NUREG/CR-6623

Vapor Explosions in a 

One-Dimensional 

Large-Scale Geometry 

With Simulant Melts 

University of Wisconsin -Madison 

U.S. Nuclear Regulatory Commission 

Office of Nuclear Regulatory Research 

Washington, DC 20555-0001
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11
NUREG/CR-6623

Vapor Explosions in a 

One-Dimensional 

Large-Scale Geometry 

With Simulant Melts 

Manuscript Completed: October 1999 

Date Published: October 1999 

Prepared by 

It-S. Park, R. Chapman, M.L Corradini 

University of Wisconsn - Madison 

1500 Engineering Drive 

Madison, WI 53706 

S. Basu, NRC Project Manager 

Prepared for 

Division of Safety Analysis and Regulatory Effectiveness 

Office of Nuclear Regulatory Research 

U.S. Nuclear Regulatory Commission 

Washington, DC 20555-0001 

NRC Job Code W6183
Abstract

In light water reactors after a prolonged lack of cooling, vapor explosions could occur when molten fuel is generated and contacts residual water coolant within the reactor vessel 

or below in the containment reactor cavity. The experimental objectives for this work 

were to obtain well-characterized data for the explosion propagation/escalation phases, while systematically investigating the effect of a comprehensive set of initial and boundary conditions on the explosion energetics; i.e., trigger strength, fuel mass, composition and 

temperature, coolant mass, viscosity and temperature and system constraint. First, a vapor explosion apparatus, WFCI, was developed which allowed for well-characterized explosion data and demonstrated reproducible explosions with a simulant fuel. Second, the explosion energetics was examined as a function of varying initial and boundary conditions for this 

tin simulant. Finally, the simulant fuel was changed to a molten iron-oxide, which was more 

prototypic of the actual molten fuel compositions and explosion energetics were reexamined 

and FCI were found to be quite weak. 

With respect to reactor safety issues, this experimental work has quite important safety implications. First, this work has provided clear evidence of the reproducibility of vapor explosion energetics for a controlled set of initial and boundary conditions. This suggests empirically that this phenomenon is predictable if one can establish and control the initial 

and boundary conditions. Second, the experiments demonstrate that geometric scaling can 

be properly specified; e.g., a rigid radial constraint for one-dimensional tests is conservative 

for energetics when compared to full-scale, while the axial constraint scale factor from test 

to prototype needs to be unity to preserve energetics. Finally and most importantly, the data suggests that once the fuel-coolant initial conditions are within an envelope for 

triggered events, the energetics is much less than thermodynamic, apparently due to the 

small amount of fuel that participates in an explosion timescale. And this envelope of 

triggerability is much smaller for a simulant molten oxide with low superheat, such as 

molten iron-oxide in our tests (corium in the KROTOS tests). This suggests that material 

scaling for reactor safety issues must preserve the same fuel composition and superheat from the test to the prototype. 

The current work has limited data at larger scales with more prototypic molten oxides. It is 

recommended that further tests could be carried out under these conditions to empirically verify our findings. Models developed from our analysis can also be used to analyze these 

experiments. Finally, it is known that the mixing conditions determine the envelope of 

explosivity for the vapor explosion. Thus, it is of fundamental interest to better measure 

the mixture local conditions just prior to the explosion to correlate with the explosion energetics; i.e., void fraction profiles, fuel volume fractions and mixing diameters. Our 

future work in vapor explosion research is specifically targeted toward this purpose. 

iii NUREG/CR-6623
Contents

1 Introduction 

1.1 Background and Motivation .............. 

1.2 Objectives and Scope ............... . . 

2 Literature Review 

2.1 Perspectives on Vapor Explosions .......... 

2.2 Past Vapor Explosion Studies ............. 

2.2.1 Large Scale Experiments ........... 

2.2.2 Experiments using Tin as a Fuel ...... 

2.2.3 Experiments using Iron-Oxide as a Fuel. 

2.2.4 Related Experimental Parameters .... 

2.3 Key Findings of the Experimental Review . . . 

3 Experimental Description 

3.1 Introduction ......................... 

3.2 Experimental Apparatus ................ 

3.2.1 Heating and Transport System ...... 

3.2.2 Test Section .................... 

3.3 Instrumentation and Data Acquisition ...... 

3.3.1 Instrumentation ................. 

3.3.2 Data Acquisition and Experimental Conti 

3.4 Initial and Boundary Conditions .......... 

3.5 General Experimental Procedures .......... 

3.5.1 WFCI-A, B, C, E, F Series ......... 

3.5.2 WFCI-D Series .................. 

3.5.3 WFCI-G Series .................. 

3.5.4 WFCI-H Series .................. 

3.5.5 WFCI-K Series ..................

•o1

4 Experimental Results and Discussion 

4.1 Introduction ................... 

4.2 Trigger Characterization Experiments. 

4.3 Conversion Ratio ............... 

4.4 Reproducibility Tests: WFCI-A Series

NUREG/CR-6623

1 

•.. • • • • • • •1 I 

. . ... .... . . . . .. 3

6 

6 

8 

8 

14 

16 

17 

19 

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47 

48 

51 

51 

53 

54 

54 

54 

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56 

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72 

73 

75

. . . . . . 

. . . . . . 

. . . . . . 

. . . . . . . . .

V
4.5 

4.6 

4.7 

4.8 

4.9 

4.10 

4.11

Spontaneous Explosion Tests: WFCI-B Series ..................... 78 

Effect of External Trigger: WFCI-C Series ....................... 79 

Effect of System Constraint: WFCI-D Series ..................... 82 

Temperature Effects of the Fuel and the Coolant: WFCI-E and F Series.. 87 

Coolant Additives: WFCI-G Series ............................ 88 

Mass Ratio Effect: WFCI-H Series ............................ 92 

Iron-Oxide Fuel Composition: WFCI-K Series ..................... 94

5 Analysis of Experimental Results 

5.1 Temperature Effects .................................... 

5.2 Propagation Speed Characteristics ........................... 

5.3 Qualitative Debris Analysis ............................... 

5.4 Thermodynamic Analysis ................................. 

5.5 Energetics of Iron-Oxide Melt ............................. 

5.6 Possible Explosion Mechanism ............................. 

5.7 Vapor Explosion Scaling for Reactor Safety Issues ................ 

6 Conclusions and Recommendations 

7 References 

Appendix A Experimental Data for the Debris 

Appendix B Melt Preparation and Delivery

NUREG/CR-6623

163 

163 

164 

165 

167 

169 

172 

174 

183 

1817 

199 

203

vi

11
List of Figures 

1.1 Typical Geometries of Fuel-Coolant Interactions .................... 5 

2.1 Schematic Illustration of the COREXIT Facility [28] ..... ............ 35 

2.2 Schematic Illustration of the MFTF Facility ...................... 36 

2.3 Schematic Illustration of the Open Geometry Facility [40] .... ......... 37 

2.4 Schematic Illustration of the EXO-FITS Facility [40] ..... ............ 38 

2.5 Schematic Illustration of the FITS Facility [40] ..... ............... 39 

2.6 Schematic Illustration of the ALPHA Facility [43,44,45]. .............. 40 

2.7 Schematic Illustration of the FARO Facility [48,49] ..... ............. 41 

2.8 Schematic Illustration of the KROTOS Facility [47,48,63] .... ......... 42 

2.9 Temperature Interaction Zone for 12 g of tin dropped through 3 cm into 

boiled distilled water . . .......... .......... ................ 43 

2.10 Schematic Illustration of the Hall's Facility [60] ..... ............... 44 

2.11 Schematic Illustration of the Baines' Facility [62] ...... .............. 45 

3.1 Overall Side View of the WFCI Facility ....... ................... 58 

3.2 Overall Three Dimensional View of the WFCI Facility ..... .......... .59 

3.3 Schematic Illustration of the Front View of the WFCI Facility Including the 

Simplified Waterloop and the Location of the Pressure Transducer ..... .. 60 

3.4 Schematic Illustration of the Furnace ...... ..................... 61 

3.5 Schematic Illustration of the Transport System ...................... 62 

3.6 Schematic Illustration of the Slide Gate and Supporting Structure ..... .. 63 

3.7 Schematic Illustration of the Magnet Trigger System ..... ............ 64 

3.8 Schematic Illustration of the Trigger System Circuitry ................ 65 

3.9 Schematic Illustration of the Level Swell Meter ...................... 66 

3.10 Typical Piston Signal Measured in the Expansion Tube ............... 67 

3.11 Overall Layout of the WFCI Facility in a Top View .................. 68 

3.12 Schematic Diagram of the Control System ....... .................. 69 

3.13 Viscosity Ratio with respect to the Polymer Concentration ............. 70 

4.1 Sequential Events of the Typical WFCI Tests ...................... 103 

4.2 Sequential Events of the WFCI-A-01 Test ..... .................. 104 

4.3 Sequential Events of the WFCI-H Series Tests ...................... 105 

4.4 Peak Trigger Pressures vs. Supplied Voltages ...................... 106 

4.5 External Trigger Pressures Generated by a Supplied Voltage of 200 V . . . 107

NUREG/CR-6623 vii
4.6 Average Time Delay of the External Trigger vs. Supplied Voltages ...... 

4.7 Product of the Peak Pressure and Impulse at the First Pressure Transducer 

with respect to Supplied Voltages ...........................

4.8 Measurement of the Work done by Energetic 

Facility.......................

4.9 Pressure 

4.10 Pressure 

4.11 Pressure 

4.12 Pressure 

4.13 Pressure 

4.14 Pressure 

4.15 Pressure 

4.16 Pressure 

External 

4.17 Pressure 

External 

4.18 Pressure 

External 

4.19 Pressure 

External

FCIs on the Slug iin WFCI

Traces of the KROTOS-21 Test [66] ................. 

Traces of the WFCI-A-01 Test .................... 

Traces of the WFCI-A-02 Test .................... 

Traces of the WFCI-A-04 Test ........................ 

Traces of the WFCI-A-06 Test ........................ 

Traces of the WFCI-B-02 Test ........................ 

Traces of the WFCI-B-03 Tests ......................

Traces of the 

Trigger of 400 

Traces of the 

Trigger of 450 

Traces of the 

Trigger of 300 

Traces of the 

Trigger of 100

WFCI-C-03 Test with a Supplied Volts 

V .......................... 

WFCI-C-04 Test with a Supplied Volta 

V .......................... 

WFCI-C-06 Test with a Supplied Volta 

V .......................... 

WFCI-C-05 Test with a Supplied Volta 

V ..........................

Lge for the

108 

109

110 

111 

112 

113 

114 

115 

116 

117

....... 118

tge for the 

tge for the 

tge for the

4.20 Conversion Ratios with respect to the Peak External Trigger Pressures . .

4.21 Pressure Traces of the WFCI-E-02 Test with the Fuel and Coolant

atures of 790 and 86 *C, respectively . 

4.22 Pressure Traces of the WFCI-E-03 Test 

atures of 491 and 93 °C, respectively . 

4.23 Pressure Traces of the WFCI-F-02 Test 

atures of 782 and 27 °C, respectively . 

4.24 Pressure Traces of the WFCI-F-03 Test 

atures of 850 and 62.5 *C, respectively 

4.25 Pressure Traces of the WFCI-F-04 Test 

atures of 871 and 72 'C, respectively . 

4.26 Time Scales with respect to the System

with the Fuel and Cool.an 

with the Fuel and Coolant 

...th th Fe .. and Cool 

with the Fuel and Coolant 

with the Fuel and Coolant 

with the Fuel and Coolant 

Constraint . .. .. .. .

Temper

Temper

Temper

Temper

Temper4.27 Pressure Traces of the WFCI-D-04 Test with a Total Constraint Mass of 

56.8 kg, < m >=6.38 ............................. ...... 

4.28 Conversion Ratios and Atl.9 with respect to Axial Constraint ........ 

4.29 Conversion Ratios and Peak Pressures with respect to Axial Constraint . . 

4.30 Conversion Ratios with respect to the System Constraint calculated by the 

Parametric Model (UWFCI Code) ................... ....... 

4.31 Conversion Ratios and Pressure Histories calculated by the Parametric Model 

(UWFCI Code) with respect to < m > of 3.3, 11.2 and 44.4 .... .....

NUREG/CR-6623

119 

120 

121 

122 

123 

124 

125 

126 

127 

128 

129 

130 

131 

132 

133

viii

1!
4.32 

4.33 

4.34 

4.35 

4.36 

4.37 

4.38 

4.39 

4.40 

4.41 

4.42 

4.43 

4.44 

4.45 

4.46 

4.47 

4.48 

4.49 

4.50 

4.51 

4.52 

4.53 

4.54 

4.55 

4.56 

4.57 

4.58 

4.59

WFCI-1 trigger peak propagation curve ................... 

WFCI-2 trigger peak propagation curve ................... 

WFCI-K-01 transient pressure histories ................... 

WFCI-K-02 transient pressure histories ................... 

WFCI-K-03 transient pressure histories ................... 

WFCI-K-04 transient pressure histories ................... 

WFCI-K-05 transient pressure histories ................... 

WFCI-K-07 transient pressure histories ................... 

WFCI-K-08 transient pressure histories ................... 

WFCI-K-09 transient pressure histories ................... 

WFCI-K-10 transient pressure histories ................... 

WFCI-K-01 pressure peak propagation curve ............... 

WFCI-K-09 pressure peak propagation curve ............... 

WFCI-K-05 slug position signal ........................

...... 148 

...... 149 

...... 150 

. ..... .. 151 

...... 152 

.. ..... 153 

. ..... .. 154 

...... 155 

...... 156 

...... 157 

.. ..... 158 

...... 159 

.. ..... 160 

.. ..... 161

4.60 WFCI-K-05 slug position as a function of time with polynomial curve fit . 162 

5.1 Temperature Interaction Zone in the Large Scale Tests ............... 177 

5.2 Conversion Ratios with respect to the Fuel Thermal Energies ........ .. 178 

5.3 Conversion Ratios with respect to the Fuel Specific Thermal Energies . . . 179 

5.4 Propagation Speeds with respect to Locations of the Test Section ..... .. 180

NUREG/CR-6623

Typical Pressure Traces in Pure Water (WFCI-G-05) ................. 134 

Typical Pressure Traces in Polymer Solution (WFCI-G-01) ............ 135 

Debris Shape of the WFCI-G-06 Test in Size between 1 to 2 mm...... ... 136 

Debris Shape of the Pure Water Test (WFCI-G-05) ................. 137 

Distribution of the Debris Mass by Comparison with the Tests of Pure Water 

(WFCI-G-05) and Polymer Solutions (WFCI-G-01 and G-04) ........ .. 138 

Distribution of the Cumulative Debris Mass Fraction ................. 139 

Sauter Mean Diameters of Collected Debries in the WFCI-G and F Series 

Tests .......... ........................................ 140 

Pressure Traces of the WFCI-H-03 Test .......................... 141 

Pressure Traces of the WFCI-H-04 Test .......................... 142 

Level Swells of the WFCI-H-03 and H-05 Tests ..................... 143 

Void Fractions of the WFCI-H-05 Test Comparing with the Visual Observation144 

Conversion Ratios with respect to the Coolant to Fuel Mass Ratio Compared to Thermodynamic Models .................................. 145 

Trigger pressure histories for WFCI-1 apparatus charged to 200 V ..... .. 146 

Trigger pressure histories for WFCI-2 apparatus charged to 400 V ..... .. 147

ix
5.5 Variation of the Chapman-Jouguet Pressures associated with the Fuel-Coolant 

Mass Ratio in the Typical WFCI Experimental Conditions .......... 

5.6 The Variation of the Chapman-Jouguet Velocities associated with the Fuel

Coolant Mass Ratio in the Typical WFCI Experimental Conditions .... 

A.1 Cumulative Debris Distribution for the WFCI-A Series ............. 

A.2 Cumulative Debris Distribution for the WFCI-B Series ............. 

A.3 Cumulative Debris Distribution for the WFCI-C Series ............. 

A.4 Cumulative Debris Distribution for the WFCI-D Series ............. 

A.5 Cumulative Debris Distribution for the WFCI-E Series ............. 

A.6 Cumulative Debris Distribution for the WFCI-H Series .............

NUREG/CR-6623 x

11

181 

182 

200 

200 

201 

201 

202 

202
List of Tables

2.1 Summary of the ANL Experiments ...

Summary of the SUW Experiments ........................... 

Summary of the WUMT Experiments ......................... 

Summary of the SNL Experiments (Pat I) .... .. .... . ..... 

Summary of the SNL Experiments (Part II) ................. 

Summary of the SNL Experiments (Part III) ................ 

Summary of the SNL Experiments (Part IV) ................ 

Summary of the ALPHA Experiments .................... 

Summary of KROTOS Experiments with Alumina ............. 

Summary of KROTOS Experiments with Corium .............. 

Summary of the FARO Experiments ..................... 

Simulant Fuels and Coolant used in FCI Studies .............. 

Large Scale Experiments with Tin (Part I) ...................... 

Large Scale Experiments with Tin (Part II) .....................

S. . . . . . . . . . . . . 21

22 

23 

24 

25 

26 

27 

28 

29 

30 

31 

32 

33 

34

WFCI series A-H experimental initial conditions and results ......... 98 

WFCI series K pressure peak propagation speed and impulse data .... 99 

Impulse data for tests K-01 and K-09 .................... 100 

Experimental Conditions of Experiments in Comparison .............. 101 

Debris Masses Collected from the WFCI-G Series Tests .... .......... 101 

Experimental Results of the WFCI-A, B, C and D Experiments ...... .. 102

NUREG/CR-6623

2.2 

2.3 

2.4 

2.5 

2.6 

2.7 

2.8 

2.9a 

2.9b 

2.10 

2.11 

2.12 

2.13

4.1 

4.2 

4.3 

4.4 

4.5 

4.6

xi
EXECUTIVE SUMMARY 

Vapor explosions have occurred in various industrial processes following some accidental 

contact of a hot liquid with a cold more volatile liquid. In light water reactors after a 

prolonged lack of cooling, this process could occur when molten fuel is generated and 

contacts residual water coolant within the reactor vessel or below in the containment reactor 

cavity. Past studies have demonstrated the explosive nature of certain liquid pairs, but have 

not systematically examined explosion energetics at larger scales as a function of controlled 

initial and boundary conditions. This has hampered basic understanding and has been a 

major deficiency in the database for modelling purposes. Without such a systematic study 

one cannot verify the interrelationships between competing effects on explosion energetics, 

given a specific set of initial and boundary conditions. 

Our objectives for this work were to obtain well-characterized data for the explosion prop agation/escalation phases, while systematically investigating the effect of a comprehensive 

set of initial and boundary conditions on the explosion energetics; i.e., trigger strength, fuel mass, composition and temperature, coolant mass, viscosity and temperature and sys

tem constraint. This objective was subdivided into three specific stages. First, a vapor 

explosion apparatus, WFCI, was developed which allowed for well-characterized explosion 

data and demonstrated reproducible explosions with a simulant fuel. Second, the explosion energetics was examined as a function of varying initial and boundary conditions for one 

simulant fuel composition. Finally, the simulant fuel was changed from a metal to a molten 

oxide, which was more prototypic of the actual molten fuel compositions and explosion 

energetics were reexamined. 

The experimental investigation was subdivided into smaller test series to better understand 

specific vapor explosion mixing and propagation behavior. First, a series of experiments 

were performed to demonstrate the reproducibility of the explosion phenomena with molten 

tin as the fuel simulant. Initial conditions similar to the KROTOS-21 experiment were 

chosen as the nominal case for these tests, for comparison to independent data. The 

explosion behavior in the WFCI facility showed good reproducibility and agreement with 

KROTOS-21. Dynamic pressures had peak values of 10 MPa and quasi-steady values of 2-3 

MPa, while explosion propagation speeds were about 200 m/s. The explosion conversion 

ratios were about 0.2 to 0.5 percent. The next three test series investigated the effect 

of the external trigger on the vapor explosion; i.e., spontaneous explosions and triggered 

explosions with specific impulse strengths. The tests showed that as the external trigger was 

reduced and then eliminated, multiple propagation events occurred during the explosion. 

The energetics of the complete interaction was relatively independent of trigger strength, but the detailed behavior of propagation process became more complex with spontaneous 

triggers. Once the trigger strength exceeded 3 MPa no multiple propagation events were 

observed and single propagation events of similar energetics were observed. 

The next test series investigated the effect of axial constraint. The WFCI facility was 

originally designed with a rigid radial constraint to maximize the energetics for any given 

xiii NUREG/CR-6623
set of mixing conditions, but the axial constraint could be varied. The axial constraint 

was varied by changing the slug mass in the horizontal expansion tube by an order of 

magnitude. Results indicated that there was an optimal degree of axial constraint to 

maximize explosion energetics. This could be explained by the competing effects of rapid 

vapor production during the propagation and quenching in surrounding coolant liquid as 

the explosion mixture expands. This effect suggests that the actual axial constraint needs 

to be used in tests when considering reactor safety issues. 

The sixth and seventh test series investigated the effect of the fuel and coolant temperatures 

on energetics. The results indicated that once the fuel temperature was above a threshold 

value its effect on energetics was of second-order importance. The same result was noted 

for a variation in the coolant temperature. This suggests that if the fuel and coolant 

temperature are large enough to satisfy the qualitative requirements of a molten fuel and 

stable film boiling at the time of triggering, then an energetic explosion can result, with 

temperature having a small quantitative effect within this envelope of conditions. This 

should be scale independent and KROTOS tests also suggest this based on the scale for 

KROTOS-21 and for compositions with alumina fuels. This is also consistent with the lack 

of energetics from our later observations in iron-oxide tests as well as the KROTOS corium 

tests. 

The eighth test series was performed to investigate the suppression effect of polymer ad

ditives by an increase in the coolant viscosity. Polymer additives suppressed spontaneous 

vapor explosions in this larger scale geometry in qualitative agreement with past small scale 

tests. This was the first time that explosion suppression was demonstrated at relatively 

large scale. Also, the post-test fuel debris generated in the absence of the explosion may 

be quite representative of the fuel debris during the mixing process. This experimental 

technique should be considered as useful for subsequent fuel-coolant mixing studies. 

In the ninth series the fuel jet diameter and the timing of the external trigger were altered 

to vary the ratio of coolant mass to fuel mass in the mixture at the time of the explosion. 

Varying this mass ratio of coolant to fuel indicated that the explosion conversion ratio 

exhibited a broad maximum in energy conversion. These values were more than an order 

of magnitude lower value than one predicts from ideal thermodynamic situations. Analysis 

again indicated that the broad maximum is created by competing effects related to the 

development of the fuel-coolant mixture prior to triggering. The location of this maximum 

relative to mass ratio is secondary to its qualitative existance. However, the value of 

this coolant to fuel ratio can be understood relative to mixing kinetics. The small energy 

conversion ratio can only be explained by the observation that only a few percent of the fuel 

mass fragments into small enough debris ( <0.1mm) to directly participate in explosion 

timescales. This conclusion is applicable for all of our test data and seems to be scale 

independent, and thus, has important implications for reactor safety issues. 

Finally, in the tenth test series, the effect of changing the fuel composition from molten 

tin to a more prototypic simulant (iron-oxide) was observed. These tests indicated that 

the triggering of energetic fuel-coolant interactions with more prototypic fuel materials and 

superheats was quite difficult and no propagating vapor explosions were observed. This 

NUREG/CR-6623 xiv

I!
observation is consistent with those of JRC staff and their coriumn tests in KROTOS at 

atmospheric pressures in which weak vapor explosions were observed. 

With respect to reactor safety issues, this experimental work using fuel simulants has yielded 

a number of results that have potentially quite important safety implications. First, this 

work has provided clear evidence of the reproducibility of vapor explosion energetics for a 

controlled set of initial and boundary conditions. This suggests empirically that this phe

nomenon is predictable if one can establish and control the initial and boundary conditions. 

Second, the experiments demonstrate that geometric scaling can be properly specified; e.g., 

a rigid radial constraint for one-dimensional tests is conservative for energetics when com

pared to full-scale, while the axial constraint scale factor from test to prototype needs to 

be the unity to preserve energetics. Finally and most importantly, the data suggests that 

once the fuel-coolant initial conditions are within an envelope for triggered events, the en

ergetics is much less than thermodynamic, apparently due to the small amount of fuel that 

participates in an explosion timescale. And this envelope of triggerability is much smaller 

for a simulant molten oxide with low superheat, such as molten iron-oxide in our tests or 

corium in the KROTOS tests. This suggests that material scaling for reactor safety issues 

must preserve the same fuel composition and superheat from the test to the prototype. 

The current work has limited data at larger scales with more prototypic molten oxides; 

i.e., larger fuel volumes than 0.5 liters, larger chamber geometries, prototypic molten oxide 

compositions and superheats. It is recommended that further tests could be carried out 

under these conditions to empirically verify our findings. Models developed from our anal

ysis can also be used to analyze these experiments. Finally, it is known that the mixing conditions determine the envelope of explosivity for the vapor explosion. Thus, it is of fun

damental interest to better measure the mixture local conditions just prior to the explosion 

to correlate with the explosion energetics; i.e., void fraction profiles, fuel volume fractions 

and mixing diameters. Our future work in vapor explosion research is specifically targeted 

toward this purpose.

XV NUREG/CR-6623
Acknowledgements

The support from the U.S Nuclear Regulatory Commission under contract for this work is 

gratefully acknowledged. In particular the suggestions by Dr. Sudhamay Basu have been 

quite helpful. In addition, the authors wish to express their gratitude to the JRC research 

staff, in particular, Drs. H. Hohmann, D. Magallon, A. Yerkess, A. Annunziato and I. 

Huhtiniemi for fruitful discussions. Much of this report could not have been done without 

the help of Dr. Rosa Marina Bilbao and her detailed knowledge of the NSRC computer 

systems. 

xvii NUREG/CR-6623
Nomenclature

A Area 

A,m Fuel volumetric surface area 

Bo Bond number 

CD local drag coefficient 

C1 local friction factor 

C0 Drift flux coefficient or Proportionality constant 

CR Conversion Ratio 

Df, Initial fuel diameter 

Dh Hydraulic diameter 

DPO Pouring diameter 

DSE Delayed spontaneous vapor explosion 

E Thermal energy 

ESE Early spontaneous vapor explosion 

F Compensation factor for the fragmentation time 

F, Fragmentation rate 

H, Depth of coolant pool 

H(t) Heaviside function 

Ja Jacob number, Cp 7 

ill Volumetric vaporization rate 

MWt Thermal Mega-Watt Power 

NE No vapor explosion 

NP Number of fuel particle 

Nf, Number of fuel fragments 

P Pressure 

Q Heat transfer rate 

R. Transfer coefficient between the vapor and vapor-liquid interface 

R. Transfer coefficient between the coolant and coolant-liquid interface 

R! Final radius of the fuel after mixing occurred 

Ra Rayleigh number, 9D 3 Ap 

S Surface Area 

SE Spontaneous vapor explosion 

T Temperature 

TE Triggered vapor explosion 

AT.b Degree of subcooling 

U Velocity 

U, Relative velocity 

V1 Initial volume of the fuel mass to be mixed 

Viet Average coolant jet velocity

xix NUREG/CR-6623
We Weber number or Mechanical equilibrium work of the bubble 

cf Specific heat of fuel 

C, Constant pressure specific heat 

e Internal Energy 

fd The fraction of the fragmentation debris energy 

fl. Enhancement factor 

g Gravity constant 

h Overall heat transfer coefficient 

hc,sat Heat transfer coefficient of the coolant 

hgb Heat transfer coefficient with a gas blanket 

h!€ Heat transfer coefficient with a finite fragmentation rate 

hqs, Quasi-steady-state heat transfer coefficient 

hvsat Heat transfer coefficient of the vapor at saturated condition 

i Enthalpy 

i!.9 Latent heat of Vaporization 

iyus Heat of fusion 

irc Heat of reaction 

k Thermal conductivity 

m Fuel mass 

rh!,. Fuel fragmentation rate 

rh" Vapor generation rate 

qd Heat flux from the fuel droplet 

qýCHF Critical heat flux 

q4et,f Net heat transfer from the fuel to vapor and coolant 

4fra Net heat transfer from the fragmented fuel to generate vapor 

rd Fuel drop radius 

t• 6Dimensionless break-up time 

ttag Lag time 

tmix Mixing time 

tm Fragmentation or mixing time 

u Velocity 

v Specific volume 

x Mass fraction 

m/o Mole fraction 

w/o Weight fraction 

NUREG/CR-6623 xx

I
Greek 

a Volume fraction vapor 

# Constaht, (kfpfcf/kcp~cc)°" 

"y Specific heat ratio 

r. Surface area source due to primary break up 

re Surface area source due to surface entrainment 

6b Gas blanket thickness 

p Viscosity 

p Density 

a Surface tension 

Superscript 

min Minimum 

1 One-step 

Subscript 

a,oo Ambient 

c Coolant 

ch Chamber 

cr Critical 

cv Interface between the coolant and vapor 

e Equilibrium 

f Fuel 

fr Fuel Fragment 

g Gas or vapor 

hm -Homogeneous nucleation 

i Initial or interface 

m, mit Melt 

max Maximum 

min Minimum 

p Fuel Particle 

r Relative 

ref Reference 

sat Saturate 

sn Spontaneous nucleation 

sp Specific 

v Vapor 

vol Volumatric

xxi NUREG/CR-6623
Chapter 1

Introduction 

1.1 Background and Motivation 

Consider the case of a hot liquid rapidly releasing its energy into a surrounding cold, volatile 

liquid under direct contact, with excessive amounts of vapor production. If this event occurs 

within a timescale short compared to vapor expansion, it can cause local pressurization similar to an explosion and threaten the surroundings by the subsequent high pressure 

vapor expansion. From a historical point of view, this process has occurred since man 

began to work with molten metals. Berman et aL[1] used the term "Thermal Explosion" to 

contrast it from a chemical explosion. Condiff[2] suggested the term "Thermal Detonation" 

as a self-sustaining thermal explosion involving propagating shock waves. In the nuclear 

industry, the terms, energetic " Fuel-Coolant Interaction (FCI)" and "Vapor Explosion", 

are widely used in severe accident safety analyses, since molten materials being considered 

are nuclear fuel and associated metals in the reactor core. 

FCIs can be produced by different modes of contact, generally categorized in three different 

types as shown in Figure 1.1; i.e., pouring mode, injection mode and stratified mode. In 

the pouring mode, the hot liquid enters into the cold liquid, the hot liquid breaks up into 

small fragments and is mixed with the cold liquid. In the injection mode, however, the 

cold liquid is injected into the hot liquid and can be vaporized within the hot liquid. In 

the stratified mode, one liquid is separated from the other due to density differences and 

by the vapor generated at the interface; i.e., both liquids exist as stratified layers. 

In the nuclear industry, the vapor explosion phenomenon has been an issue in safety analyses for almost three decades. If a complete and prolonged failure of normal and emergency cooling systems in a nuclear reactor occurs, the reactor core could melt due to the fission 

product decay heat. There may be two possible modes of contact after the molten corium 1 is formed: the pouring and stratified modes of contact. In the pouring mode, the molten 

corium drops into the residual water coolant either inside or outside the reactor pressure vessel (RPV), and can result in energetic FCIs. The stratified mode, however, can be 

established during refilling of emergency water to the molten corium residing inside or 

outside the RPV. Experimental results[3] have suggested that the pouring mode of contact 

produces more energetic FCIs than the stratified mode of contact. 

1Material formed after the core melt, which includes fuel, partially oxidized cladding material and 

structural material

1 NUREG/CR-6623
After the contact between both liquids, a significant fraction of the thermal energy in 

the hot liquid is transferred into the cold liquid and in turn, converted into destructive 

mechanical energy due to the explosive vapor production and expansion. If the FCI thermal 

to mechanical work conversion is efficient, it may threaten the integrity of the containment 

by missile generation 2 or dynamic shock pressures, thereby leading to release of radioactive 

materials to the environment[5]. 

This type of explosive interaction has occurred in other industrial processes[6]; molten 

metal-water contact in the metal-casting industry[7], LNG (Liquefied Natural Gas) spilling 

over water during its transportation[8], leakage of cooling water onto molten salts, "smelt", 

in the paper industry[9], and lava flow into sea water[10]. 

A vapor explosion is a phenomenon which combines several different physical processes which occur in sequence. From several decades of research into the vapor explosion phe

nomenon, these phases have been conceptually identified as four different stages [11]: (1) 

Mixing, (2) Triggering, (3) Explosion propagation and (4) Expansion. 

In the mixing phase, if a hot liquid pours into a cold liquid, the hot liquid is broken down 

into smaller sizes due to fragmentation processes as it interpenetrates the coolant; separated from the cold liquid by a stable vapor film (i.e., the surface temperature of the hot liquid 

needs to be sufficient to maintain this film). In the triggering phase, if this metastable 

vapor film collapses locally due to a disturbance, very rapid heat transfer occurs due to the 

direct contact between the hot and cold liquid, and produces local high pressures due to the 

rapid vaporization process. This local explosion provides a trigger source to generate more 

fuel surface area and vapor generation in the adjacent mixture. These reactions produce a 

spatial propagation of the explosion as a shock wave passes through the melt-water mixture. 

With this shock passage in the mixture, fuel fragmentation and heat release occur near the 

shock front and reinforce the shock wave strength. During the expansion phase, high heat 

transfer rates from hot to cold liquid produce rapid vapor volume increases which may 

cause mechanical damage against the surrounding system constraint. 

During a postulated severe accident in a light water nuclear reactor (LWR), vapor ex

plosions may occur at two different locations; inside of the reactor pressure vessel, RPV 

(In-Vessel) and outside of the RPV (Ex-Vessel). For the case of the in-vessel vapor explo

sion, Theofanous et aL[12] estimated that an excessive explosion energy release could cause 

the R.PV to rupture in the lower head region. In addition, a liquid slug is accelerated up ward by the explosion expansion. Following some dissipation of energy as the slug passes 

through the upper internals, it impacts the RPV upper head and loads the head bolts. 

This may endanger the integrity of the containment since the vessel head (or portions of it) 

may become a missile with sufficient energy to penetrate the containment. There are other 

possible hazards from the vapor explosions that indirectly threaten containment: e.g., the 

explosion causes fuel dispersal into the containment and generates dynamic pressures that 

could induce a local RPV wall failure, significantly altering melt entry into the containment. 

'This specific sequence is termed t mode failure in WASH-1400 [4] 

NUREG/CR-6623 2

11
For an ex-vessel vapor explosion, the damage potential differs somewhat from an in-vessel 

event. Because the explosion occurs in a much larger volume, beneath the vessel in the 

LWR, the likelihood of direct containment failure due to missile generation is very small. 

In this case the major concern for an ex-vessel explosion centers on dynamic pressures produced in the water pool adjacent to the fuel-coolant mixture. If such pressures can 

be coupled to the containment pressure boundary or critical structures, then significant damage may occur. However, such a process is quite geometry dependent and is now being 

examined in safety studies. 

In most accident situations for current LWRs, the pouring mode of contact is the pre dominant geometric condition, since fuel melting can only occur in the absence of water; 

thus requiring a pouring of fuel into coolant for the FCI to occur. Since the vapor explo sion phenomenon'is a highly transient, nonequilibrium and multi-phase problem, there are 

many difficulties in uncovering its basic mechanisms from either experimental or theoret

ical approaches. From an experimental point of view, numerous small scale experiments for investigating the fundamental mechanisms causing a vapor explosion have been carried 

out [13]. These experiments provided some ideas to develop better models of the basic 

mechanism of the vapor explosion [14, 15, 16]. 

1.2 Objectives and Scope 

Attempting to explain large scale behavior using these small scale data, however, remains 

an issue since there is no well-defined scaling relationship for the vapor explosion process. In 

this context large scale experiments are defined as those in which spatial propagation occurs 

in the fuel-coolant mixture. The fuel mass is approximately equal to the coolant mass in 

the FCI with length scales much larger than the mixing length scale; i.e., length scales 

much greater than 1cm and fuel masses greater than a kilogram. It is almost impossible 

to carry out prototypic scale experiments, with fuel masses in the tons. Because of these 

facts, many researchers have been working on developing more mechanistic models of the 

vapor explosion instead of building prototypic experimental facilities. In theory, using a 

computer model for parametric and sensitivity studies with controlled input data is easier 

than conducting a large scale experiment, if the computer models can be believed; i.e., 

verified and validated at different scales. However, in order to develop such models, at 

least some well-controlled larger scale experiments are required in which the initial and 

boundary conditions are known and the key dependent quantities are measured. 

Such a situation provides the impetus for the present investigation of large scale vapor explosion phenomena under well-controlled experimental initial and boundary conditions. 

The main objectives of our present study are as follows: 

1. to obtain well-characterized data for the explosion propagation/escalation phases us

ing simulant fuel compositions,

3 NUREG/CR-6623
2. to investigate the effect of particular initial and boundary conditions (e.g., triggering, 

fuel mass, temperature and composition, coolant mass, temperature and additives as 

well as system constraint effects) on the explosion energetics and 

3. to use well-understood methods of analysis to understand our experimental results. 

In section two, previous experimental and theoretical studies are reviewed. These reviews 

support the necessity of this work and indicate its appropriate direction. The experimental 

facility designed for accomplishing the objectives is described in section three. The experi

ments are discussed in section four, and subdivided into a number of smaller test series to 

investigate specific initial and boundary conditions. The first four test series examine the 

reproducibility of the vapor explosion under well-controlled conditions, and the influence of 

an external trigger on the explosion behavior. The next test series determines the effect of 

the system constraint on the explosion energetics. The next four experiments demonstrate 

the effect of the fuel and coolant temperatures on FOI behavior as well as the ability to 

suppress the explosion with coolant additives. Finally, the fuel composition is changed to 

a more prototypic simulant, iron-oxide, to see its effect on explosion energetics. In sec

tion five, the discussion of the experimental results is presented, along with our supporting 

analyses. Finally, the conclusions for this work and recommendations are given in section 

six. Some detailed experimental data are presented in Appendices.

NUREG/CR-6623 4

11
of 

Dropping

Stratification 

SHot liquid 0 Cold iquid 

Figure 1.1: Typical Geometries of Fuel-Coolant Interactions

NUREG/CR-6623

Injection.

5
Chapter 2

Literature Review 

2.1 Perspectives on Vapor Explosions 

From 1950 to 1960, accidental pouring of molten materials [7, 17] into water resulted in 

energetic FCIs causing massive structural damage in the metal-casting industry. From the 

middle of the 1950's a new improved ingot continuous-casting method was introduced in 

the aluminum industry which cooled the hot metal very rapidly with water. Accidental 

contact of the water and molten aluminum also resulted in vapor explosions as the num

ber of hazardous events increased. Such vapor explosion hazards have been noted in a 

number of other process industries over the last fifty years; e.g., paper pulping, ferroalloys 

manufacturing, and liquified natural gas. 

In the nuclear industry, several incidents associated with this phenomenon in test reactors 

were reported[18]; e.g., BORAX-I (1954), SL-1 (1961), SPERT (1964), etc. One notable 

accident occurred at the SL-1 test reactor[18]. The SL-1 reactor was a 3 MWt aluminum 

clad-aluminum alloy fueled prototype natural circulation boiling water reactor. Rapid 

steam production caused a high pressure buildup which accelerated the body of the water 

upwards above the core. The water slug hit the vessel lid, collapsed the extension tube 

housings around the control rods and bulged the pressure vessel. A major reason for the 

destructive potential of this FCI was that a significant amount of the premixed aluminum 

melt was oxidized on this explosive timescale and this enormously increased the explosion 

yield. 

In the 1970s, the vapor explosion process was intensively studied for the fast breeder reactor, 

in which uranium oxide and sodium were fuel and coolant, respectively. While research was 

continued on vapor explosions for the safety of the fast breeder reactor, a comprehensive 

risk assessment to estimate the likelihood of containment failure in Light Water Reactors 

(LWR) was conducted in 1975 and reported in WASH-1400[4]. This study focused on 

two specific reactor designs; the Surry PWR and the Peach Bottom BWR-Mark I. The 

hazard from a vapor explosion was considered in this assessment. It was determined that 

the containment could be threatened by three possible damage mechanisms; (a) dynamic liquid phase pressures on structures, (b) static overpressurization of the containment by 

steam production, and (c) solid missile generation from the impact of a liquid slug accel

erated by the vapor explosion. In this analysis, the primary concern was a direct failure of 

containment caused by an energetic FCI causing missile generation (designated "a-mode" 

failure). 

NUREG/CR-6623 6

1!
In WASH-1400, the possibility of a large-scale vapor explosion which would threaten the 

containment of a light-water reactor was estimated to be about 10-2 per reactor year with 

the likelihood of water availability and triggering of the explosion as the major uncertainties. 

Conservative results were obtained from this analysis because it was conducted under some 

bounding assumptions. Those assumptions were: first, all the fuel inventory in the core was 

melted and well-dispersed with half of the water in the lower plenum. Second, the remaining half of the water was above this fuel-coolant mixture when it exploded. And third, the 

piston of water was driven upwards through a clean vessel without internal structure and 

then impacted coherently on the top of the vessel. 

Corradini et al. [19] developed a discretized Monte Carlo probablistic analysis to estimate 

the WASH-1400 vapor explosion failure mode with models for fuel-coolant mixing and 

explosion, i.e., expansion work and associated dissipation and lower head failure models 

to estimate the a mode failure probability. Given a core melt, the probability of a mode 

failure was estimated to be less than 10- per reactor year. In 1985, in NUREG-1116[20], 

the Steam Explosion Review Group (SERG) estimated the probability of the a mode failure. 

This group of experts performed independent analyses and estimated that the conditional 

probability of the a mode failure was much less likely than in WASH-1400 (10-2 , 10- 4 /yr 

as upper bound given a core melt). This group also recognized that these estimates were 

founded on the judgment that the amount of fuel-coolant mixture was limited and/or 

the explosion yield was less than the maximum thermodynamic values. Included in their 

findings was the consensus recommendation that fundamental experiments needed to be 

performed at larger scales to characterize fuel-coolant mixing and measure explosion yield 

as well as the effect of mixing on yield. 

Theofanous et al. [12] performed a comprehensive analysis of a mode failure and found 

the upper bound value to be much lower than past estimates. Recently, Turland et al. [21] 

investigated a methodology for quantifying the conditional probability of the a mode failure 

in the Sizewell B PWR. They estimated that the probability was approximately 10- and 

the effect of the system pressure elevation on the probability of this mode of failure was 

modest. Theofanous et al. [22] updated their original risk assessment[12] and concluded 

that even vessel failure by vapor explosions might be regarded as physically unreasonable. 

Almost ten years after SERG-1 the same group of experts convened in 1995 for SERG-2 

[24, 23]. Their consensus opinion was that the probability of a mode failure was <10-4 per 

reactor year given a core melt accident and was essentially resolved from a risk perspective. 

Other FCI issues were identified by the experts, specifically those that related to the ener

getics of vapor explosions and their impact on lower head failure and structural integrity. Further experimental research and associated analysis was recommended to address these 

issues. 

This historical perspective on this phenomenon indicates that one needs to better under

stand the energetics of the explosion as a function of the initial and boundary conditions. 

Specifically, one needs to be able to correlate the expected energetics to variations in these 

conditions. There have been past experiments and analyses that could be examined with 

this purpose in mind. A comprehensive summary is presented below to aid the reader for

7 NUREG/CR-6623
this particular purpose.

2.2 Past Vapor Explosion Studies 

2.2.1 Large Scale Experiments 

In this work, we provide the current status of large scale experiments and organize the 

data based on the research group as well as on various fuel simulants, e.g., tin, salt, ther

mite, corium, etc., for a pouring mode of contact. More comprehensive reviews for other 

topics have appeared in several review papers[5, 13, 25, 26, 27]. The review of small scale 

experiments is brief and limited to experiments performed with tin as the simulant fuel 

material. 

2.2.1.1 Argonne National Laboratory 

Recently, Spencer et al. [28] reanalyzed two experimental series, designated as CWTI[29] 

and CCM[30]. These experiments were performed in the COREXIT facility. The facility 

allowed for FCI studies of 1-10 kg of corium and consisted of a containment cell, the 

experimental apparatus, instrumentation, control systems, optical photography and flash 

X-ray as shown in Figure 2.1. The CWTI tests focused on FCIs occurring in the ex-vessel 

reactor cavity and the associated containment pressurization from high and low pressure 

melt ejections into the cavity region. The CCM tests were performed to investigate the 

coarse mixing and melt jet breakup for an in-vessel geometry. The detailed experimental 

conditions and results are summarized in Table 2.1. Major conclusions from the work were: 

1. energetic FCIs did not occur in these tests, rather vigorous vapor generation during 

the FCI was observed under all circumstances; 

2. the CWTI test series indicated that the presence of water exhibited a mitigative effect 

on containment pressurization for the range of initial conditions studied; 

3. the CCM test series clearly showed that fuel jet breakup occurred at the leading edge 

of the jet as well as along its upstream surface due to hydrodynamic instabilities 

induced by relative velocity and vaporization. 

The latter experimental observations were used in the development of the THIRMAL mix

ing model for the fuel-coolant interaction. 

NUREG/CR-6623 8

1i
2.2.1.2 Winfrith

Four different series of large scale experiments were performed at Winfrith by Fry using 

corium thermite et aL. [31], SUW[32], WUMT[33] and MIXA series[34, 35]. Both the SUW 

and WUMT series of experiments were conducted in the MFTF facility and MIXA in the 

MIXA facility. 

The MFTF facility was a large pressure vessel with a volume of 1.7 m3 , equipped with the 

thermite charge container as shown in Figure 2.2. To release the melt from this thermite 

container, two different methods were applied; free and restricted release modes. In the free 

release mode, the melt was released freely into the surrounding water, and in the restrict 

release mode, the melt was constrained by an open catchpot and ejected with the end cap 

of the charge container. 

The SUW series of experiments focused on the investigation of the effects of melt mass, 

ambient pressure and subcooling of water. A total of twelve tests in the SUW series were 

performed including nine tests with the melt mass of 24 kg and three tests with the melt 

mass of 8 kg as shown in Table 2.2. Spontaneous explosions and triggered explosions were 

observed in all the experiments, but two (SUW-02 and SUW-10). Some experiments were 

triggered by the impact of the catchpot on the bottom of the vessel; i.e., impulsive trigger 

from this impact. Bird[32] observed in this series of tests that the conversion ratio tended to 

increase with decreasing water subcooling and increasing system pressure even though the 

likelihood of an explosive interaction decreased. Note that the unusual contact geometry 

makes these tests difficult to interpret for reactor safety considerations. 

In the WUMT tests, 24 kg of melt was poured into water by gravity through a circular orifice 

with a diameter ranging from 40 to 100 mm. Two experiments produced a spontaneous 

explosion in water at 80 K subcooling and saturated water. As shown in Table 2.3, no 

explosive interactions occurred in any other tests. These tests provided little insight into 

explosion energetics. 

In the MIXA series of tests, the facility was designed for studying the mixing phase of 

FCIs. In particular, the corium melt poured through a droplet former installed at the top 

of the mixing vessel, which produced various diameters of melt droplets before melt-water 

contact. In the MIXA series of tests, Denham et a. [34, 35] reported five experiments as 

shown in Table 2.3. Melt and water masses were about 3 kg and 80 kg respectively. There 

were no spontaneous explosions in any of these series of tests, even with water subcooled 

by 20 K. 

2.2.1.8 Sandia National Laboratories 

Fuel-coolant interaction data are available from many large scale experiments conducted 

by Sandia National Laboratories (SNL). These tests focused on estimating the FCI energy

9 NUREG/CR-6623
conversion ratio, determining the triggering behavior and explosion threshold, and iden

tifying the effects of various parameters on FCIs. Tables 2.4, 2.5, 2.6 and 2.7 show the 

detailed experimental conditions and results. 

The purpose of this Open Geometry test series [36, 37] was to estimate the conversion 

ratio of vapor explosions at large scale geometry. About sixty experiments were conducted 

in the open-geometry facility (OG) with minimal instrumentation as shown in Figure 2.3. 

The two sets of experiments were performed with different fuel materials (thermitically 

generated Fe - A120 3 and Corium). The thermite melt consisted of 55 w/o Fe and 45 

w/o A120 3 , at a theoretical (maximum) temperature of approximately 3100 K. These tests 

resulted in spontaneous explosions with conversion ratios ranging from 0.1 to 1.5 %, with 

small correlation to initial and boundary conditions. In the second set of experiments[37], 

with a prototypic reactor material of Corium-A+R 1 , no energetic explosions were observed. 

Corradini[38] explained the reason for the lack of an explosion for the corium test as the 

solidification of the molten Corium due to relatively low superheat of molten fuel. In 

this series of tests, however, the precise data measurement was limited due to the lack of 

instrumentation and the facility geometry. 

The tests performed in the EXO-FITS facility focused on instrumentation development 

and pre-tests for the FITS series. Even though the experimental data were limited due to 

the lack of instrumentation, it provided qualitative information on the effects of fuel com

position, the characteristics of spontaneous explosions and selected explosion propagation 

data. 

The MD series of experiments[39] was performed primarily to study melt delivery methods. 

A total of thirteen experiments as shown in Table 2.4 were conducted. A 0.6 to 5.11 kg 

quantity of Fe-A1203 melt was poured into subcooled water at an ambient pressure of 

0.083 MPa. In highly subcooled conditions, the spontaneous explosions showed random 

characteristics. In most of the spontaneous explosion cases, the propagation velocities were 

estimated by visual observation and ranged from 300 to 550 m/s, with shock pressures as 

high as 34 MPa. The conversion ratio of the explosions were estimated to be quantitatively 

greater than the nominal 1.0 % in the Open-Geometry tests. The MDC tests[40] were 

also perfosrmed to test various compositions of corium melt with a fuel mass from 4 to 

20 kg injected into a water mass from 77 to 276 kg. A wide variation in the explosivity 

was observed with no pattern observed. This again suggests that reproducibility in test 

conditions was lacking and may have caused scatter in the explosion data. 

Five tests in the MDF series of tests[40] were conducted for examining a new iron oxidic 

simulant (Fe 30 4). The oxidic melt was produced by mixing appropriate quantities of iron 

powder and potassium perchlorate followed by a thermite reaction. There were no spon

taneous explosions, although an external trigger (0.635g of explosive) was applied in three 

tests. Berman et al. [40] observed that the melts were dispersed before entering the water 

153 w/o U0 2, 16 w/o ZrO2, 2 w/o NiO 27 w/o SS and 2 w/o Mo 

NUREG/CR-6623 10

1!
and individual single drops were locally interacted. They suggested that energetic explo sions were dependent on void fraction in the mixture (i.e., melt dispersion) and explosive 

propagations were not triggered if local triggers were weak. 

The CM series tests[40] were designed to investigate the coarse mixing behavior of FCIs 

with extensive visual observation. For this coarse mixing study, tests shown in Table 2.5 

were mostly conducted at nearly saturated water conditions with subcooling from 1 to 

9 K, without an external trigger. All twelve experiments produced violent eruptions at 

approximately 20 to 70 ms after the melt contacted the surface of the water. Multiple 

explosions were observed in some of the spontaneous triggered tests and provided damage 

to the water chamber. In these tests, the difference between the eruption and explosion 

was defined by the duration time of the interaction; the eruption lasted for an extended 

period of time longer than 50 ms with little pressurization. 

Two RC tests[40] were conducted by replacing the weak constraint lucite interaction water 

chamber with a steel pipe (0.6 m diameter and 25 mm thick) to estimate the effect of the 

rigid radial constraint on FCIs. The Fe-Al20 3 fuel was poured into subcooled water. The 

first experiment produced a violent surface eruption as observed in the CM tests. In the 

second test, however, the largest recorded vapor explosion in the EXO-FITS tests occurred. 

Rough estimates of the conversion ratio for this test ranged from 0.8 to 14 percent. This 

strong explosion was explained at that time by the radially confined geometry enhancing 

liquid-liquid contact by driving the liquids together as they approached the wall, thereby 

providing more fragmentation. It is interesting to note that these test conditions were not 

replicated again in any Sandia experiments. Detailed experimental conditions and results 

are shown in Table 2.5. These results were an impetus to consider the current experimental 

series. 

Based on experiences from the EXO-FITS tests, the FITS series of tests[39, 40, 41] were 

performed in the chamber shown in Figure 2.5. The chamber gas pressure, water phase 

pressures were measured, debris was collected and analyzed from the resulting fuel-coolant 

interaction and the explosion process was observed. The purposes of these tests were to 

determine the trigger behavior, explosion threshold and estimate of the conversion ratio as 

a function of ambient pressure, fuel composition, and other initial conditions. 

The FITS-A test series was performed with similar conditions as some of the earlier EXO

FITS experiments with iron-alumina thermite and weak wall constraint (MD-15 and MD

16). In these tests the effect of the system ambient pressure was demonstrated. It was found 

that a system pressure increase suppressed explosions (FITS4A). However, the explosion suppression did not persist against an external trigger with a trigger pressure of 4 MPa 

(FITS5A). 

The FITS-B test series was primarily performed to investigate the effect of the coolant to 

fuel mass ratio on spontaneous vapor explosions, at an ambient pressure of 0.083 MPa and 

a coolant temperature of 300 K. The mass of the fuel was increased from 2 kg (FITS-A 

series) to 18.7 kg. The mass ratio was controlled by altering the dimension of the interaction 

chamber and water depth. The conversion ratios of these tests were calculated by using

11 NUREG/CR-6623
the estimated coolant slug kinetic energies from observations. Among these tests, two 

experiments (5B and 6B) were performed at temperatures near water saturation and the 

energetics was small. Most experiments were spontaneously triggered at the surface of 

the coolant and/or the base of the chamber in a subcooled coolant. For the range of the 

mass ratio, Mo/Mf, from 3 to 15, conversion ratios estimated from the coolant slug kinetic 

energies were relatively unaffected by this variation. Once again data scatter was large and 

this inhibited a clear understanding of the effect of initial conditions on energetics. 

The FITS-C series tests were performed to study the effects of fuel composition and system 

pressure on energetics of FCIs, the resultant debris formation and hydrogen production. 

The experiments were conducted in an inert nitrogen environment to sample post-test 

hydrogen content. No external trigger was applied as shown in Table 2.7. The FITS1C test 

was conducted with similar conditions to the FITSB series of tests with an iron-alumina 

fuel. None of these series of tests, except two, produced explosions. Fletcher[27] suggested 

that both low-melt superheat and dispersion of the melt before contact with the coolant 

reduced the likelihood of triggering. 

The FITS-D series[42] investigated the effects of the coolant to fuel mass ratio, water 

subcooling and ambient pressure and measuring of the hydrogen generation. In this series 

as shown in Table 2.7, no external trigger was provided. Increases in ambient pressure 

again suppressed the explosion. Only one explosive interaction (FITS5D) was observed. 

This test was performed at atmospheric pressure with highly subcooled water and produced 

a double explosion separated by about 3 ms. In past experiments double explosions were 

observed with a delay time of about 100 ms. 

2.2.1.4 JAERI 

In the 1990's the ALPHA program was initiated at JAERI (Japan Atomic Energy Re

search Institute) in Japan to investigate phenomena that may threaten containment in

tegrity during postulated severe accidents. One experimental program examines the vapor 

explosion using a chamber type facility, named ALPHA as shown in Figure 2.6. The AL

PHA facility[43, 44, 45] simulates the containment with an inner volume of 50 m3 and a 

weak-walled interaction chamber. The melt is generated in the melt generator by a ther

mite reaction with iron-oxide and aluminum. This melt is identical to what was used in 

the Sandia tests. All detailed conditions and results are shown in Table 2.8. The results 

from over twenty experiments are qualitatively similar to what had been observed in the 

Sandia FITS tests. Similar data scatter can also be noted and this may again be due to 

the lack of precise control on the constraint and initial conditions. 

NUREG/CR-6623 12

1i
2.2.1.5 JRC-Ispra

At the European Joint Research Center (JRC) at Ispra, two different sets of experiments (KROTOS and FARO) related to fuel-coolant interactions are being conducted. These 

experiments began over fifteen years ago applied to fast reactors. Since 1990, these tests 

have focused on light water reactor safety. These test facilities are unique because they have 

the capability to use prototypic reactor materials; i.e., urania and other high temperature 

oxides. 

The FARO facility shown in Figure 2.7 was designed to provide an experimental data base 

on molten fuel jet and water quenching and mixing phenomena. These tests are performed with about150 kg of corium in prototypical conditions; e.g., early test series shown in 

Table 2.10. The tests simulate a corium pour into a water pool in the lower plenum of a 

reactor pressure vessel (RPV), or in the reactor cavity, its debris settling on the chamber 

bottom and its long term quench behavior. 

The KROTOS facility shown in Figure 2.8 was built to obtain experimental information 

on explosive fuel-coolant interactions for various fuel-coolant pairs. Several different types of simulant fuels were utilized; e.g., tin, aluminum oxide and corium (mixture of urania 

and zirconia). These tests mainly focused on studying the effects of initial and mixing conditions on the energetics of the FCIs. These experiments are complimentary to the 

experiments performed in our WFCI facility. 

In these tests[47, 48, 49], the objective was to prepare and deliver a known molten fuel 

material into a water pool under controlled conditions. To accomplish this purpose the 

facility was designed to produce an explosion propagation in a one-dimensional geometry. The melt is prepared in a radiation furnace to a desired temperature, the melt crucible is 

then delivered to a location where melt release produces a fuel jet pour of 30 mm diameter 

into the test section. At a prescribed time the diaphragm to a pressurized gas volume is 

ruptured at the chamber bottom and a shock propagates vertically upwards through the 

mixture to trigger the explosion. Dynamic pressures and ambient pressures are recorded 

during the explosion transient to ascertain the explosion energetics. Debris is collected 

in the surrounding containment vessel (free volume of 275 liters). Two different inner 

diameters of the explosion test tube have been used (95 mm and 200mm with a tube 

length of about 1.2 meter) for the various experimental conditions as shown in Table 2.9. 

About ten years ago the KROTOS tests utilized molten tin as the simulant. These were 

prepatory experiments for high temperature oxide melts. As the heating furnace was up

graded, molten aluminum-oxide was used as the fuel simulant. Most recently, molten corium 

is the fuel simulant and is quite prototypical of the molten core material; i.e., 80 % urania 

and 20 % zirconia. Since this transition to molten oxide fuels, over two dozen tests have 

been conducted to investigate the explosivity of alumina and corium and to understand 

the effect of certain initial conditions. The experimental results can be summarized in the 

following manner:

13 NUREG/CR-6623
1. Molten alumina melts produced explosive interactions under all conditions except 

for nearly saturated water. These explosions were both triggered and spontaneous. 

The general characteristics were supercritical pressures (20 - 100 MPa) with large propagation velocities (about 400 - 600 m/s) with narrow pressure pulse widths (1 

msec) and energy conversion ratios of a few percent (1 - 3 %). 

2. The effect of altering the fuel temperature or coolant temperature (below saturation) 

had little effect on energetics for alumina. Modest increases in ambient pressure 

below 2 bar also did not affect energetics. Nearly saturated water appeared to cause 

much larger vapor formation and thus may have substantially increased the void in 

the mixture. 

3. Molten corium melts did not produce an explosion under atmospheric pressure condi

tions. Only when the ambient pressure rose above 2 bar could the FCI be externally 

triggered and be energetic enough to become a propagating explosion. Peak pressures 

were much lower (20 - 40 MPa) and associated pulse widths (<1 msec) and conversion 

ratios were quite small ( 0.02 - 0.05 %). The reason for this behavior is not totally clear, but is attributed to low melt superheat for corium and large void fractions in 

the mixture at the time of triggering. An increase in ambient pressure seems to be 

sufficient to reduce the mixture void to allow the explosion to be triggered. 

2.2.2 Experiments using Tin as a Fuel 

As a simulant fuel, tin has been widely used for investigating fuel-coolant interaction phe

nomena because it has a relatively low melting point of 231.9 0C, low toxicity and produces 

vapor explosions at subcritical pressures, when it contacts water. Initial FCI experiments 

were performed at small scales where small masses of tin were dropped into water to iden

tify regions of spontaneous interactions (Figure 2.9). In Table 2.11, tin and other simulant 

fuels used in past experiments are compared [51] with respect to their mechanical and 

thermal properties. Tin has a similar energy content per unit volume as aluminum, while 

alumina and corium are also quite similar. However, we know from empirical evidence that 

aluminum and alumina result in much more energetic explosions. 

Hall et al. [60] investigated vapor explosions in a long tube geometry as shown in Figure 2.10. 

The main test tube was 0.85 to 1.0 m in length and a 25 mm inner diameter; the outer 

tube had a 50 mm inner diameter. The test vessel was filled with water at 85 to 95 0C and 

tin as fuel material at 600 to 750 °C was poured into the top of the inner tube. They 

observed self-sustaining shock propagations and escalation. All detailed conditions and a 

brief summary of results are shown in Table 2.12. 

Briggs[61] investigated the metal/water interaction including aluminum/water and tin/water 

as the first stage of the THERMIR experimental program. These experiments as shown in 

Table 2.12, had no external trigger device because of the scoping nature of the tests. Three 

experiments (8,9, and 10) with a molten tin mass of 2 kg produced localized interactions 

NUREG/CR-6623 14

1!
with peak pressures of 0.2 - 0.3 MPa.

Fry and Robinson[31] continued the experimental studies in the THERMIR facility to 

investigate explosion propagation. In their experiments, explosions were initiated by an 

external trigger using a mechanical impact or detonator. Molten metal up to 16 kg at 

about 800 0C was poured into water at a temperature between 44 and 86 *C. The first 

five experiments in Table 2.12 produced coherent explosions at high melt temperatures. 

These tests suggested that coherent explosions might occur for a large mass system (several 

kilograms) even when the interface temperature was below the homogeneous nucleation 

temperature. 

More recently Baines[62] conducted an experiment to investigate vapor explosion propaga

tion and work yields in a constrained system. This apparatus (shown in Figure 2.11) was 

equipped to measure the explosion work output, which was the expansion tube attached 

horizontally to the upper part of the vertical test section. The vertical test tube (,I m 

long) and the horizontal cylinder (-3 m long) were made of stainless steel tubing with an 

inner diameter of 29.5 mm. The tin (-1 kg) is heated to about 800 0C and pours into 

the water in the test section. After a preset time delay (-2 s) to allow the tin to almost 

reach the bottom of the test section, the gate valve at the top of the test section closes 

and a steel plate simultaneously cuts off the flows of tin into the funnel. A fraction of a 

second later, molten tin reaches the cold water in the triggering section, which causes the 

vapor film to collapse, initiating a steam explosion. The resulting mixture then expands to 

cause the magnet piston in the horizontal cylinder to move. He observed that the explosion 

efficiencies were low with work yields less than 0.4 % of the thermal energy in the fuel, 

propagation velocities were about 100 m/s and the explosion mixtures produced were rel

atively "weak" with the tin volume fractions in the range of 0.08 to 0.14. This experiment 

is of note because it is the only test apparatus which took great care in measuring the 

energetics under controlled conditions. 

A total of twenty-two tests for a tin-water vapor explosion were performed in the KROTOS 

facility. Only the KROTOS-21 test[63], however, was reported in the open literature and 

used as a benchmark data for the verification of several computer models. In this test, a 

test section with an inner diameter of 95 mm was chosen for investigating one dimensional 

explosion propagation eliminating the explicit radial distribution of the fragmented melt. 

As shown in Table 2.13, about 6.5 kg of the melt was poured into the water and explosive 

interactions were initiated by the gas trigger with a magnitude of 12 MPa. The propagation 

of the coherent explosion with a peak pressure of 6.5 MPa moved upward with velocities 

ranging from 150 to 270 m/s. The post-test analysis showed that approximately 0.85 

kg, i.e. 13 % of the tin mass, had fragmented to diameters below 250 pm. No direct 

measurement of work yield was available, although estimates could be made based on 

impulse and chamber pressurization of 1-2 kilojoules.

15 NUREG/CR-6623
2.2.3 Experiments using Iron-Oxide as a Fuel

Iron-oxide is not commonly used as a fuel simulant in vapor explosion experiments. It has 

a fairly high melting temperature (1550 0C), and though it is chemically inert in water, it 

is very corrosive to containers holding it in its molten state. The cases where it has been 

used have been small scale single droplet experiments [58, 59, 66, 67, 68]. where the fuel 

was prepared in the form of Fe 30 4 . These experiments are considered small scale because 

the mass of the fuel is small when compared to the mass of the coolant, thus there is no 

spatial propagation of the explosion through the fuel-coolant mixture. The fuel droplet 

was prepared by melting a small piece of iron foil with a heat source like a laser at above 

1850 *C. At these high temperatures, the iron oxidizes in air to a composition of magnetite 

similar to Fe 3O4 , which represents its equilibrium state. This small droplet of material is 

usually held by surface tension to a wire made of iridium, and is easily dropped on demand. 

This method is reliable and easy to perform for repeated experiments, since it eliminates 

the need for a crucible and large furnace, but it is always limited to very small masses 

(<0.1gm), and hence is only used for small scale experiments. 

At Sandia National Laboratories, many single droplet experiments were performed, with 

iron-oxide as the principal fuel composition. Nelson and coworkers [58] performed a series 

of experiments with small iron-oxide droplets (< 0.1 g and 3mm in diameter) at temper

atures close to 2000 0C, dropped into water at 30-50 *C. The coolant test section was 

transparent, which allowed filming of the experiment with a high-speed camera. As the 

droplets contacted the water, a stable vapor film could be seen forming around the fuel 

droplet. This vapor film was destabilized by a pressure pulse, thereby forcing liquid-liquid 

contact and initiating the local vapor explosion. Triggering of the explosion was achieved by 

electrically vaporizing a thin gold wire, and pressure traces were acquired with submerged 

pressure transducers. It was found that multiple explosions would occur from the single 

droplet. The high-speed film allowed measurement of the size of the vapor bubble that was 

created as a result of the explosion. Conversion ratios of the explosions were found to be in 

the range of 1 - 3 %, and peak pressures were estimated to be above 10 MPa near the drop 

surface. If the trigger pulse was less than 0.2 MPa, no explosion would occur, and with 

trigger pulses of 0.4 MPa or greater, a prompt explosion would always occur. For trigger 

pressure peaks in the range of 0.2 - 0.4 MPa, the drops would undulate, then show delayed 

cyclic explosions. Based on debris sizes of less than 250 microns, it was determined that 

the explosion occurred due to the interaction of the coolant with liquid fuel, with no evi

dence present that solidified fuel participated in the explosion. Electron microscope images 

of the debris confirmed this, with no evidence of sharp edges on the particles that would 

have indicated that a solid fuel particle had fragmented. All the particles were spherical in 

shape, indicating that they were in a liquid state during the interaction. 

At the University of Wisconsin, Kim [67] performed a series of single droplet experiments 

very similar to Nelson with 0.1 g or less of iron-oxide at 1900 0C, as the fuel dropped 

into water at 15 *C. Pressure histories at three different locations in the container were 

measured, and the entire experiment was filmed with a high-speed camera. In addition to 

baseline experiments and investigations into the effect of variation in trigger pressure, the 

NUREG/CR-6623 16

11
effect of changing the coolant viscosity on the explosion energetics was to be observed by adding a cellulose gum solution to the water. Conversion ratios for the baseline experiments ranged from 1- 6 %. The results showed that an increase in coolant viscosity reduced the 

conversion ratio of the explosions, and in some cases, suppressed the explosion altogether. Also, an increase in trigger pressure seemed to offset the effects of the viscosity increase, but only to a limited extent. Even with a 0.4 MPa trigger pulse, if the viscosity was larger than 150 centipoise (cp), the conversion ratio was very low (< 0.5 %). Peak pressures were 

higher with a lower viscosity, with a maximum of about 2 MPa for a 40 cp solution, and 

a maximum of about 0.5 MPa for a 240 cp solution. This suggests that the addition of 

an additive to the coolant which increases viscosity could help suppress or mitigate vapor 

explosions. 

Baker [68] added to the work of Kim by using the same apparatus to investigate the effect of 

surfactant additives (substances that affect surface tension) to the coolant on the energetics of single droplet vapor explosions with iron oxide. Two different surfactants, ethoxilated

nonyle-phenole and sodium-dodecyl-benzene, were investigated. The peak pressure data 

was inconclusive with values ranging from 0.1 to 1.2 MPa. Also, the conversion ratio 

data indicated the possibility of a slight decrease in the explosion intensity with higher concentrations of either surfactant. Conversion ratios were in the range of 0.5 to 12 % at 

low concentration, but dropped to around 0.2 % at 10 wppm. The particle debris size data 

(Sauter mean diameter), which is used as a measure of fuel participation in the explosion, 

was inconclusive for ethoxilated-nonyle-phenole, but showed some evidence of a suppressive effect for sodium-dodecyl-benzene. 

2.2.4 Related Experimental Parameters 

Coolant Additives 

To reduce the potential hazard of the FCIs, many methods have been investigated [7, 66, 67, 69]. In the aluminum industry in late 50's the causes and prevention of these explosive 

interactions were studied by Long[7]. He performed large scale tests with aluminum and 

various combinations of materials. In these experiments, a sudden discharge of 22.7 kg of 

molten aluminum through an 89 mm diameter hole into a clean degreased steel container 

partially filled with water at 12.8-25.6 °C, always produced an explosion. Based on his 

results, Long suggested that the cause of these explosive interactions was the rapid vapor ization of water entrapped between the melt and container surface. In order to reduce the 

amount of water entrapment, he coated the inside wall with grease, oil or a bituminous 

paint and found that the explosion did not occur with the coated container surface. 

Nelson et al. [70] investigated the reason for these non-explosive interactions. They sug gested the hypothesis that the initiating action at a wet surface was caused by a thin 

layer of liquid water enclosed beneath the molten metal and solid surface. Nonwettable 

("coated") surfaces did not have the thin layer of water between the melt and surface, and 

offered little or no assistance to triggering during the interaction.

17 NUREG/CR-6623
Recently, in the nuclear industry, numerous efforts in suppressing or mitigating the FCI risk 

have been made [66, 67, 69, 71, 72]. Most of the work has focused on the effect of additives 

in the coolant, which changed the characteristics of the coolant, resulting in suppression of 

explosive interactions. As one of many possible additives, water-soluble polymers, which 

change coolant properties such as viscosity and heat transfer characteristics, were previously 

investigated. Kotchaphakdee et al. [73] studied the nucleate pool boiling of dilute aqueous 

polymer solutions on a heated flat plate. Paul et al. [74, 75] measured nucleate boiling 

curves for aqueous solutions of nine different drag-reducing polymers on a platinum wire. 

They found that nucleate boiling heat flux for the polymer solutions was substantially 

increased. 

Most of the past work focused on the investigation of the boiling heat transfer characteristics 

in dilute polymeric solutions. One of the early studies for investigating the threshold 

viscosity of polymeric solutions was performed by Flory et al. [76]. They prepared a dilute 

polymeric solution by the addition of carboxymethylcellulose and dropped molten lead, tin 

and bismuth into the solution. They observed that the interaction between the molten 

metal and solution was substantially reduced or totally eliminated at a viscosity ratio of 

approximately five. The viscosity ratio is defined in this study as the ratio of the viscosity 

of the solution to that of pure water. Nelson et al. [70] performed experiments by pouring 

12 g of molten tin at 923 K into 1.8 kg of aqueous glycerol or cellulose gum solution. They 

found that there was a threshold solution viscosity near a viscosity ratio of 15, above which 

spontaneous explosions no longer occurred. Also they performed one field-scale experiment 

in which 50 kg of molten Fe-A120 3 were poured into 190 kg of an aqueous solution of 

cellulose gum at the viscosity of about 0.080 Pa s. They observed that only a gentle 

breakup of some of the melt occurred. Kim[671 investigated the effect of external trigger 

strength on the explosion of molten iron oxide drops in aqueous solutions of cellulose gum 

and associated efficiencies. He applied external trigger pressures of 200 and 400 kPa and 

varied the viscosity ratio from 40 to 240. He observed that the explosion efficiency decreased 

from about 6 % to 0 % as the viscosity ratio of the solution increased from 40 to 240. Kim 

noted that the threshold for suppression was dependent on trigger strength. Recently, 

Dowling et al. [69] examined the ability of a dilute aqueous solution of polyethylene oxide 

(PEO) to suppress spontaneous explosions. Twelve grams of molten tin at temperatures 

from 600 to 1000 0C were poured into a solution-filled container at 25 0C with a viscosity 

ratio from 1.01 to 2.0. They showed that spontaneous explosions were markedly suppressed 

and entirely eliminated when the polymer solution was twice as viscous as pure water. 

Note that all this previous work has been conducted at small scales. Thus, it is important 

for us to verify that these suppressive effects can be realized at large scales. The analysis 

performed in conjunction with these tests suggest scale independence, but this has not 

been empirically observed. This will be one of the important parameters investigated in 

this experimental body of work. 

Radial Constraint 

The geometry surrounding the fuel-coolant mixture plays an important role in FCI ener

getics. Early research on explosion propagation recognized the role of the radial constraint. 

NUREG/CR-6623 18

1!
One of the experiments, Board et al. [16], produced a coherently propagating explosion us

ing two types of chambers with a different degree of constraint; one is a 'V' shape channel 

(or trough) and another is a narrow channel. Board produced a coherently propagating explosion with a narrow channel which had a higher degree of radial constraint. Previously mentioned, the RC tests performed by Sandia investigated the effect of the radial constraint 

on the vapor explosion. These tests produced the most energetic explosion recorded by the 

Sandia researchers. 

Recently, Frost et al. [77] investigated more systematically the degree of geometrical con

straint required to sustain vapor explosion propagation. In their experiments an array of 

melt droplets were injected into a narrow channel as a radial constraint (or confinement). The explosive interaction produced a shock propagation with a speed of about 100 m/s. 

Without the radial constraint, incoherent sequential explosions of the drops still occurred. 

If there was no confinement, the propagation velocity was decreased by an order of magni tude to 5-10 m/s. They found a high degree of radial constraint was required to sustain 

a propagating explosive interaction. Also, they observed interactions initiated in a radially unconstrained cylindrical geometry always failed to propagate after a short distance. In 

the KROTOS tests and our WFCI tests[78], this high degree of radial constraint has been 

maintained resulting in reproducible propagating vapor explosions. Conceptually, the pres

ence of a strong radial wall constraint in these one-dimensional experiments is necessary 

to simulate a "slice" of a larger scale fuel-coolant mixture which would provide the radial 

confinement in larger scale explosions that approaches a rigid wall constraint. In reality, 

at larger scales the compressibility of the surrounding mixture would reduce the rigidity of 

the radial constraint. The relationship between the scale of experimental geometries and 

fuel-coolant mixtures is a crucial factor in deciding the degree of geometrical constraint 

needed to bound the effect of larger scales and sustain propagations. It is a believed that 

a strong radial constraint maximizes the energetics for a particular fuel-coolant mixture at 

any given geometric scale. 

2.3 Key Findings of the Experimental Review 

* One of the key experimental results needed to understand the energetics of FCIs is the 

resulting work output. The data associated with precise measurement of this work 

output are extremely limited. Over the last two decades each experiment has used its 

own method to estimate the work output and because of experimental imprecision has 

led to large uncertainties. Such uncertainties in measuring the work output precludes 

any direct data comparison on key effects. This is the major improvement we plan 

to make with this work. 

e In FCI experiments, various materials are used as a simulant of the core melt. Even 

with a large amount of experimental data associated with various simulant mate

rials and variations in initial and boundary conditions, it is difficult to extrapolate their data to prototypic conditions. Again this is because of insufficiently defined 

measurements and differing definitions of work yield.

19 NUREG/CR-6623
9 Large amounts of experimental data are available from small and large scale experi

ments both triggered and spontaneous. However, experimental data associated with 

the systematic effects of the trigger on energetics in FCIs is very limited This is es

pecially the case in large scale tests where most experimental data are not obtained 

with well controlled measurements. 

* The study of vapor explosion suppression, using some additive, in large scale sys

tems is nonexistant. Historically, only the suppressive effect of ambient pressure are 

known. Most previous work has been performed on small scale geometries (single 

drop experiments). The effects of polymer additives on the FCI looks promising and 

it is necessary to verify scale effects.

NUREG/CR-6623 20

II
Table 2.1: Summary of the ANL Experiments

Test Fuel/Coolant Test-section Results 

Facility & Fuel TI/AT, b M,/M. H, Dp. Trigger Exp.;' Ppk V, a CR Remarks 

Run No. 9C kg cm cm Type MPa Type MPa n/s % % 

ANL : Chamber Type 

CWTI Pa : 0.1 MPa 

CWTI-09 Corium 2800/6 2.18/11 31 2.20 - NE3 0.245 

CWTI-10 Corium 2800/75 1.31/11 32 2.54 - NE 

CCM Pa : 0.1 MPa 

CCM-1 Corium 2800/43 2.15/37 106 2.54 - NE 0.139 - - ID : 0.21m 

Single jet CCM.2 Corium 2800/1 11.15/22 63 2.02 - NE 0.386 - - ID : 0.21m 

Four jets CCM-3 Corium 2800/0 3.34/39 110 2.54 - NE 0.389 - - ID : 0.21m 

Single jet CCM-4 Corium 2800/37 9.24/38 107 5.08 - NE 0.430 ID : 0.21m 

Single jet CCM-5 Corium 2800/45 11.34/525 107 5.08 - NE 0.246 ID : 0.76m 

CCM-6 Corium 2800/0 12.89/510 107 5.08 - NE 0.417 ID : 0.76m 

2

Explosion 

w sNo Explosion
Table 2.2: Summary of the SUW Experiments !'3 

C: C.

4Free Release 

5Restricted Release

Test Fuel/Coolant Test-section Results 

Facility & Fuel T,/AT,,b Mj/MC -He Dp. P. Trigger Exp. Ppk CR Remarks 

Run No. 9C kg cm cm MPa Type MPa Type MPa % 

SUW: Chamber Type, M./Mj: 63-188 

01 U0 2+Mo 3300/78 24/1500 0.1 - - SE 2.8 FR4 

02 U0 2+Mo 3300/87 24/1500 0.4 - - NE - FR 

03 U0 2+Mo 3300/80 24/1500 0.1 - - SE 1.6 FR 

04 U0 2+Mo 3300/80 24/1500 0.1 - - SE 2.3 RIt 5 

05 U0 2+Mo 3300/61 24/1500 0.1 - - SE 3.2 RR 

06 U0 2+Mo 3300/31 24/1500 0.1 - SE 3.1 RR 

07 U0 2+Mo 3300/0 24/1500 0.1 - SE 4.3 RR 

08 U0 2+Mo 3300/60 24/1500 0.5 - SE 2.8 RR 

09 U0 2+Mo 3300/60 24/1500 1.0 - SE 17.2 3.0 RR 

10 U0 2+Mo 3300/60 8/1500 0.1 - SE 1.1 RR 

11 U0 2+Mo 3300/60 8/1500 1.0 - - NE - RR 

12 U0 2+Mo 3300/60 8/1500 1.0 - - SE - RR

t•3
Table 2.3: Summary of the WUMT Experiments

Test Fuel/Coolant Test-section Results 

Facility & Fuel T,/AT, . M,/Mc 1I1 Dp, P. Trigger Exp. Vp CR Remarks 

Run No. 0C kg cm cm MPa Type MPa Type m/s % 

WUMT : Chamber Type, M,/Mi: 0.33-7.5 

01 U0 2 +Mo 3300/0 24/806 50 7 0.1 - - NE 

03 U0 2+Mo 3300/80 24/180 50 7 0.1 - - SE 

04 U0 2+Mo 3300/0 24/180 50 4 0.1 - - NE 

05 U0 2+Mo 3300/0 24/72 20 10 0.1 - - NE 

06 U0 2 +Mo 3300/0 24/20 50 4 0.1 - - NE 

07 U0 2+Mo 3300/0 24/8 20 10 0.1 - - NE 

08 U0 2+Mo 3300/0 24/180 50 10 1.0 - - NE 

09 U0 2+Mo 3300/0 24/180 50 10 0.1 - - SE 

MIXA: Chamber Type, V.: 5 m/s, Mc/Mj: 27 

Dimension: 0.37 m Vessel Side 

01 U0 2+Mo 3300/0 2.84/82 60 0.1 - - NE 

04 U0 2+Mo 3300/0 2.75/82 60 0.1 - - NE 

05 U0 2+Mo 3300/20 3.00/82 60 0.1 - - NE 

06 U0 2+Mo 3300/0 3.00/82 60 0.1 - - NE 

07 U0 2+Mo 3300/20 3.00/82 60 0.1 - - NE

t0

'Water mass was calculated by the dimension of the chamber given in ref. [33, 35]
z•

Table 2.4: Summary of the SNL Experiments (Part I)

Test Fuel/Coolant Test-section Results 

Facility & Fuel Tf/ATzb M//Mc H, Dp. Trigger Exp. Ppk Vp COR Remarks 

Run No. C kg cm cm Type MPa Type MPa m/s % 

EXO-FITS : Open Chamber Type 

MD V. : 5.3-6.4, Em: 3.3, P. : 0.083 

07 Fe-A1203 2800/- 4.80/75 45.7 - - SE - Base7 

08 Fe-A120 3 2800/90 4.70/258 71.0 - - SE >34 415 >1 Base 

09 Fe-A120 3 2800/91 0.60/56 61.0 - - NE 

10 Fe-A120 3 2800/90 1.31/56 61.0 - - NE 

11 Fe-A1203 2800/89 4.70/267 71.1 - - SE >34 365 >1 Above Base 

12 Fe-A120 3 2800/89 1.46/42 45.7 - - NE 

13 Fe-A1203 2800/89 1.69/66 30.5 - - NE 

14 Fe-A1203 2800/90 4.70/264 71.1 - - SE 555 Above Base 

15 Fe-A120 3 2800/90 1.88/90 43.2 - - SE 249 Base 

16 Fe-A120 3 2800/90 1.85/90 43.2 - - SE 429 Base 

17 Fe-A120 3 2800/82 2.85/152 53.3 - - NE 

18 Fe-A1203 2800/80 2.74/152 53.3 - - SE 9 313 Base 

19 Fe-A1203 2800/74 5.11/224 61.0 - - SE 17.5 427 Base

7

Spoutaneous Explosion occurred at the Base of the Chamber
Table 2.5: Summary of the SNL Experiments (Part II)

Test Fuel/Coolant Test-section Results 

Facility & " Fuel Tfi/aTub Mj/Me He Dpo Trigger Exp. Pph Vp rer/ Tee5 Remarks 

Run No. _C kg cm cm Type MPa Type MPa m/s mI

EXO-FITS : Open Chamber Type

V. : 2.4-5.9, M /M : 6-55 

Fe-AI2 0 3 2800/9 18.5/109.7 122 

Fe-AI2 0 3 2800/4 18.0/109.3 122 

Fe-A120 3 2800/3 " 18.0/437.0 122 

Fe-A12 0 3 2800/3 18.9/218.5 61 

Fe-A12 0 3 2800/4 7.6/218.7 61 

Fe-A12 0 3 2800/3 4.0/218.5 61 

Fe-A12 0 3 2800/73 18.5/169.6 46 

Fe-AI20 3 2800/2 18.6/218.4 61 

Fe-AI2 03 2800/3 18.6/218.6 61 

Fe-A12 03 2800/1 18.4/109.3 31 

Fe-Al 2 O3 2800/1 18.7/218.6 61 

Fe-AI2 0 3 2800/69 18.5/112.9 31 

V, : 3.3-3.8, Me/Mj: 11-24 

FeOy -/69 -/66.10 36 

FeOV -/69 9.00/100.9 36 

FeOV -/69 10.0/131.7 36 

Fe.O, -/4 9.00/218.6 61

33 

53 

28 

18 

23 

26 

19 

20 

20

- NE 

- NE 

- NE 

NE 

NE 

NE 

SE 

SE 

SE 

SE 

NE 

SE

CM 

01 

02 

03 

04 

05 

06 

07 

08 

09 

10 

11 

12 

OM 

01 

02 

03 

04

30/- Eruption 

73/- Eruption 

43/- Eruption 

18-89/- Eruption 

27/- Eruption 

22-159/- Eruption 

301 43/69,503 

37/216 

65/105 

43/112,311 

52/- Eruption 

37/69,125 

193 47 

785 141 

332 19-360 Multiple 

Explosions

RC Ve : 5.8, Me/Mj: 6 

01 Fe-AI20 3 2800/69 19.0/111.7 46 

02 Fe-Al2 03 2800/64 18.5/111.6 46

- NE - - 86/- Eruption 

- SE 76 1100 56/180

8Event Time of Eruption/Spontaneous Explosion after Melt Entry

S - SE 

S - SE 

S - SE 

S - SE

t4 

C)
Table 2.6: Summary of the SNL Experiments (Part III)

t4

Test Fuel/Coolant Test-section Results 

Facility & Fuel T,/ATub M//M, Hlc Dp. Trigger Exp. PPk CRT Remarks 

Run No. 0C kg cm cm Type MPa Type MPa % 

FITS : Chamber Type 

FITS-A Ve - , M,/M,:42-53 

1A Fe-A120 3 2500/85 1.94/90 43.2 - - SE Mild Explosion 

2A Fe-A12 03 2500/82 2.87/152 53.3 - - SE < 1 Surface Explosion 

3A Fe-A12 0 3 2500/81 5.30/226 61.6 - - SE 1-2 Above Base 

4A Fe-A1203 2500/80 4.29/226 61.0 - - NE - Pa: 1.02 MPa 

5A Fe-A120 3 2500/81 5.38/226 61.0 TE 1-2 P,: 1.09 MPa, 

0.64g PETN detonator 

FITS-B Ve :5.4-7.2, Mc/Mi:1.5-15, P.: 0.083 MPa 

1B Fe-A120 3 2500/70 18.7/226 61.0 5 - - SE 3.7 Double Explosion 

(5.1) at surface and base 

2B Fe-A120 3 2500/70 18.6/113 30.0 20 - - SE 4.4 Single Explosion 

3B Fe-A12Oa 2500/67 18.6/57 30.0 22 - - SE 5.3 Single Explosion 

413 Fe-A120 3 2500/69 18.7/226 61.0 18 - - SE 7.0 Double Explosion 

(9.0) at surface and base 

5B Fe-A120 3 2500/1 14.5/- 37.0 - - - NE 

6B Fe-A120 3 2500/1 18.7/63 30.0 - - - NE - Weak Interactions 

71B Fe-A12 0 3 2500/78 18.7/28 15.0 - - - SE 0.5 

7B11 Fe-A1203 2500/79 18.7/- 15.0 22 - - SE - EXO-FITS facility 

8B Fe-A12 03 2500/81 18.7/284 77.0 19 - - SE 7.4 Double Explosion 

(9.2) at surface and base 

9B Fe-A12 03 2500/80 18.7/170 46.0 20 - - SE 5.5 Single Explosion
Table 2.7: Summary of the SNL Experiments (Part IV)

Test Fuel/Coolant Test-section Results 

Facility & Fuel T,/AT. b M,/M l H. Dp, Trigger Exp. PpLk CR Remarks 

Run No. C kg cm cm Type MPa Type MPa % 

FITS : Chamber Type 

FITS-C Ve :, M,/Mi: 5-15 

1C Fe-A1203 2500/69 17.1/115 31.0 - - SE P.: 0.083 MPa, 

Surface Explosion 2C C-A+R 2500/72 16.0/227 61.0 - - SE Pa: 0.083 MPa, 

Surface Explosion, Base 

3C C-A+R 2500/68 11.5/128 38.0 - - NE P.: 0.083 MPa, 

Mild interaction 

4C Fe-A120s 2500/67 19.0/115 31.0 - NE Pa: 0.550 MPa, 

Mild interaction 

5C Fe-A1203 2500/69 19.5/115 31.0 - NE Pa: 0.520 MPa, 

Mild interaction 

FITSoD Ve :, Ml/M1 : 5-15 

OD Fe-A1203 2400/0 17.8/182 51.0 - - NE Pa: 0.085 MPa 

2D Fe-A1203 2400/169 19.0/95 66.0 - - NE P.: 1.100 MPa, 

Mild interaction 

2DR Fe-A12 03 2400/158 18.7/95 66.0 - NE Pa: 1.100 MPa, 

Mild interaction 

3D Fe-A1203 2400/37 18.9/16.2 15.0 - - NE Pa: 0.700 MPa, 

Mild interaction 

5D Fe-A12 03 2400/83 19.2/383 66.0 - SE Pa: 0.083 MPa 

8D Fe-Al 2O3 2400/0 19.5/21.3 15.0 - - NE Pa: 0.083 MPa

z
Table 2.8: Summary of the ALPHA Experiments

z4

9Dispersion Device

Test Fhel/Coolant Test-section Results 

Facility & Fuel T/,AT.ub M4/Me He Dpa frigger Exp. Ppk Vp a CR Remarks 

Run No. 'C kg cm cm Type MPa Type MPa m/s % % 

ALPHA Chamber Type 

Dimension: 3.9m D x 5.7m H, Mc/Mj: 10-80 

STXOO1 Fe-AI20 3 10/785 2450/20 100 20 - - NE Pa: 0.1 MPa 

STX002 Fe-Al 2 03 20/785 2450/16 100 20 - - SE Pa: 0.1 MPA 

STX003 Fe-A120 3 20/784 2450/19 100 20 - SE 1.10 Pa: 0.1 MPA 

STXOO5 Fe-AI 20 3 20/774 2450/27 100 20 - - SE P.: 0.1 MPa 

STXOO6 Fe-AI2 0 3 20/774 2450/25 100 20 - - NE Pa: 0.1 MPa, 

D.D9 

STXOO8 Fe-Al203 20/774 2450/186 100 20 - NE Pa: 1.6 MPa 

STX009 Fe-AI20 3 20/774 2450/16 100 20 SE 1.50 Pa: 0.1 MPa 

STXO1O Fe-Al2 O3 10/774 2450/24 100 20 - - SE 2.20 Pa: 0.1 MPa 

STX011 Fe-AI 20 3 20/774 2450/27 100 20 - - SE Pa: 0.1 MPa, 

D.D 

STX012 Fe-A1203 20/774 2450/184 100 20 - - NE Pa: 1.6 MPa 

STX013 Fe-AI20 3 10/774 2450/11 100 20 - - NE Pa: 0.1 MPa 

STX014 Fe-A12 0 3 20/816 2450/99 100 20 - - NE Pa: 0.1 MPa 

STX015 Fe-AI2 0 3 20/774 2450/171 100 20 - - NE P,: 1.0 MPa 

STX016 Fe-AI2 0 3 20/174 2450/22 100 20 - - SE 0.86 Pa: 0.1 MPa 

STX01T Fe-Al 203 20/174 2450/87 90 20 - - SE 0.66 Pa: 0.1 MPa 

STXO18 Fe-A12 0 3 20/174 2450/90 90 20 - - SE 20-25 3.33 Pa: 0.1 MPa 

STX019 Fe-AI2 0 3 20/392 2450/92 90 20 - - SE 42-50 5.67 Pa: 0.1 MPa 

D.D 

STX020 Fe-AI2 0 3 20/774 2450/92 100 20 - - NE 58-68 - P.: 0.1 MPa 

D.D 

STX021 Fe-AI203 20/697 2450/92 90 20 - SE 63-80 4.05 Pa: 0.1 MPa 

D.D in Water

00
TESTID

Melt: Atumint 

Mass [kg] 

Temp [K] 

Brake Disk 

Jet Diam [min 

Coolont:Water 

Mass [kgJ 

Depth [m) 

Temp [C] 

Fall Distance [mW 

Press [bar] 

Ves'l ID [mm] 

Gas Trigger

K26 K27 K28 K29 K30 K38 K40 K41 K42 K43 K44 K49 K50 K51 

1.0 1.43 1.43 1.45 1.52 1.53 1.47 1.43 1.54 1.5 1.5 1.74 1.57 1.8 

2573 2573 2573 2573 2573 2665 3073 3073 2465 2625 2673 2415 2200 2475 

yes yes yes yes yes no no no no no no no no no 

30 30 30 30 30 30 30 30 30 30 30 30 30 30 

7.2 7.2 7.2 7.2 7.2 34.5 34.5 34.5 34.5 34.5 34.5 34.5 34.5 34.5 

1.08 1.08 1.08 1.08 1.08 1.11 1.11 1.11 1.11 1.11 1.11 1.11 1.11 1.11 

60 90 80 20 20 21 17 95 20 22 90 20 87 95 

0.46 0.46 0.46 0.46 0.46 0.46 0.46 0.46 0.46 0.46 0.46 0.46 0.46 0.46 

1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 2.0 1.0 3.5 1.0 1.0 

95 95 95 95 95 200 200 200 200 200 200 200 200 200 

yes no yes no no no. no no no no yes no no no

RESULTS

Penet.Melt Jet 

Steam Explosion 

Max.Press. [MPa] 

Cony. Ratio [%] 

Total Debris [kg] 

Debris<0.25 mm [%] 

Debris<0.1 mm [%M

T5 TO TI TO TI T3 <T1 TO T2 TI <T2 T3 TO TO 

yes no yes yes yes yes yes no yes yes yes yes no no 

>26 -- >50 >104 >106 67 83 -- 57 117 65 82 ... . 

0.35 - 0.85 0.72 1.1 1.45 0.9 -- 1.9 1.3 2.6 2.2 1 .. . 

<1 1.1 1.36 1.5 1.43 1.5 1.45 1.43 1.5 1.5 1.5 1.7 1.55 1.8 

31 0.1 62 62 88 60 35 -- 60 47 75 60 -. 

22 .. 40 39 73 35 25 -- 35 30 50 35 -. ...

Table 2.9: Summary of KROTOS Experiments with Alumina

4 

0•
TETIDz 

C

Melt: U02ZrO2 

Mass [kg] 

Temp [1K 

Brake Disk 

Coolant:Water 

Mass [kg] 

Depth [m) 

Temp [C] 

Press [bar] 

Ves'l ID [mm] 

Trigger Press [MPa] 

Trigger Time [msec] 

RESULTS 

Avg. Jet Velocity [n/s] 

Melt Jet Penetration 

Max. Level Swell [mm] 

Mixing Press. Inc. [bar] 

"Steam Explosion 

Max. Press. [MPa] 

Prop. Speed [n/s] 

Conv. Ratio [%] 

Total Debris [kg] 

Mass Mean Diam [mm] 

Debris<0.1 mm [%]

NOTES: Jet Dlam.[mm] = 30; Fall Hr. [m] = 0.44; Free Vol. [cu.m] = 0.23

K32 K33 K35 K36 K37 K45 K46 K47 K52 W53 

81/19 81/19 79/21 79/21 79/21 80/20 79/21 79/21 80/20 80/20 

3.03 3.2 3.1 3.03 3.22 3.08 5.05 5.15 2.62 2.62 

3063 3063 3023 3025 3018 3105 3088 3023 3023 3023 

no no yes no no no no no no no 

7.1 7.7 7.7 7.7 34.2 36.3 34.1 34.1 34.1 34.1 

1.08 1.08 1.08 1.08 1.1 1.14 1.1 1.1 1.1 1.1 

78 25 90 21 23 96 17 18 17 17 

1 1 1 1 1 1 1 1 2 3.6 

95 95 95 95 200 200 200 200 200 200 

-- - 13 13 13 14.5 15.5 15 14.5 15.2 

.. . 980 990 977 1305 1020 1620 1090 1100 

. . - - 6.7 4.4 6.1 6.7 2.7 3.0 

T5 T4 -- TS T4 T6 T3 T4 T5 T4 

S.- 300 >500 >500 >500 156 169 

2.3 1.4 1.7 1.3 0.6 1.8 0.7 0.8 0.3 0.6 

no no no FCI no no yes FCI yes yes 

S. . . .. . . 20 16 17 35 

-.. . . . . 373 <200 321 454 

0.. . . . . . 0.04 0.01 0.02 0.05 

2.6 2.8 1.42 2.8 2.9 2.8 5.0 5.15 2.43 ... 

2.0 3.0 3.0 1.1 1.4 1.7 0.5 1.22 1.0 0.55 

1.0 1.4 1.4 9.0 1.0 2 24 7.0 14. 22.

9

Table 2.9: Summary of KROTOS Experiments with Corium

Trigger Vol.[cc] = 15
Table 2.10: Summary of the FARO Experiments

Test Fuel/Coolant Test-section Results 

Facility & Fuel TI/AT.,,b Mu/M. H. Dp, Pa L.S.IU a- Remarks 

Run No. VC kg cm cm MPa m % 

FARO 

L-06 U0 2+ZrO 2

1 1 2923/0 18/120 87 10 5.0 0.13 0.13 NE 

L-08 U0 2+ZrO 2 3023/0 44/255 100 10 5.0 0.41 0.29 NE 

L-14 U0 2+ZrO 2 3073/0 125/623 205 10 5.0 0.60 0.23 NE 

L-11 U0 2+ZrO 2+Zr 12 2823/0 151/608 200 10 5.0 0.80 0.29 NE 

C.0 

C: 

, 

1 Level Swell 

1180 w/o U0 2, 20 w/o ZrO2 

D 12 76 w/o U0 2 , 19.2 w/o, Zr0 2 , 4.1 w/o Zr
Table 2.11: Simulant Fuels and Coolant used in FCI Studies 

Properties UO2/ZrO 2/SS Fe/A1203 UxAly A120 3 Al Pb Sn H 2 0 

C' Tf,ref, K 2800 2800 1000 2800 1000 1000 1273 300 

Tmit, K 1700-2700 1700-2300 90541230 2313 933 600 505 

w/o13 , % 75/12/13 55/45 30/70 100 100 100 100 100 

m/o 4, % 51/17/32 69/31 5/95 100 100 100 100 100 

pf, kg/m 3 7660 3830 3620 3400 2690 10000 6507 997 

cpf, kJ/kg K 0.526 1.06 0.86 1.4 1.16 0.15 0.257 4.179 

pfCpf, MJ/ma K 3.997 4.06 3.113 4.76 3.12 1.50 1.672 4.166 

kf, W/m K 8.5 16.5 150 8 230 15 32 0.613 

Of, mPa/s 5 5 3 3.9 2 2 0.87 0.855 

orf, N/m 0.5-1.5 0.5-2 1.0 0.5 0.9 0.5 0.53 0.072 

in,,, kJ/kg 323 600 318 1000 396 25 59.5 P' 3.4 4 12.5 3.86 • 26 2.8 4.58 1 

ir., kcal/g-mol fuel 13.4 5.3 92 - 90 >0 >0 

ESp16 , MJ/kg 1.65 3.28 0.95 4.50 1.24 0.134 0.310 

Evo0 1

17, GJ/m 3 12.6 12.5 4.3 15.3 3.33 1.34 2.01 

'

3 Weight Fraction 

"14Mole Fraction 

"

5Constant, (kjpjc/ /kepcpc,) 0

° 5 

1

6Total Energy per Mass 

17Total Energy per Volume
Table 2.12: Large Scale Experiments with Tin (Part I) 

Test Fuel/Coolant Test-section Results 

Facility & TM 1T.ub /M e He Dpo Trigger Exp. Ppk Vp CR Remarks 

Run No. kg cu cm Type MPa Type MPa m/s % 

Hall [60], Tube Geometry 

2 750/5-15 0.5/0.523 85 DTis - TE 0.5 300-600 0.13-0.2219 Strong Tube 

4 660/5-15 0.5/0.523 85 IST20 - TE 0.45 300-600 0.15-0.24 Strong Tube 

5 720/5-15 1.2/0.523 85 IST - TE 2.4 300-600 0.06-0.09 Strong Tube 

7 720/5-15 1.2/0.35-0.4 85 IST - TE 16.5 300-600 0.06-0.09 Glass Tube 

8 6205-15/ 1.2/0.35-0.4 85 IST TE 6 300-600 0.06-0.1 Glass Tube 

THERMIR by Briggs [611 

008 597/93 2.0/20-6021 35 - NE 0.2 - Local Interactions 

009 602/90 2.0/20-60 20 - NE 0.2 - Local Interactions 

010 605/91 2.0/20-60 35 - NE 0.2 - Local Interactions 

044 503/88 20.0/20-60 20 - NE - Local Interactions 

with Frag'ed Debris 045 797/87 20.0/20-60 20 NE 0.05 - Mild Interaction 

THERMIR by Fry et al. 31) 

061 786/56 -/- 20 MB2 2 TE 0.37 35 Open Rect.Vessel 062 786/41 16/- 20 MB SE 0.90 47 Open Rect.Vessel 

096 800/24 -/- 45 - - SE 1.05 109 Narrow Vessel 

107 800/15 6/- 45 DT TE 1.70 81 Narrow Vessel 

109 800/14 6/- 45 DT TE 3.40 120 Narrow Vessel 

084 300/3 Coherent Interaction, Mf : 5 kg 087 306/2 Coherent Interaction, V,/V, : 4 

095 302/4 Triggered Coherent Interaction, V./Vf : 4 

097 300/4 Incoherent Interaction, Ve/V : 4 

100 302/5 Incoherent Interaction, Vc/V! : 4

18

Detonator, approximately 0.1 g of PETN 

"19

Conversion ratios were calculated by the ratio of the observed impulse (150 - 250 J) to the internal thermal energy of the fuel 2 0 lnduced spontaneous triggering due to 20 OC of water at the bottom of the test section 2 1

These numbers were roughly estimated by the sizes of the test section. 

22Mechanical blow

z 

-S 

0
Table 2.13: Large Scale Experiments with Tin (Part II) 

Fuel/Coolant STest Test-section Results 

Facility & T /AT bu M,/M H. 1 Dp, Trigger Exp. Ppk Vp a CR Remarks 

Run No. C kg cm cm Type MPa Type MPa m/s % % 

b- Baines [62] Tube Geometry, Dim: Im H x 2.95cm ID and 3m H of Expansion Tube CT04 670/10 - 10013 IST - TE 6.0 40024 No constraint 

CT05 700/21 - 100 IST - SE 4.0 1600 No constraint 

CT06 730/16 100 IST - SE 3.3 200 No constraint 

CT07 820/7 - 100 IST - TE 3.0 110 21 0.39 No constraint 

CT08 790/7 0.53/0.8725 100 IST - TE 9.0 240 23 0.63 Partial Constraint 

CT09 840/8 0.53/0.87 100 IST - SE 4.0 240 19 0.41 Partial Constraint 

CT10 830/7 0.46/0.87 100 IST - TE 3.5 140 17 0.49 Partial Constraint 

CT1l 830/6 0.62/0.87 100 IST - TE 3.0 160 30 Partial Constraint 

CT12 820/6 0.46/0.87 100 IST - TE - 130 0.19 Full Constraint 

CT13 820/10 0.66/0.87 100 IST - TE 4.0 120 0.32 Full Constraint 

CT14 800/6 0.57/0.87 100 IST - TE 3.3 70 20 0.18 Full Constraint 

800/11 0.78/0.87 100 IST - TE 3.8 150 10 0.21 Full Constraint 

CT16 830/15 0.82/0.87 100 IST - TE 3.5 180 10 0.38 Full Constraint 

KROTOS Tube Geometry, M,/Mj: 1.04 (0.77 assuming 20 % void fraction) 

K-21 1080/13 6.5/6.75 110 >5 Gas 12 TE 6.5 270 20-50 0.054 

23Assumed from the length of test section 

24Choosed largest velocity for propagation speed from the paper reported by Baines 

"25Calculated based on the test section volume
X-RiAYI ODERA"-R 

,M & CEL 

VIWFR 

YnI 

FC 

Figur 2.1 Scheaticmustrtionof th CORXIT aciitG[8

35 NUREG/CR-6623
Figure 2.2: Schematic Illustration of the MFTF Facility

NUREG/CR-6623 36

11
PIPE

GENERATOR

CRO -...---. CANGBL II II 

THREADED ROD 

TANK~ SUPPORT RIND BS U 

NONEYCOM3 BL=C 

Figure 2.3: Schematic ilustration of the Open Geometry Facility [40]

37 NUREG/CR-6623
r1

Camera 

View

L

Chamber

ire Transducers

i

Camera View

TOP VIEW

SIDE VIEW

Figure 2.4: Schematic Illustration of the EXO-FITS Facility [40]

NUREG/CR-6623 38

I

f-
Figure 2.5: Schematic Illustration of the FITS Facility [40]

39 NUREG/CR-6623
Conceptual Diagram of 

Melt Drop Steam Explosion Experiments 

Figure 2.6: Schematic Illustration of the ALPHA Facility [43,44,45]

NUREG/CR-6623 40

1!
Lower electrode 

Release Tube 

Closing Disc (W) 

Syl. y2 detectors 

Release tube 

(+ - 50rmm)

Mirror system dnive 

W-disc: 

GLUV~elance 

Protection valve SOI 

Main Inthers 4vl5--me u02 

3" (* - 327S=- Doubl.e num) rapbm) 710 ds 

30-..-Safesty vat"e (10 m PS) 

(0-Melt catcher 

F(vilu7he u tp s0.=28 m3r) to 802 

"•--• ~Hirnged-llap 

(4•z~ s100 mmn) 

TERMOS vesse 

+k-t A 710 mam) 

, Water container i (4'ex - 500 nun. 

#int -: 470 mam) 

A1203 bk4 r ~(thickness 105 mmr) 

13S- a. B ottc. Plate oam a"-- (debris catcher) 

-250 

Figure 2.7: Schematic Illustration of the FARO Facility [48,49]

41 NUREG/CR-6623
cn
0" 100 W x 

x 

w 80 O 

x 0 

•.0 0 o9 o' o I 

40 0 

00 0 80g 0 0 

z 0 0 0 

Z 0 0 0 1 40- 9O 0 0 

0 040 0 690 90 08 00 100 

09 

080 0 

~20- 0 ~ 00 0 

HORIZONTAL 

BOUNDARY 

400 600 800 1000 

INITIAL TIN TEMPERATURE,*C 

Figure 2.9: Temperature Interaction Zone for 12 g of tin dropped through 3 cm into boiled 

distilled water: 0 indicates an interaction: X indicates no interaction [52]

43 NUREG/CR-6623
/

-v CGas tire

Outer 

vessel 

Inner tube 

2-8 cm dia. 

85 cm Long.

Detonator

in thin tube 

expts. 2.3.

/

Tin at - 700 (: 

Water at - 90°C 

Pressure 

transducers 

Cold water behind 

p~lythene

S~diaphragm expts. 4,5. 

Figure 2.10: Schematic Illustration of the Hall's Facility [601

NUREG/CR-6623 44

J;
Motion tin in

Simit plate 

Fun -Horizontal cylinder 

ktB leo v Overflow Water 

lypous lull 

fell 1cliift 

full of walt,,, U Rubbr Cel 

pistonI 

• n l. • .__.Pressure 

tIrnsductri 

Figure 2.11: Schematic Illustration of the Baines' Facility [62]

45 NUREG/CR-6623
Chapter 3

Experimental Description 

3.1 Introduction 

The Wisconsin Fuel-Coolant Interaction Facility (WFCI) has been constructed at the Uni

versity of Wisconsin to support FCI research as originally sponsored by the US Nuclear 

Regulatory Commission (NRC), in particular to perform fundamental experiments in en

ergetic fuel-coolant interactions. The facility was designed with insights and improvements 

from past experimental facilities of Baines [62] at CEGB and Hohmann [128] at JRC Ispra 

in which controlled vapor explosion tests were conducted. The experimental test chamber 

was designed as an approximate one-dimensional geometry (L/D>10). It allows for con

trolled initial and boundary conditions and for measurements of fuel-coolant mixing and 

explosion propagation/escalation data such as mixture level swell, spatial and temporal 

shock pressure histories, explosion propagation velocities, expansion work and fragmented 

fuel debris distributions. 

3.2 Experimental Apparatus 

The WFCI facility was originally housed on campus in a below-ground bunker, outside 

the Engineering Research Building. The roof of this bunker was protected by heavy steel 

grates that allow for pressure relief, yet provide a barrier to the environment. At this site 

initial tests were performed including the WFCI-A series. Subsequently, the apparatus was 

moved to the larger remote laboratory, which was the Physical Science Laboratory Tan

talus building, about fifteen miles south at the University of Wisconsin Kegonsa Research 

Campus. The building is a below-ground vault with concrete walls, floor and ceiling. While 

the ambient conditions for the experiments at the previous site were difficult to maintain 

for consistency, the new site provided constant temperature and humidity. The general ex

perimental arrangement is given with side views in Figure 3.1, with the major components 

identified. Figure 3.2 provides a three dimensional view. The facility includes the furnace, 

transfer vessel, upper funnel test section and slide gate, explosion test tubes and quench 

tank and trigger. Figure 3.3 is a front view of the facility which helps to better identify 

the water loop and bypass loop. 

NUREG/CR-6623 46

ii
3.2.1 Heating and Transport System

Furnace 

The furnace as shown in Figure 3.2 consists of a cylindrical Kanthal Al filament heater 

embedded in a low "K" factor vacuum cast ceramic fiber insulator. The insulator provides minimum heat loss, high temperature capability and rigid structure. This tube furnace 

is then encased in a steel outer jacket with support brackets for mounting at a particular 

location with the ability to rotate or move in place. The furnace is mounted in a vertical 

position but can be rotated by a motor (Masterflex variable-speed pump motor, Model 

7018) with its own solid-state controller through remote operation to a nearly horizontal 

position. The bottom end section is closed with a round alumina plate insulator and the 

retort for the fuel fits through the top end section. The overall dimensions of the furnace 

are: outer diameter of 0.41 m; inner diameter of 0.15 m; length of 0.56 m; heating volume 

of about 8.3 liters. The temperature rating of the furnace is 1200°C with a heated length of 

0.46 m. The single-phase electric power supply consumes 7.4 kW at 208 V. A melt crucible 

(retort) is placed in the furnace to hold the fuel during the melting process as shown in 

Figure 3.4. This crucible is made of a ceramic tube (I-MUL900) housed in a stainless steel 

tube. The stainless tube is supported from below by an alumina brick and from above 

by a steel ring attached to a pouring tube which transfers the melt to the transfer vessel. 

At the top of the stainless steel tube is attached a round ceramic collar to reduce the 

heat loss which results from direct metal-to-metal contact. The furnace is heated in an 

inert argon atmosphere with temperatures being monitored by a K-type thermocouple; the 

temperature is controlled by a Single Zone Temperature Control System (ATS series 3821). 

For the iron-oxide experiments in the final test series, this furnace could not be used. Rather 

a furnace with a higher operating temperature (>1600 *C) was needed. We designed and 

fabricated the furnace box and control system with assistance from regional manufacturers. 

To accomplish this, a resistance heater fabricated from molybdenum-silicon-carbide was 

employed with an operating temperature up to 1800 *C. The insulation material was similar 

to that used in the current furnace; i.e.alumina and silica preformed sheets and cylindrical 

walls. Finally, to hold and deliver the molten iron-oxide, the crucible and retort was 

fabricated from solid blocks of boron nitride. Although this was an expensive approach, it 

provided the most reliable containment of this highly corrosive melt. The final result was 

a furnace with a similar design to our tubular furnace, but with a discharge hole in the 

retort. The detailed design of the furnace and its operating characteristics and difficulties 

are described in the Appendices. 

Water Loop 

To achieve the desired temperature, the water is heated in a boiler and then circulated by a

47 NUREG/CR-6623
pump (Grundfos Model UP15-42SF) through the bypass loop as shown in Figure 3.3. The 

boiler consists of a 60 liter capacity aluminum tank and four 0.8 kW rod heaters. To control 

the water temperature, one thermocouple is installed in the middle of the water loop. An 

automatic temperature controller (Omega Model CN9000A) is used for maintaining the 

water temperature at a preset temperature. 

Transfer Vessel 

Because of the furnace construction, the tin furnace required a vessel to receive the molten 

tin and transport it to the WFCI chamber. The oxide furnace was equipped with a blast 

shield and required no transfer vessel. Once the tin melt is at the desired temperature the 

furnace is tilted to a horizontal position and the molten tin pours into the transfer vessel. 

This transfer vessel is then transported to the test section as shown in Figure 3.2. Figure 3.5 

shows the transfer vessel and the spring-loaded bottom plug. The transfer vessel is made 

of alumina (I-COR909) with a spring-loaded stainless steel plug at its bottom. It contains 

a thermocouple for the measurement of the melt temperature. Fine ceramic powders are 

introduced for preventing molten fuel leakage from the gap between the transfer vessel 

bottom hole and the bottom plug. An alumina paper (Zircar-APA-2), with a thickness of 

1.3 mm, is also attached to the inside wall of the transfer vessel to prevent crack development 

from sudden thermal shock between the hot molten fuel and cold ceramic vessel wall while 

molten fuel is pouring from the furnace into the transfer vessel and being transferred from 

the furnace to the test section. This paper maintains its rigidity by an alumina hardener 

(Zircar-A17401) and sustains up to 1650 °0. The transfer vessel can be moved by a motor 

on a transport rail system to a position over the funnel above the test chamber. At 

the appropriate moment, a pin holding the spring loaded bottom plug is removed by a 

pneumatic piston, the plug falls away and a melt jet pours into the test chamber. 

3.2.2 Test Section 

Funnel 

The funnel as shown in Figure 3.2, is a stainless-steel type 304 schedule 160 with an outer 

diameter of 0.32 m, an expanded section of the test chamber. The funnel can hold a volume 

of about 0.038 m3 of water. As the melt pours into the test section any water displaced 

from the test tube, due to the fuel or vapor production, is collected in this funnel and the 

level swell meter (discussed in next section) records the overall water level swell. 

Slide Gate 

The slide gate system is composed of a top and bottom stainless steel 304 flange, 25.4 mm 

thick and 0.38 m long stainless steel 304 plate with an 87 mm diameter hole and an air 

NUREG/CR-6623 48

11
cylinder to pull the slide gate plate. The schematic of this system is shown in Figure 3.6. 

Teflon dry lubricant is applied on the entire surface of the slide gate to allow the slide gate 

plate to move smoothly and prevent the surface of the slide gate from scratching. After a 

designated time to allow the fuel and coolant to mix, the stainless steel slide gate separating 

the explosion tube and the funnel is closed, isolating the mixture from the environment: 

at this time the fuel-coolant mixture is isolated and resides in the explosion tube. 

Test Tube System 

The test tube system shown in Figure 3.2 is composed of two tubes; one is a stainless steel 

cylinder Type 304, 101.6 mm NPS, Schedule 160 (called the I-tube) with an outer diameter 

of 114 mm, an inner diameter of 87 mm, and a height of 914 mm, and has a volume of 

0.055 m3. The other is a stainless steel tube (called the T-tube) with the same material 

and dimensions except a different height of 546 mm, which contains a water volume of 

0.033 m3. The total volume of the test tube system is 0.087 m3 . Along the test tube 

are mounting locations for high-range quartz pressure transducers. Near the bottom of 

the I-tube, a stainless steel bypass pipe, shown in Figure 3.3, is also provided to allow for 

heating of the water in the explosion tube and bypass of water flow as fuel-coolant mixing 

occurs. This bypass piping, if open during testing, can also provide for the water coolant 

to flow out of the mixing region as the fuel enters the chamber. This creates an additional 

path for coolant up and around the fuel-coolant mixing region during the FCI and before 

the vapor explosion. 

Trigger Device 

The trigger device shown in Figure 3.7 is a movable piston located in the lower explosion 

tube end plate. The sudden motion of this piston generates a shock wave (up to -4 MPa 

for 170ps) in the water which then propagates into the fuel-coolant mixture to trigger the 

explosion. The piston is set into motion by the mechanical impact of an electromagnetic 

hammer from below. Initially the trigger system is prepared by charging a capacitor bank 

(0-500 VDC) at a preset voltage. If an external trigger signal is provided at a preset time, 

a sudden discharge of the capacitor bank initiates. This sudden discharge generates the 

electromagnetic force that causes the hammer movement. The trigger circuitry as shown 

in Figure 3.8 is used to provide the fast switching of high voltage and current from the 

capacitor bank. The capacitor discharge voltage, the hammer geometry, and the spacing 

between the piston and the hammer can be varied to tailor the initial pressure shock wave 

pulse height and width. 

The pressure shock generated by the trigger device is produced by applying a mechani

cal impact to the trigger piston using an electromagnetic force generated by the sudden 

discharge from the capacitor bank. The sudden motion of the electromagnet produces 

an impact to the trigger device via a stainless steel rod which is directly attached to the

49 NUREG/CR-6623
trigger piston disk. Motion of the magnet is actuated by an electrical circuit, shown in 

Figure 3.8, which consists of a variable DC power supply (0,500 VDC), a capacitor bank 

and a trigger circuit. In order to obtain a high current to the magnet, the capacitor bank is 

initially charged at a desired DC voltage and the fast switching to the magnet is made by a 

Silicon Controlled Rectifier (SCR) shown in Figure 3.8. The strength of the pressure shock 

is controlled by the applied DC voltage. The relationship between the charged voltage and 

corresponding pressure shocks will be discussed in a later section. 

In order to measure temperatures in the test tubes and detect the melt during the mix

ing phase of the fuel-coolant interaction, a total of eight exposed K type thermocouples 

with 0.02 inch diameter stainless steel sheath are mounted on the test tubes. All eight 

thermocouples, as well as three for the transfer vessel, and those at the bottom of the test 

section and water loop are connected with the 16 Channel Thermocouple FET Multiflexer 

(HP E1353A) which has a scanning rate up to 13000 channel/s and a built-in thermister 

reference junction. The signals from the multiflexer are digitized by the Digital Multimeter 

(HP E1326B) with 14-bit resolution, a maximum 14 kHz scanning speed and 1 Mbyte non

volatile memory. All the temperature measurement systems are housed in the HP E1301A 

B-Size VXI Mainframe. 

Expansion Tube 

Once the explosion is triggered the subsequent expansion is directed down the stainless steel 

expansion tube which is composed of three identical tubes (one with a different flange) in 

series (Type 304, 4 inch NPS, Schedule 160, class 600 flange, class 1500 flange, 87 mm inner 

diameter, 1.22 m long), as shown in Figure 3.2. At the end of the expansion tube, a rupture 

disk (a 1 mil thick Kapton film which bursts at about 0.23 MPa differential pressure) is 

attached to hold the coolant in the system during the mixing phase. The rupture disk 

can be easily ruptured to release the expanded volume of coolant in the system for the 

expansion phase. A movable piston, which has a magnet at the center, is initially located 

in between the test tube and the expansion tubes. As the vapor explosion pushes the piston 

down the expansion tube, ten coils, wound outside the tube, sense the location of the piston 

as a function of time. The distance between the adjacent coils is evenly divided in 0.41 m 

segments. Water and/or lead blocks can be placed within the expansion tube to vary the 

inertial constraint of the system for the vapor explosion process. The explosion expansion 

products are captured in the downstream quench tank. 

Larger Scale Test Section - VWFCI-2 

Two different test sections were used for the WFCI final test series (series K). The first 

test section is the same as that described above, referred to in this section as WFCI-1. The 

other test section is larger and has an inside diameter more than twice that of the smaller 

section. It will be referred to as WFCI-2 throughout this report. Both test sections are 

NUREG/CR-6623 50

11
rated to 3000 psig and are made of type 304 stainless steel pipe (NPS schedule 160). This 

larger test section was constructed to examine the effect of the lateral scale on the FCI, 

specifically with the iron-oxide melt. 

The WFCI-2 test section is made of 8 inch NPS schedule 160 type 304 stainless steel pipe 

with inside diameter 6.82 in (173 mm) and wall thickness 0.91 in (23.1 mm). It is similar 

in geometry to the smaller diameter tube and is divided into two pieces, an I-Tube and a 

T-Tube, that are secured with NPS class 1500 type 304 stainless steel flanges. The I-Tube 

is 36.0 in (914 mm) tall and is the lower of the two pieces. It is mounted directly above the 

piston and base flange and also has a flanged opening near the bottom for the water return 

loop. The WFCI-2 I-Tube has a 21.5 L volume. Five pressure transducers (P1 through 

P5, numbered from bottom to top) are mounted along its length on the same side as the 

expansion tube opening. The T-Tube is 21.5 in (546 mm) in height, is mounted on top of 

the I-Tube, and facilitates the connection of the expansion tube. The WFCI-2 T-Tube has 

a 12.8 L volume. Three pressure transducers (P6 - P8) are mounted along its length on 

the side opposite the expansion tube opening. 

3.3 Instrumentation and Data Acquisition 

3.3.1 Instrumentation 

The main physical parameters measured in the WFCI experiments are the transient pres

sures, temperatures and level swell histories as well as piston-slug displacement during the 

fuel-coolant interaction. Temperatures of the furnace, melt and water are monitored by 

K-type thermocouples located at the furnace, water loop, magnet trigger piston and test 

tubes. A video camera records the whole sequence of the experiment. Since the phenomena 

in this study are of a fast transient nature, a high speed, computer-based data acquisition 

system is used for digitizing and storing the analog signals. The following are details of the 

instrumentation and data acquisition system. 

Transient Pressure and Temperature 

The test section has been designed to accommodate up to eight high sensitivity piezo

electric pressure transducers (PCB series 112A03) for pressure signals during the explosion 

phase as well as the mixing phase. These are flush-mounted on the test tube. All locations 

and distances of transducers are shown in Figure 3.3. Each transducer is connected to a 

charge amplifier (PCB Model 462A) for converting electrostatic charge (1.0 pC/psi nomi

nally) into a voltage output. The charge amplifier has multi-range (unit/volt) capability, 

thus a higher resolution can be obtained in the digitization by adjusting the full voltage 

scale. To attenuate high frequency noise (>180 kHz), each charge amplifier has a plug-in

51 NUREG/CR-6623
filter (PCB Model 476A07). Low noise coaxial cables (PCB 003A35) are connected between 

the transducers and charge amplifiers. All pressure transducers have response frequencies 

greater than 10 kHz and the signals are recorded via a high speed A/D buffer interface (up 

to 1 MHz). 

Several thermocouples (Type K) are used to monitor the melt temperatures in the furnace 

and transfer vessel, as well as the water temperature in the test tube. In addition, ther

mocouples are placed in the explosion tube at various heights to detect the leading edge of 

the fuel jet as it pours through the explosion tube and mixes with water. Thermocouples 

for detection of the fuel jet during the mixing phase are mounted on the test tube and 

are selected by response time. One of these thermocouples is used to control the water 

temperature in the test tube and connected to a temperature controller (Omega Model 

CN9000A) which maintains the water at a preset temperature. 

Level Swell Meter 

A level swell meter, shown in Figure 3.9, is used to measure the level swell in the funnel 

above the test section during the fuel-coolant mixing phase. The integral vapor volume 

fraction during mixing can be calculated from the signal measured from the level swell 

meter. The level swell meter is composed of two copper tubes with an outer diameter of 

13 mm attached parallel to each other on the funnel side wall. The level swell with time 

is measured by the resistance changes between the two copper tubes. When the working 

voltage is supplied, however, the voltage generated by the level meter slowly decreases with 

time and also with the coolant temperature because of a slow degradation of the copper 

rod surface with time and temperature while electric power is supplied. The level swell, 

however, can be estimated by measuring the proportionality constant of the relationship 

between the water levels and the measured voltages from the initially known water level at 

the initial temperature and its corresponding voltages. 

In the current test series the inductive level meter was not always utilized because weak 

explosions were triggered with residual tin melt at the water surface above the slide gate. 

This altered the signal and at times destroyed the metering system. When no energetics 

interactions occurred we were able to obtain data which will be subsequently discussed. To 

improve on its reliability and accuracy, the meter should be repositioned so it can be used 

in future tests. The level meter (OMEGA LD300-250) is intrinsically AC operating LVDT 

with -10" linear stock, 0.25% full scale linearity and 2 ms frequency response. This meter 

incorporates an AC LVDT signal conditioner (OMEGA LDX-3) which provides ±5 VDC 

analog output. 

Slug Displacement 

The location of the piston as a function of time is measured by installing electrical coils 

NUREG/CR-6623 52

11
along the expansion tube that sense the magnetic core piston passing each coil location at 

a particular time as shown in Figure 3.10. This figure shows the series of signals generated 

when the piston passes through the series of coils. As the piston approaches the coil at 

point 1 as shown in Figure 3.10, the magnitude of the signal is positively increased. If the 

piston is just passing the location of the coil, the magnitude of the signal is rapidly falling 

down to zero as indicated at point 2. Finally as the piston begins passing away, point 3, 

the magnitude of the signal is rapidly increased in the negative value and then approaches 

zero asymptotically. Therefore, the exact time for the piston passing the coil is measured 

at point 2 as shown in Figure 3.10. From these data and the fact that all the explosion 

expansion products are directed through this tube, the explosion expansion work can be 

calculated. 

3.3.2 Data Acquisition and Experimental Control 

Figure 3.11 provides the overall facility layout and identifies the data acquisition systems 

and experimental control. A 486 PC based CAMAC system (LeCroy) is used for digitizing 

and storing the fast transient signals from the instruments. It includes four 12-bit digitiz

ers (LeCroy 8200) with a 32 K memory module (LeCroy 8800A). Each digitizer has four 

input channels with 1 MHz maximum sampling rate per channel. The A/D system, T/C 

amplifiers, trigger control, pressure transducer amplifiers are located in the off-site control 

room and the signals are transferred to the off-site computer. In addition the computer and 

off-site timing control circuit are used to automatically control the operation of any test. 

Once the melt is remotely poured into the transfer vessel and the transfer vessel is remotely 

moved into position above the test section, the experiment enters an automatic sequence 

controlled by electronic timing circuitry as shown in Figure 3.12. As the melt pours into 

the funnel, the timing circuitry provides a sequence of events to close the slide gate, isolate 

the bypass loop, and initiate two simultaneous trigger pulses. One of the trigger pulses is 

for actuating the external trigger system; the other is for triggering the data acquisition 

system to begin recording the mixing process and the explosion event. 

Timer 

Two sets of timer blocks (called TA and TB) are connected between the power lines to 

actuate the specified automatic experimental sequence. Each set of timer blocks has two 

timers (called TA1, TA2, TB1 and TB2) as shown in Figure 3.12. The timers in the first 

block, TA1 and TB2, count a preset time (1 s) from reception of the arrival signal of the 

transfer vessel from the limit switch on top of the funnel. After one second, the bottom 

plug beneath the transfer vessel is opened by the TA1 which allows melts to be poured 

from the transfer vessel into the test tube. At the same time the second block of timers is 

triggered by the TA2. The first timer in the second block, TB1, starts counting the preset

53 NUREG/CR-6623
time from reception of the TA2 signal and provides electric power to the slide gate system 

to be closed. Finally, the second timer, TB2, provides a signal to the external trigger and 

the data acquisition systems to be actuated. Each timer (335B Shawnee Timer) is a digital 

set and solid state timer with 10 ms resolution. 

3.4 Initial and Boundary Conditions 

Sixty experiments in separate test series have been conducted with tin as the simulant fuel 

material and water as a coolant to investigate the energetics of fuel-coolant interactions. 

In addition, we finished testing in the WFCI facility with a series using molten iron-oxide 

as the fuel. Each set or series of tests focused on particular initial or boundary effects on 

FMIs. The WFCI-A series tests were conducted for the repeatability of the WFCI facility 

with all conditions fixed. The WFCI-B series of tests were performed to investigate the 

spontaneous explosion conditions at the same experimental conditions as the WFCI-A series 

but without an external trigger. As one of the key boundary conditions, the effect of the 

external trigger strength on the energetics of FCIs was investigated in the WFCI-C series. 

The effect of a system constraint, another important boundary condition, was studied in 

the WFCI-D series. To study the effects of the initial conditions such as fuel temperature 

and coolant temperature, the WFCI-E and F series, respectively were conducted with 

varying temperatures. In the WFCI-G series, the effect of one coolant property, coolant 

viscosity, was investigated. For this test, water as a coolant was mixed with a certain 

amount of polymer to alter its viscosity. The WFCI-H series was performed by varying the 

mass of the fuel and other jet parameters to observe the effect of the fuel to coolant mass 

ratio on FCI energetics. Finally, we performed a series of twelve tests (WFCI-K series) 

with iron-oxide as the molten fuel. The purpose of these tests was to observe explosion 

energetics when a high temperature molten oxide is used at prototypic melt superheats. In 

the prepartion for this final test series a number of oxide melt delivery and melt preparation 

tests were conducted, but will not be reported in detail here. 

3.5 General Experimental Procedures 

3.5.1 WFCI-A, B, C, E, F Series 

Every series of tests has the same procedure to investigate the fuel-coolant interaction 

phenomena with changing specific initial and boundary conditions. After the set-up of the 

WFCI facility, water is heated up to a preset temperature (typically 86 0C) in a boiler. 

The pump circulates the heated water from the boiler to the test section through the water 

NUREG/CR-6623 54

11
loop. About one hour before the actual test the entire system (including water and test 

tube) is heated up to a preset temperature (about 3 hours from starting circulation of the 

water), the furnace is turned on to start heating the melt (tin) up to a preset temperature 

(normally 1100 0C). The argon gas is supplied to the retort in the furnace before the 

furnace is turned on in order to produce an inert environment. It takes about 1 hour to 

achieve 1100 °C for the molten tin. While the water and tin are heating, all instruments 

are prepared in stand-by condition. After all the instruments are set in stand-by and the 

preset temperatures of water and fuel are achieved, the video camera recorder is turned 

on to film the whole sequence of the experiment. First, the furnace is turned to pour the 

molten fuel into the transfer vessel via manual operation from the remote site control area. 

The complete furnace tilting process is indicated by a signal from the furnace limit switch 

located above the furnace. When completion of the furnace tilting process is assured, 

the transport process starts with the turning of another manual switch in the remote site 

control area. At the top of the funnel is another limit switch which indicates the arrival 

of the transfer vessel in the proper position. When the limit switch is activated by the 

arrival of the transfer vessel, the rest of the experimental sequence (including opening of 

the transfer vessel to pour the transported molten fuel, closing of the slide gate, closing of 

the bypass valve, triggering the external trigger system and the D/A systems) are governed 

by the timers described in the previous section. 

For all series of tests, the experimental procedures are similar but preparations are a little 

different for specific tests due to the different purposes of the experiments. In particular 

the following three series of tests, the WFCI-D, G and H will be discussed in more detail. 

3.5.2 WrFCI-D Series 

The system constraint for a typical series of tests is provided by filling up the expansion 

tube with water. The degree of the system constraint or slug mass is defined by the initial 

total mass in this tube. Since the vapor expands due to the explosion in the expansion tube, 

it is expected that the mass of the slug in the expansion tube would affect the explosive 

interaction. To adjust this degree of system constraint in the D test series, two different 

methods are introduced to vary the total mass of the slug: the first is to change the mass 

of water, installing the rupture disk in conjunction with the test tube and the expansion 

tube. This method provides the minimum degree of system constraint in this series of 

tests. The second is to add lead slugs with a diameter of 73 mm and a length of 100 mm 

in the expansion tube, as a more massive slug for the explosion with the slug length being 

maintained.

55 NUREG/CR-6623
3.5.3 VWFCI-G Series

Based on previous studies [69], Polyethylene-oxide (PEO) with an average moleculax weight 

of 4x106 , was selected as the polymeric additive in this study. Polymeric solutions with 

polymer concentrations -of 400, 600, and 800 wppm are prepared by slowly adding the 

polymer in the form of a powder into distilled water while the water was stirred using 

a motor-driven stirrer. The stirring continues for another 16 hours at slower speed until 

the polymer dissolves completely. The viscosity of the solution is measured with a size 50 

Cannon-Fenske type capillary viscometer. The viscosity ratios, 77,, of resulting solutions 

are shown in Figure 3.13 as a function of polymer concentration at 25 *C. Also in this 

figure, computed values using an empirical equation [129] are compared with the measured 

data. For each experiment, about 12 liters of polymeric solution is prepared for the desired 

concentration at room temperature. The polymer solution is poured into and stored in the 

test section at a temperature of 24 0C. 

Table 4.1 indicates that two different molten tin masses of 2.4 and 5 kg were used to examine 

the effect of mass scaling on the FCI suppression by polymeric solution. An external trigger 

of magnitude 3 MPa was provided only in the WFCI-G-02 test to examine the effect of 

an external trigger on the FCI with polymeric additives. Two tests with pure water were 

conducted as a control to compare the effects of polymeric additives. 

3.5.4 WFCI-H Series 

In this series of tests, to control the amount of the fuel mass mixed with the coolant during 

the mixing phase, four different methods are possible: change the fuel pouring time, control 

the bypass valve opening during the mixing phase, change the pouring diameter or change 

the inner diameter of the test section. Among these, the first three methods were initially 

employed. In the final test series (K series), we constructed the WFCI-2 explosion tube 

and thus, the inner diameter of the test section was also varied. 

The first method was introduced in the WFCI-H-02 and H-03 tests. In these tests, the 

pouring time is defined by the time period between the time of the bottom plug opening 

and that of the slide gate closing. The pouring time was increased from 1.5 sec to 1.9 sec in 

this test series. In the WFCI-H-04 test and most of the subsequent tests, the bypass valve 

was opened to allow more molten fuel to be mixed with water during the mixing phase of 

the vapor explosion. The bypass valve was then closed when the slide gate was closed. This 

would occur since the water would have an alternate flowpath in addition to counterflow 

upward through the fuel-coolant mixture. Finally, the fuel jet pour diameter was varied 

from 20 to 45 mm over these tests to produce a wider range of fuel pour rates. 

NUREG/CR-6623 56

1!
3.5.5 WFCI-K Series

The general procedure for experimental preparation was similar for these molten oxide 

experiments as in the case of the molten tin experiments. However, certain aspects of the 

melt preparation are distinctly different and need to be briefly explained. Up to the point 

of iron-oxide melting the process is quite similar. Approximately 1 hour into the water 

heating, the furnace was readied for operation. Pneumatic cylinders were activated that 

hold a safety catcher pan in place underneath the furnace and a radiation shield closed at 

the bottom to prevent heat loss. The power circuit was activated, and the power level was 

manually controlled to prevent large current changes through the elements. The power was 

ramped up to maintain a 120 amp current at all times. It took approximately 2 hours to 

reach 1700 0C. During this waiting period, the areas around the test section were covered 

with fiberglass cloth to protect from damage that could be caused by high temperature 

ejected molten debris. Two video cameras were positioned to tape record the experiment, 

and the data acquisition system was tested to insure all the desired data would be captured 

at the right time. 

When the melt was close to the desired temperature, the trigger capacitor bank was charged 

and the water heating circuit (if loop heating was used) was deactivated by shutting off the 

pumps and the heaters. Then the water level in the test section was lowered to a point near 

the bottom of the funnel. This was done to limit the amount of water the fuel jet must 

travel through to penetrate the test section, and it allowed for greater water level swell 

without overflowing. Then the rod that held the slug in place was removed, and when the 

furnace temperature reached 1700 °C, the test area near the test section was evacuated. 

The furnace catcher pan and shield were controlled remotely, and moved out of the way. 

Then the timing circuit was activated, and the test was initiated. Immediately after the 

test, the water was drained out of the test section and kept for processing. Then the test 

section and furnace were completely disassembled for cleaning and repair. The collected 

water was filtered through a 25 micron sieve to capture as much of the small debris as 

possible. Other remnants of iron oxide were removed from the bottom of the test section 

and scraped off the internal surfaces of the test section. All the debris was consolidated 

based on the location it was recovered from and weighed.

57 NUREG/CR-6623
Figure 3. 1: Overall Side View of the WFCI Facility

z.t4 )

. . I .. " . .. I.
Furnace

Transport Vessel

Sfide Gate 

Assembly

Expansion Tube

Coil

Funnel 

T-Tube 

I-Tube 

jriger Device

Figure 3.2: Overall Three Dimensional View of the WFCI Facility

59 NUREG/CR-6623
Transfer 

Vessel

Level Meter 

Slide 

-. Gate

Wat•er Heater ___ 6& 

6 &2 

Pump Bypass Vatve Trigger 

Figure 3.3: Schematic Illustration of the Front View of the WFCI Facility Including the 

Simplified Waterloop and the Location of the Pressure Transducer (Unit:Inch)

NUREG/CR-6623 60

1!
Top 

Side

Elbow

/Retort

Figure 3.4: Schematic Illustration of the Furnace (Unit: Inch)

NUREG/CR-6623

\ 2

61
Top View 

I

I 

Side View

Transport 

Vessel

Plug

Pneumatic 

Piston

Figure 3.5: Schematic Illustration of the Transport System 

NUREG/CR-6623 62

11

00 0 

0 

.0 

0 

0 0 0 0
Top View

0

O 

0

0 

0 0

0 

0

LTest Tube 

,/ Flange

Pneumatic

Side View

Figure 3.6: Schematic Illustration of the Slide Gate and Supporting Structure

NUREG/CR-6623

/ 

Slide 

Gate

m

63
Trigger

e

Capacitor 

Discharge

Figure 3.7: Schematic Illustration of the Magnet Trigger System

NUREG/CR-6623 64

11

P isto-n
OISCR 2N6405 CECC5558) 

R (j>741 

(2)741 

@3SCR ECG5547 

10 F Units 

Resistance ohmq 

Capadance micro F

00 

01 

C+)

IN4007 100

*-12IV DC

+12V DC

+12V DC

CA'

450V DC

C&3
In The Funnel

Water

Figure 3.9: Schematic Illustration of the Level Swell Meter

NUREG/CR-6623 66

1l

Copper Tuibes
0 200 % 60 800 1000 1200 

Thie(ms=c)

1400 1600 1800

IC 

-14

0. 

5.

D" 

0 20 40 k i 0 0 I8 0 120 I40 160 Ii O 200 220 240 

mc (ol) 

Co~il

Magnet Z

I Piston

Figure 3.10: Typical Piston Signal Measured in the Expansion Tube 

67 NUREG/CR-6623

2D

10 

10' 

20

.1

-2
Figure 3.11: Overall Layout of the WFCI Facility in a Top View

NUREG/CR-6623 68

1!
Figure 3.12: Schematic Diagram of the Control System

69 NUREG/CR-6623
* Poly(ethylene oxide) solution at 25 OC.. . . 

2.75 

S... . . ... ,,,:,•.. . .. ... ;......... ..... 

: calculation 

2.50 .--. measured : .......S.........i ..........i .. ..i .. 

2.25 

2.00 . 

......... ....... i ......... .... .. , ............. ..... 

1.75

1.50 1 : . . . . . . . .

1.25 

1.00

0 100 200 300 400 500 600 700 800

Concentration, wppm 

Figure 3.13: Viscosity Ratio with respect to the Polymer Concentration 

NUREG/CR-6623 70 

II

0

0 

A 

0o 

o

/

I f:1s
Chapter 4

Experimental Results and Discussion 

4.1 Introduction 

The WFCI experiments involved several test series ; trigger characterization and series A to 

H and K. The first three series of tests, i.e., triggering and series A and B, were performed 

as reference experiments. Among them, the triggering experiments were conducted to 

characterize the external trigger by varying the voltage charge to the capacitor in the 

external trigger system. The strength of the external trigger is determined by the amount 

of this voltage as well as pulse shape. The WFCI-A series was performed to verify the 

reproducibility of the WFCI facility by repeating a group of tests with the same initial 

conditions. Those initial conditions were intentionally similar to those of the KROTOS-21 

experiment [63]. To obtain well controlled experimental data, it was also necessary to verify 

the possibility of spontaneous explosions under similar initial conditions. For this purpose, 

the WFCI-B series was conducted at the same initial conditions without the external trigger 

to examine the possibility of the spontaneous explosion. 

The WFCI-C series was designed to determine the effect of trigger pulse strength on the 

vapor explosion behavior. In the WFCI-D series, the effect of the axial constraint on 

the vapor explosion was examined by varying the system constraint mass. The effects of 

fuel and coolant temperatures were investigated in the WFCI-E and WFCI-F test series, 

respectively. The WFCI-H series was conducted to examine the effects of the coolant to fuel 

mass ratio. To vary this coolant to fuel mass ratio, the molten fuel mass and jet diameter as 

well as its pouring time were varied. In addition, the test section bypass valve was opened to 

allow more fuel-coolant mixing during the fuel pour into the test-section. In contrast to the 

above experiments, the WFCI-G series of tests was conducted to investigate the possibility 

of vapor explosion suppression in a large scale geometry by changing a particular property 

of the coolant with an additive. 

Finally, the WFCI-K series was performed with a different molten fuel composition, molten 

iron-oxide. The reason for this change is directly related to material scaling issues for these 

tests. Molten tin was known to produce vapor explosions when poured into water. Thus, 

the intent of the experiments using this fuel was to provide benchmark data to characterize 

explosion energetics under well-controlled conditions. However, it was recognized that the 

use of molten tin as a fuel simulant for the prototypic fuel melt in a severe accident had

71 NUREG/CR-6623
limitations; e.g.low internal energy, high melt superheat and metallic fuel properties. Our 

material scaling considerations suggested that a molten oxide with modest superheat and 

similar properties to corium would be a better simulant. Thus, WFCI-K series was designed 

to deliver a similar volume of fuel as in the previous tests to the WFCI test section but 

with molten iron-oxide as the fuel simulant. 

Initial and boundary conditions for all of the test series are shown in Table 4.1. A typical 

event timing is shown in Figure 4.1. In most experiments the slide gate was closed 1.5 s 

after the molten fuel pour began, then the trigger was activated less than 0.5 sec after slide 

gate actuation. However, this time sequence was slightly altered, if needed, in each exper

iment; for instance, WFCI-A-01 test and WFCI-H series as shown in Figures 4.2 and 4.3, 

respectively. Time zero is defined as the time when the control system simultaneously 

activates the trigger device and the data acquisition system. The pressure history plot is 

composed of eight different plots representing the pressure from the bottom to the top of 

the test section. 

4.2 Trigger Characterization Experiments 

To examine the characteristics of trigger pressure pulse, a large number of trigger pulse 

tests were performed. The external trigger system, composed of the capacitor bank, the 

mechanical hammer device and the piston in the test chamber, provided the external per

turbation to the fuel-coolant mixture resulting from the sudden discharge of the capacitor 

bank. Various strengths of the external trigger are achieved by varying the supplied voltage 

of the capacitor bank; i.e., 100 to 400 V. All experiments were conducted in the test section 

filled with water at 15 to 20'C. 

Figure 4.4 shows the relationship between the trigger pressure at different locations and 

the supplied voltage. The peak pressures are in a range of 2.0 to 4.0 MPa with respect to 

the supplied voltages of 100 to 400 V. The attenuation of the peak pressure is fairly linear 

up to 0.8 m from the bottom of the test section. 

Figure 4.5 shows typical trigger pressure histories for capacitor charge-up voltages of 200 

DCV. The first pressure transducer, located 0.12 m above the bottom of the test section, 

detected the peak pressures of 3.2 MPa. The rise time of the trigger pressure is usually 

less than 100 Its and the full width at half maximum of the pressure is approximately 100 

jts. The fast Fourier transform analysis of these pressure signals shows that the dominant 

frequencies range from 0.2 to 0.4 kHz and harmonic behaviors were clearly observed result

ing from the reflection of the pressure waves against the rigid boundaries. The reflection 

waves generate several "wiggles" after the primary peak pressure propagation. These trig

ger pulses can be easily distinguished from the explosion pulses which have wider pulse

NUREG/CR-6623 72
width, slower pulse rise time and more noise resulting from the wave reflection. 

The appearance of peak pressures at the first pressure transducer are delayed approximately 

from 6 to 14 ms due to the supplied voltage after the external trigger system is triggered 

as shown in the Figure 4.6. The propagation speed of the trigger pulse is about 1500 m/s 

which corresponds to the sound speed of water. This speed is weakly dependent on the 

supplied voltage as expected. 

The impulses of the trigger pressure measured at the first pressure transducer range from 

0.8 to 1.6 kPa-s. The impulses do not vary with the supplied voltages because most trigger 

pressures have characteristics similar to each other except for the magnitude of the peak 

pressures. Instead of characterizing the trigger pulses with their impulses, the product of 

the peak pressure and impulse may provide a more reasonable indicator for the strength 

of the external trigger, as noted by Berman et al. [1], because this product represents the 

energy density of the external trigger pulse. Figure 4.7 shows that the product of the 

peak trigger pressure and its impulse ranges from 3 to 6 MPa kPa-s. This energy density 

is equivalent to 2 to 4 Joules per liter of water and is negligible compared to the energy 

density of the molten tin fuel simulant; i.e., 2 MJ per liter. Also note that this figure of 

merit increases with the increase of the supplied voltage up to 250 V and increases more 

slowly as the voltage increases to 400 V. 

4.3 Conversion Ratio 

The conversion ratio has widely been used as a prime measure for the energetics of a fuel

coolant interaction. We discuss its general concepts here, since it can be directly determined 

from the data, and is integral to our understanding of the FCI. It is defined by the ratio of 

the work done by the fuel-coolant mixture, W, to the initial total fuel thermal energy, Ef 

given by 

CR= (4.1) 

The initial total fuel thermal energy Ef is usually given as 

Efj = ms {cp,liq (Tri - T,,p) + cp,.ol (Trap - T.) + ifs, (4.2) 

where m! is the total fuel mass, cp is the specific heat of molten material involved in the 

explosion, Tfi is the initial molten fuel temperature, T, is the coolant temperature, T ..p 

is the melting temperature of the fuel and if,, is the specific latent heat of fusion of the

73 NUREG/CR-6623
molten material. If the fuel can undergo chemical reaction (fuel oxidation), then there 

would be an appropriate term added for E1 ,j. 

In order to determine the value of CR, it is necessary to estimate how much work has been 

done by the fuel-coolant mixture. Several experimental methods have been introduced 

to estimate the work output from the fuel-coolant interaction. In single drop or small 

scale experiments, the fuel-coolant mixture expansion during the explosive interaction was 

observed by high speed cinematography and the work was estimated by graphically inte

grating the pressure and bubble volume traces. In this case, the work for the expansion of 

the fuel-coolant mixture can be calculated by 

W PdV (4.3) 

assuming that the expansion process is reversible. 

For large scale tests, however, it is very difficult to visualize the total fuel-coolant mixture 

expansion. Some of the EXO-FITS tests performed at SNL determined the work output 

using the expansion velocity of the fuel-coolant mixture from the analysis of high speed 

film data. Because of large uncertainties in the analysis of the film data, the work outputs 

obtained by this method gave only rough estimates. In the OG (Open-Geometry) series of 

tests as one of the experiments performed at SNL, crushable aluminum honeycomb block 

was placed underneath the interaction chamber. The work output was estimated by the 

degree of deformation of the honeycomb when the interaction energy was driven downward 

by the explosion. In the FITS series of tests, the work output was estimated by the chamber 

pressurization using equation 4.3. However, since the estimation was largely varied with 

the assumption of the relationship between the pressure and volume and the expansion 

process [133], this also provided only a rough estimate. 

In this work, the work output is more directly determined by measuring the expansion 

speed through the expansion tube during the explosion. Figure 4.8 shows the schematic 

diagram for measuring work done by energetic FCIs in an interaction volume on the slug in 

the expansion tube. The slug as shown in Figure 4.8, is defined by the entire water in the 

expansion tube including the magnet embedded piston. When the interaction volume starts 

to increase due to FCIs, the piston moves and a part of the initial slug escapes from the 

expansion tube. Therefore, the slug mass, m.(t), in the slug control volume decreases. The 

position of the slug as a function of time in the expansion tube is provided by measuring a 

signal of the piston as the piston passes each of the ten electrical coils wound outside the 

tube. The first coil is located at the initial position of the piston to detect the initial time 

of the piston movement and the rest of 9 coils are located at the expansion tube. These 

ten data points are fitted to a polynomial function, x8(t), and the slug acceleration, a,(t), 

NUREG/CR-6623 74

1t
as a function of time is obtained by differentiating x.(t) twice with respect to time. From 

this position of the slug, the mass of the slug, m.(t) is also determined as illustrated in 

Figure 4.8. Hence, the work done on the slug at any time t after the slug starts to move, 

is given by the following equation, 

W-(t) = I F(t')v(t')dt' =ojmo(t')v(t')a,(t')dt' (4.4) 

where F,(t) is a net force acting on the slug. The maximum work done on the slug is 

obtained by calculating the above equation until t = ta when the acceleration of the 

piston reaches zero. Therefore, the conversion ratio is obtained by the following equation 

using the measured maximum work done on slug as the work output, neglecting work losses 

by friction and drag, w w. 

CR = E, 1 (4.5) 

4.4 Reproducibility Tests: WFCI-A Series 

The WFCI-A test series was performed to investigate the reproducibility of vapor explosion 

under nearly identical initial conditions in the WFCI facility. The initial conditions chosen 

are quite similar to those of the KROTOS-21 test. The reason for this choice is to obtain a 

comparison between similar experimental facilities, with similar fuel and coolant conditions 

to determine if the explosion is reproducible and to understand any observed differences. 

The tests involved an initial molten fuel mass of five kilograms at a temperature of 1120 

to 1300 K poured into the test section filled with distilled water through a 38 mm orifice. 

The actual mass of the fuel below the slide gate in the test tube was found to be an average 

of 3.5 kg after the experiments. The water in the test section had a mass of 8.6 kg and a 

temperature of 358 to 360 K with a depth of 1.48 m below the slide gate. An additional 

water mass of 23.4 kg as an inertial constraint was provided in the expansion tube and 

separated from the explosion tube by the plug. In all the tests, the bypass loop was closed 

just before triggering while the slide gate was closed 1.5 s after the fuel pour began. 

Table 4.1 shows that both spontaneous and triggered vapor explosions were produced in this 

series of tests even though experimental conditions of all experiments were initially set to 

be equal. Some of tests in this series were performed outside the building at temperatures 

below 20 °C as shown in Table 4.1. This cold weather caused some variability in the 

experimental conditions and may have contributed to experimental scatter. The coolant 

temperature varied only one degree and the fuel temperature at melt delivery varied about 

one hundred degrees. This latter variation may have contributed to the spontaneous trigger 

delay; i.e., spontaneous versus triggered explosions. Another inconsistency was caused by

75 NUREG/CR-6623
the location of the measurement of the coolant temperature just above the piston trigger. 

The thermocouple was positioned below the recirculation line inlet to the test tube and 

thus, may be at a lower temperature than the rest of the test tube due to temperature 

stratification. The inside coolant temperature was assumed to be equal to the temperature 

measured at the outlet of the water circulation loop. Thus, in the case of the spontaneous 

explosions, the coolant temperature may have been lower than the reported temperature, 

since the melt may have penetrated below the recirculation line (about 120 mm from the 

bottom). Even though spontaneous explosions occurred, the slide gate did close completely 

in all tests and isolated the explosion from the environment, allowing the explosion to 

propagate through the explosion and expansion tubes. 

In all experiments, most explosion characteristics were similar to each other. First, a double 

explosion was observed; i.e., the explosion was triggered at the bottom of the mixture and 

then a second explosion propagation wave moved downward growing in strength through 

the mixture as the fuel-coolant mixture began to expand. Additional propagations may 

have been present, but could not be distinguished from the pressure data. This observation 

should not be that surprising since the explosion is not just constrained by the mixture 

constituents as in all past FCI tests; i.e., the fuel-coolant mixture expansion is controlled 

by the slug constraint which can be independent of the coolant depth. The phenomenon 

is considered a double explosion, since the explosion continued to escalate in strength as it 

propagated up the explosion tube and back down from the upper slide gate. 

This explosion escalation is similar to observations for the KROTOS-21 test (shown in Fig

ure 4.9) as were the overall transient pressure histories. The results from both spontaneous 

and triggered explosions indicated that measured peak pressures range from 2 to 8 MPa, 

but the pressure plateaus (analogous to a Chapman-Jouguet (C-J) pressure) observed after 

the explosion peak are always in a similar range of 2-,3 MPa. These magnitudes of the 

quasi-steady pressures are very similar to those measured in the KROTOS-21 test. The 

explosion propagation velocities of about 200 m/s are also quite similar to each other. The 

conversion ratios for the WFCI-A series of tests are in the range of 0.2 to 0.7%. These 

results are also comparable to the KROTOS-21 test with a low conversion ratio of less 

than 0.2% 

In the WFCI-A-01 test, the slide gate was closed 1.5 s after the melt pour began, then the 

trigger was actuated 0.5 s after slide gate actuation, while in subsequent tests this latter 

time was reduced to 0.4 s to assure explosion triggering before any spontaneous trigger may 

have occurred (Figures 4.1, 4.2). In these experiments it was found that the explosion was 

quite sensitive to the timing of the external trigger. In all these tests the external trigger 

produced a pressure pulse of about 3 MPa over a time span of 0.1-s0.2 ms. In the first test 

the time delay was 500 ms after the closure of the slide gate which isolated the explosion 

chamber from the environment. This allowed a spontaneous explosion to occur about 80 

ms before the trigger was provided, as shown in Figure 4.10. In this case, the explosion 

NUREG/CR-6623 76

I1
occurred near the PT3 transducer, about 0.45 m above the bottom of the test section. The 

pressure wave propagated upward and downward simultaneously. In the upward direction 

the wave has an average propagation velocity of 250 m/s. The downward propagating 

explosion occurred at the top of the test section with a peak pressure of 4.2 MPa at t =-78 

ms and proceeded downward. 

In all subsequent tests the time delay was 400 ms and this successfully allowed for external 

triggering; e.g., as in the WFCI-A-02 test as shown in Figure 4.11. As a typical case of 

a triggered explosion, the explosion in the WFCI-A-02 shown in Figure 4.11 is induced 

by the external trigger shown at the PT1 with about 2.9 MPa at t =7.9 ms. The actual 

explosion due to the external trigger occurred at PT3 and escalated with peak pressures 

of 2.5 MPa to 3.9 MPa and velocity of 333 m/s. After about a 10 ms delay a downward 

propagating explosion occurred and propagated from the PT7 to PT1 with an average propagation velocity of 180 m/s. This explosion shown in this test is stronger than the 

first propagation. This might be explained considering that a more appropriate mixing 

condition for the onset of the second explosion is achieved by the first explosion. This will 

be discussed in further detail in a later chapter. 

In the WFCI-A-03 test the vapor explosion occurred just before the trigger and may have 

been caused by the slide gate impact upon closure, triggering the explosion about 50 ms 

after the slide gate closed. The pressure signal was not properly measured in this test but 

available information suggests the same trend of an upward and then downward propaga

tion occurred. In the WFCI-A-04 test, as shown in Figure 4.12, the vapor explosion was 

triggered spontaneously at about 100 ms. Such a delayed spontaneous explosion was caused 

by the failure of the external trigger which was identified by the pressure signal recorded 

at the first pressure transducer. There was no trigger pressure-like signal in the pressure 

transducer of PT1. However, the first spontaneous explosion occurred at the bottom of the 

test section. The pressure traces also clearly show the same explosion behavior. 

In the WFCI-A-06 test, eight pressure transducers were mounted in the test section and 

provided a much better spatial resolution of the explosion pressure behavior as shown in 

Figure 4.13. The traces show more clearly the explosion behavior; i.e., upward and then 

downward propagation. The first explosion occurred at about 0.8 m above the test section 

and propagated in both directions. The second explosion propagated from the top to 

bottom of the test section escalating its strength. 

Our goal was to produce energetic FCI events under controlled conditions for this first set 

of experiments. The results indicate qualitative and quantitative agreement with the one 

published KROTOS experiment using molten tin, KROTOS-21. In addition, the explo

sion behavior is consistent between the six experiments whether spontaneous or externally 

triggered. The WFCI-B test series further examines spontaneous explosions and the obser

vation of multiple propagation events.

77 NUREG/CR-6623
4.5 Spontaneous Explosion Tests: WFCI-B Series

To further characterize the spontaneous explosions observed in WFCI-A, additional ex

periments in the WFCI-B series were performed without any external trigger under the 

same initial and boundary conditions as the WFCI-A series. In the first experiment of the 

WFCI-B series, the spontaneous explosion occurred about 1200 ms after the D/A system 

was triggered. Even though the pressure signal could not be recorded because of an over

flow of the of D/A data storage, the movement of the magnet piston in the expansion tube 

was measured and provided an estimate of the explosion timing. The average slug expan

sion speed was about 10 m/s which was a lower speed than any observed in the previous 

WFCI-A series. After the first experiment, it was decided to decrease the sampling rate in 

the D/A system in order to catch the pressure signal while measuring a wider time range. 

The disadvantage of this choice is a loss in the time resolution. In the second experiment of 

WFCI-B series, explosion pressures were successfully measured. A spontaneous explosion 

occurred about 660 ms after the D/A system was triggered, which was earlier than in the 

WFCI-B-01 test. 

Figure 4.14 illustrates the characteristics of WFCI-B-02 pressure signals. Once again an 

upward explosion propagation and then a downward explosion propagation is observed. In 

Figure 4.14, the first peak appears about 663 ms after the D/A system is triggered at PT1 

transducer. Considering a falling speed of the molten fuel estimated from the WFCI-A-06 

test, the leading edge of the fuel had already reached the bottom of the test section and 

started to accumulate on the chamber base. This pressure peak propagated through the 

test tube up to the top transducer with an average velocity of 216 m/s. Following this 

relatively mild explosion (-s2 MPa), another larger explosion (-•5.7 MPa) occurred at the 

top of the test section, and then propagated down with an average velocity of 94 m/s. 

This second explosion occurred in the same manner as in the previously discussed WFCI-A 

series. The pressure histories show that the magnitudes of the peak pressures at the top 

and bottom are much larger than those at the tube mid-section. This may be explained 

by the fact that a larger amount of vapor was present in the middle of the test section 

due to the fuel-coolant mixing and the upward propagation. Therefore, the pressure wave 

propagating downward from the top was attenuated to some extent. 

In the WFCI-B-03 test, shown in Figure 4.15, (equipped with eight pressure transducers 

in the test section) an explosion occurred spontaneously near the location of the PT3 

transducer. The pressure wave travelled with propagation velocities of approximately 660 

m/s and 156 m/s in the upward and downward directions, respectively. In this test the 

behavior of the pressure was very similar to those in the WFCI-A tests and WFCI-B-02 

test even though the exact time of the spontaneous explosion was different. In these three 

experiments, the average expansion velocities were measured as 10.2, 12.1 and 17.6 m/s 

respectively. 

NUREG/CR-6623 78

11
In the WFCI-B tests, the kinetic energies of the spontaneous explosions were somewhat 

smaller than in the other tests (shown in Table 4.1), even though similar quantities of 

fuel were initially used in each test. The conversion ratios ranged from about 0.15 to 0.45 

percent. These explosions from these spontaneous events occurred relatively late; usually 

later than 0.5 s after the time of external trigger is normally applied in other series of tests. 

Since the external trigger was not used in this test series, the fuel cooled for a longer period of time than in other experiments, causing the fuel temperature to be lower at the time of 

the explosion. If the WFCI-A and WFCI-B test series are taken together we have a data 

set in which qualitatively similar behavior is exhibited under spontaneous and externally 

triggered circumstances. 

4.6 Effect of External Trigger: WFCI-C Series 

To examine the fuel-coolant interaction phenomena, a controllable trigger was used to 

synchronize the interaction and aid in reproducility. Also, since the external trigger pressure 

is an important initial condition to the energetic FCI, it is important to investigate its 

effect on the vapor explosion. A total of six experiments were performed with different 

magnitudes of trigger DC voltages of 100, 200, 300, 400 and 450 V which corresponded to 

trigger pressures of about 1.3, 2.4, 3.2, 4.0 and 4.3 MPa, respectively. 

First, the WFCI-C-01 test was conducted with a supplied voltage of 200 V in the trigger 

system, equivalent to about 2.4-3.0 MPa of trigger pressure. The initial conditions for 

this test were the same as the WFCI-A series, and it also showed the similar double ex

plosion behavior. The first pressure peak propagated upward with an average velocity of 

210 m/s and the average expansion velocity was measured at about 15.8 m/s which was 

comparable with the results of the WFCI-A series. WFCI-C-02 was similar to WFCI-C-01 

from inspection of pressure histories, however, all of the expansion data was lost due to a 

malfunction of that data signal. 

The WFCI-C-03 experiment shown in Figure 4.16 was performed with a charge voltage 

of 400 V. The clear difference in this experiment to the previous five externally triggered 

tests was that one single explosion occurred propagating upward through the fuel-coolant 

mixture. One can detect a pressure signal propagating downward from the slide gate, but 

in this case it is not growing in strength and quickly diminishes and disappears. Also, 

the energetics from this single upwardly propagating explosion is similar in magnitude to 

the cumulative energetics from the double explosions in these past triggered events. Note 

that the external trigger pressure appeared at t =5.8 ms with a peak pressure of 4 MPa. 

This pressure trace shows that the leading edge of the molten fuel had not reached this 

region. However, the pressure trace from the PT3 transducer indicates that the explosion had been induced by the external trigger pulse near this location. The data from the

79 NUREG/CR-6623
propagation velocity between PT1 and PT3 pressure transducers show that the average 

velocity was 224 m/s. This average velocity also indicated the occurrence of an explosion 

between those transducers. The explosion propagated with escalating magnitude of the 

propagation velocity and the pressure peak in an upward direction up to the location of 

the PT5 transducer. In the upper region of the test section, however, the propagation 

velocity decreased noticeably. This test gave the first indication that the presence of the 

double explosion is a strong function of the initial trigger strength. 

The WFCI-C-04 test was conducted with a supplied voltage of 450 V. In addition, four ad

ditional pressure transducers were installed to provide better spatial resolution to pressure 

histories. The WFCI-C-04 experiment also resulted in a single explosion. The pressure 

histories are shown in Figure 4.17. The first and second pressure signals show that the 

leading edge of the melt had not reached the second transducer, PT2. The external trigger 

pressure pulse propagated upward with a velocity of about 1400 m/s (approximately equal 

to the sonic velocity of single phase water) and a peak pressure of 4.3 MPa. An explosion 

then occurred near the location of the PT3 transducer. The explosion pressure was small 

at this location, similar to the previous experiment. Post-test debris suggested that the 

leading edge of the melt stream had mostly solidified and a small part of the melt was finely 

fragmented. This indicates that the small peak in the explosion pressure at this location 

resulted from a small amount of fuel participation and the relatively low fuel temperature 

of 880 °C. 

Although the external trigger pulse had already diminished at this location, this small 

local explosion provided a sufficient trigger to the upper mixture. The magnitude of the 

explosion increased in the upper region and the propagation velocity of the explosion in 

the test section was 210 m/s. In this test, the explosion behavior at the slide gate was 

more clearly recorded by additional pressure transducers installed at this location. The 

peak explosion pressure occurred near this location at 17 MPa. The reason for this can be 

understood by the observation of the distribution of the fuel debris after the experiment. 

Approximately 30 to 50 % of the fuel debris was found in the T-tube with a shape of packed "sand" just below the slide gate. This suggests that the fuel-coolant mixture was located 

at this position or had moved to this position together and became packed in this region 

during the explosion expansion. Thus, the relatively 'rich' fuel-coolant mixture contributed 

to the larger pressurization at this particular location. 

The WFCI-C-06 test was performed with a supplied voltage of 300 V to find the transition 

from single to double explosions. Figure 4.18 shows that the pressure traces of the explosion 

had characteristics of a single explosion with a pressure reflection. The trigger pressure at 

the bottom is similar as in previous tests and propagates with a velocity of 1300 m/s. The 

explosion was induced at the second pressure transducer, however, a little earlier than in 

the WFCI-C-04 test. A sharp pressure wave with rise times of 20 to 40 As propagated 

with an average speed of 240 m/s. Compared to past double explosion cases, the second 

NUREG/CR-6623 80

]1
peak in this test appeared as a wide and low pressure pulse. It did not have any explosion characteristics and is interpreted as a reflected wave from the initial explosion at the top of the test tube. This reflected wave had an average propagation velocity of 68 m/s. The 

empirical conclusion is that external triggers near this value result in an upward propagating 

single explosion. 

The WFCI-C-05 test was repeated with the same conditions as the WFCI-C-02 test, in 

which the supplied voltage was 100 V, however, the piston signal in the expansion tube 

was not measured. Figure 4.19 shows the pressure traces induced by the weakest trigger in this series of tests. The explosion occurred near the PT3 transducer, propagated with 

a velocity of 190 m/s and peak pressure of more than 17 MPa. As expected, a downward 

explosion propagation also occurred at the top of the test section. The weak trigger pulse provides an initial behavior similar to that of the WFCI-C-06 test, in which the second 

explosion propagated and induced a stronger explosion. The detailed analyses on this 

subject axe discussed in the next section. 

In the WFCI-C series of tests, the conversion ratio was in the range from 0.25 to 0.38 

percent, as shown in Table 4.1. These are similar to the conversion ratios obtained from 

the WFCI-A series. Figure 4.20 shows the relationship between the trigger pressures and 

the measured conversion ratios. In this figure, the pressure signals at the PT1 transducer 

were taken as the peak trigger pressure since the molten fuel had not reached that location 

when the external trigger was actuated. In this figure, the WFCI-A-02, WFCI-A-05, and all 

WFCI-C series of tests were included because their mixing conditions were similar to each 

other and all tests produced triggered explosions. The remaining WFCI-A and WFCI-B 

explosions were spontaneous. The figure shows that the conversion ratio tends to increase 

when the peak trigger pressure increases from 1.75 to 2.36 MPa. However, the conversion 

ratio decreases with peak pressures increasing from 2.36 to 4.24 MPa. The maximum 

conversion ratio was observed at the peak trigger pressure of 2.3 MPa. Because the data 

base is sparse, it is not our intent to draw any major conclusions from the work. However, 

we must emphasize that energetics of spontaneous explosions and weak externally triggered 

explosions can be affected by the trigger magnitude. At some point for any fuel-coolant 

combination the explosion energetics would become independent of the trigger and this 

is the region where we need to operate for our experiments. The presence of the double 

explosion is another issue the trigger affects and which we must also consider. 

In these tests, the presence of double shock wave propagations through an axial interaction 

zone with the presence of the rigid slide gate at the top of the test section was observed. 

It is interesting to note that, in the KROTOS tests, similar multiple wave propagations have also been observed even without a solid boundary at the top of the test section. 

Post-test analyses of the KROTOS tests show that about fifteen percent of the total debris 

was packed in a region near the top of the fuel-coolant mixture, presumably caused by the 

mixing during fuel pouring and subsequent freezing during the FCI. Thus, one can consider

81 NUREG/CR-6623
the fuel in this region to act as a "porous plug" of melt providing a "semi-rigid" boundary 

at the top of the mixture. This boundary not only can dissipate the explosion impulse and 

expansion, but also acts as a point of shock reflection for downward FCI propagation. 

The WFCI-C experiments confirm that the characteristics of the explosion are strongly 

affected by the external trigger. If the external trigger pressure is weak or absent, in the 

case of spontaneous events, the fuel jet may not be sufficiently fragmented by a single 

propagation event through the mixture. A more energetic explosion would be expected by 

any subsequent pressure reflection back into the fuel-coolant mixture, if a proper combina

tion exists of unfragmented molten fuel, mixture void fraction and relative mixture length 

scales, such as the fuel mixing diameter. From the dozen experiments which exhibited 

this behavior in these three test series, it is not possible to ascertain what is the precise 

combination of conditions, but it clearly exists. If the trigger pressure is sufficiently strong, 

the initial explosion propagation event causes sufficient fuel fragmentation in the mixture 

to allow the pressure to grow to a level that fragments enough fuel and creates mixture 

conditions where no subsequent propagation is possible. From an energetics standpoint, 

it is important to also note that the explosion energetics is not noticeably different for 

explosions with single or double propagation events. The source of any pressure reflection 

in a fuel-coolant interaction has been addressed above and is not relevant to our main point 

here. 

In subsequent tests the trigger pressure is planned to be set at about 3 MPa to minimize the 

chance of multiple propagations. This trigger level also addresses the independence of the 

trigger from any subsequently observed energetics. The presence of these double explosions 

only emphasizes the fact that the fuel-coolant mixture is metastable and potentially capable 

of sustaining subsequent explosive propagations, even after the first explosion propagation, 

if the conditions are conducive for them to occur. This has never been clearly documented 

from past FCI data. 

4.7 Effect of System Constraint: WFCI-D Series 

In the WFCI-D test series, the explosion expansion was channeled through a lateral expan

sion path just below the closed slide gate. This geometry is similar to the design used in 

the past by Baines to establish a system constraint to measure energetics. This constraint 

is easily controlled and allows one to vary the constraint independently of the fuel and 

coolant masses in the mixing region. In the stratified tests performed by Bang [3], it was 

noted that the expansion of fuel-coolant mixture was controlled by the constraint of the 

slug mass and by the coolant depth. 

During an energetic FCI the explosion impulse and the mechanical work potential depend on 

NUREG/CR-6623 82

1!
the system constraint. Hydrodynamically, the constraint was classified by Cho et al. [130] 

as acoustic or inertial. The system constraint was considered in developing their parametric 

models for an acoustic or inertial system, based on two different timescales for potential mechanical work output. The first timescale is the time for relief of a shock wave in the 

system. During this time interval (so called the "acoustic unloading time") a shock wave 

propagates within the fuel-coolant mixture in the explosion chamber. This represents the 

time for the shock pressure front to travel through the mixture and system constraint 

above it to the nearest free surface and reflect as a rarefaction wave travelling back to the 

interaction zone. This time is approximately equal to 21/c, where 1 is the distance from 

the interaction zone to the nearest free surface and c is the velocity of sound in the media. 

The second timescale is the time for the inertial expansion of the fuel-coolant mixture. In 

this period, the mass of the liquid column and the fuel-coolant mixture itself is accelerated 

upwards by vapor pressure in the expanding mixture. 

From a thermal point of view the timescale for explosion expansion, as related to the sys

tem constraint, could be divided into two time periods. The first period is for heat transfer 

between rapidly fragmenting fuel and the local coolant causing vapor production. The 

second period involves heat transfer from the vapor to the surrounding coolant liquid with

out significant vapor production, either because fuel fragmentation has ceased or from fuel 

quenching. These periods are not as separable as the hydrodynamic time periods. These 

two time periods are conceptually plotted with the mechanical timescales in Figure 4.26. 

During the first time period, since the vapor production rate is directly proportional to the 

fuel fragmentation rate, more energetic FCIs may be expected as the time increases. This 

is because as more time is made available, larger vapor pressures driving the explosion can 

result, with minimal slug expansion. This phenomenon was observed in stratified geometry 

experiments [131]. In that test, slug masses of 1 and 4 kg were used as the lower and 

upper limits, respectively. The results showed that the maximum mechanical energies of 

the explosion for the upper limit case was twice as large as in the lower limit case. If the 

time continues to increase into the second period, however, less energetic interactions are 

anticipated. Under these circumstances the vapor production could significantly diminish 

or revert into vapor condensation as the fuel quenches, as the slug expansion becomes ap preciable and the coolant bulk heating begins to dominate. The system constraint directly 

affects these timescales since it controls both the acoustic and inertial timescales and thus 

affects these thermal time periods for fuel-coolant energy exchange. 

Figure 4.27 shows the pressure history for the WFCI-D-04 test as an example of this 

concept. This data illustrates the typical pressure history of the WFCI-D experiments. The plot presents the pressure histories measured with eight pressure transducers from the 

bottom to the top of the test section. It clearly shows pressure propagation and escalation 

from the bottom of the test section. The bottom two figures show the typical pressure history of the external trigger. The external trigger provided a pressure of about 3 MPa

83 NUREG/CR-6623
with a pulse width of about 0.1 to 0.2 ms at its half height to all WFCI-D series experiment. 

The trigger shock traveled with a speed of 1530 m/s from the first transducer to the second 

one. However, the shock speed dropped to about 290 m/s and the shape of the shock 

was changed at the third transducer location. This indicates that no melt had reached the 

location of the second transducer and the leading edge of the fuel melt was located between 

the second and the third pressure transducers about 1.1 m below the slide gate. After that 

point, the shock generated by the explosive interaction with less than 0.1 ms peak rising 

time traveled with a speed ranging from 290 m/s down to 80 m/s. These speeds can be 

explained by the condition of the mixture; i.e., high vapor void fractions as one moves up 

the explosion tube. 

To gain an understanding on the effect of the axial constraint on energetics the six WFCI-D 

tests along with selected tests from WFCI-A, B and C series are compared. The degree 

of axial constraint for fuel-coolant interactions is characterized by the total initial mass of 

the slug which can be made up of water or water interspersed with high density lead slugs 

(73 mm diameter and 100 mm long each) in the expansion tube. The degree of an axial 

constraint is represented by a dimensionless parameter, < m >, defined by the following 

expression, 

< M >= M_..i (4.6) 

where mr,i is the total initial mass of the slug in the expansion tube as shown in Figure 4.8 

and rn0 ,i is the total initial mass of water in the interaction zone. In this experiment, < m > 

ranges from approximately 0 to 8. The experimental conditions of the WFCI-D series and 

other series in the WFCI experiments are summarized in Table 4.1. 

The WFCI-B series tests are included even though no external trigger was provided since 

spontaneous explosions also occurred over a wide range of mixing times and these data 

illustrate the effect of the fuel quenching time on the FCI energetics. 

Specific comparison to past data [60, 62, 63] was chosen because of their similarity in 

geometry, materials, and experimental conditions as well as their completeness in data 

measurement. For the KROTOS-21 test [63], since the facility is not equipped with a 

system for direct measurement of work output, the conversion ratio for the KROTOS-21 

test is estimated from the TEXAS simulation [117], which produced the maximum work 

output of 1.1 kJ. The total fuel mass of 5.4 kg measured after the experiment and called 

'the probable-mixed mass' was used for the estimation of the total thermal energy. 

The conversion ratio for Hall's tests [60] was determined from the measured impulses of the 

explosion pressures which ranged from 150 to 250 J. These corresponds to the estimated 

conversion ratios ranging from 0.06 to 0.25 %. In the Baines' tests [62] the estimated 

conversion ratios were ranging from 0.18 to 0.38 %. 

NUREG/CR-6623 84

1!
For Hall and KROTOS experiments, their axial constraints, < m >, are assumed to be zero 

to be consistent with the definition applied in the WFCI test; i.e., the slug mass is defined 

as the slug mass above the actual fuel-coolant mixture. For Baines's experiments, however, 

< m > is obtained by the total initial mass of water as a slug in his expansion tube divided 

by the total initial mass of water in the interaction tube. Table 4.4 summarizes the ranges 

of their experimental conditions. 

The conversion ratios of this series of experiments are plotted in Figure 4.5 with respect 

to the dimensionless parameter, < m > as an indicator of the axial constraint, including 

previous other test results [63, 60, 62] for comparison. In this figure, the duration time 

of the piston movement, At,, in the expansion tube is plotted with respect to the axial 

constraint, < m >, as well. The data of the duration time, At,, shown in Table 4.6 are 

linearly fitted with their standard deviation of about 50 ms. It indicates that the average 

slug velocity in the expansion tube linearly decreases as the axial constraint increases. 

All data for the conversion ratio are fitted to the second order polynomial function and 

show that the conversion ratio increases with an increase of the axial constraint up to 

< m > of approximately 5, then decreases, although At. keeps increasing with < m >. 

Figure 4.29 shows the relationship between the peak pressure of vapor explosions and the 

axial constraint comparing with the conversion ratio. Note that peak pressures obtained 

from the WFCI-A, B and C-01 are not used in this figure because the location of the 

pressure transducers for obtaining maximum pressures and preset maximum ranges of the 

transducers are quite different from the WFCI-D series. The figure shows that the peak 

pressures increase up to approximately < m > of 3.5 and decrease with the axial constraint 

increase. The maximum peak pressure occurs at a slightly lower axial constraint than 

the maximum conversion ratio. However, the general trend of the peak pressure with 

respect to the axial constraint is almost identical to that of the conversion ratio. It clearly 

indicates that the peak pressure of vapor explosions is closely related to the energetic of 

vapor explosions (i.e., conversion ratio). 

A possible reason for the effect of axial constraint can be linked to the two thermal time 

periods previously discussed. Those two time periods related to the axial constraint are 

a time period for heat transfer between the fragmented fuel and the adjacent coolant 

causing rapid vapor production, and a time period for heat transfer between the fuel-vapor 

mixture to the bulk coolant causing vapor condensation and fuel quenching (especially in 

highly subcooled coolant) without significant vapor production. Those time periods were 

conceptually plotted for comparison with the mechanical time scales suggested by Cho et 

al., in Figure 4.26, and were previously discussed. 

In general, the axial constraint directly affects these time scales since it controls both the 

acoustic and inertial time scales through its variation and its effect on fuel-coolant energy

85 NUREG/CR-6623
exchange. If the axial constraint increases, both acoustic and inertial times will increase, 

leading to an increase of the time duration of the slug movement. It provides more time 

for fuel-coolant mixture energy exchange and increases the FCI energetics. However, as 

the time duration increases further, vapor condensation and fuel quenching play more 

significant roles and result in the conversion ratio decrease. 

In this experiment, the turning point occurs within a range of 500 msec for At,. Because of 

low subcooling of the bulk coolant in this series of tests, the vapor condensation may play 

a less important role than fragmented fuel quenching on conversion ratio decrease. One 

should note that qualitatively, such behavior would always be expected but it would occur 

at quantitatively different values for different fuel compositions. 

Figure 4.30 shows the maximum conversion ratios calculated by the UWFCI code [105] 

with respect to the degree of the axial constraint, < m >. This calculation was done, 

not to validate the code as much as to demonstrate that this qualitative effect should be 

demonstrable in a straightforward manner. The computer model calculation shows fairly 

similar qualitative trends of the axial constraint effect. In this figure, calculated conversion 

ratios for the axial constraint are maximized at < m > of about 10, which are larger 

than that observed in the experiments; i.e., < m > of about 5. The maximum conversion 

ratio decreases from 2 to 1 % as the volume ratio of the vapor to water in the mixing 

zone increases from 20 to 50 %. It shows that the conversion ratio is strongly affected by 

the initial mixing conditions, i.e., void fraction, decreasing with the void fraction increase. 

Although the conversion ratios calculated are about an order of magnitude higher than 

those obtained, the general trend with respect to the axial constraint is nearly identical to 

the experimental results. 

The time histories of the conversion ratios and system pressures for < m > of 3.3, 11.2 and 

44.4 are shown in Figure 4.31. The figure shows that the conversion ratio starts to increase 

earlier as the axial constraint is smaller. It is because the conversion ratio is calculated by 

the slug movement as in the experiments. The higher the axial constraint, the later the 

slug starts to move. It also illustrates that the maximum conversion ratio varies with the 

axial constraint and does pass through a maximum value. The system pressure, however, 

increases with the increase of the axial constraint without exhibiting any maximum as in 

the experiments. 

The completion of the WFCI-D test series represents the final test series which examines 

the boundary conditions of the FCI on its energetics. We have seen that the trigger (or 

lack of it) and the system constraint do have a marked effect on the explosion propagation 

and energetics. We next examine how the fuel and coolant initial conditions affect the 

explosion process. 

NUREG/CR-6623 86

11
4.8 Temperature Effects of the Fuel and the Coolant: WFCI-E and F Series 

To investigate the temperature effects on the vapor explosion, two series of tests were 

performed, the WFCI-E and WFCI-F for fuel and coolant temperatures, respectively. The 

test conditions for these experiments are shown in Table 4.1. The test series are similar to 

what have been reported in the past for other fuel-coolant pairs. Also similar to past data, 

once the intial conditions have been found that produce the explosion, the effect of the 

absolute temperature variation on energetics is not large. Note that the fuel in all of these 

cases had significant superheat. In these tests, the general behavior was that explosions 

were observed for most of the test conditions, with the most notable change being that 

spontaneous explosions became more frequent and unpredictable. As the temperature 

of fuel or coolant decreased, further explosions became more incoherent; i.e., vapor film 

stability required for fuel-coolant mixing would not be satisfied. 

In the WFCI-E tests, the fuel temperatures were varied from 490 to 790 °C, with an 

other conditions similar to the previous WFCI-A series of tests. There are two tests that 

are of particular note. The WFCI-E-02 test, as shown in Figure 4.21, had an initial fuel 

temperature of 790 0C and produced a triggered double explosion. In this case though 

the second explosion, with a wide pulse width, can be identified as the combination of two 

pressure pulses; one is a typical downward propagation and the other is a reflected wave 

from the bottom of the test section superimposed on it. This reflected wave propagates 

upward with a sonic velocity of single phase water but diminishes near the PT5 transducer. 

Once again we would expect that the largest local void fraction in the mixture would occur 

near this point. In the WFCI-E-03 test (Figure 4.22), a spontaneous explosion occurred 

about one second before the system was to be triggered near the top of the test section. 

It was the first explosion that occurred before the slide gate was closed. No damage was 

noted because it happened near the top of the test section where there was minimal liquid 

constraint before the explosion vented to the ambient. The likely reason for this explosion 

was that fuel entrapped water between the structure and funnel. Both steel frames were 

bent up to 450 by the impact. The work done by this structure against the steel frame was 

estimated as at least 2 kJ, based on the frame deflection analysis. This demonstrates that 

the damage resulting from an impact of a missile generated by an explosion is more severe 

than that due to the explosion expansion in the ambient air without missile generation. 

In the WFCI-F series of tests, a total of three tests were conducted to investigate the 

coolant subcooling effects on the vapor explosion, varying the coolant temperatures from 

27 to 72 *C. All other conditions were the same as in the WFCI-A series. In all tests, 

spontaneous explosions occurred before the slide gate closed. Small amounts of molten fuel 

penetrated below the slide gate because of these early FCIs, as shown in Table 4.1. The 

WFCI-F-02 test was performed with a coolant subcooling of 73 0C. As shown in Figure 4.23, 

several pressure spikes with relatively small magnitudes appeared near the top of the test

87 NUREG/CR-6623
section about 1.25 s before the system was triggered. Such early FMIs probably occurred 

since the vapor films around the fuel could not maintain their stability at these fuel-coolant 

conditions. These events occurred near the top of the test section preventing pouring of 

the molten fuel into the test section. Only about 1 kg of fuel was collected in the test 

section. As the coolant subcooling decreased (higher water temperatures), the pressure 

signals changed from incoherent spikes to an explosive behavior as shown in Figures 4.24 

and 4.25. 

4.9 Coolant Additives: WFCI-G Series 

A group of experiments were conducted to investigate the ability of a polymeric solution 

to suppress vapor explosions. This series of tests was denoted as the WFCI-G series. Five 

experiments were conducted with polymer solutions of concentration from 400 to 800 wppm. 

The viscosity ratios measured just before the experiments with polymeric concentrations 

of 400, 600, and 800 wppm were 1.5, 2.0, and 2.5 respectively. Two other experiments were 

performed with pure distilled water at the temperature of 24 0C to provide reference cases 

of explosions in the polymer solution. In the two cases of pure water experiments, i.e., the 

WFCI-G-05 and WFCI-F-02, spontaneous interactions occurred as the melt entered the 

water, as shown in Figure 4.32. The early local explosions in this highly subcooled water 

limited the amount of melt entering the explosion tube and subsequent energetics. 

In the cases of polymer solutions, however, none of the five tests showed initial eruptions or 

energetic spontaneous explosions as shown in Figure 4.33. This figure shows that multiple 

pressure "spikes" with magnitudes of several hundreds kilopascals occurred during the 

pouring period of the molten tin into the water. It is believed that these spikes result from 

the vapor film collapse around the melt drops. But it was shown that such vapor film 

collapse in polymer solutions did not initiate coherent vapor film collapse of adjacent melts 

and eventually the energetic events were suppressed. In the single drop tests, the same 

order of magnitude of pressure spikes was treated as the explosion. Even if these spikes are 

pressures resulting from the explosion, one may say that explosions are suppressed since 

there was no escalation and propagation of the explosion. 

Such suppression of the FMIs due to the coolant viscosity increase can be explained by two 

possible effects. First, the fragmentation mechanism during the melt and water contact 

is altered. In general, one possible melt fragmentation mechanism has been described by 

the Taylor instability between the vapor and water interface. When the melt contacts 

with the water, the melt is surrounded and stabilized by the vapor film. The vapor film, 

however, becomes unstable due to the Taylor instability of the interface between the vapor 

and water. Taylor fingers develop and water jets penetrate through the vapor film to the 

molten fuel surface [14], [57]. If the viscosity of the coolant is high enough, however, the 

NUREG/CR-6623 88

11
instability may not grow sufficiently fast to break up the melt.

Second, the boiling heat transfer characteristics are also changed. As mentioned previously, 

the polymer solution shifts the boiling curve, for example, increasing the nucleate boiling 

heat flux in general. Recently, Bang [132] investigated the boiling heat transfer characteris

tics of a highly subcooled polymeric solution (PEO and water). He used a heated stainless 

steel ball with a diameter of 22 mm. He observed that the film boiling over the sphere 

in the polymer solution at 30 °C sustained for a period of time relatively longer than in 

pure water. The minimum film boiling temperature rapidly decreased increased polymer 

concentration up to 300 wppm, and did not change from this value when the concentration 

was further increased. Bang also observed from the temperature-time trace at the center 

of the sphere and synchronized visual observation that the stable vapor film was sustained 

even after the temperature of the sphere surface reached the solidification temperature of 

tin. From this observation the tendency of no energetic event in the WFCI-G series experi

ments with polymer solutions may be explained; the surfaces of melt jets with stable vapor 

films were solidified and then quenched before the vapor films were destabilize enough to 

cause energetic interactions. Either of these effects of weaker hydrodynamic instability 

development or more stable film boiling, may be the root cause of the suppression of the 

energetic FCI. 

In order to investigate the stability of vapor film in polymer solutions under an external 

disturbance, WFCI-G-02, identical to the WFCI-G-01 test but with an external trigger, 

was conducted. An external trigger with a magnitude of 3 MPa peak pressure was applied. 

No explosion was observed in this case. It is noted that in the cases of pure water at 85 °C, 

the molten tin exploded without the external trigger or with a trigger of a magnitude from 

3 to 4.5 MPa [65], [78]. Since the vapor fims around the melt are stable in the polymer 

solutions, there may be a potential for the melt to explode energetically if the external 

trigger is strong enough. 

In this series of experiments, the concentration of polymer was changed from 400 to 800 

wppm, corresponding to viscosity ratios of 1.5 and 2.5. For small scale experiments [69], 

at a viscosity ratio of 2 the explosion was completely suppressed. Based on this empirical 

information, the viscosity ratio was reduced to 1.5. Even at this ratio (WFCI-G-04) no 

explosions occurred. This confirms the belief that small scale single droplet results cannot 

be directly extrapolated to these larger scale tests. In order to examine the mass scale effect 

(WFCI-G-06), the molten tin mass was increased to twice the previous value. This test 

with tin mass of 5 kg and polymer concentration of 800 wppm also showed no explosion. 

For this test series, the debris analysis provides some unique insights. Thus, this debris

89 NUREG/CR-6623
analysis is presented here with comparisons to these other results.

WFCI-G Debris Analysis 

The post-test debris analysis showed that in the polymer solutions, the melt was broken 

into nearly spherical particles ranging from a few tenths micrometers to centimeters as 

shown in Figure 4.34. The shape of the melt fragments in this series of tests was closer 

to spherical than those observed in Dowling's single drop tests. In the case of pure water, 

however, the debris appeared to be larger and more arbitrarily shaped, indicating some fine 

fragmentation at the time of the FCI as shown in Figure 4.35. Table 4.5 shows that about 

70% of the molten tin entered the test section with no explosion. However, in the cases 

with spontaneous explosions, only 20 to 35% of the molten tin had entered. Some sort of 

surface eruption at the early stage of the mixing phase in two spontaneous explosion cases 

limited the amount of molten tin entering the test section. 

Figures 4.36 and 4.37 show that debris mass less than 0.5 mm in size in a spontaneous 

explosion case (WFCI-G-05 test) is much larger than those in other tests (WFCI-G-01 and 

G-04) because of the explosion and its associated fine fragmentation. The debris masses of 

the non-explosion cases are bounded by the highest polymer concentration case (WFCI-G

01) and the lowest polymer concentration case (WFCI-G-04). 

Figure 4.38 shows the relationship between the Sauter mean diameter and the viscosity 

ratios of the polymer solutions in this series of experiments. The Sauter mean diameter, 

D.m, [88], is defined in general as follows: 

jDmax pDd 

DM - 10Dao V() (4.7) 

Jovm'D 2p(D)dD 

where p(D) is the probability of a debris particle having diameter between D and D + dD. 

Since the different sizes of debris have been separated by a set of sieve bins with mesh sizes 

ranging from 25 to 8000 pm, the above equation was modified to the following by assuming 

that all debris particles in a certain sieve bin, i, have a perfectly spherical shape with only 

one average diameter, D1: 

E Dj~p(Dj)AD, 

Dam = i,<N (4.8) 5 Djp(DA)ADA 

i<N, 

NUREG/CR-6623 90

11
where N, is the total number of sieve bins used in the process, and ADi is the mesh size 

difference between the sieve bin numbers of i and i + 1. The debris particle probability, 

p(Di), was calculated from the mass of the debris at the sieve bin number i divided by the 

total mass of the debris. As indicated in equation (4.7), the Sauter mean diameter, SMD, 

is proportional to the ratio of the volume to surface area of the debris particle. Therefore, it 

represents a characteristic diameter of particles with the proper surface area to volume for 

heat and mass transfer. One should note that because of the method of averaging, the SMD 

tends to be larger than the mean diameter based on particle number or particle surface 

area. SMD values are biased by the large particles in the fuel debris. Figure 4.38 shows the 

SMD in these entire tests with respect to the viscosity ratios. At the viscosity ratio of one, 

the SMD for two tests conducted at the pure water are almost identical, having diameters 

of about 3 mm. Compared with other tests performed in dilute polymer solutions, these 

diameters are approximately four times smaller. In fact, the SMD values for pure water 

tests will be somewhat smaller than the values shown in this figure because this value was 

calculated by assuming all debris having a spherical shape. These small SMDs for the pure 

water tests are indicative of the explosive heat transfer between molten fuel and water due 

to vapor explosions and its fine fragmentation. 

This figure also shows that the Sauter mean diameters at higher viscosity ratios seem 

to be independent of the viscosity ratios ranging up to 2.5. The SMD of WFCI-G-02 

conducted with an external trigger of 3 MPa was also similar to that of non-triggered tests. 

It indicates that there was no fine fragmentation of molten fuel due to the external trigger. 

The trigger was not strong enough to initiate the explosive interactions between molten 

fuel and solutions. For the WFCI-G-06 test with a doubled mass of molten fuel, its Sauter 

mean diameter was also similar to those of other tests. 

These debris size distributions are indicators that identify the characteristics of vapor ex

plosions (e.g., its energetics). Therefore it is valuable to know the debris size distribution 

at the mixing phase since this provides the initial mixing conditions of the fuel-coolant mix

ture and eventually determines FCI energetics. However, most debris data were obtained 

after explosions. Therefore, it is almost impossible to predict the debris size distribution 

for the mixing phase from these data. It is our hypothesis that the non-exploded debris 

size distributions obtained in dilute polymer solutions such as Figure 4.36 and 4.37 may provide a possible fragment size distribution of the molten jet of fuel in the mixing phase 

of the FCI phenomena. Hence, these debris data may be useful to predict the size of the 

jet breakup in the mixing phase of an FCI.

91 NUREG/CR-6623
4.10 Mass Ratio Effect: WFCI-H Series

The WFCI-H series was designed for investigating the effect of mass on the vapor explosion, 

by varying the coolant to fuel mass ratio. Also, during this series, our level swell meter was 

tested and verified by visual observation and filmed by a video camera. Different methods 

to control the mass ratio were introduced; variation of the fuel pouring time and a change 

in the fuel pouring diameter. We also controlled the position of the explosion tube bypass 

valve to affect the mixing process. 

During the WFCI-H tests, the pouring time was varied from 1.5 to 2.0 s, to allow the trigger 

to be actuated 0.1 s after the slide gate closed, as shown in Figure 4.3. The explosion then 

occurred at a lower axial position, near the second pressure transducer instead of the third 

transducer as shown in the WFCI-A series. Different fuel masses can be introduced into 

the test section by variation in the pouring time. Finally, it was known that in our test 

section at low water subcooling and long pour times (large fuel mass) one could develop 

conditions in which fuel penetration would be hampered by vapor production. Under these 

conditions we would open the bypass valve to allow for faster penetration of the fuel jet 

into the water pool, enhandng mixing and changing the fuel-coolant mixture volume. 

In this test series, the level swell meter described in chapter three was installed and verified 

by visual observation in the WFCI-H-03 and H-05 tests. The original level swell meter 

signals are shown in Figure 4.41. The voltage signal decreases with coolant level increase, 

due to the variation of the electric resistance between two copper rods. About 1.6 s before 

the trigger, the signal from the level swell meter indicates the increase of the water level 

resulting from the coolant splashing when the fuel jet initially contacts the surface of the 

water. Such transient surface movements of the coolant are restabilized about t =-1 s, the 

voltage decreased gradually while approaching asymptotic levels. 

In the WFCI-H-03 test, violent oscillations of the level swell signal as shown in Figure 4.41 

after the onset of external trigger resulted from the spontaneous explosion occurring above 

the slide gate. The difference between the initial voltage signals resulted from the operation 

time of the level swell meter. However, both signals show very similar behavior while 

maintaining a constant voltage difference. 

Figure 4.42 is a plot of integrated average void fractions converted from the voltage signals, 

and from the images captured with the video camera. The upper and lower limits for the 

level swell signals represent the error range of the signal interpretation with a maximum 

range of ± 13 %. The lower limit for the level swell was calculated by assuming that the 

vapor was distributed only within the range of the test tube diameter, from the bottom of 

the test section to the surface of the water. However, the upper limit case was estimated 

by assuming uniform distribution of the vapor in the test section and funnel.

NUREG/CR-6623 92
The signals agree well with the visual observation until t =-0.8 s; at later times values 

from the level meter gradually increase while the visual observation indicates that the level 

remains constant at a void fraction of about 25 %. The constant water level measured from 

the visual observation is due to the slide gate closing about 400 ms before the trigger. The 

reason for the asymptotic increase of the level swell meter signal is the resistance change 

due to the suspension of molten fuel debris in the coolant. After the slide gate is closed, 

the signal tends to remain constant but indicates higher void fractions than from the visual 

observation. From the level meter results, the average void fraction for the typical WFCI 

tests is in the range of 20 to 30 %. 

As the fuel is poured into the coolant chamber and subsequently mixes with the water, a 

certain coolant to fuel mass ratio is established in the mixture. This ratio determines the 

local vapor production rate and can have a significant effect on the void fraction in the 

fuel-coolant mixture and the subsequent explosion escalation and propagation. From an 

experimental viewpoint, this ratio is indicative of the integral quantities because it only 

determines the integral quantity of fuel and coolant in the test section and not the local 

mixture concentrations of fuel and vapor volume fractions in the liquid. 

Thermodynamic analyses [95, 134] show that, theoretically, the volume ratio which max

imizes the work output from an explosion, is near one. For the case of tin as a fuel, this 

volume ratio corresponds to a mass ratio of approximately 1/7. If the mass ratio is smaller 

than this, the conversion ratio decreases with increasing mass ratio. This happens because 

the mixture is 'coolant lean' with too little coolant available as the working fluid for heat 

transfer from the fuel. In contrast, as the mass ratio increases beyond the maximum point, 

the conversion ratio tends to decrease, because the mixture is 'fuel lean', due to the rela

tively large amount of coolant liquid which quenches vapor production. It is interesting to 

note that the observed effects for axial constraint and time scales for heat transfer represent 

dynamic analogues to this thermodynamic explanation. 

The experimental results as shown in Figure 4.43 indicate that the conversion ratios increase 

with increasing mass ratio in the range from 0.2 to 1.0. This corresponds to a mixture 

volume to fuel volume ratio of 1.5 to 7.0. A broad maximum of the conversion ratio of 

less than 1.0 % appears at a mass ratio of about 1 to 3 (mixture volume to fuel volume 

ratio of about 7 to 20). Finally, the conversion ratio is seen to decrease when the mass 

ratio (or volume ratio) increases further. These experimental results show that the optimal 

point for the mass ratio was shifted from the theoretical point of 1/7 to an actual value 

of 1. This might be expected because the kinetics of the mixing process would increase to 

large values of this ratio to accommodate the fuel jet penetration into the coolant along 

with vapor production and coolant liquid outflow. These results include other tests similar 

to the WFCI tests; i.e., tests conducted by Baines [62] and Hall et aL.[60]. These tests 

were chosen because of their similar geometry and experimental conditions as discussed 

previously in the WFCI-D series.

93 NUREG/CR-6623
In Figure 4.43, the conversion ratio for the WFCI series and the above two experiments 

can also be compared with the Hicks-Menzies' thermodynamic model [95] and Board-Hall's 

detonation model [134]. The interesting point to note is that all the data is far below the 

thermodynamic values expected for the tin-water FCI pair. The Hicks-Menzies and the 

Board-Hall models both indicate theoretical conversion ratios of about 30 percent. This 

point will be revisited in the experimental analysis as to the implication of these data. 

4.11 Iron-Oxide Fuel Composition: WN CI-K Series 

A group of experiments were conducted in the K-series with iron-oxide as the fuel simulant. 

The purpose was to investigate the likelihood of an energetic FCI with a fuel simulant that 

is more prototypic of molten fuel materials; i.e., molten urania, zirconia and zirconium 

with modest superheats. The initial conditions were similar in all tests to the past WFCI 

tests, except for the change in fuel composition and the use of a larger test section in the 

last two tests. Table 4.2 summarizes the experimental results for all series K iron-oxide 

experiments. The trigger strength varied from the original WFCI test section (designated 

WFCI-1) to the larger test section (designated WFCI-2), only because of specific mechanical 

limitations. The larger piston trigger assembly in WFCI-2 was not able to generate the 

same trigger peak as in the smaller one, because the same solenoid was used for both 

designs. Thus, even though the capacitance of the trigger charging circuit was doubled for 

the WFCI-2 tests, the actual acceleration of the larger piston was smaller. This resulted in 

a delivered trigger impulse of 8.5 millinewton-sec compared to 14 millinewton-sec in WFCI

1. In addition to these WFCI-K experiments, melt generation and delivery development 

tests were performed. Smaller masses of melt were dropped into a pit of sand to test the 

reliablility of the delivery method and repeatability of the melt jet geometry. These tests 

are not reported in this topical report (Series I and J). Two other tests in the series were 

excluded because of equipment malfunctions. In test K-06, the melt was not successfully 

delivered to the test section due to a failure to remove the catcher pan from beneath the 

furnace. Test K-I1 was successful, but the data acquisition system failed to record any 

data. Since the rupture disk did not break and the slug did not move, there is no way to 

tell if an explosion occurred. 

Since the trigger signal is an integral part of every experiment, it is important to understand 

its characteristics, specifically with test section size. Past data suggest changes in the 

magnitude and duration of the trigger signal may affect the experimental results, so care 

was taken to limit the variation in trigger input. There was some variation between the 

triggers for tests with WFCI-1 and WFCI-2, and this will be illustrated. It was desired that 

all the triggers be of the same amplitude and duration for every experiment to eliminate 

its effect as an experimental variable on the explosivity for the oxide fuel. 

NUREG/CR-6623 94

1!
Figure 4.44 shows a typical set of pressure peaks generated by the trigger for WFCI-1, and 

Figure 4.45 shows the same for WFCI-2. The WFCI-1 trigger clearly shows a sharper and 

more defined peak. This is due to the fact that the WFCI-1 piston is smaller and has a 

larger travel distance, so it can reach a higher velocity when struck from below, and hence a 

larger peak (as mentioned above), even with a lower charging voltage. All WFCI-1 triggers 

were charged to 200 V with two capacitors, while the WFCI-2 triggers were charged to 400 

V with four capacitors. The magnitude of the peak pressure and the width of the peak 

are measured directly from the data. Summation of the pressure values generates the area 

under the peak, which when multiplied by the surface area of the transducer (= 0.236 cm2), 

has units of N-s as the impulse. The WFCI-1 trigger has a peak amplitude of around 3.0 

MPa, a peak width of 150 microseconds, and an impulse of 0.008 N-s. The WFCI-2 trigger 

has a peak of near 1.5 MPa, a similar pulse width and impulse of 0.005 mN-s. Though the 

peak characteristics differ, the propagation speeds of these pulses are similar, as expected. 

The data for shock arrivals at various positions are plotted in Figures 4.46 and 4.47. The 

propagation speeds are 1441 and 1431 m/s, respectively, and constant along its length. 

The pressure data taken during each experiment provides information about the pressure 

at eight points along the test section tube. Table 4.2 gives the maximum pressures, the 

peak impulses, and propagation speeds for the pressure data taken in all experiments and 

two trigger runs, and Figures 4.48 to 4.56 show the transient histories for each experiment. 

It was from this data that information about the peak pressure, impulse, and propagation 

speed was determined. 

Peak pressures give a fairly qualitative measure of interaction strength. Impulses are cal

culated using an integration technique where the pressure values are added over a given 

time span, and are dependent upon the choice of integration region. The time period is 

constant for all of our comparisons. The propagation speed is estimated by determining 

the time that a given peak arrives, which is characterized by the pressure value attaining 

a certain level above the background noise level (chosen to be 0.5 MPa). These values are 

only used for rough comparisons between experiments. For experiments K-02 to K-08, the 

pressure histories are quite similar. In all cases, the trigger peak can be seen from the P1 

trace, while the P2 shows a lower value, and P3 through P8 show no pressure signals at all. 

The traces for test K-10 are similar to those for tests K-02 through K-08, where the trigger 

can be seen, and the amplitude markedly decreases as the pressure pulse propagates. Some 

evidence of the trigger can be seen up to P5. The pressure data for experiments K-01 and 

K-09 show the peak pressures propagating upwards and gaining in amplitude, while the 

propagation speeds and the associated impulses decrease. In addition, the peaks tend to 

become wider at the top of the test section, and peaks can be seen apparently coming down 

from the top in traces P7 and P6. 

The amplitudes of the pressure peaks are tabulated in Table 4.2 and give some insight 

into the strength of the interactions that occurred. In tests K-03 and K-05, the pressure

95 NUREG/CR-6623
peak amplitudes of 5.25 MPa and 5.02 MPa at the P1 transducer are larger than the 

known trigger amplitude for WFCI-1 of 3.0 MPa, which would indicate a weak interaction 

occurring there. Tests K-02, -04, -07, and -08 all have maximum pressures less than the 

trigger pressure, which indicates that there was no interaction. Experiments K-01 and 

K-09 yielded the largest maximum pressures: 6.53 MPa for K-01 and 9.70 MPa for K

09, with these occurring near the top of the test section. Test K-10 had a maximum 

pressure of 4.36 MPa at transducer P1, and the pulse tapered off as it traveled upwards. 

The impulses of the pressure peaks can also give some insight into the energy released 

during the interactions (or explosions). For a given timespan encompassing the FCI, data 

are numerically integrated resulting in a number that is multiplied by the surface area of 

the transducer face ( 0.236cm 2). This number is only used as a guide to indicate which 

interactions were stronger relative to others in this test series. The impulses are the largest 

for tests K-01 and K-09 with values of 0.73 N-s and 0.16 N-s, respectively. 

A closer examination of tests K-01 and K-09 will now be made, since there are multiple 

peaks on the pressure plots, and the impulses given in Table 4.3 do not differentiate between 

the different peaks. Table 4.3 gives detailed impulse calculation information for each peak. 

In experiment K-01, the impulse is fairly constant for the first four transducers, and the fifth 

loses its signal after peak arrival. Transducers P6 and P7 show peaks traveling upwards, 

and peaks traveling downwards. The impulse of the upward peaks are 0.03 N-s and 0.024 

N-s, for P6 and P7, respectively. For its downward propagation, they are 0.03 N-s and 0.015 

N-s, for P7 and P6, respectively. There was no data collected for transducer P8. For test 

K-09, the impulses are very small to begin with, then get large very quickly from P4 to the 

top. The largest occurs at P7 traveling upwards, and has an impulse of 0.13 N-s. Traveling 

downwards, the peaks at P6 and P5 have a 0.1 N-s and 0.075 N-s impulses, respectively, 

then the impulses get smaller after that. The propagation speeds are calculated by noting 

the time of the initial rise of each pulse, and performing a simple linear slope calculation 

between each transducer. In tests K-01 peaks, so the position of each transducer is plotted 

as a function of the time that the peak arrives (Figures 4.57 and 4.58). The slope of 

this curve is the peak propagation speed. Both K-01 and K-09 show that the initial peak 

propagation speed is 509 m/s, and the speed drops off as the peak rises to a minimum of 

94 m/s for K-01 and 27 m/s for K-09 (both at transducer P7). In addition, in test K-01, 

it is estimated that the pressure peak was traveling upwards from P7 to the slide gate 

and back down at a rate of 168.9 m/s, then had a speed of 183.6 m/s downward to P6. 

The propagation speeds of the downward moving peaks in experiment K-09 are difficult to 

measure due to the number of peaks present in P6. 

The conversion ratio is the parameter used to represent the work done by the fuel-coolant 

interaction. As noted previously, test K-05 was the only test in which the slug moved 

from the interaction, and the resulting conversion ratio was determined to be 0.058 percent 

Clearly, the energetics from the FCIs with iron-oxide are quite weak compared to the 

NUREG/CR-6623 96

11
molten tin tests, even though the available thermal energy is much larger. Let us consider a 

few visual observations before we address the actual existence of explosions for the WFCI-K 

series in the next section. 

Consider several visual clues that aided in the determination of whether or not an explosion 

occurred, and where it occurred. A visual record of each experiment was taken with two 

different video cameras, and from these it was noted if there was a visible interaction 

in the funnel. This would indicate the occurrence of a delayed spontaneous explosion or 

interaction. In addition, the rupture disk at the end of the expansion tube can be used 

to determine if an interaction occurred inside the test section that caused an increase in 

pressure. On several occasions the rupture disk did break, and corresponding pressure peaks were noted, but the explosion (or interaction) was not strong enough to move the 

slug down the tube. From this information, it was determined that in tests K-05, K-09, 

and K-10, there were fuel-coolant interactions strong enough to break the rupture disk, and 

in tests K-04, K-05, and K-09, there were delayed spontaneous explosions that occurred in 

the funnel after the slide gate closed (Table 4.2).

97 NUREG/CR-6623
Table 1: WFCI Series A - H Experimental Initial Conditions and Results

h'nltfa Conditlons _e"ulb 

Eqmh,.s" uM&a uaW ICOWlNW .blmit Il r.q gger Slug ftntai W&Imcdnm 

Nua* Mass Term Trmp Tamp RsUe Votge Atass Eww Work M 7pv 

ft) (cI (cQ (C) M ) O) Wc) OU)J ) I( 

A-01 3.4 940 85.0 2S.0 1 200 23.4 963.35 2.49 0.258 ESE 

A-02 3.73 980 85.0 3.0 1 200 23. 1079.89 2.05 0.190 TE 

A-03 3.45 940 85.0 13.5 1 200 23.4 963.3 6.75 0.701 ESE 

A-04 3.30 955 850 15.0 1 200 23.4 934 22 0.452 DSE 

A5 358 940 85.0 23.0 1 200 23.4 999.66 325 0.325 TE 

063.50 8 84.5 24.0 1 200 23.4 .8 3.15 0.351 ESE 

"9-01 3.87 97S 85.0 23.0 1 0 23.4 1115.45 1.730.155 OSE 

3-€2 2.?a 9 85.0 23.0 1 0 23.4 822.79 2.49 0.303= OM 

9-03 3.84 970 83.0 15.0 1 0 23.4 1103.84 5.04 0.4•7 ESE 

C-l 3.31 978 85.0 24.0 1 200 23.4 958.59 3.66 0.383 TE 

C002 3M 937 85.0 24.0 1 100 23.4 921.72 - - SE 

03 3.48 983 85.0 24.0 1 400 23A 1010.2 2.50 0.257 TE 

881 83.0 10.0 1 450 23. 944.57 2.48 0.263 TE 

0.05 3.09 913 84.0 13.0 1 100 23. 842.19 .67 0.317 TE 

- 3.19 923 88.5 24.0 1 300 23.4. 873.S9 2.43 0.278 TE 

D-01 3.19 901 84.3 '24.0 1 200 . 0.53 859.356 O3 0.097 TE 

D002 3.31 911 84.6 24.0 1 200 42.2 899.94 3.78 0.420 TE 

0-03 3.51 877 84.1 24.0 1 200 42.2 924.1 4.9c 0.530 TE 

004 3.19 901 88.5 24.0 1 200 58.8 857.47 - - "E 

D-05 2.91 819 82.9 24.0 1 200 71 723.55 2.50 0.345 ESE 

D-06 3.01 889 83.6 24.0 1 200 71.2 802.13 2.40 0.299 TE 

E01 3.85 87.0 24.0 1 200 23.4 815.8 3.44 0.422 ESE 

-02 3.6 790 88.0 24.0 1 23.4 879.96 2.65 0.301 TE 

.-03 0.72 491 93.0 24.0 1 200! 23.4 - - - ESE 

F-02 0.99 782 27.0 24.0 1 0 23. - - - ESE 

F-03 1.5 850 62-5 24.0 1 200 23A - - - ESE 

PF-04 1.3 871 72.0 24.0 1 2o0 23.4 - - ESE 

G-01 1.62 -84 25.0 24.0 2.5 0 23.4 454.03 - - NE 

0.02 1.65 820 25.0 24.0 2.5 200 23.4 .437.9M - - NE 

0-03 1.70 875 33.0 24.0 2.0 0 23. 469.02 - - 'NE 

G044 1.70 853 28.0 24.0 1.5 023. 461.59 - - NE 

0-05 0.82 850 27.5 24.0 1 0 23. - - - ESE 

0 63.7 880 27.5 24.0 2.5 0 23.4 1003.$7 - NE 

M-01 1.01 855 82.0 24.0 1 200 23.4 - - - ESE 

H-02 87.5 24.0 1 200 23. 908.89 - - E 

H-03 4.2 920 88.0 24.0 1 200 23.4 1151.07 4.83 0.416 TE 

H-04 4.48 86.0 24.0 1 200 23.4 122679 2.69 0.219 7! 

ES!.- Early SPon==ou Explosian 

DSE Delayed Span== e xlosin mu 

TE a Triggered Explsio 

NE = No Explosion

Table 4.1: WFCI series A-H experimental initial conditions and results

NUREG/CR-6623

11

98
Table 4-2: Pressure Peak Propagation Speeds and Impulses for WFCI Experiments 

_ _ _ _ WFC1.l "rrfo WFCI.2 Trd K-01 K102 K.03 K.04 .KO5 K.07 K-0O8 IK-9 K-10 

Peek Pressure (MPe) 2.5 1.7 2.4 1.3 5.3 3.0 5.0 2.4 2.0 9.7 4.4 

Impulse (MRa-$) .1 4.6 55.8 19.5 9,7 22.6 30.4 25.6 18.5 141.8 68.0 speed.• ..... one" .. .. . --. .. - . ..... 

Peak Pressure (MPa) 2.5 1.6 1.5 0.4 1.3 0.4 1.2 1.6 

Impulse (mrN.S) 7.6 4.4 45.1 22.6 2.4 16.1 28.2 11.7 10.9 107.7 45.1 

Speed jis 1402.3 • 1435.7 509.6 730.1 254.8 238.6 473.23 552.1 225.5 509.61 565.2 

Peak Pressure (MPe) 2.6 1,2 2 -.. ...... - 0.7 1.1 

Impulse (mIN-e) 7.6 4.4 57.7 11.? 0.5 5.4 23.1 9.4 8.0 26 45.2 

smm m 1402.3 1423.3 17..1. ... - ....... ... .. -44.5 . 560.3 

Peak Pressure (MPS) 2.2 1.4 2.8 ..... 3.3 1.0 

Ifmpulse (mM..) 6.8 4.4 34.6 15.3 0.4 4.6 20.9 8.8 8.1 59.1 42.4 

speed (J) 1402.3 1331.5 114.2 -. -. ... .-. -- 70.5 ... 

Peak Pressure (MP&) 2.3 1.5 1.2 . 9.. 9.1 

Impulse (mPN-S) 7.8 4.4 6.4 10.5 0.0 4.6 18.3 7.7 8.6 148.7 33.8 

Speed (n*1 1529.8 1641.4 132.5 .. - -.... 50.2 ... 

Pressure (MP&) 2.5 1.4 4.7 -.... 6 'r SPeak 

Impulse (MN-.e) 7.3 4.4 61.6 14.2 0.0 5.3 14.7 8.8 8.6 149., 27.6 

speed (mis) 1528.0 1346.8 128.6 ... .... ... ... 

Peak Pressure (A#Pe) 2.2 1.3 6.5 - -. .... 3.81. 

Impulse (maN-) 3.1 4.0 73.2 14.7 0.0 4.3 24.5 8.4 8.6 160.4 34.5 

S &M) 138. 1419.8 93.8 1. 28.61 .. 

en Peak Pressure (MPe) 19 1.3 ... ... .. 8.3 ... 

Impulse (MI.e) 2.9 3.9 --- 11.9 0.0 3.6 24.8 6.4 8.4 114.9 14.5 

Speed (m/st 1270.01 1567.7 .. . ... ...- ... ... 105.3 ---

€.0

Table 4-2: WFC1 series K pressure peak propagation spee4 and Impulse data.
Table 4-3: Impulse Data for Tests K-01 and K-09

NUREG/CR-6623 100

1!

K-01 K-49 

P1-up Impulse (mN.s) 27.9 5.4 

T7me Range (fms) 18-21 17-21 

P2-un Impulse (mN-s) 25.8 6.2 

Time Range (fms) 18-21 17-21 

P3 - un Impulse (mN.s) 30.8 0.0 

Time Range ,ms) 18-21 17-21 

P - ui Impulse (MrN-s) 27.2 18.5 

Tme Range (fms) 18.5-20.5 19-23 

PS up Impulse (nNos) 2.0 35.4 

Time Rane, (fms) 19-21 21-25 

P6-U Impulse (mN-s) 30.8 35.9 

Time Range (fms) 20-22 23-27 

IT.up Impulse (mN.s) 22.4 130.2 

Time Range (ms) 20-22 25-31 

El Impulse (mN.s) - 103.6 

Time Range (ins) - 25-31 

P, - down Impulse (raN.s) 31.4 

Time Range (ms) 24-26 

PS- down Impulse (MN.s) 14.7 103.1 

Time Range (ms) 24-28 27-31 

PS - down Impulse (mN-s) - 71.5 

Time Range (ms) - 29-33 

P4 - down Impulse (inN-s) - 0.0 

_Time Range (ms) - 29-331
Table 4.4: Experimental Conditions of Experiments in Comparison

H,+ (in) H4~, (in) 

D (mm) 

Tc (0C) 

Tf (°C) 

mf (kg) 

m, (kg)

Hall 

0.85 

0 

28 

85-95 

600-750 

0.5- 1.2 

0 

0

K-21 

1.1 

0 

95 

85 

1080 

6.5 

0 

0

Baines 

1 

3 

30 

90 

800 

0.5 -0.8 

2.4 

3.4

WFCI 

1.5 

3.9 

87 

83 -, 87 

819 ,,i 911 

3 -, 3.5 

0.5 -' 71 

0.06 - 8

0.06 8

Table 4.5: Debris Masses Collected from the WFCI-G 

Series Tests

WFCI 

Series 

G-01 

G-02 

G-03 

G-04 

G-05 

G-06 

F-02

M0 

(kg) 

2.4 

2.4 

2.4 

2.4 

2.4 

5.0 

5.0

Mit1 

(kg) 

1.60 

1.66 

1.70 

1.70 

0.82 

3.67 

0.99

(kg) 

0.52 

0.70 

0.57 

0.62 

1.35 

1.21 

3.79

St•j 

(Kg) 

2.12 

2.36 

2.27 

2.32 

2.17 

4.88 

4.78

Remarks 

NE 4 

NE 

NE 

NE 

SE5 

NE 

SE

1Collected mass in the test-section tube 

2Collected mass outside of the test-section tube 

3Total recovered mass 

'No Explosion 5Spontaneous Explosion

101 NUREG/CR-6623
Table 4.6: Experimental Results of the WFCI-A, B, C and D Experiments

tezp 

ms 

9.66 

9.01 

8.21 

8.58 

-307.7 

8.58 

8.21 

12.04 

1189.2 

663.4 

-67.8 

9.31 

13.97

At, 

ms 

107.1 

416.1 

376.1 

n/a 

622.2 

636.1 

358.0 

283.6 

392.4 

332.0 

228.8 

260.0 

319.5

Remark 

TE 

TE 

TE 

TE 

ESE 

TE 

TE 

TE 

DSE 

DSE 

ESE 

TE 

TE

S peak pressure was exceeded the maximum range of pressure transducer 

b four pressure transducers (instead of eight in present study) were installed 

C estimated by the incompletely collected debris due to the early spontaneous explosion

NUREG/CR-6623

11

Test 

Series 

D-01 

D-02 

D-03 

D-04 

D-05 

D-06 

A-02 

A-05 

B-01 

B-02 

B-03 

C-01 

C-05

CR 

0.068 

0.365 

0.442 

n/a 

0.317 

0.274 

0.184 

0.324 

0.153 

0.297 

0.453" 

0.384 

0.309

Pk 

MPa 

9.4 

16.6 

>17.02 

8.8 

n/a 

10.0 

> 7.0w 

> 5.1a 

n/a 

5. 7b 

> 4.0' 

> 3.8' 

>17.08

uprop 

m/s 

203 

217 

271 

203 

n/a 

203 

282 

222 

n/a 

216 

157 

210 

190

102
-2900 -1900 -400 0 

Time (msec) 

Figure 4.1: Sequential Events of the Typical WFCI Tests

NUREG/CR-6623

T-Cýdble Plug S-Gate 

Arival OpeLn0a

103
PI c1ig 

en

S-Gate

Ue

T-Cricible 

Anival

I J I I

-3000 -2000 -500 

Tie (mSec)

0

Figure 4.2: Sequential Events of the WFCI-A-01 Test

NUREG/CR-6623 104

1!

U

a6•

g•
cn 

rT4 

4-4 

0 

(a 

-4.2 

tko ... 4 

rZ4
5 

4.5

,.,3.5 

S3 

2.5 

2 

1.5

I F

0.5 [

0

p I II I i I 

0 A A 

aIy4 

M 

M 

A 

I I I I I I I I I

0 50 100 150 200 250 300 350 40 450 500 

Voltage (VDC 

Figure 4.4: Peak Trigger Pressures vs. Supplied Voltages 

NUREG/CR-6623 106

1!

41
a

En

2 

4 

3 

2 

1 

0 

3 

4 

3 

2 

I 

0 

S 

4 

3 

2 

I 

0 

S 

3 

2 

I 

0 

-1

TDmdE (MS4SEC)

5. 

P--0-23.dat"- I 

9pt3-0-23.d t:..... 

2

17

0 10 12 14 16 18 20 

TINIR (IVS ECFigure 4.5: External Trigger Pressures Generated by a Supplied Voltage of 200 V 

107 NUREG/CR-6623
15 1 1 

14 Delay Time * 

13 

12 

10 

I

71 

6 

S10 

0 50 1O0 1.50 200 2.50 :300 350 400 450 500 

Voltage (D 

Figure 4.6: Average Time Delay of the External Trigger vs. Supplied Voltages

NUREG/CR-6623 108
E 

r 

C 

0. 

V 

0

10

9[

8I

5

4

31

2F

I .

n

I I 

PeaklmpluseT e 

e 

8 8 

o 8 

o0 

8 

I I I I I I t I

0 50 100 150 200 250 300 350 400 450 

Voltage (VDC)

500

Figure 4.7: Product of the Peak Pressure and Impulse at the First Pressure Transducer 

with respect to Supplied Voltages 

109 NUREG/CR-6623

7

I
Slide Gate

Magnet Piston Coils 

.../....Z........... IJ ...........

.. ....... ... .t ............. .. ....... 

Slug Control Volume, m.(t) 

------ Time =t I0.o...... ..... ...... .. a....... 

.. .( .• . .. .. ..... .....

External Interaction Volume 

Trigger

Figure 4.8: Measurement of the Work done by Energetic FCIs on the Slug in WFCI Facility

NUREG/CR-6623 110

i1
:1

TINME (MS) 

Figure 4.9: Pressure Traces of the KROTOS-21 Test [66]

NUREG/CR-6623

5 

0 

0 

0

5 

0 

5 

0

"Ia.

5 

0 

0 

0

0

ill
. . .... .... 

Dpm3-mia- .d..t -

o .-a-~ .d1 

"'IDutmy-.4Ix .dat 

Si 

0 2 

Figure 4.10: Pressure Traces of the WFCI-A-01 Test 

NUREG/CR-6623 112 

i i I i i i ; _ •1!
S

a 

S

C

0

z 

C 

S 

C

z 

C 

S 

C 

S

0

"riMZ IF--CP-SSSCEA.•)

Figure 4.11: Pressure Traces of the WFCI-A-02 Test

113 NUREG/CR-6623

dad 

m-a 

In 

a
S

Cp 

0 

S 

...... ... ... .. ... . . . .. . . . . . . 

"'D u my-mix .dotf*" -

0 

Fgr 4.2 Pressure4dat Trcso-h-WC--4Ts 

S1 

0 

Si l l l l l l i 

C p 2 n- 4d t" 

o 

1100) 105 1 10 115S 120 125 ]1.304• 135 1.44 145 ]I.SO 

Figure 4.12: Pressure Traces of the WFCI-A-04 Test 

NUREG/CR-6623 114 

~1!
C 

Sn 

Sn 

a

flrr.IF CN4S5C)1

Figure 4.13: Pressure Traces of the WFCI-A-06 Test

115 NUREG/CR-6623
S 

0 

-eh~ ~ .L4b-2 ~~ .. .............. 

0 

S 

........ 0 ...... .. .. . ...... 

&Sa0 ~ 45 70 ýSoa ý5O '700 71 0 

•a-vr.1 "C[M :Y-3=-a¢> 

Figure 4.14: Pressure Traces of the WFCI-B-02 Test 

NUREG/OR-6623 116

1!
S 

00 . . .. .... ... .. . . .. . 

S 

40 

"".. ... . . '..

o 

-SO -75-7 -05 -- so -55 -SO -. Cs -40a -SS -30 

""mS--lt: -0p .dMe" 

Figure 4.15: Pressure Traces of the WFCI-B-03 Tests 

117 NUREG/CR-6623
0 

. .. .. ... ... . ........... 

........................ .d.. - ------

0 

5 ais=Da 0.:>2 35: _0 d 0 d&Z so 

NU C-R-O3pod3 t1 -

S 

im0 

"Dpu1y-cSp.dt" -

S 

0 

0 S 10 1 5 20[ 2S5 30 3 S 40 4.S1. SO 

Figure 4.16: Pressure Traces of the WFCI-C-03 Test with a Supplied Voltage for the 

External Trigger of 400 V 

NUREG/CR-6623 118 

1i
5

0........ . . . . 

SS 

S0 

... . ...... .. . .. .......... ... ............. .. . .. . . ...... ........... . 

...... .................... ... .................. ............... .. .... ........................ . 

0 

0 10 is 20 25 30 35 443 4S SO 

Figure 4.17: Pressure Traces of the WFCI-C-04 Test with a Supplied Voltage for the 

External Trigger of 450 V 

119 NUREG/CR-6623
5

.. . . .I ......... . .. . ................ ............. . ............................. ..... 

- .. . ...... . ......... .... . .... ...................... 

o i3 

| !xP ! | 

0 

S 

0 

S 

0 S 10 iS 20 25 30 35 4O 4S SO 

Figure 4.18: Pressure Traces of the WFCI-C-06 Test with a Supplied Voltage for the 

External Trigger of 300 V 

NUREG/CR-6623 120

I!
S 

0

5 

G0

1

C . .. ... ... 

. .. **........ . .... .... .. .... ... . . 

=0os 

Figure 4.19: Pressure Traces of the WFCI-C-05 Test with a Supplied Voltage for the 

External Trigger of 100 V 

121 NUREG/CR-6623
0.45

0.40

0.35

0.30

6 0.25

0.20

0.15 1

*1 -

.5 2.0 2.5 3.0 3.5 4.0 4.5

Peak Trigger Pressure (MPa) 

Figure 4.20: Conversion Ratios with respect to the Peak External Trigger Pressures

122

I!

WrSrVS' ME IM ) I STS(SM 

WrS :Weak Trigge-Strength 

ITS Intermediate Trigger Stretgth 

STS Strong Trigger Strength 

DE Double Explosion 

SB Single Explosion 

0

NUREG/CR-6623

i
z

......... 

........... .. i. 

• ~~ .. .. .......... .. .. . .. .. . .. .. . .. . .. .. . .. .. . .. . 

o> . ... ................................. 

0 a 1• is =0 =a MC> aS"o s .4o .4-s sa 

Figure 4.21: Pressure Traces of the WFCI-E-02 Test with the Fuel and Coolant Tempera

tures of 790 and 86 °C, respectively 

123 NUREG/CR-6623
S 

S 

0 

mD--0SId..t 

S 

"S 

"pm1-e-OS.dai" -

S 

o 'K, -1i0 1 2 -10 0 B -9. -94 -9 0 -0 ""N4 p.4-S -3.a= 

Figre4.2:PrssrTrcsothWFIE0TetwtthFuladClntemr

turs o49 an 3C esetvl 

NURE/Ci6 2i12 

"'pml- -O3.d t'" -
2 

1 .3 

1 

0.3 

1.3 

0-s

1Lm40f002.-=0 

1 .3 

-'o ~jp [ Cr.I 4t 

S0.3 

S1.3 

tures of72ad2 epc ivl 

1p2m5 -f--0)2.d°

1.3 

"L I "~pr2- r-02.dat 1 .3 

1.3 

-1o - __- 1000 - B.. . _ .... ~ .. 0 .- 20 

Figure 4.23: Pressure Traces of the WFCI-F-02 Test with the Fuel and Coolant Tempera

tures of 7/82 and 27 °C, respectively 

125 NUREG/CR-6623
pm8-f-OS..4..- -

S 

0 .......... 

S 

.d ..... "". . m S- -4 t 

0 

o 

S"'~pm3rt-f-0S.dat" -

S 

-. 4 -Soopa2- f-03. dmat -

S 

S 

0 

Figure 4.24: Pressure Traces of the WFCI-F-03 Test with the Fuel and Coolant Tempera

tures of 850 and 62.5 00, respectively 

"NUREG/CR-6623 126 

51
tue of 87 an-2*,rsetvl 

1 0 

"pm -g-c-4..d.t,~ -

0 ~ -B B0 -4 -- ~ m O -70 -~ 

Figure 4.25: Pressure Traces of the WFCI-F-04 Test with the Fuel and Coolant Tempera

tures of 87/1 and 72 °C, respectively 

127 NUREG/CR-6623
Thermal Effect of Constraint

3offing/

Mechanical Constraint

fI !

10 100 1000

Time (ms)

Figure 4.26: Time Scales with respect to the System Constraint 

NUREG/CR-6623 128

11

Time 

Scales

I

1

wtial
15 20 25 30 as

Figure 4.27: Pressure Traces of the WFCI-D-04 Test with a Total Constraint Mass of 56.8 

kg, < m >=6.38

129 NUREG/CR-6623

o 

O 

S 

0 

S 

0 

0 

C 

0 

0

0
I 0 . I I . I . I . I . I . , . 0 0 

0.4

Z 0.3

0.2-

0.1-

0.4

-1 I * I • I. . . .

9

-600 

-Soo _ 

-400 

-300 

-200

100

9

< m >, Axial Constraint 

Figure 4.28: Conversion Ratios and At,I., with respect to Axial Constraint 

NUREG/CR-6623 130

1!

I

-Fitted Data for CR ...... Fitted Data for A•t 

0 ,wFCI 

A KROTOS-21 

V Baines 

o Hall

• D-01

v

0.5 700

|D
A S 

-lu-s

A A

A 

S 

S

0

0 i •- 3 4 5 6 7 - m -, Axial Constraint

S

*-14 

18 

16 

-14 

-12 

-10 

-8 

-6 

-4 

-2 

.0

9

aI,

Figure 4.29: Conversion Ratios and Peak Pressures with respect to Axial Constraint 

131 NUREG/CR-6623

9 

Cý

0.8 

0.7

0.6 

0.5 

0.4 

0.3 

0.2 

0.1 

0.0

-1

. . . . . . . . . . . . . . . . . 20
onU "

0

2.2

2.0

1.

1.6

1.4

S1.2

" 1.0-. 

• 0.3

0.6

0.4

0.2-

................. 

Fi- 4..CV,.--3gCw 

V1V-50% S...... 

Figure 4.30: Conversion Ratios with 

Parametric Model (UWFCI Code) 

NUREG/CR-6623

" 1*0i1 

< NI>.

"-2'0 "2 5 " 3!0 "35 " 40 

Axial Constraint 

respect to the System Constraint calculated by the 

132

I!
Time (ms) for Conversion Ratio

1 

5- • 

0.8 

2 

1 / 0.4 

,~0.0 

0.0 0.1 0.2 0.3 0.4 0.5 

Time (ms) for Pressure 

Figure 4.31: Conversion Ratios and Pressure Histories calculated by the Parametric Model 

(UWFCI Code) with respect to < m > of 3.3, 11.2 and 44.4

133 NUREG/CR-6623
O.S

0 

0-s 

0 

0 

-' 0.51 

S0 

0

NUREG/C

.. ..... .: ..... ; .•. . . -• 

0 

0).5 

0. 

-00-00-800 -600Q -. 4400 -m00 a 

Figure 4.32: Typical Pressure Traces in Pure Water (WFCI-G-05) 

R-6623 134

1!
0.5 

0 

0.5 

0 

0.5 

0 

- 0.5 

0 

-,0.5 

Im 

0

p.. 1-B.}1 - -

-1000CM -043C -. 60>0 -. 40-0 -2400 0 200 .400

Figure 4.33: Typical Pressure Traces in Polymer Solution (WFCI-G-01)

135 NUREG/CR-6623

..... ................ ......... .. .......... ..•_... • .......... 

-- . - - --- 

pm& AR)1 d 

.S. . . . . . . . E ...L ., .,Jh t~ ,J• J., 

' 4..

0 

0-s

C>
Figure 4.34: Debris Shape of the WFCI-G-06 Test in Size between 1 to 2 mm

NUREG/CR-6623 136

1!
Figure 4.35: Debris Shape of the Pure Water Test (WFCI-G-05)

137 NUREG/CR-6623
'40 

4-D 

44.4 

V) Vi0 

. . ...... 

* ~ *~ ~ . ~ ~.. .... I .~~** ...

'o

co -0 

C4 

rZU
10, 0 ........... .. S.. ...... ... ... i . ... .. ........ i ........... . . .i. . . S... ... ... .... . l... ... ......... I .................. I..... 

10, 

!! !! ! ! ! ..! .......... !! i !! !. .• •i !! ! 

lop 1!!!!!!!!!• 0-!!!!!!!!!

U, 

6) ifl'

E

10-2 

10 

1o-

"'...-: - DEB-G-OI (r, 2..

5

) 

....... A DEB-G-02 (nr.2.5) 

-- V D-EB-G-03 (n--2.0) 

-=0-- DEB-G.04 (vr- 1.5) 

-- -- DEB.-GO (N1-1.0) 

SDE-F-02 (N- io)

lo2 I13

Debris Size (an) 

Figure 4.37: Distribution of the Cumulative Debris Mass Fraction

139 NUREG/CR-6623

....... • ... !! i

I

......... ... ;w ........ ..................

I!
I *, -. -

14, 

125 

10, 

8

6

4

2-

13

n- -..

0.5

* I 

0

Spontaneous 

Explosions / . I 

A

1.0

0-01,0-03,.0-04 

0-02 (Trigger) 

G-06 (Doubled mass) 

C-05 (Pure water) 

F-02 (Pure water)

1.5 2.0 2.5

Viscosity Ratio, 1, 

Figure 4.38: Sauter Mean Diameters of Collected Debries in the WFCI-G and F Series 

Tests 

NUREG/CR-6623 140

1I

I 

Is I

3.0
"20 25 50 35 .40 45 50 "farIME C(ISSC)

Figure 4.39: Pressure Traces of the WFCI-H-03 Test

141

5

0 

S

0

0

3

C 

Baa 

a 

Baa 

a

S 

0 

S 

0

S 

0 

0

0

NUREG/CR-6623
0 

0 

0 

S 

0 

S 

0

S 

0 

S 

0

TIhcM (NSS•C)

Figure 4.40: Pressure Traces of the WFCI-H-04 Test

142

1!

C 

= 

1*3 

a 

153 

a

NUREG/CR-6623
12000 

"lev-h-03.dat 

"Iev-h-05.dar 10000 

II I t "6~ 8000 .mA 

- I t 

~6000 __ 

\ I I: ', 

4000 

-2000 , 

-2000 -1500 -1000 -500 0 500 1000 

"Time (msec) 

Figure 4.41: Level Swells of the WFCI-H-03 and H-05 Tests

143 NUREG/CR-6623
0.9 . Avmgae (LM) 0 .8 ... ...- ... 

L4wer UmitLm 

0.7 

0 0.6 

S 0.4--

0. 

0 3 4 ..... ..... ., .. .... ..... ... .. . • . • . - . .

o . .... ... i 

0.2 -- i : e *. 

-0 -0-. . -2 -1.8 -1.6 -1.4 -1.2 -1 -0.8 -0.6 -0.4 -0.2 0

Time(sec) 

Figure 4.42: Void Fractions of the WFCI-H-05 Test Comparing with the Visual Observation 

NUREG/CR-6623 144

i!
SHicks-Menzies

........ Board-Hal 

U WFCI-Expt 

* HalI-Expt A Baines-Expt S al-6pt•

E

I.. 

aLA U

lb

0.5 1.0 1.5 2.0 2.5 

Mass Ratio of the Coolant to Fuel

0

3.0

0 2 4 6 8 10 12 14 16 18 20 

Volume Ratio of the Coolant to Fuel

Figure 4.43: Conversion Ratios with respect to the Coolant to Fuel Mass Ratio Compared 

to Thermodynamic Models

NUREG/CR-6623

jul-I

10

0

I

- - - i I I l • • ! I

0.1 

0.01 

0.0

145
5 

4 

3 

2 P8 

0 

4 18 20 22 24 26 28 • 

3 

2 P7 

1 

0 

3 

S 2 P6 

A 1 

* 0 

5P 

4P 

2 1 

0 

4. is 20 22 24 26 28ZD 

3 

2 P4 

0 

5 I -- -I I4 

4 18 20 22 24 26 28 

3 

2 P3 

1 

0 

4 18 20 22 24 26 28 

0 _ 

18 20 22 24 26 28 30 

TIME (msec) 

Figure 44: Trigger pressure histories for WFCI-I apparatus charged to 200 V 

NUREG/CR-6623 146

V
5 

4 

3 2 P8 

0 

5 I 

S26 2097 2098 2099 21 DO 

3 

2 P7 

1 

0 

Z 5 26 209'7 2098 2099 21 00 

3 

2 P6 

A 1 

* 0 

2 26 2097 5P 2098 2099 2100 

2 P1 

0 

5 

41(6 2097 2d98 2099 21300 

3 

2 P4 

3 

2 1P2 

2? 6 2097 2098 2099 2100 

3 

2 P2 

0 

4W026 2d97 2098 2d99 210DO 

TIME (msec) 

Figure 45: Trigger pressure histories for WFCI-1 apparatus charged to 400 V 

147 NUREG/CR-6623
Cm 

10 

0.. 

0

0 

a 

0e 

ai

a CS* 

(No) uOl1lSOd 

Figure 4.46: WFCI-1 trigger peak propagation curve 

NUREG/CR-6623 148

o. as 

a• 

a~ 

a 

a

I

g

A
WFCI-2 Trgger Peak Propagation

y1- 143.138x - 300033.8491 R"" 1.000

I 2 . 0 II

20M.1 2096.2 2096.3 2096.4 209.5 . 2096.6 2096.7 2096.8 209.9 2097.0 2097.'

Tim. (ms)

160

140 F

Co

I 

'1 

'-I 0 

0 

0

120 

100 

U 

8e0 

40 

20 

* 0

t4 

0i
18 21 24

TME (msec) 

Figure 48: WPU-K-01 transient pressure histories 

NUREG/CR-6623 150

I!

8 

6 

4 

2 

0 

8 

6 

4 

2 

0 

8 

6 

4 

2 A 

, 0 

* 8 

8 

0 

CIO 8 

S 6 

S 4. 

V I 2 0 

0 

8 

6 

4 

2 

0 

8 

6 

4 

2 

0 

8 

6 

4 

2 

0

P8 

P7 

P6 

P5 

P4 

P3 

P2 

P1

27
6 

4 

2 P8 

0 

81 1 1 

6 18 21 24 

4 

2 P7 

0 

8 i II 

6 18 21 24 

4 

2 P6 

A * 0 

8 

t 1821 24 2 

6 

S4 

"4 P5 94 2 

* 4 

0 

S8II I 

618 S6 21 24 

4 

2, P4 

o2 

0 

6 18 21 24 

4 

2 P3 

0 

81 

18 21 24 

4 

2 P2 

0 

6 18 21 24 27 

4 

2 P1 

1821 24 27 

TIME (msec) 

Figure 49: WFCI-K-02 transient pressure histories 

151 NUREG/CR-6623
8 

6 

4 

2 

0 

8 

6 

4 

2 

0 

8 

6 

4 

2 A 

* 0 

* 8 

g 4 6 

i 2 

0 

4 

2~ 6 

•i 4 

2 

0 

9 

6 

4 

2 

0 

8 

6 

4 

2 

0 

8 

6 

4 

2 

0

18 21 24 

TIME (msec)

Figure 50: WFCI-K-03 transient pressure histories 

NUREG/CR-6623 152

1!

PS 

P7 

P6 

P5 

P4 

P3 

P2

PI

27
.18 2 I 

18 21 24 

isI 2 

18 21 24 

I 21 24 

18 21 24 

Is 21 24 

I I 

18 21 24 

I I I 

18 21 24

21

TIE (msec)

24

Figure 51: WFCI-K-04 tansient pressure histories 

153 NUREG/CR-6623

A 

U

P8

P7 

P6

P5

P4

P3 

P2 

PI

18 27
6 

4 P8 

2 

0 

$ 1 I I 

18 21 24 6 

4 P7 

2 

0 

$ SI I 

18 21 24 6 

4 P6 

2 

A 

3 0 

3 

18 21 24 

CS 6 

P5 

So2 

618 21 24 442 

0 

81 

18 21 24 

6 

4 P3 

2 

0 

S I I 6 18 21 24 

4 P3 

2 

8 1 I 618 21 24 

4 P1 

2 

0 

18 21 24 27 

TME (msec) 

Figure 52: WFCI-K-05 transient pressure histories 

NUREG/CR-6623 154

11
6 

4 

P8 2 

0 

8 4 I I 

6 21 24 

4 2 P:7 

0 

1821 24 2 

6 

4 

2 P6 

I 0 

I* 8! 

6 18 21 24 

4 P5 

0 

21 24 17 

4 

4 P4 2 

0 

8 I I I 

6 18 21 24 27 

4 

P3 

2 

0 

6 18 21 24 

4 

2 P1 

0 

18 21 24 27 

.TIME (msec) 

Figure 4.53: WFCI-K-07 transient pressure histories 

1.55 NUREG/CR-6623
8. 

6 

4 P8 

2 

0 

81 i 

6 18 21 24 2 

4 

2 3P7 

0 

18 21 24 6 

4 

.2 P6 

I 0 

I 8 ] I 6 S6 18 21 24 

2 P5 

0 

8 I 

6 18 21 24 

4 P4 

1 2 

v 0 

8 1 I I 618 21 24 

6 

4 P3 2 

0 8 Il I 

6 8 21 6 24 

4 P2 

2 

0 

8 I I 

6 18 21 24 

4 P1 2 

0 

18 21 24 27 

TIME (msec) 

Figure 4.54: WFCI-K-08 transient pressure histories 

NUREG/CR-6623 156

1f
PI

17 22 27 32 

TIME (msec) 

Figure 4.55: WFCI-K-09 transient pressure histories

NUREG/CR-6623 157

A 

I 

C', 

C', 

V

P8

P7 

P6 

P5 

P4 

P3 

P2
8 

6 

4 P8 

2 

S I I I 

is 30 1932 1934 1936 1938 4 10 

6 

4 P7 

2 

0 . I II 

i0 1932 1934 1936 1938 40 

6 

4 P6 

2 

0 

'I l 

i s 30 1932 1934 1936 1938 U 10 

4 P5 

SII I I 

1932 1934 1936 1938 1 

6 

I 4 P4 

III I • 01932 2P 1934 1936 1938! 0 

III I 

is 30 1932 1934 1936 1938 I0 

6 

4 P3 

2 

1930 1932 1934 1936 1938 1940 

TIME (resec) 

Figure 4.56: WFCI-K-10 transient pressure histories 

NURiEG/CR-6623 158

1.
WFCt-K-01 Pressure Peak Propagation 

140 

120 

100 

"0 

- * 

cI 

S e0 -d 

40 

20 

S0 . . . . .. .:. . . .. .. . . . . .. . . . . . . . . ., 

17.5 1.0 18.5 19.0 19.5 20.0 20.5 21.0 21.5 

Thme (ms) 

CA*
(UM) UOtUsod 

Figure 4.58: WFCI-K-09 pressure peak

NUREG/CR-6623

o a 0 IV CmI 

propagation curve

160

11

ta 

0 

I

06 

Ct 

9

Coj

0 

Uq
V/PCI-K-OS Slug Position KIstory

Thme (ms)

CA' 

to 

0 0 0

t43

900
aa 

t 

1. 

z 

+ 

x ID 

a l C 

0. 

So 

*(w) UUMd 

Figure 4.60: WFCI-K-05 slug position as a function of time with polynom 

curve fit 

NUREG/CR-662316 

106

P!

jal
Chapter 5

Analysis of Experimental Results 

5.1 Temperature Effects 

The initial temperatures of the fuel and coolant may have a noticeable affect on the mixing phase of the FCI through the coolant vaporization during mixing and the mixture void 

fraction. This in turn can affect the ability to trigger the FCI event. Based on the results 

from the WFCI-A, B, C, E, F and H series, the relationship between temperature and 

the likelihood of spontaneous explosions was investigated and shown in Figure 5.1. In 

this figure, other available experimental data with tin, [60, 62, 63] which have similar 

geometrical scales are also included and compared with the temperature cut-off line of the 

single drop tests [53]. 

Note that temperature effects on the vapor explosion are dependent on the fuel mass in

volved in an experiment. As shown in Figure 5.1, the cut-off temperature for a spontaneous 

explosion in the large scale tests is higher than in smaller scale tests. The spontaneous ex

plosions in large scale tests occurred at coolant temperatures more than 20 °C higher than 

that in the single drop tests. Such a scale dependency may be explained by considering 

that in large scale tests, fuel in greater quantities, can have a range of diameters and length 

scales during mixing. Because of the natural oscillations occurring during film boiling 

around the fuel surface and considering the wide array of fuel diameters in the mixture, an 

FCI can be easily triggered by any perturbation, whether it be an external pressure source 

or some relative velocity or pressure fluctuation. 

Second, the effect of coolant subcooling on the FCI is more profound than the effect of 

the fuel temperature. Note that this observation for tin is predicated on the fact that 

the fuel superheat is quite large, >500 degrees.This is because the net vapor produced 

and vapor film stability strongly depend on the net energy transfer at the coolant vapor

liquid interface. As the coolant bulk liquid temperature decreases (subcooling increases) 

the net amount of vapor produced during the mixing phase is significantly reduced, due 

to the large temperature gradients near the vapor-liquid interface which promote rapid 

vapor condensation. This in turn causes the vapor film surrounding the fuel particles to 

be relatively thin and more easily destabilized due the fluctuations encountered in the 

large scale boiling processes. Thus, an increase in subcooling would promote spontaneous 

interactions. As shown in Figure 5.1, fuel temperatures above 800 °C have little effect on the

163 NUREG/CR-6623
likelihood of spontaneous explosion. However, modest changes in the coolant temperature 

can induce early FCIs and spontaneous explosions. 

Finally, we note in Figure 5.2, the relationship between the explosion conversion ratio and 

different fuel thermal energies. The figure shows that conversion ratios of 0.2 to 0.7 % are 

not dependent on fuel energy within the variation of the fuel thermal energy from 0.8 to 

1.25 MJ for a tin simulant. Thus, for a particular fuel simulant, once the temperature is 

above some threshold which satisfies the criterion for molten fuel and the ability to mix 

with the coolant without early spontaneous FMIs, variation in the fuel thermal energy does 

not have a first-order effect on energetics. This threshold is related to some minimum 

amount of melt superheat. In contrast, a different fuel simulant with sufficient superheat, 

but larger thermal energy per unit volume, would be expected to produce a vapor explosion 

with a a larger conversion ratio. Figure 5.3 illustrates this point for molten alumina melt 

compared to tin. We will revisit this later when we investigate the question of scaling. 

5.2 Propagation Speed Characteristics 

The propagation velocity is measured using the pressure traces from the eight axially dis

tributed pressure transducers. For most of the WFCI tests, the propagation behavior of 

the explosion pressures is shown in Figure 5.4. In this figure, the positive value of the 

propagation speeds represent the upward propagation from the bottom to the top of the 

test section. The negative values represent the reflected explosion wave propagation from 

the top to bottom of the test section. In the explosions observed in most of the WFCI tests, 

average propagation speeds ranging from 800 to 1600 m/s were measured at 0.2 m above 

the bottom of the test section. This shows that the leading edge of the fuel had not reached 

this position during the mixing period. The trigger pressure wave propagated through sin

gle phase water with near sonic velocity. However, the propagation speed quickly drops to 

about 200 to 400 m/s as the wave encounters the mixture of fuel and coolant liquid/vapor. 

The explosion first occurred at or near the leading edge in a location between 0.3 to 0.7 

m from the bottom and escalated upwards even where the local void fraction was high; 

i.e., more than 25 %. Possibly because of this high void fraction near the top of the test 

section, the explosion pressure peak propagates with a velocity of about 100 m/s. In some 

of the double explosion cases, the second explosion occurs near the upper rigid boundary 

and propagates downward to the lower rigid boundary. The propagation velocity for the 

second explosion ranges from about 100 m/s or less at the top to about 600 m/s at the 

bottom. As a typical case for a double explosion, propagation speeds for WFCI-C-06 test 

are plotted with a solid line in Figure 5.4. 

In particular, the WFCI-H-04 test shows no indication of single phase sonic velocity behav

ior at the bottom of the test section. This indicates that the leading edge of the molten fuel 

NUREG/CR-6623 164

V
had already reached the bottom and generated some amount of vapor. This was expected 

because the pouring time was increased by 400 ms and the bypass valve was opened in 

this test. Due to the external trigger, the explosion occurred at the bottom and escalated 

upward. The maximum escalation speed was reached relatively early, as compared with 

other experiments. Comparing this with the WFCI-H-03 test, in which only the pouring 

time was increased, the fluid near the bottom of the test section remains as pure water 

during the interaction. This clearly shows that fuel penetration was hindered by vapor 

upflow. 

Based on Park et al.'s analysis [87], the void fraction at the flooding region is given by 

following equations, __1 

',, Co (5.1) 

where, 

a= 1.7 [ Ljp.1Dh D] (p.,- p.).(.2 (5.2) 

In this model, the coolant hydraulic diameter, Dh is given by 

4 (A,,, - -f.. ) 

Dh = pHmi) (5.3) 

pu + 1 4,rnf~ 

where P,. is the chamber perimeter, Hmt is the depth of the pool and mf is the fuel mass 

mixed in the pool. For the WFCI-H-02 test conditions, Dh is about 60 mm. If the drift 

flux coefficient is 1.4, a void fraction at the flooding conditions in the present facility would 

be about 66 percent. The average vapor velocity, U,7, at the flooding condition can be 

calculated using Wallis' flooding criterion [88] given by 

j*=j P. (5.4) 

" = iv -=,,U, = Dh (pc-p,) 

This equation gives an average flooding vapor velocity of greater than 20 m/s. For our 

conditions this limit would not be reached but the mixing process was still hindered by the 

vapor upflow. 

5.3 Qualitative Debris Analysis 

The post-test debris collected after an FCI is another observable used to characterize vapor 

explosion energetics. Whether the explosion occurred or not, debris size distribution is 

also of importance in other aspects of FCI analysis. The debris distribution is generated

165 NUREG/CR-6623
by fragmentation processes during two different time scales in the fuel-coolant interaction, 

the mixing phase, and the explosion phase. In the mixing phase, the characteristic size of 

the debris is governed by hydrodynamics related to phase relative velocities. The WFCI

G series of tests showed that the shape of the debris resembled a sphere with a Sauter 

mean diameter on the order of a few millimeters. However, in the explosion phase, fine 

fuel fragments may also be produced by other fragmentation processes in sub-millisecond 

times. The characteristic size of the debris in the explosion phase is much finer than. in 

the mixing phase such as sub-millimeter sizes. The shape of the debris may also be very 

arbitrary. 

The debris from the explosions are collected within the test tube, the expansion tube and 

the quench tank. Less than a fifth of the initial fuel mass charge for our tests remains as 

a crust within the furnace, transfer crucible and the upper portion of the test tube above 

the slide gate. In many tests spontaneous explosions occurred in the funnel above the 

slide gate, but this FCI did not directly affect the interaction below as long as the gate is 

closed. Thus, this debris was not collected and analyzed. The explosion debris had general 

characteristics similar to all past molten tin tests [63, 62]: i.e., porous tin "plugs". They 

were found at various locations within the explosion tube with fine fragments intermingled 

and fused with the coarser porous plugs. In addition, discrete particles of fine tin fragments 

were swept down the expansion tube with the steam/water mixture, eventually settling in 

the quench tank. Different shapes of debris were found in different locations. In most tests, 

one piece of a long, round, porous plug was found in the I-Tube. The I-Tube debris was 

fused together and seemed to have resolidified after the explosion. The tin debris which 

was presumably near the leading edge of the fuel jet did not exhibit as many fine particles 

within its porous plug or as separate debris. This may be interpreted as being caused by 

some partial solidification as it cooled and penetrated the water pool. One can hypothesize 

that although this was probably the sight for the FCI triggering, a smaller portion of the tin 

melt at this location underwent rapid fragmentation during the explosion. In the T-tube, 

however, the shape of the debris resembled densely packed "sand." This may be more 

indicative of finely fragmented debris which participated in the explosion. In all tests the 

debris were removed from the tube and separately collected, weighed and measured. 

These "plugs" seem to be remnants of the initial fuel jet which entered the water, spread 

radially and mixed with the water, and then broke into these plugs as the explosion propa

gated through the mixture and expanded, thereby, fragmenting portions of the fuel mixture. 

The fuel locations of these plugs seem to have no relationship to where they were formed 

since the multiphase explosion expansion would transport them to various locations where 

they are eventually quenched and become lodged in the test section. However, it is again 

important to note here that the plugs in the T-tube are just below the slide gate and would 

act a point of shock reflection for the upward propagating shock wave from the escalat

ing explosion. Such plugs were also noted in the KROTOS tin tests and seem to have

NUREG/CR-6623 166
contributed to the observed behavior of shock reflection in KROTOS.

The quantitative debris analysis of all the loose debris (no "plugs") was performed with 

mechanical sieves down to 25 pm. All of the debris distributions for this work are listed 

in the Appendix. As one can see, only a small fraction of the debris is at sizes which could 

be interpreted to be small enough to rapidly quench in an explosion timescale of a couple 

of milliseconds; i.e., less than 5 % below 150 pm. Even accounting for the fact that some 

of these particles were trapped within the porous "plugs" leaves one with the qualitative 

conclusion that only a small fraction of the fuel rapidly fragmented in the propagation 

process which initially drives the explosion. This is consistent with the analysis of TEXAS 

for the tin test KROTOS-21 [117], which indicated that a fuel mass less than 0.1 kg was 

needed to drive the explosion. This is also consistent with the concept that the explosion is 

actually more efficient when one considers the smaller masses that may be actually involved. 

5.4 Thermodynamic Analysis 

In thermodynamic analyses, it has been historically assumed that all of the fuel and coolant 

are ideally mixed together and participate in the FCI. This is one of the key reasons for 

the large pressure and work output estimated by such models. However, it is possible to 

re-examine this assumption by using these experimental results. In this section, the exper

imental results obtained in the present work are analyzed to predict the minimum amount 

of fuel participation during the vapor explosion. This is done by using a thermal detonation 

model which employs the explosion pressures and propagation velocities obtained from the 

experiments to predict the mixture conditions which produced them. 

The thermal detonation model proposed by Board and Hall [11] has the advantage that the 

explosion propagation behavior is obtained without knowing any detailed rate processes. 

In their model, a shock wave is generated by the vapor explosion in an one-dimensional 

geometry and propagate as a "quasi-steady" state through uniformly mixed fuel and coolant 

as designated '1'. At the downstream point, '2', the interaction between the fuel and coolant 

occurs and the fuel and coolant reach thermal and mechanical equilibrium. This satisfies 

the one-dimensional steady-state conservation balance of mass, momentum and energy. The mass, momentum, and energy balance across the shock front is shown with subscript 

1 denoting upstream conditions,and 2 denoting quasi-steady interaction conditions; 

mass balance: 

pIul = p2u2 (5.5) 

momentum balance: 

pI + pJU2 P2 + P2U (5.6)

167 NUREG/CR-6623
energy balance: 

S~(5.7) il + -U- = i2 + -•-(.7 

12 2 

For a homogeneous mixture, the properties are obtained by the mass averaging, as follows: 

v E xivi (5.8) 

i 

i = E iii (5.9) 

i 

where at state 1 the fuel, coolant, and vapor are at different temperatures and at state 2, 

the fragmented fuel, participating coolant and any remaining vapor are in thermodynamic 

equilibrium. Combining the above equations, the equilibrium Hugoniot condition for state 

2 is derived for given initial conditions at state 1. The pressure and velocity at state 2 are 

given by 

PI 2 +P 2 ( (V2 - V1) = el - e2 (5.10) 

U2 = - 1 P; (5.11) 

In this analysis, however, the initial mixing conditions were estimated by this model using 

experimentally measured explosion propagation behaviors such as the explosion pressure 

peak, P.,, and the propagation velocity, u,. The minimum mass of the fuel involved in 

the explosion is obtained by using the calculated initial conditions and measured explosion 

work output. The reader should note that this is an estimate based on the assumption of 

quasi-steady state behavior and that the C-J chemical dentonation analogy is applicable 

to the vapor explosion process. 

Figure 5.5 and 5.6 are the C-J pressures and propagation speed calculated from the above 

equations with respect to the ranges of initial conditions, i.e., the coolant to fuel mass 

ratios and void fraction which are representative of the WFCI experiments. For the current 

experimental conditions, the measured C-J pressure and propagation speed are equal to 

approximately 3 MPa and 200 to 300 m/s, respectively as shown in Table 4.1. From these 

data and two figures, the corresponding initial void fraction and the coolant to fuel mass 

ratio can be estimated to be about 20 % and 0.33, respectively. The measured work is 

in the range of 2 to 5 kJ and the calculated isentropic maximum work ranges from 40 to 

60 kJ/kg. Thus, the minimum mass of fuel involved is calculated by the measured work 

divided by the calculated specific thermodynamic work. Corresponding fuel masses of 40 

to 120 g are obtained from this estimate. These masses are about 1.0 to 3.6 % of the total 

fuel mass injected of 3 to 4 kg. 

NUREG/CR-6623 168

1!
The large difference in the explosion energy conversion between actual measurements and 

ideal conditions might be explained by considering significant thermal dissipation of the 

fuel thermal energy during the expansion process. Using this hypothesis, one estimates 

that a large amount of vapor (more than 50 %) would be produced, while more fuel could 

participate in the interaction under such circumstances. Considering a propagation velocity 

of 200 to 300 m/s as observed in the experiments with a relatively high void fraction, say 60 

%, one can estimate a mass ratio of 0.1 from Figure 5.6. However, under these conditions, 

the possible C-J pressure is near the critical pressure of the water coolant. Larger void 

fractions can be assumed to reduce this explosion pressure, but such volume fractions are 

unphysically high. This again suggests that a small fraction of fuel participating in the 

FCI seems to be more plausible than a large fraction of fuel participation with large void 

fraction. Most post-test debris analyses in the WFCI tests indicates that a large fraction 

of molten fuel in the I-tube did not take part in the interaction. 

The fuel debris distribution for the WFCI-A and C test series have debris sizes of 60 to 

660 pm. The associated mass to this size range corresponds to similar mass fractions of 

few percent. This result is also consistent with KROTOS-21 analysis [117] which indicated 

a fuel mass on the order of 100 g was needed to drive the explosion. It suggests that 

the explosion is actually more efficient with small fuel masses driving the interaction, and 

the remainder of the fuel and coolant may not participate over the explosion propagation 

time-scale. The same analysis has been also applied to recent KROTOS tests [135] for high 

temperature molten oxides and similar conclusions were reached. 

5.5 Energetics of Iron-Oxide Melt 

Most WFCI experiments (series A to H) used a molten fuel simulant, tin, which was known 

to easily produce energetic explosions. Our major contribution using this simulant was to 

demonstrate explosion reproducibility and provide an extensive data base on the effects of 

initial and boundary conditions on explosion energetics. Based on analysis we expect the 

insights gained are applicable to other molten fuel-coolant fluid pairs. In contrast to these 

past tests, the WFCI-K test series was focused on using a molten fuel simulant that was 

more prototypic of the high temperature molten oxide that would be present in a severe 

accident. Our choice of iron-oxide was based on proven energetics at small scales, the ability 

to melt and deliver kilogram quantities and the conclusion that it was a good simulant for 

reactor materials. This approach allows us to gain some understanding of FCIs with molten 

oxides at prototypic superheats as well as generate a more extensive data base on materials 

scaling. Note that no previous large scale experiments have been attempted with this fuel 

simulant.

169 NUREG/CR-6623
The WFCI-K test series results were ambiguous as to whether an energetic explosion oc

curred. There are three pieces of data that give us information about whether an explosion 

occurs during any experiment. The first is the pressure data that is taken during the test. 

This provides detailed information about what occurred inside the test section during the 

interaction, and is examined in several ways to see if the interaction has explosive qualities. 

For example, sharp pressure peaks that escalate and propagate through the test section 

suggest an explosive interaction. The second indication is the breakage of the rupture di

aphragm at the end of the expansion tube, due to the initial movement of the water slug as 

work is being done on the surroundings. The final evidence of an explosion is the acceler

ation of the slug down the expansion tube. Given a measureable slug velocity, this allows 

measurement of the slug kinetic energy and calculation of the conversion ratio. All three 

of these data indicators were always present for vapor explosions when tin was poured into 

water. For this series of experiments with iron-oxide, no single test had all three of the 

above characteristics, but several tests had at least two. This suggests that fuel-coolant 

interactions did occur in K-01, -05 and -09, but none were vapor explosions. Let us consider 

each source of data. 

Pressure Histories and Impulses 

The pressure histories give the most insight into the characteristics of each interaction, 

because these data are quantitative and recorded at a high sampling rate to capture explo

sive propagations. Seven of the ten series K experiments performed exhibited very similar 

pressure traces, with rather weak signals from the trigger occurring at transducer P1, and 

very weak or no signal for all others. It is most likely that the pressure signals are absent 

at higher levels in the test section due to a large increase in void from boiling during fuel

coolant mixing, which limits the pressure increase due to the compressibility of the vapor. 

Thus, in the absence of an FCI, the trigger pressure impulse is quickly dissipated in the 

multiphase medium. This effect is most prominent in tests with the WFCI-1 test section, 

which has a smaller inside diameter. There is less water per unit length of test section, 

and thus a greater likelihood of vapor production producing a larger local void fraction. In 

fact, since the iron-oxide (assuming a composition of FesO4) begins to solidify at 1556 0C, 

substantial boiling will occur even after the fuel has solidified. This effect seemed to be 

independent of water temperature with modest subcooling. Tests K-02, -03, -04, and -10 

all had uniform water temperature (using loop heating) ranging from 70 to 85 °C. Tests 

K-05, -06, and -07 all had nonuniform temperature gradients set up with hot water in the 

funnel ( 80 °C) and cooler water at the bottom of the test section (25 - 47 0C). This was 

done in an attempt to reduce the void fraction inside the test section, thus allowing an 

interaction to occur, and to suppress an FCI in the funnel. All these cases exhibited a 

similar type of characteristic pressure traces. The magnitude of the peak pressures give 

NUREG/CR-6623 170

1!
some indication of which tests had the strongest fuel-coolant interactions. Tests K-01 and 

K-09 had peak pressures of 6.53 and 9.70 MPa respectively. This suggests to us that the 

mass of melt required to generate an explosive pressure trace need not be very large, since 

test K-01 had about 15 % of the melt mass delivered to the test section. 

The shock propagation speeds give some insight into whether an interaction is explosive. In most of the tests, shock propagations were only measured between P1 and P2, and 

the values were always much less than the sound speed in water, indicating that there is 

voiding in this region, but not necessarily an explosive interaction. Test K-01 also has a 

moderate propagation speed at the bottom (510 m/s) which quickly drops as the peaks 

propagate upwards to about 100 m/s between transducers P6 and P7. Subsequently, the 

:speed increases, as the time for the first P7 peak to travel up to the slide gate and back 

down indicates a speed of 169 m/s. Then the speed increases further to 184 m/s from P7 

to P6, which indicates that some sort of explosive interaction had occurred locally and was 

propagating downward. Test K-09 also exhibits similar characteristics, with a large speed 

a the bottom (510 m/s) which drops as it propagates upward (30 m/s from P6 to P7) then 

increases again at the top (105 m/s from P7 to P8). The difficulty in resolving any clear 

shock propagation behavior and the fact that the speeds were relatively low again suggest that local FCIs occurred, but give no clear indication of a propagating explosion. 

The impulses tended to give results similar to the peak pressure data. Test K-10 has 

a moderately large impulse of 0.069 N-s, and this occurs at the bottom (P1). The two 

tests with the largest impulses are K-01 and K-09, with values of 0.073 N-s and 0.16 N-s 

respectively. These maximum impulses occur near the top of the test section (P7) in both 

cases. In addition, tests K-01 and K-09 exhibit some evidence of an increase in impulse 

in the upper portion of the test section. In test K-01, the impulse decreases slightly from 

around 0.03 N-s as the pressure peak propagates upward, to around 0.02 N-s when it reaches 

the top. Then as the peak travels back down, an impulse of 0.03 N-s is measured. There 

must have been additional energy release due to a fuel-coolant interaction, because if the 

peak simply reflected off the top surface (the slide gate) its impulse would be the same or 

slightly smaller. In the data for test K-09, there is also clear indication of an increase in 

impulse at the top. The impulses are very low at transducers PI to P3, then they increase 

to around 0.035 N-s at transducers P4 and P5. Then the impulse jumps to 0.13 N-s at 

P7 and 0.1 N-s at P8. Traveling downward, the impulse is still 0.1 N-s at transducer P6. 

These impulses clearly indicate that an interaction occurred in this region with a rather 

large energy release compared to all other experiments. 

Work Output and Efficiency 

Experiment K-05 was the only test that provided a measurement of the conversion ratio,

171 NUREG/CR-6623
0.056 percent. The slug was forced down the tube at a significantly rapid speed to have its 

position recorded by the instruments. Based on the available internal energy of the fuel, 

the conversion ratio is quite low, especially when compared to values attained for other 

materials such as tin or alumina (0.3 to 3.0 % for these simulants respectively). The data 

for the slug movement is typical of other slug movements with tin. The low conversion ratio 

can be interpreted not only because of a high internal energy of the fuel, but also because 

the fuel is at a low superheat (100 to 200 0C) when compared to other fuels (e.g. tin 

and alumina with superheats greater than 300 0C). This means that the fuel can solidify 

quickly, which can prevent any further fuel fragmentation, after longer mixing or upon 

pressure reflections in the chamber. 

Another issue that needs to be considered is the effect of the contamination of the melt 

by the erosion of the boron nitride crucible. The Appendix discusses in some length the 

results of our debris analysis and the associated inspection of the boron nitride crucible, the 

delivery tube, and the test chamber region above the slide gate. Post-test debris analysis did 

not to provide much quantitative data into FCI energetics directly, but provides qualitative 

insights and detailed morphology of the melt composition. The most important conclusion 

from these analyses seems to point to the fact that as much as 10 percent of the non-oxidic 

melt by weight delivered to the test section is boron. This suggests that the melt upon 

delivery is iron-oxide with small amounts of boron-oxide intermixed. The net effect of 

this melt contamination by crucible material is that the liquidus temperature of the fuel 

is decreased. The exact amount of decrease is difficult to know, but an estimate can be 

given. Based on past work with iron-oxide, one can estimate the decrease in the melting 

temperature by the product of the percentage of the minority composition and the melt 

temperature difference. For iron-oxide and boron-oxide the maximum melt temperature 

difference is 1100 0C, and the minimum is about 950 *C, from liquidus to solidus for iron

oxide and assuming the boron-oxide has a discrete melting point (note: it has a 'softening 

point' of 450 0C). Thus, the depression in liquidus point of 1556 0C, would be at most 

about 100 °C, increasing the initial superheat to about 200 *C. This is not a substantial 

effect given the initial melt temperature and the rapid cooling that seems to take place 

during the mixing phase. In addition, the thermophysical properties of the melt would not 

be seriously affected by this minority melt component. It is still our view that the melt 

superheat is small enough to be the major reason for the limited energetics observed during 

the WFCI experiments. 

5.6 Possible Explosion Mechanism 

The results of the WFCI experiments seem consistent with the rapid fragmentation process 

resulting from the vapor film collapse process. Kim [14] proposed that the model for 

the fuel fragmentation is due to a jet impingement mechanism, in which the fuel outer 

NUREG/CR-6623 172

11
surface mixes with or entraps coolant below its surface. This local microscale mixing 

causes further pressurization and drives further fuel droplet fragmentation in the near 

vicinity. This process can be cyclic, with the fuel surface progressively fragmented by this 

collapse-jetting process. 'However, in a large scale explosion, such as that observed in these 

experiments, the collapse process would be operative for some time span after local film 

collapse. Following this time span local pressure differences caused by coolant vaporization and fuel fragmentation would equilibrate and only hydrodynamic fragmentation would take 

place. 

Recent single droplet (or small scale) data by Ciccarelli et al. [56] seem to provide some 

visual confirmation of this process. They suggest that upon film collapse the molten fuel 

droplet expands into a "starfish" shape. This may be due to local coolant jets impacting 

and perhaps touching or penetrating the fuel surface at local spots. Because the liquids 

are essentially incompressible, this coolant jet impingement also expels fuel outward as "projections". In a large scale explosion these expelled fuel particles would be fragmented 

in the coolant and quickly quenched within it. This would generate vapor at the local 

pressures, sustaining and escalating the explosion event. 

This conceptual picture is similar to Kim's in that each one views the explosion fuel frag mentation process as a surface phenomenon in which there is a time span over which it is 

operative. The explosion propagation leaves the fuel-coolant mixture partially "reacted or 

exploded." The local void of the fuel-coolant mixture determines the initial extent of film 

collapse and associated fuel-coolant contact. Fuel thermal energy and fuel droplet surface 

area determine the associated local pressurization as the surface of the fuel disintegrates due 

to this "microscale" fuel-coolant mixing, coolant vaporization and fuel quencing process. The local pressures and the induced particle velocities then transmit this event spatially through the fuel-coolant mixture and the escalation continues based on the volume frac

tion of the fuel-coolant mixture. Such an explanation would be consistent with the limited 

fine fragmentation energetics data taken from the experiments. Also large scale explosion 

mixture might be able to sustain multiple propagations through the mixture, provided it is 

held together by a sufficient inertial constraint and the associated vapor void behind each 

propagation is not too large to retard further shock propagation, vapor film collapse and 

heat transfer with the molten fuel particles remaining in the mixture. Thus, the extent of 

the fuel participation in an explosion is really controlled by the initial fuel-coolant mixing conditions, the inertial constraint and the fuel thermal energy. Scaling to reactor conditions 

implies one must primarily consider these conditions.

173 NUREG/CR-6623
5.7 Vapor Explosion Scaling for Reactor Safety Issues

Most experimental investigations, including this work, have been performed with various 

simulants used for the actual molten reactor fuel material; e.g., using less toxic or more 

readily available lower temperature metals or inert oxides in smaller geometries. In order to 

relate observed experimental behavior to postulated reactor scale behavior, it is necessary 

to establish appropriate scaling for the vapor explosion; i.e., geometric as well as material 

scaling. In the first section of this work, we established the reactor safety issues for pos

tulated in-vessel and ex-vessel conditions; i.e., how to predict energetics, work output and 

dynamic pressures, under a variety of fuel-coolant contact conditions. From a theoretical 

point of view, direct use of governing equations to scale the phenomenon is not possible 

because the details of vapor explosion dynamics, involving mixing, fuel fragmentation and 

explosion propagation, still have key uncertainties. Rather, it would be useful to consider 

which initial and boundary conditions are of greatest importance to energetics and what 

would be a proper scaling approach. A complete discussion of this was presented at the 

1997 CSNI workshop by one of the current authors (Appendix). Let us try to apply these 

principles to what has been observed in these experiments. 

In this work, the following four different initial and boundary conditions were examined 

with consideration for the geometric, materials and thermodynamics aspects of the FCI 

phenomenon. 

* The effect of the trigger; 

@ The system constraint specifically the degree of axial constraint; 

* The effect of fuel temperature with substantial superheat and water subcooling; 

* The effect of changing fuel composition to a molten oxide with limited superheat; 

* The effect of the coolant to fuel mass ratio or volume ratio. 

The effect of the trigger would be highly variable for fuel-coolant mixing and explosion 

propagation in a real situation. Our approach for the trigger effect in these experiments 

was to understand its effect for our test conditions, and then eliminate it from any energetics 

considerations. The first four WFCI experimental series characterized the trigger, showed 

reproducibility of energetics for spontaneous and externally triggered events and then found 

the external trigger range that could be used in subsequent tests for consistency of test 

results without further concern that the trigger affected the data. The WFCI test program 

was successful in these efforts. If FCI triggering was the main issue, it would be necessary 

to identify the cause of the trigger in a real situation, while investigating the trigger effect 

in a systematic manner for the actual materials. This was not our main objective. 

NUREG/CR-6623 174

1!
The system constraints for the fuel-coolant mixture can be divided into two different con

straints; radial and axial constraints. For the radial constraint, our previous discussions for 

WFCI-D test indicated that in small scale tests [77] the rigid radial boundary provided re

producible results for explosion propagation and escalation more easily. In contrast, under 

weak radial boundary conditions local interactions occurred incoherently without propaga

tion. To achieve controllable experiments, one needs to use a rigid boundary for a given fuel-coolant mixture. This rigid boundary would allow experiments, large in scale but small 

in relation to the reactor scale, to behave in a manner similar to a "one-dimensional slice" 

of a larger fuel-coolant mixture. Because the reactor scale environment would exhibit some 

radial compressibility, a rigid radial boundary would tend to maximize energetics. For the 

axial constraint, as discussed in a previous section, the degree of constraint affected the en

ergetics of the vapor explosion. As the degree of the acoustic constraint was increased, the 

energetics of the vapor explosion also increased. A further increase of the inertial constraint 

can maximize the explosion energetics due to a competition between vapor production and 

fuel quenching during the propagation-expansion. Therefore, it is necessary to scale the 

experiments with an axial constraint that mimics expected reactor scale conditions coupled 

with rigid radial boundaries. 

The results of the WFCI-E and WFCI-F test series suggest a scaling process that is more 

qualitative than quantitative. Results indicate that once fuel and coolant temperatures 

are within an envelope of fuel-coolant mixture conditions where the fuel remains molten 

(sufficient superheat) and the vapor film boiling process is stable (proper subcooling), a 

vapor explosion can result and constituent temperatures have a second order effect on 

energetics. The implication from this experimental observation needs to be taken with the 

result that the molten iron-oxide tests were at modest superheats and did not result in 

vapor explosions. This leads one to conclude that experimental values of fuel superheat 

and coolant subcooling must also mimic reactor scale conditions to be scaled appropriately. 

These conclusions require us to focus on material scaling and the issue of fuel composition. 

It should be recognized that it is almost simulate the thermal behavior of one fuel material 

composition with another. Thus, any simulant used for the actual fuel material introduces 

distortions in the materials scaling of the phenomenon, that must be understood to the 

extent that the mixing or energetics is affected. In fuel property table in chapter 2 we 

assumed that most fuels had a modest superheat above its melting point, as has been 

predicted from severe accident simulations. This is in accord with the scaling argument 

just made. The major exception are the low-melting-point liquid metals, which has been 

historically used at higher temperatures to compensate for their low energy content near 

their melting point. Inspection of the table indicates quite clearly that iron-oxide and the 

corium melts are quite analogous in their specific energy, energy density and latent heat. 

Also, alumina, a chemically inert fuel simulant, contains more specific energy per unit mass 

than any other fuel melts, while tin contains much less specific energy per unit mass. This

175 NUREG/CR-6623
suggests that iron-oxide may be a good thermodynamic match for corium, while tin and 

alumina would bound its energetics. 

The final condition to consider is the coolant to fuel ratio in the fuel-coolant mixture. The 

WFCI-H experimental series provides two important observations about this ratio. First, 

there appears to be a broad maximum in explosion energetics for a range of coolant to fuel 

mixture conditions. Thus, from a scaling perspective once other experimental conditions are 

specified a range of mixtures can be considered to produce similar energetics as long as 'fuel 

lean' or 'coolant lean' mixtures are avoided. The second observation is that the explosion 

conversion ratio is quite low for all the tests, far below calculated thermodynamic values. 

The reason for this observation is the limited amount of fuel participating, and it may be 

necessary to verify this observation at different geometric scales and material compositions. 

Note that alumina tests in the KROTOS facility corroborate this observation.

NUREG/CR-6623 176
I

70- T n 

'60- for Single Drop Tests 

•-40' 

_ -Single Drop 

8 30 1 TE (WFCl) -- • SE (WFCI) • 

"20 * WE (Hall) Ax TE (K-21 ) 

10 V TE (Baines) 

0 . v SE (Baines) 

200 300 400 500 600 700 800 "960 2000 1100 

Initial Fuel (Tin) Temperature (C) 

Figure 5.1: Temperature Interaction Zone in the Large Scale Tests

177 NUREG/CR-6623
1.0- a I 1 4 a 

"M Triggered Explosions (TE) 

0.9- 0 Early Spontaneous Explosions (ESE) 

A Delayed Spontaneous Explosions (DSE) 

0.8

0 .7- 0 

-0.6

.2 0.5

00 0.4- 0I 

0.3- An m 

0.2- U 

0.1 .. . .. 

0.7 0.8 0.9 1.0 1.1 1.2 1.3 

Fuel Thermal Energy (MJ) 

Figure 5.2: Conversion Ratios with respect to the Fuel Thermal Energies 

NUREG/CR-6623 178

lV
10

0

1

WFCI (Sn)

L r -3 WFCI KROTOS (Sn) (Sn,AI20 3 Tests) 

0 

3

0 .1 . 

0.1 1 10 

Fuel Thermal Energy per Unit Mass (MIJ/kg)

Figure 5.3: Conversion Ratios with respect to the Fuel Specific Thermal Energies

NUREG/CR-6623 I

179
0

a

SWTCI-A-02 FW. BW 

*WFCI-A-05 FW, DW 

A WFCI-B-O2 FW. SW 

A WFCl-C-03 FW 

v WFCI-C-04 FW 

-1-- WFcI-C-06 FW. BW

vVVFCI-E-O2 FW, BW ~0 WVFCI44.O3FW, 1W 

••.,•o 0-- WVM-H-04 F 

- .

0:2 0:4 0:6 00:8 - 1:0 1:2 

Location from the Bottom of the Test Section (m)

Figure 5.4: Propagation Speeds with respect to Locations of the Test Section

NUREG/CR-6623

I OLu I

i

1400

1200

1000

800 

600

400

200

0

-200

-400-

-AciNJ -4ý

0.0 1.4

l

180
" 10 10 ;;ý i,. .

1 I.CM I 

S..... .. ..................... ....... 

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 

Void FRacon 

Figure 5.5: Variation of the Chapman-Jouguet Pressures associated with the Fuel-Coolant 

Mass Ratio in the Typical WFCI Experimental Conditions

181 NUREG/CR-6623
,Mc/=5. 

I0 D 

10 . 1 P t I I 

0 0.1 02 0.3 0.4 0.U 0.6 0.7 0.9 1 

Void Fracion 

Figure 5.6: The Variation of the Chapman-Jouguet Velocities associated with the Fuel

Coolant Mass Ratio in the Typical WFCI Experimental Conditions

NUREG/CR-6623 182

P.
Chapter 6

Conclusions and Recommendations 

Vapor explosions are a potential hazard in light water reactors after a prolonged lack of 

cooling allows reactor core materials to melt and contact residual water coolant within the 

reactor vessel or below in the containment reactor cavity. Past studies have demonstrated 

the explosive nature of certain liquid pairs, but have not systematically examined explosion 

energetics at larger scales as a function of controlled initial and boundary conditions. This 

has hampered basic understanding and has been a major deficiency in the database for 

modelling the vapor explosion. 

Our objectives for this work were to obtain well-characterized data for the explosion prop

agation/escalation phases, while systematically investigating the effect of a comprehensive 

set of initial and boundary conditions on the explosion energetics: 

e trigger strength, 

e system constraint, 

e fuel mass, composition and temperature, and 

* coolant mass, viscosity and temperature. 

This objective was subdivided into three specific tasks. First, a vapor explosion apparatus, 

WFCI, was designed and fabricated for well-characterized explosion data and demonstrated 

reproducible explosions, with tin as the simulant fuel. Second, the explosion energetics was 

examined as a function of varying initial and boundary conditions for this simulant fuel. 

Finally, the simulant fuel was changed to iron-oxide, a fuel that was more prototypic of 

actual fuel compositions and explosion energetics were reexamined. 

The experimental investigation was subdivided into smaller test series to better under

stand specific vapor explosion mixing and propagation behavior. First, the WFCI-A test 

series was performed to demonstrate the reproducibility of the explosion phenomena. Ini

tial conditions similar to KROTOS-21 were chosen as the nominal case for these tests, 

for comparison to independent data. The explosion behavior in the WFCI facility showed 

good reproducibility and agreement with KROTOS-21. Dynamic pressures had peak val

ues as high as 10 MPa and quasi-steady values of 2-3 MPa, while explosion propagation

183 NUREG/CR-6623
speeds were in the range of 200 m/s. The explosion conversion ratios were about 0.2 to 

0.5 % for both spontaneous and for externally triggered explosions. The next test series, 

WFCI-B and C investigated the effect of the external trigger on the vapor explosion; i.e., 

spontaneous explosions compared to triggered explosions with specific impulse strengths. 

The tests showed that as the external trigger was reduced and then eliminated, multiple 

propagation events occurred during the explosion i.e., double propagations first upward 

and then downward. The energetics of the complete interaction was relatively indepen

dent of trigger strength, but the detailed behavior of propagation process became more 

complex with spontaneous triggers. Once the trigger strength exceeded 3 MPa no multiple 

propagation events were observed and single propagation events of similar energetics were 

observed. Energetics for all of these vapor explosions were less than one percent. 

The WFCI-D series investigated the effect of axial constraint. The WFCI facility was 

originally designed with a rigid radial constraint to maximize the energetics for any given 

set of mixing conditions, but the axial constraint could be varied. The axial constraint 

was varied by changing the slug mass in the horizontal expansion tube by an order of 

magnitude. Results indicated that there was an optimal degree of axial constraint to 

maximize explosion energetics. This could be explained by the competing effects of rapid 

vapor production during the propagation and vapor condensation and fuel quenching in 

surrounding coolant liquid as the explosion mixture expands. This effect suggests that the 

axial constraint needs to be similar in FCI tests when considering energetics for reactor 

safety issues. 

The sixth and seventh test series, WFCI-E and F investigated the effect of the fuel and 

coolant temperatures on energetics. The results indicated that once the fuel temperature 

was above a threshold value its effect on energetics was of second-order importance. The 

same result was noted for a variation in the coolant temperature. This suggests that if the 

fuel and coolant temperature are large enough to satisfy the qualitative requirements of a 

molten fuel and stable film boiling at the time of triggering, then an energetic explosion 

can result, with temperature having a small quantitative effect within this envelope of 

conditions. This should be scale independent and KROTOS tests also suggest this based 

on the scale for KROTOS-21 and for compositions with alumina fuels. 

The WFCI-G test series was performed to investigate the suppression effect of polymer 

additives by an increase in the coolant viscosity. Polymer additives suppressed spontaneous 

vapor explosions in this large scale geometry in qualitative agreement with past small scale 

tests. This was the first time that explosion suppression was demonstrated at a larger 

scale. Also, the post-test fuel debris generated in the absence of the explosion may be quite 

representative of the fuel debris during the mixing and fuel quenching process. Use of the 

polymeric solution should be considered as a technique for mixing studies. 

In the WFCI-H series the fuel jet diameter and the timing of the external trigger were 

NUREG/CR-6623 184

1!
altered to vary the ratio of coolant mass to fuel mass in the mixture at the time of the 

explosion. Varying this mass ratio of coolant to fuel indicated that the explosion conversion 

ratio exhibited a broad maximum in energy conversion. These values were more than an 

order of magnitude lower value than one predicts from ideal thermodynamic situations. 

Analysis again indicated that the broad maximum is created by competing effects related 

to the development of the fuel-coolant mixture prior to triggering. The location of this 

maximum relative to mass ratio is secondary to its qualitative existence. However, the 

value of this coolant to fuel ratio can be understood relative to mixing kinetics. The small 

energy conversion ratio can only be explained by the observation that only a few percent of the fuel mass fragments into small enough debris ( << 1mm) to directly participate in 

the explosion. This conclusion is applicable for all of our test data and seems to be scale 

independent and has important implications for reactor safety issues. 

Finally, in the WFCI-K test series, the effect of changing the fuel composition from molten 

tin to more prototypic simulant (iron-oxide) was observed. These tests indicated that the 

triggering of energetic FCIs with more prototypic fuel materials and superheats was quite difficult. Fuel-coolant interactions were empirically observed in four of the twelve tests, but no propagating vapor explosions were observed. This observation is consistent with 

those of JRC staff and their corium tests in KROTOS at atmospheric pressures in which 

weak fuel-coolant interactions were observed. The reason for these weak FCIs seems to be 

that the melt superheat was low and mixing and quenching was efficient. This serves as 

validation of the temperature effects previously discussed. 

With respect to reactor safety issues, this experimental work using fuel simulants has yielded 

a number of results that have potentially quite important safety implications. First, this 

work has provided clear evidence of the reproducibility of vapor explosion energetics for a 

controlled set of initial and boundary conditions. This suggests empirically that this phe

nomenon is predictable if one can establish and control the initial and boundary conditions. 

Second, the experiments demonstrate that geometric scaling can be properly specified; e.g., 

a rigid radial constraint for one-dimensional tests is conservative for energetics when com pared to full-scale, while the axial constraint scale factor from test to prototype needs to 

be the unity to preserve energetics. 

Finally and most importantly, the data suggests that once the fuel-coolant initial conditions 

are within an envelope for triggered events, the energetics is much less than thermodynamic due to the small amount of fuel that participates in an explosion time scale. And this 

envelope of triggerability is much smaller for a simulant molten oxide with low superheat, such as molten iron-oxide. This observation is somewhat of an enigma at the present time 

since iron-oxide and corium both exhibit 'weak vapor explosions' where the energetics is far 

below what was observed in WFCI for tin and in KROTOS for alumina. It is hypothesized that the mixing process for these fuels with modest superheat is efficient enough to cause the 

local void fraction in the mixture to be large and the fuel to be near solidification; i.e., both

185 NUREG/CR-6623
conditions hampering explosion triggering and propagation. This suggests that material 

scaling for reactor safety issues must preserve the same fuel composition and superheat 

from the test to the prototype. 

The current work has limited data at larger scales with more prototypic molten oxides; 

i.e., larger fuel volumes than 0.5 liters, larger chamber geometries, prototypic molten oxide 

compositions and superheats. It is recommended that further tests could be carried out 

under these conditions to empirically verify our findings. Models developed from our anal

ysis can also be used to analyze these experiments. Finally, it is known that the mixing 

conditions determine the envelope of explosivity for the vapor explosion. Thus, it is of fun

damental interest to better measure the mixture local conditions just prior to the explosion 

to correlate with the explosion energetics; i.e., void fraction profiles, fuel volume fractions 

and mixing diameters. Our future work in vapor explosion research is specifically targeted 

toward this purpose.

NUREG/CR-6623 186

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Chapter 7

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University of Virginia, (1981). 

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poration, Washington, USA, (1992).

NUREG/CR-6623 198

]i
Appendix A 

Experimental Data for the Debris

199 NUREG/CR-6623
U 

a 

0 

a 

0 

I 

U 

a 

2

100 

10 

1 

0.1 

(nnil

<25 25 38 75 150 300 600 1000 2000 4000 >4000 

Debris Size (micro) 

Figure A.I: Cumulative Debris Distribution for the WFCI-A Series

inn

4 10 

T 

= 

0.1 

d .0

<25 25 38 75 150 300 600 1000 2000 4000 >4000

Debris Size (micro) 

Figure A.2: Cumulative Debris Distribution for the WFCI-B Series 

NUREG/CR-6623 200

i

WFCI-A-01 

WFa-A-02 

WFCI-A-03 -9--

WFCI-A-04 -N

WFCI-A-05

. , I I I p p p

..- fWFI-B-O1 '••" 

_.-'•'/ /WFCI-B-02 -+

WFC-B-03 -9...

i a i i I I I I I I I

, o

I | | Q | • •
S0 

.100 

F 

CA 

10 

0.I P 

U 

S0.01

<25 25 38 75 150 300 600 1000 2000 4000 >4000

Figure A.3: 

1o0

10 0!0 

V 

0 

0.0

Debris Size (micro) 

Cumulative Debris Distribution for the WFCI-C Series

<25 25 38 75 150 300 600 1000 2000 4000 >4000 

Debris Size (micro) 

Figure A.4: Cumulative Debris Distribution for the WFCI-D Series

NUREG/CR-6623

)

201
100

0 

°0

10 

I 

0.I 

0.01

Figure A.5: 

100 

10 

! 

0.1 

0.01

<25 25 38 75 150 300 600 1000 2000 4000 >4000

Debris Size (micro) 

Cumulative Debris Distribution for the WFCI-E Series

<25 25 38 75 150 300 600 1000 2000 4000 >4000

Debris Size (micro) 

Figure A.6: Cumulative Debris Distribution for the WFCI-H Series

NUREG/CR-6623 202
Appendix B

Melt Preparation and Delivery 

Introduction 

A key part of the WFCI series K experiments is the delivery of about a kilogran of molten 

iron oxide to a test section filled with water. This is accomplished with a specially designed 

furnace mounted directly above the test section. The firnace is constructed of several concentric 

cylindrical layers of insulation made of alumina (A120 3) and silicon dioxide (SiO%) to create a 6 

inch thick layer on all sides. The inner cavity is approximately 13 inches tall and 12 inches in 

diameter, and is heated to over 1700 EC with eight radiative heating elements. Placed in the 

center of this cavity is a cylindrical crucible made of boron nitride (BN). This material was 

chosen because it seemed to resist erosion by molten iron oxide better than other materials tested, 

was reasonably affordable (when compared to platinum or iridium), and was easily machinable. 

The melt is delivered through a hole in the bottom of the crucible that is plugged from above by 

a pneumatically actuated plunger (also made of BN). The crucible used in experiment K-01 was 

5 inches in diameter with 0.5 inch thick walls. A second crucible (used in experiment K-02) was 

larger with a 6 inch diameter. The walls in this second crucible were designed to extend the 

crucible lifetime by being much thicker in the regions that would be in contact with the molten 

fuel. This crucible was used for six experiments before it was replaced with a new one. The 

bottom portion of the crucible (where the greatest amount of erosion occurred) was cut away and 

replaced. 

203 NUREG/CR-6623
After tests K-01 and K-02, significant damage to the bottom ofthe firnace chamber was 

discovered. During test K-01, the smaller (5 inch) crucible cracked open, and much ofthe molten 

iron oxide leaked out onto the insulation, which caused serious damage to three upper layers.* In 

test K-02, the melt bubbled over the top of the crucible, and again did serious damage to the 

firnace. A detailed chemical analysis was performed on the residues with the hope of 

understanding what caused the melt to bubble or foam over in test K-02 so fi-nace damage will 

be limited in future tesis. Another reason was to give a quantitative measure of the amount of 

boron removed from the crucible after each melt, which could aid increasing the life of the 

cruc•ble and preventing a Mure (such as in test K-01) which could cause severe furnace damage. 

Samples were taken from three different locations in both tests. Since the greatest 

amount of residue was a very hard rocky substance of heterogenous composition located on the 

floor of the furnace, samples were taken there (designated KIF and K2F). The samples taken 

were removed from the top layer, and were expected to contain iron from the melt leakage, 

aluminum and silicon from the furnace insulation, and possibly some boron from the crucible. 

Samples of the small gray colored pieces that remained inside the crucible were taken as well 

(designated KIC and K2C). These would be expected to contain only iron and boron. The third 

set of samples came from the melt that was dropped into the test section (designated KIT and 

K2T). These samples also would be expected to contain only iron and boron. 

Several different types of chemical analyses were performed on these samples. A 

quantitative analysis of iron and boron composition using inductively coupled plasma - atomic 

emission spectroscopy (ICP-AES) was performed on all six samples at the UW Chemistry 

Department. In addition, an X-ray powder diffraction analysis was done at the UW Chemistry 

NUREG/CR-6623 204

i!
Department on sample K2F to look for particular iron and boron compounds. Finally, an electron 

microprobe analysis and scanning electron microscope imaging was performed on the upper 

surface of sample K2F at the UW Geology Department to determine the structure and to also 

give a quantitative compositional analysis of the different regions of the heterogenous furnace 

residue. 

Atomic Emission Spectroscopy Analysis 

Atomic emission spectroscopy is a method used to determine amounts of specific 

elements in a sample by resolving the intensity of the light emitted at certain wavelengths when 

an element is burned. A Miation grating in combination with a phototube detector is used for 

this purpose. The diffraction grating is rotated to scan a region ofwavdengths that corresponds 

to the values of known emission peaks for a particular element. The amplitude of the peaks 

correspond to the intensity of the emitted light, which is proportional to the amount of the 

element present in the sample. To determine the quantitative amount of the element in the 

sample, a standard solution with a known amount of the element must be analyzed at several 

different concentrations. This allows the generation of a relation between concentration and light 

intensity, which is used to determine the concentration of the element in the unknown samples. 

An ICP-AES analysis requires that the sample be in liquid form for injection into the 

plasma generated inside the instrument This proved to be difficult, since the samples were 

composed mostly of ceramic materials that do not dissolve easily. The method chosen for the 

sample preparation was a hydrofluoric acid (HF) fision. Fr•st the samples were powdered for 

easier dissolution, then they were placed in a platinun crucible with hydrafluoric acid and heated 

until no visible residue remained. The liquid was then diluted with water. Sodium tetraborate

205 NUREG/CR-6623
(Na2B40-XlOH2O) was used as the standard solution for the boron analysis since boron nitride 

would not dissolve in a HF fusion. Pure hematite (Fe2O3) was used as the standard in the iron 

analysis. Boron only had a single resolvable peak while iron had five, therefoe the iron data was 

in the form of a range of values, since each peak gives a slightly different result 

X-Ray Powder Diffraction Analysis 

This method of analysis simply involves scattering x-rays off the sample and measuring 

the corresponding scattering angle and peak intensity at that angle This resolves molecules 

instead of atoms, and is not quantitative. It can be used as a fingerprint to indicate the presence 

of a particular compound in an unknown sample. A large database of difiraction data is available 

for almost every compound known, but cross referencing is not available, so knowledge of what 

compounds might be in the sample is necessary. In this case, some of the compounds found to 

be present in a powdered sample of K2F were Fe2O3, Fe3O4, A120 3, SiOz B20 3, and AIFeO3. 

Electron Microprobe Analysis 

A scanning electron microscope can produce very detailed images by scanning a'surface 

and bombarding it with electrons. If the backscattered electrons are evaluated, a detailed image 

resuts. If the corresponding x-rays are analyzed, some indication of the composition of the area 

given which can also used to measure the quantitative elemental composition of a sample at a 

particular point. The regions scanned are on the order of microns in size so to get a 

representation of the entire sample, a combination of both of the above methods must be used. 

The K2F sample to be analyzed was first mounted in a resin and polished smooth. Then an 

image close to the upper surface of the sample was taken. Four distinct compositions in this area 

were observed. Then the instrument was set to search for iron, silicon, and boron over the same 

NUREG/CR-6623 206

11
area. Boron is an element that is not able to be evaluated quantitatively, so all other elements 

known to be present were measured, and whatever remained was assumed to be boron. Further 

quantitative analyses of each composition were then done 

Condusions 

These analyses have given some insight into the chemical interactions that occur inside 

the firne upon heating. These interactions are responsible for eroding the crucble to a point 

where is cracks open (K-01) as well as causing the melt to bubble over the top ofthe crumble and 

leak onto the fimace floor (K-02). They will be addressed individually. 

The erosion of the boron nitride is an unavoidable consequence of heating iron oxide in 

air at such a high temperature. The ICP-AES results show that between 14 and 22% of the 

material dropped into the test section is boron. This is supported by the fact that residues inside 

the crucible are between 14 and 21% boron (ICP-AES) and residues on the furnace floor are 10 

to 301h boron (SEM). This means that for every kilogram of melt, approximately 200 grams of 

boron, or 460 grams of BN could be lost. From earlier observation of the smaller crumble, 

approximately 1/8 inch of the wall was eroded each time in the region of the molten liquid. This 

is a significant amount, which motivated the design of the second crucible to have thicker walls. 

It is not completely understood why the iron oxide bubbled over the top of the crucble 

in test K-02. It is highly unlikely that the melt reached boiling temperatures, but it is known that 

gases are given off as the iron oxide is heated. Fe2O3 is the equilibrium state between iron and 

oxygen at normal temperatures and pressures, but at over 1650 EC, the equilibrium state shifts 

and oxygen will be given oWE This off-gassing might be responsible for the melt bubbling out. 

This hypothesis is supported by the fact that the x-ray diffmcton found evidence of Fe30 4 in the

NUREG/CR-6623 207
flrnace residue. Another explanation would be that any oxygen given off will form BA03 with 

the crucible material, and nitrogen gas will be given off, which again could cause bubbling. This 

hypothesis is supported by the fact that BA was also found in the fiunace floor residue. In 

either case, little can be done to prevent the interaction. The possibility of using magnetite 

(Fe3sO) instead ofhematite as a fuld was considered since it was thought that choosing a fuel that 

was closer to the stoichiomfetricaly stable state at 1700 EC would limit oxygen generation. This 

was rejected after small scale tests showed no improvement. It is possible that the magnetite 

becomes hematite during heating, which would then lead to no reduction on oxygen production 

once melting occurred. The other possibility is that the primary reason for the foaming of the fuel 

is the nitrogen off-gassing. 

It was decided that the only way to prevent serious Runace damage was to attempt to 

prevent the molten fuel from contacting the firname floor. A V thick boron nitride plate 

(Carborundum, Grade A) was placed underneath the crucible to catch any iron oxide that bubbled 

over the top. In addition, for the last few tests, and expendable alumina ring was placed on top 

of the plate as a type of wail that would slow the melt even more. These improvements only 

marginally helped, and damage to the furnace was almost always inevitable. 

NUREG/CR-6623 208

II
NRC FORM 335 U.S. NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER 

NRCM11) PAssined by NRC, Add Val. Supp., Rev.. 

3201.3202 BIBLIOGRAPHIC DATA SHEET adAdnm Numer% N any.) 

(See fmncos nm etem) 

2. TITLE AND SUBTITLE ( , NUREGICR-6623 

Vapor Explosions in a One-Dimensional Large Scale Geometry with Simulant Melts 

3. DATE REPORT PUBUSHED 

MONTH I YEAR 

October 1999 

4. FIN OR GRANT W6183 NUMBER 

5. AUTHOR(S) 6. TYPE OF REPORT 

H.S. Park, R. Chapman, and M.L Corradini Technical 

7. PERIOD COVERED prxhow Dates) 

April 1994 to September 1999 

8. PERFORMING ORGANIZATION - NAME AND ADDRESS (INRC, provide ,sion, Office orRegion, U.S. NucdearRegaetoryCwm,, nd maini •adss; idd otVcaoIr, ovid ne and maifing addr-s.) 

Nuclear Safety Research Center 

University of Wisconsin - Madison 

Madison, Wisconsin 53706

anr d a W. I "" * O PdU,=1%, ffNRC, ype "name as &bowl ifdoore p ARC Dmwon Ofe orRogion, U.S. NucmearaVtng and mef a~dret) j, 

Division of Systems Analysis and Regulatory Effectiveness 

Office of Nuclear Regulatory Research 

U.S. Nuclear Regulatory Commission 

Washington, D.C 20555-0001 4nl -Q00 ID 1.r-rT0% 6 Lifnt

S. Basu, NRC Project Mana.cer

11. ABSTRACT (2wo wors ark=) 

In light water reactors after a prolonged lack of cooling, vapor explosions could occur when molten fuel is generated and contacts residual water coolant within the reactor vessel or below in the containment reactor cavity. The objectives for this work were to obtain well-characterized expenmental data for the explosion propagation/escalation phases using different melt simulants and to investigate the effects of a comprehensive set of initial and boundary conditions on the explosion energetics; i.e., bigger strength, fuel mass, composition and temperature, coolant mass, viscosity and temperature, and system constraint 

This experimental work has yielded a number of results that have potentiallyimportanti safety implications. First, it has provided evidence of the reproducibility of vapor explosion energetics for a co.rtrolled set of initial and boundary conditions. Second, the experiments have demonstrated that if geometric scalinp is properly specified, it is possible to extrapolate the results of laboratory scale experiments to reactor scale predictions. Finally and most importantly, the experimental data suggests tat once thefcoolant.initial conditons ar within anenvelope fr triggered events, the energetics is much less than thermodyami lm, apparenty due to the small amount of fuel that participates in the explosion time scale. This envelope of tggerabity is much smaler fr a simulat molten oxide with low superheat, such as molten iron-oxide in the tests described in the report and corum in the KROTOS tests.

12. KEY WORDS/DESCRIPTORS (List wrd orwatses M usas at mches ear inbnguge nmpor) I& AVAILABUJTY STATEMENT 

fuel-coolant interactions (FCI), steam explosions, vapor explosions, simulant melts, molten tin, molten unlimited Iron oxide, corium, zirconium-containing melt, chemical augmentation, explosion 14. SECURLM CLASSFICATION propagation/escalation, explosinn ssn~mAtiem_ explosivity, thermodynamic limit, triggering "f "n5 envelope, initial and boundary conditions 0t'Wer strength, fuel mass and composition, superheat, subcooling, unclassied viscosity, system constraint, severe accident, nuclear reactor safety 

(Thi Repot 

unclassified 

15. NUMBER OF PAGES 

16. PRICE 

NRC FORM 33.5(2-M)1

This form was lectroncally prduced by Elite Federal Fonms kic.

V • b !1•1%1 •VJ;g 'I
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on recycled%, 

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