it took 15 to 30 minutes to reach a quasi-steady state in a low power range and more than 60 minutes near the critical power.

Results and discussions

Figures 5 and 6 show the temperature measurements made at a gap size of 2 mm and atmospheric pressure. Those are projections of the hemispherical surface of the copper shell. The boundary and the center of the circular area refer to top end of the gap and the lowest bottom of the copper shell, respectively. Small circles shown in radial directions represent thermocouple locations. The readings from those thermocouples are interpolated to give iso-thermal lines. Figures 5a-d show temperature variation measured at heat fluxes of 32, 42, 52 and 60 kW/m2,

re-spectively. Through the present paper, heat flux refers to the average heat flux over the whole outer surface of the copper shell. Temperatures of the heater and copper shell reached a quasi-steady state after 15 ~ 30 minutes since the heater power changed. Quasi-steady state values were used to make these plots. The surface temperature was

the azimuthal and downward directions with an increase in heat flux. Considering the visual observations of VISU-II experiments, this indicates that a local dryout region develops and expands with an increase in heat flux. At a quasi-steady state corresponding to each heat flux, the local dryout region continues to exist with just a little movement of its boundary. In the VISU-II experiment, it was observed that the interface between wet and dryout region fluctuated in a distance of several centimeters while the majority area of dryout region remained out of contact with water. A similar situation was observed in the CHFG test The temperature readings started to fluctuate in a range of 5 °C just before they jumped up from nearly saturated values to much higher values. The fluctuation lasted for a while but no more fluctuation was observed after the temperature reached a high value. During a transient period caused by an increase in heater power, the dryout region expands and its front passes the locations where thermocouples are embedded. When the front passes a thermocouple, it is believed, the dryout front fluctuation causes the temperature fluctuation. The dis-tinguishably higher temperature region, than the saturated value, is regarded as a dryout region in the CHFG test even though the temperature does not exceed the minimum film boiling temperature for a pool boiling condition. Whalley

Fig. 5a-d. Temperature variation. Gap size: 2 mm; Avg. q" (kW/m2); a 32. b 42;

below 290 °C. The temperature of the surface remained below 290 °C until the whole surface dried off as can be seen in Fig. 6. The conduction through the copper shell and the heater itself of the CHFG rig seems to cause the temperature of such a stable dryout region to remain be-low 290 °C. That is, the remaining heat that was not re-moved in the local dryout region, moves to the wetted region by conduction to be removed. Another noticeable point is the non-contact of water with the dryout region. The temperature of the isotherm line-crowded region (dryout region) corresponds to a transient boiling regime of a pool boiling curve. It is believed however that water is never in contact with the surface of the dryout region while it is in a quasi-steady state because there was no sign of water contact like a sudden temperature decrease or fluctuation. It seems that the hydrodynamic characteristics in a hemispherical narrow gap, which are different from those of a pool boiling case, caused the non-contact at the temperature corresponding to transient boiling regime of a pool boiling case.

The high temperature region always started from the upper left edge. The reason for this is speculated to be because the gap size around this area is smaller compared with the other region. Although the curved components of the facility were machined using a CNC machine, the gap size can be slightly non-uniform because of miss-align-ment or thermal expansion. The measuremiss-align-ments of actual gap sizes by the ultra-sonic technique (UT) under the same condition as the experiments showed that the devi-ation in gap sizes from the design value was within the instrument's inherent error. It can therefore be said that the gap size was reasonably uniform.

Figure 6 shows the temperature variation with time in seconds when the heat flux was fixed at 68 kW/m2.

Tem-peratures over the whole surface increase by itself, even though the heater power is maintained at a fixed level. This means there is no steady state at this heat flux, which is a different situation from that shown in Fig. 5. In the lower part of Fig. 6a, there is a large area without isotherm lines. The temperature of this region remains slightly higher than 100 °C and the area keeps shrinking with time. That is, the wetted region shrinks and the dryout region ex-pands with time. The velocity of the dryout expansion and temperature increase rate of the dryout region get larger with time. This is because the local heat flux at the wetted region increases due to extra heat transferred from the dryout region by conduction. When the dryout region expands to the bottom of the copper shell, the temperature increase of that location is so fast that the heater power should cut off immediately for higher protection. We de-fined this heat flux with which the dryout region under-goes self-expansion as the critical power in the geometry of hemispherical narrow gaps.

The present definition of critical power is different from the usual definition of CHF used in other experiments. If experiments are carried out with a specimen like a thin plate, pipe and wire, the temperature jumps quickly as

enough, such as the present facility (copper 200 kg), the temperature increases slowly because it needs a lot of heat to be heated up. Even in the local dryout region, the temperature increases slowly and it is limited. This is be-cause the remaining heat which was not removed in the local dryout region, moves to the wetted region to be re-moved. Some experimental results related to such a heater mass (thickness) effect on the maximum pool boiling heat flux are available in the literature. Peyayopanakul and Westwater [16] reported that for test sections of copper cylinders >25 mm in thickness, maximum heat flux was independent of wall thickness and the pool boiling curves were all quasi-steady state and equivalent to those ob-tained in steady-state tests. Depending on the thickness, the quenching time varied from 7 hr to 2 min in their experiments. El-Genk and Glebov [17] employed down-ward-facing curved copper sections to test the effect of test section thickness. Their results showed that the maximum boiling heat fluxes were independent of wall thickness

>19 mm. In the present facility, the thickness of copper shell only is 25 mm. Considering the fact that a heater block is situated inside the copper shell, the present heated section is believed to be thick enough. The reason that the occurrence of local dryout is not chosen as a critical value like a usual CHF in the present experiments is as follows: (i) The present experimental facility does not allow the visual observation of the flow pattern, and it is therefore hard to tell whether the flow is in a stable film boiling regime only with temperature readings, (ii) The occur-rence of local dryout is induced by CCFL phenomena, which is highly dependent on the geometry of the test section. Therefore global parameters rather than local ones seem to be better to characterize the boiling behaviour, (iii) Even though local dryout occurs, all the generated heat finally cools down due to conduction, as mentioned before. Therefore the temperature does not increase mo-notonously but is limited by a certain value. However, above a certain heat flux, temperature increases continue. It is therefore reasonable in the present geometry that critical power should be defined as the heat flux which exceeds the maximum cooling capacity through a hemi-spherical gap so the dryout region undergoes self-expan-sion and leads to a global dryout

Figure 7 shows the temperature variation at critical power under atmospheric pressure and the gap size is 1 mm. The global dryout occurred at the heat flux of 60.3 kW/m2, which is lower than that measured with a gap

size of 2 mm. The local dryout started to develop at the same location as the case where the gap size is 2 mm. When the gap size was 1 mm, the dryout region expands faster in the azimuthal than downward direction. As can be seen in Fig. 7c, the bottom surface is still wet although the top end of the gap is dried off. In Fig. 7d, the dryout region finally expands to the bottom and the temperature over the whole surface increases to over 200 °C.

The predictions by Chang & Yao [10] and Monde et al. [ll]'s CHF correlations, Eqs. (1) and (2), are compared

Fig. 7a-d. Self-expansion of dryout. Gap size: 1 mm; Elapsed time at q" = 60.3 kW/ m2. a 0. b 1000. c 1500. d 2500

compared. They carried out CCFL experiments in narrow-gap annular passages and presented their experimental data. The inner diameter of the outer pipe of the facility was 100 mm. Various sizes of the inner pipe were used to make the gaps of 0.5, 1.0, 2.0 and 5.0 mm. Since they did not suggest any CCFL correlation, we did some regression analysis to develop the following CCFL correlations: ; '1 / 2 + 0.23;*1/2 = 0.32 for 2 mm gap jl1/2 + 0.35;f1/2 = 0.35 for 1 mm gap where, jl = • (3) (4) 250, 200 \ " I 150 X 100 o 50 2 mm 1 mm 0.5 mm Chang & Yao Monde et al. Koizumi(1 mm) Koizumi(2 mm)

20 40 60 80 100 120

In order to present them in Fig. 8, superficial velocities are changed into corresponding heat flux that can produce the same mass flow of steam. Koizumi et al.'s CCFL cor-relation seems to be close to the present measurements in terms of value and pressure trend, while Chang & Yao and Monde et al.'s correlations predict much higher figures. The reason is thought to be that, in the present experi-ments, CCFL prevents water from penetrating into the top end of the gap and global dryout occurred at lower heat fluxes. Compared with those two empirical CHF correla-tions, the pressure effect of the present results seems to be quite small. As those two CHF correlations were developed based on the data measured under atmospheric pressure, the pressure trend predicted by them might not be correct Especially, Monde et al. [ll]'s correlation shows monot-onous and rapid increase in CHF with system pressure. It does not seem to be appropriate to use Eq. (2) at an ele-vated pressure. The gap size effect on critical power in hemispherical narrow gaps is also shown in Fig. 8. The increase in gap size from 0.5 mm to 1 mm almost doubles the critical power but that from 1 mm to 2 mm affects critical power just a little. The measurements with the gap size of 2 mm were only made at atmospheric pressure because of the lack of heater capacity. Further works are needed to quantify the effect of gap size on the critical power.