NUREG/IA-0481

Evaluation of TRACE Spacer Grid Model with

FLECHT-SEASET Reflood Test

Prepared by:

Byung-Gil HUH, Ae-Ju CHEONG, Kyung Won LEE

Korea Institute of Nuclear Safety

62 Gwahak-ro, Yuseong-gu

Daejeon, 34142, Korea

K. Tien, NRC Project Manager

Division of Systems Analysis

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

Manuscript Completed: December 2016

Date Published: August 2017

Prepared as part of 

The Agreement on Research Participation and Technical Exchange

Under the Code Applications and Maintenance Program (CAMP)

Published by

U.S. Nuclear Regulatory Commission

NUREG/IA-0481

Evaluation of TRACE Spacer Grid Model with

FLECHT-SEASET Reflood Test 

Prepared by:

Byung-Gil HUH, Ae-Ju CHEONG, Kyung Won LEE

Korea Institute of Nuclear Safety

62 Gwahak-ro, Yuseong-gu

Daejeon, 34142, Korea

K. Tien, NRC Project Manager

Division of Systems Analysis

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

Manuscript Completed: December 2016

Date Published: August 2017

Prepared as part of 

The Agreement on Research Participation and Technical Exchange

Under the Code Applications and Maintenance Program (CAMP)

Published by

U.S. Nuclear Regulatory Commission

iii

ABSTRACT

The effects of spacer grid model of TRACE V5.0 patch 4 were assessed for the Full-Length 

Emergency Core Heat Transfer Separate Effects and Systems Effects Tests (FLECHT-SEASET) 

that were the typical reflood heat transfer tests. The FLECHT-SEASET test section was modeled 

in the VESSEL component of TRACE and the 161 heated rods in 17x17 assemblies were 

modeled as a single HTSTR component. The injected flow rates and temperatures were provided 

as a function of time by a FILL component connecting to the bottom of the lower plenum. The 

BREAK component was used to set the pressure boundary at the top of the test section. The 

main parameters of the spacer grid were defined by the experimental design data and eight eggcrate grids were modeled in the VESSEL component of TRACE. The calculations for eight tests of 

FLECHT-SEASET revealed that when the spacer grid model was used, the rod temperatures 

decreased and the rods were quenched at an earlier time in most other tests. In addition, as the 

reflood rate increased, the lower peak rod temperature and the earlier quenching time were 

predicted. When the test pressure was lower, the higher rod temperature and the later rod 

quenching were predicted since the liquid approached a relatively lower saturation temperature 

faster. When the subcooling degree was higher, the reduced degree of quenching time due to the 

spacer grid was further decreased because the higher subcooling degree enhanced the heat 

transfer rate. Sensitivity studies were performed to identify the effect of the grid locations and the 

difference from the spacer grid model of RELAP5. In this study, the effect of the spacer grid model 

in TRACE is shown well to simulate the FLECHT-SEASET reflood heat transfer tests. However, 

since the droplet breakup and the grid rewetting models were not fully implemented yet, there 

were some limitations in quantitatively predicting their effects. The comparison with the RELAP5 

revealed that the current RELAP5 version had some errors in implementing the spacer grid 

model, and the effect of the spacer grid of TRACE could have been over-estimated for the rod 

temperature behaviors as compared with RELAP5.

v

TABLE OF CONTENTS

Page

ABSTRACT.....................................................................................................................................iii

LIST OF FIGURES........................................................................................................................ vii

LIST OF TABLES............................................................................................................................xi

EXECUTIVE SUMMARY...............................................................................................................xiii

1 INTRODUCTION..........................................................................................................................1

2 TEST FACILITY AND MODEL DESCRIPTION...........................................................................3

2.1 Description of the FLECHT-SEASET Experiment..............................................................3

2.1.1 Spacer Grid Model in TRACE.................................................................................6

3 EVALUATION OF THE SPACER GRID MODEL FOR FLECHTSEASET REFLOOD TESTS........................................................................................................9

3.1 TRACE Modeling................................................................................................................9

3.2 Evaluation for Spacer Grid Model.....................................................................................13

4 SENSITIVITY ANALYSIS..........................................................................................................99

4.1 Effect of the Grid Location................................................................................................99

4.1.1 Run No. 31805.....................................................................................................99

4.1.2 Run No. 31302...................................................................................................105

4.2 Comparison with Spacer Grid Model of RELAP5...........................................................109

5 CONCLUSIONS.......................................................................................................................129

6 REFERENCES........................................................................................................................131

vii

LIST OF FIGURES

Page

Figure 2-1 FLECHT-SEASET Test Facility................................................................................4

Figure 2-2 FLECHT-SEASET Test Bundle Cross Section.........................................................5

Figure 3-1 TRACE Nodalization for FLECHT-SEASET Facility...............................................11

Figure 3-2 Axial Power Profile of Heated Rod.........................................................................12

Figure 3-3 Liquid Inlet Flowrate – Run No.31805 ....................................................................15

Figure 3-4 Liquid Inlet Temperature – Run No.31805 .............................................................16

Figure 3-5 Heater Rod Temperature at 0.6 m – Run No.31805 ..............................................16

Figure 3-6 Heater Rod Temperature at 1.2 m – Run No.31805 ..............................................17

Figure 3-7 Heater Rod Temperature at 1.8 m – Run No.31805 ..............................................17

Figure 3-8 Heater Rod Temperature at 1.98 m – Run No.31805 ............................................18

Figure 3-9 Heater Rod Temperature at 2.4 m – Run No.31805 ..............................................18

Figure 3-10 Heater Rod Temperature at 3.0 m – Run No.31805 ..............................................19

Figure 3-11 Heater Rod Temperature at 3.3 m – Run No.31805 ..............................................19

Figure 3-12 Heater Rod Temperature at 3.5 m – Run No.31805 ..............................................20

Figure 3-13 Quench Front Profile – Run No.31805...................................................................20

Figure 3-14 Differential Pressure for Entire 12 ft – Run No.31805 ............................................21

Figure 3-15 Differential Pressure at 10~11 ft Elevation – Run No.31805..................................21

Figure 3-16 Vapor Temperature at 1.8 m – Run No.31805.......................................................22

Figure 3-17 Vapor Temperature at 3.0 m – Run No.31805.......................................................22

Figure 3-18 Heat Transfer Coefficient at 1.98 m – Run No.31805 ............................................23

Figure 3-19 Heat Transfer Coefficient at 2.4 m – Run No.31805 ..............................................23

Figure 3-20 Heat Transfer Coefficient at 3.0 m – Run No.31805 ..............................................24

Figure 3-21 Liquid Inlet Flowrate – Run No.31504 ....................................................................25

Figure 3-22 Liquid Inlet Temperature – Run No. 31504 ............................................................26

Figure 3-23 Heater Rod Temperature at 0.6 m – Run No.31504 ..............................................26

Figure 3-24 Heater Rod Temperature at 1.2 m – Run No.31504 ..............................................27

Figure 3-25 Heater Rod Temperature at 1.8 m – Run No.31504 ..............................................27

Figure 3-26 Heater Rod Temperature at 1.98 m – Run No.31504 ............................................28

Figure 3-27 Heater Rod Temperature at 2.4 m – Run No.31504 ..............................................28

Figure 3-28 Heater Rod Temperature at 3.0 m – Run No.31504 ..............................................29

Figure 3-29 Heater Rod Temperature at 3.3 m – Run No.31504 ..............................................29

Figure 3-30 Heater Rod Temperature at 3.5 m – Run No.31504 ..............................................30

Figure 3-31 Quench Front Profile – Run No.31504...................................................................30

Figure 3-32 Differential Pressure for Entire 12 ft – Run No.31504 ............................................31

Figure 3-33 Differential Pressure at 10~11 ft Elevation – Run No.31504..................................31

Figure 3-34 Vapor Temperature at 1.8 m – Run No.31504.......................................................32

Figure 3-35 Vapor Temperature at 3.0 m – Run No.31504.......................................................32

Figure 3-36 Heat Transfer Coefficient at 1.98 m – Run No.31504 ............................................33

Figure 3-37 Heat Transfer Coefficient at 2.4 m – Run No.31504 ..............................................33

Figure 3-38 Heat Transfer Coefficient at 3.0 m – Run No.31504 ..............................................34

Figure 3-39 Liquid Inlet Flowrate – Run No.31203 ....................................................................35

Figure 3-40 Liquid Inlet Temperature – Run No. 31203 ............................................................36

Figure 3-41 Heater Rod Temperature at 0.6 m – Run No.31203 ..............................................36

Figure 3-42 Heater Rod Temperature at 1.2 m – Run No.31203 ..............................................37

Figure 3-43 Heater Rod Temperature at 1.8 m – Run No.31203 ..............................................37
viii

Figure 3-44 Heater Rod Temperature at 1.98 m – Run No.31203 ............................................38

Figure 3-45 Heater Rod Temperature at 2.4 m – Run No.31203 ..............................................38

Figure 3-46 Heater Rod Temperature at 3.0 m – Run No.31203 ..............................................39

Figure 3-47 Heater Rod Temperature at 3.3 m – Run No.31203 ..............................................39

Figure 3-48 Heater Rod Temperature at 3.5 m – Run No.31203 ..............................................40

Figure 3-49 Quench Front Profile – Run No.31203...................................................................40

Figure 3-50 Differential Pressure for Entire 12 ft – Run No.31203 ............................................41

Figure 3-51 Differential Pressure at 10~11 ft Elevation – Run No.31203..................................41

Figure 3-52 Vapor Temperature at 1.8 m – Run No.31203.......................................................42

Figure 3-53 Vapor Temperature at 3.0 m – Run No.31203.......................................................42

Figure 3-54 Heat Transfer Coefficient at 1.98 m – Run No.31203 ............................................43

Figure 3-55 Heat Transfer Coefficient at 2.4 m – Run No.31203 ..............................................43

Figure 3-56 Heat Transfer Coefficient at 3.0 m – Run No.31203 ..............................................44

Figure 3-57 Liquid Inlet Flowrate – Run No.31302 ....................................................................46

Figure 3-58 Liquid Inlet Temperature – Run No. 31302 ............................................................46

Figure 3-59 Heater Rod Temperature at 0.6 m – Run No.31302 ..............................................47

Figure 3-60 Heater Rod Temperature at 1.2 m – Run No.31302 ..............................................47

Figure 3-61 Heater Rod Temperature at 1.8 m – Run No.31302 ..............................................48

Figure 3-62 Heater Rod Temperature at 1.98 m – Run No.31302 ............................................48

Figure 3-63 Heater Rod Temperature at 2.4 m – Run No.31302 ..............................................49

Figure 3-64 Heater Rod Temperature at 3.0 m – Run No.31302 ..............................................49

Figure 3-65 Heater Rod Temperature at 3.3 m – Run No.31302 ..............................................50

Figure 3-66 Heater Rod Temperature at 3.5 m – Run No.31302 ..............................................50

Figure 3-67 Quench Front Profile – Run No.31302...................................................................51

Figure 3-68 Differential Pressure for Entire 12 ft – Run No.31302 ............................................51

Figure 3-69 Differential Pressure at 10~11 ft Elevation – Run No.31302..................................52

Figure 3-70 Vapor Temperature at 1.8 m – Run No.31302.......................................................52

Figure 3-71 Vapor Temperature at 3.0 m – Run No.31302.......................................................53

Figure 3-72 Heat Transfer Coefficient at 1.98 m – Run No.31302 ............................................53

Figure 3-73 Heat Transfer Coefficient at 2.4 m – Run No.31302 ..............................................54

Figure 3-74 Heat Transfer Coefficient at 3.0 m – Run No.31302 ..............................................54

Figure 3-75 Liquid Inlet Flowrate – Run No.31701 ....................................................................56

Figure 3-76 Liquid Inlet Temperature – Run No. 31701 ............................................................57

Figure 3-77 Heater Rod Temperature at 0.6 m – Run No.31701 ..............................................57

Figure 3-78 Heater Rod Temperature at 1.2 m – Run No.31701 ..............................................58

Figure 3-79 Heater Rod Temperature at 1.8 m – Run No.31701 ..............................................58

Figure 3-80 Heater Rod Temperature at 1.98 m – Run No.31701 ............................................59

Figure 3-81 Heater Rod Temperature at 2.4 m – Run No.31701 ..............................................59

Figure 3-82 Heater Rod Temperature at 3.0 m – Run No.31701 ..............................................60

Figure 3-83 Heater Rod Temperature at 3.3 m – Run No.31701 ..............................................60

Figure 3-84 Heater Rod Temperature at 3.5 m – Run No.31701 ..............................................61

Figure 3-85 Quench Front Profile – Run No.31701...................................................................61

Figure 3-86 Differential Pressure for Entire 12 ft – Run No.31701 ............................................62

Figure 3-87 Differential Pressure at 10~11 ft Elevation – Run No.31701..................................62

Figure 3-88 Vapor Temperature at 1.8 m – Run No.31701.......................................................63

Figure 3-89 Vapor Temperature at 3.0 m – Run No.31701.......................................................63

Figure 3-90 Heat Transfer Coefficient at 1.98 m – Run No.31701 ............................................64

Figure 3-91 Heat Transfer Coefficient at 2.4 m – Run No.31701 ..............................................64

Figure 3-92 Heat Transfer Coefficient at 3.0 m – Run No.31701 ..............................................65
ix

Figure 3-93 Liquid Inlet Flowrate – Run No.31108 ....................................................................67

Figure 3-94 Liquid Inlet Temperature – Run No. 31108 ............................................................67

Figure 3-95 Heater Rod Temperature at 0.6 m – Run No.31108 ..............................................68

Figure 3-96 Heater Rod Temperature at 1.2 m – Run No.31108 ..............................................68

Figure 3-97 Heater Rod Temperature at 1.8 m – Run No.31108 ..............................................69

Figure 3-98 Heater Rod Temperature at 1.98 m – Run No.31108 ............................................69

Figure 3-99 Heater Rod Temperature at 2.4 m – Run No.31108 ..............................................70

Figure 3-100 Heater Rod Temperature at 3.0 m – Run No.31108 ..............................................70

Figure 3-101 Heater Rod Temperature at 3.3 m – Run No.31108 ..............................................71

Figure 3-102 Heater Rod Temperature at 3.5 m – Run No.31108 ..............................................71

Figure 3-103 Quench Front Profile – Run No.31108...................................................................72

Figure 3-104 Differential Pressure for Entire 12 ft – Run No.31108 ............................................72

Figure 3-105 Differential Pressure at 10~11 ft Elevation – Run No.31108..................................73

Figure 3-106 Vapor Temperature at 1.8 m – Run No.31108.......................................................73

Figure 3-107 Vapor Temperature at 3.0 m – Run No.31108.......................................................74

Figure 3-108 Heat Transfer Coefficient at 1.98 m – Run No.31108 ............................................74

Figure 3-109 Heat Transfer Coefficient at 2.4 m – Run No.31108 ..............................................75

Figure 3-110 Heat Transfer Coefficient at 3.0 m – Run No.31108 ..............................................75

Figure 3-111 Liquid Inlet Flowrate – Run No.32013 ....................................................................77

Figure 3-112 Liquid Inlet Temperature – Run No.32013 .............................................................77

Figure 3-113 Heater Rod Temperature at 0.6 m – Run No.32013 ..............................................78

Figure 3-114 Heater Rod Temperature at 1.2 m – Run No.32013 ..............................................78

Figure 3-115 Heater Rod Temperature at 1.8 m – Run No.32013 ..............................................79

Figure 3-116 Heater Rod Temperature at 1.98 m – Run No.32013 ............................................79

Figure 3-117 Heater Rod Temperature at 2.4 m – Run No.32013 ..............................................80

Figure 3-118 Heater Rod Temperature at 3.0 m – Run No.32013 ..............................................80

Figure 3-119 Heater Rod Temperature at 3.3 m – Run No.32013 ..............................................81

Figure 3-120 Heater Rod Temperature at 3.5 m – Run No.32013 ..............................................81

Figure 3-121 Quench Front Profile – Run No.32013...................................................................82

Figure 3-122 Differential Pressure for Entire 12 ft – Run No.32013 ............................................82

Figure 3-123 Differential Pressure at 10~11 ft Elevation – Run No.32013..................................83

Figure 3-124 Vapor Temperature at 1.8 m – Run No.32013.......................................................83

Figure 3-125 Vapor Temperature at 3.0 m – Run No.32013.......................................................84

Figure 3-126 Heat Transfer Coefficient at 1.98 m – Run No.31701 ............................................84

Figure 3-127 Heat Transfer Coefficient at 2.4 m – Run No.31701 ..............................................85

Figure 3-128 Heat Transfer Coefficient at 3.0 m – Run No.31701 ..............................................85

Figure 3-129 Liquid Inlet Flowrate – Run No.32114 ....................................................................87

Figure 3-130 Liquid Inlet Temperature – Run No. 32114 ............................................................87

Figure 3-131 Heater Rod Temperature at 0.6 m – Run No.32114 ..............................................88

Figure 3-132 Heater Rod Temperature at 1.2 m – Run No.32114 ..............................................88

Figure 3-133 Heater Rod Temperature at 1.8 m – Run No.32114 ..............................................89

Figure 3-134 Heater Rod Temperature at 1.98 m – Run No.32114 ............................................89

Figure 3-135 Heater Rod Temperature at 2.4 m – Run No.32114 ..............................................90

Figure 3-136 Heater Rod Temperature at 3.0 m – Run No.32114 ..............................................90

Figure 3-137 Heater Rod Temperature at 3.3 m – Run No.32114 ..............................................91

Figure 3-138 Heater Rod Temperature at 3.5 m – Run No.32114 ..............................................91

Figure 3-139 Quench Front Profile – Run No.32114...................................................................92

Figure 3-140 Differential Pressure for Entire 12 ft – Run No.32114 ............................................92

Figure 3-141 Differential Pressure at 10~11 ft Elevation – Run No.32114..................................93
x

Figure 3-142 Vapor Temperature at 1.8 m – Run No.32114.......................................................93

Figure 3-143 Vapor Temperature at 3.0 m – Run No.32114.......................................................94

Figure 3-144 Heat Transfer Coefficient at 1.98 m – Run No.32114 ............................................94

Figure 3-145 Heat Transfer Coefficient at 2.4 m – Run No.32114 ..............................................95

Figure 3-146 Heat Transfer Coefficient at 3.0 m – Run No.32114 ..............................................95

Figure 3-147 Variation of Peak Rod Temperature at Elevation z = 1.98.....................................98

Figure 3-148 Variation of Quenching Time at Elevation z = 1.98................................................98

Figure 4-1 Node Change for Sensitivity of Grid Location.......................................................101

Figure 4-2 Heater Rod Temperature at 1.2 m – Run No.31805 ............................................102

Figure 4-3 Heater Rod Temperature at 1.98 m – Run No.31805 ..........................................102

Figure 4-4 Heater Rod Temperature at 2.4 m – Run No.31805 ............................................103

Figure 4-5 Heater Rod Temperature at 3.3 m – No.31805....................................................103

Figure 4-6 Quench Front Profile – Run No.31805.................................................................104

Figure 4-7 Heat Transfer Coefficient at 1.98 m – Run No.31805 ..........................................104

Figure 4-8 Heater Rod Temperature at 1.2 m – Run No.31302 ............................................106

Figure 4-9 Heater Rod Temperature at 1.98 m – Run No.31302 ..........................................106

Figure 4-10 Heater Rod Temperature at 2.4 m – Run No.31302 ............................................107

Figure 4-11 Heater Rod Temperature at 3.3 m – No.31302....................................................107

Figure 4-12 Quench Front Profile – Run No.31302.................................................................108

Figure 4-13 Heat Transfer Coefficient at 2.4 m – Run No.31302 ............................................108

Figure 4-14 RELAP5 Nodalization for the FLECHT-SEASET Test.........................................110

Figure 4-15 Heater Rod Temperature at 1.2 m – Run No.31805 ............................................111

Figure 4-16 Heater Rod Temperature at 1.98 m – Run No.31805 ..........................................112

Figure 4-17 Heater Rod Temperature at 3.0 m – Run No.31805 ............................................112

Figure 4-18 Heater Rod Temperature at 1.2 m – Run No.31805 ............................................113

Figure 4-19 Heater Rod Temperature at 1.98 m – Run No.31805 ..........................................113

Figure 4-20 Heater Rod Temperature at 3.0 m – Run No.31805 ............................................114

Figure 4-21 Heater Rod Temperature at 1.2 m – Run No.31805 ............................................116

Figure 4-22 Heater Rod Temperature at 1.98 m – Run No.31805 ..........................................116

Figure 4-23 Heater Rod Temperature at 2.4 m – Run No.31805 ............................................117

Figure 4-24 Heater Rod Temperature at 3.0 m – Run No.31805 ............................................117

Figure 4-25 Heater Rod Temperature at 3.3 m – Run No.31805 ............................................118

Figure 4-26 Heater Rod Temperature at 3.5 m – Run No.31805 ............................................118

Figure 4-27 Axial Rod Temperature Profile at 150 sec – Run No.31805.................................119

Figure 4-28 Axial Rod Temperature Profile for Sub-models at 150 sec – Run No.31805.......119

Figure 4-29 Variation of Heat Transfer Enhancement for Laminar Enhancement Factor........120

Figure 4-30 Heater Rod Temperature at 1.2 m – Run No.31302 ............................................121

Figure 4-31 Heater Rod Temperature at 1.98 m – Run No.31302 ..........................................121

Figure 4-32 Heater Rod Temperature at 2.4 m – Run No.31302 ............................................122

Figure 4-33 Heater Rod Temperature at 3.0 m – Run No.31302 ............................................122

Figure 4-34 Heater Rod Temperature at 3.3 m – Run No.31302 ............................................123

Figure 4-35 Heater Rod Temperature at 3.5 m – Run No.31302 ............................................123

Figure 4-36 Heater Rod Temperature at 1.2 m – Run No.31805 ............................................125

Figure 4-37 Heater Rod Temperature at 1.98 m – Run No.31805 ..........................................125

Figure 4-38 Heater Rod Temperature at 3.0 m – Run No.31805 ............................................126

Figure 4-39 Heater Rod Temperature at 1.2 m – Run No.31805 ............................................126

Figure 4-40 Heater Rod Temperature at 1.98 m – Run No.31805 ..........................................127

Figure 4-41 Heater Rod Temperature at 3.0 m – Run No.31805 ............................................127
xi

LIST OF TABLES

Page

Table 3-1 Test Matrix for TRACE Evaluation............................................................................... 13

Table 3-2 Peak Temperature and Quenching Time at Elevation z = 1.98................................... 97

Table 4-1 Peak Temperature and Quenching Time at Elevation z = 1.98 for Run No. 31805.. 100

Table 4-2 Peak Temperature and Quenching Time at Elevation z = 1.98 for Run No. 31302.. 105

xiii

EXECUTIVE SUMMARY

TRACE implemented spacer grid models in 2010 that enhanced the heat transfer downstream of 

spacer grids, but a systematic assessment of those models has not been performed for various postCHF heat transfer tests such as FLECHT-SEASET, THTF, etc. In particular, the FLECHT-SEASET 

tests are considered representative reflood experiments because the facility is relatively large and well 

instrumented. In this study, the spacer grid model of TRACE was evaluated for the FLECHT-SEASET 

test by the TRACE V5.0 patch4.

The Full-Length Emergency Core Heat Transfer Separate Effects and Systems Effects Tests 

(FLECHT-SEASET) was conducted to identify the thermal hydraulic phenomena of forced and gravity 

reflooding in a 161-rod bundle without flow blockage. The FLECHT-SEASET test section was 

modeled in the VESSEL component of TRACE. The test section was divided into 16 axial nodes: one 

node for the lower plenum, fourteen nodes for the heated section and one node for the upper plenum, 

and there were two cells between each gird. The 161 heated rods in 17x17 assemblies were modeled 

as a single HTSTR component. The injected flow rates and temperatures were provided as a function 

of time by the FILL component, which was connected to the bottom of the lower plenum. The BREAK 

component was used to set the pressure boundary at the top of the test section. Eight egg-crate grids 

without a mixing vane were located in the bottom of every other node. The grid parameters were 

determined by the experimental design data and the general fuel data. 

Eight tests were selected for analyses, and tests covered a range of the flooding rate from 2.10 

cm/sec to 15.50 cm/sec, the subcooling temperature from 5 o

C to 79 oC and the upper plenum 

pressure from 0.13 MPa to 0.41 MPa. The initial rod power at the peak location was 2.3 kW/m (0.7 

kW/ft) in all the tests. As would be expected, the rod temperatures decreased and the rods were 

quenched at an earlier time in most tests if the spacer grid model was used. In addition, as the reflood

rate increased, a lower peak rod temperature and earlier quenching time were predicted. In tests with 

a high reflood rate, the change in the peak temperature due to the spacer grid was not large, which 

resulted from a short heat up period and a faster increase in the liquid level by the high reflooding rate. 

As the test pressure decreased, a higher rod temperature and later rod quenching were predicted 

since the liquid approached the relatively lower saturation temperature faster. The use of the spacer 

grid model at a lower pressure showed relatively bigger differences for rod temperature, but not for 

quenching time. With a higher subcooling degree, the decreased amount of quenching time due to the 

spacer grid also reduced because the high subcooling degree enhanced the heat transfer rate.

Sensitivity studies were performed to identify the effect of the grid locations and the difference from the 

spacer grid model of RELAP5. When the locations of spacer grids were changed to the top of every 

other node, the rod temperature was higher and the rod quenching was delayed at a high power than 

in the case in which the spacer grids were located in the bottom of the node. It would be more 

reasonable to predict the experimental data, which would have more conservative results. The 

RELAP5 code currently implemented the KNF reflood model and the spacer grid model. This spacer 

grid model considered three sub-models: single-phase heat transfer enhancement, grid rewet, and 

droplet breakup. However, the current RELAP5 version(Version 3.3jz~3.3kl)may have some errors in

implementingthe KNF reflood and spacer grid model. It may be because there are some problems 

with using the KNF reflood model (Option 40) and the spacer grid input (43000000 cards). Therefore, 

the developmental version of KNF was used to identify the effect of sub-models, and then the droplet 

breakup model had the biggest effect on the rod temperatures among the three models. The effect of 

the droplet breakup model was more significant at the higher elevation since the droplet velocity and 

the number of entrained droplets could be larger at the higher elevation. When the RELAP5 results 

were compared with the TRACE results, the effect of the spacer grid model in TRACE was more 

significant even though the single-phase heat transfer enhancement was only implemented in 
xiv

TRACE. These results may be due to the modeling characteristics and the use of the laminar 

enhancement factor in TRACE. The convective heat transfer enhancement may be excessively large 

if the laminar enhancement factor is large.

In conclusion, the effect of the spacer grid model in TRACE was shown well to simulate the FLECHTSEASET reflood heat transfer experiments. However, there were some limitations in quantitatively 

predicting the effect of the droplet breakup and the grid rewetting models. The comparison with the 

RELAP5 revealed that the current RELAP5 version had some errors in implementing the spacer grid 

model, and the effect of the spacer grid of TRACE may have been over-predicted for the rod 

temperature behaviors as compared to RELAP5. Therefore, in future studies the current RELAP5 

needs to be modified to correct the errors for the spacer grid model and the TRACE code should be 

improved to implement the droplet breakup and the grid rewetting models. 
1

1 INTRODUCTION

A spacer grid is generally installed in most commercial pressurized light water reactor (PWR) fuel 

assemblies. The spacer grid was developed by evaluating, among other things, the fuel rod vibration 

performance, the fretting wear resistance, the heat transfer enhancement, and the pressure drop 

characteristics. In particular, the spacer grid can influence the analysis for a large break loss of coolant 

accident (LOCA), is the most severe accident of a PWR because it affects the thermal-hydraulics in 

the core during the blowdown and reflood phases. After initiating a large break in the cold legs, the 

two-phase mixture formed by the flashing flows through the core and discharges into the break. The 

reflood phase typically starts by increasing the subcooled water from the bottom of the core. Large 

amounts of steam are generated by interacting between the subcooled water and hot fuels. This 

steam increases the core pressure and then hinders the reflooding of water. Various heat transfer 

regimes such as single-phase vapor convection, nucleate boiling, transition boiling, and film boiling 

also appears during this phase [1]. Therefore, the spacer grid can affect the heat transfer mechanism 

of the core during these phases. However, the effect of the spacer grid has not been evaluated 

properly in various accident analyses, and it has just been considered by adding the loss coefficient at 

the location of spacer grid. 

The spacer grid was originally designed to maintain a geometrical configuration of the fuel bundles 

and support fuel rods laterally and vertically. However, the spacer grids affected the fluid dynamics and 

the heat transfer in the core. The spacer grids can create an obstacle to the fluid flow in the core and 

then increase the overall pressure losses. The spacer grids can also decrease the flow area by 

contracting and accelerating the flow, and the mixing vanes of the spacer grid typically promotes 

turbulence and induces a strong swirling flow in the core, which increases the local heat transfer 

downstream of the spacer grid. In addition, if the spacer grids are quenched, their surfaces are 

covered with liquid film and provide an additional interface area between liquid and vapor. Finally, the 

spacer grids can also break up the entrained droplets into smaller droplets and therefore the 

downstream vapor temperature decreases due to the enhancement of the evaporative heat transfer of 

the smaller droplets.

The TRACE code is a thermal-hydraulic system code and was developed by USNRC for a realistic 

analysis of thermal-hydraulics transients in pressurized water reactors [2]. TRACE implemented the 

spacer grid model in December 2010 [3]. The spacer grid models in TRACE consists of the singlephase convective enhancement model, the pressure loss model, the droplet breakup model, and the 

spacer grid rewet model, but the droplet breakup model and the grid re-wetting model have not been 

fully implemented in the current TRACE. The systematic assessment of those models has not been 

performed sufficiently with various post-CHF heat transfer tests, such as FLECHT-SEASET, THTF, 

etc., though the effect of the spacer grid of TRACE was recently evaluated for RBHT [4xx]. In 

particular, the FLECHT-SEASET tests have been considered as representative reflood experiments 

because the facility is relatively large and well instrumented [5]. 

In this study, calculations using TRACE V5.0 patch4 code [2] released in April 2014 and a comparison 

with experimental data were performed for FLECHT-SEASET reflood heat transfer tests to assess the 

effect of the spacer grid model of the TRACE code. Some tests were also evaluated using RELAP5

code with the spacer grid model and their results were compared to those of TRACE. 

A brief description for FLECHT-SEASET facility and the spacer grid models of TRACE is made in 

Chapter 2. The comparative assessment of the TRACE spacer grid model against eight FLECHTSEASET tests is described in Chapter 3. Chapter 4 describes the results for the sensitivity analysis, 

such as the effect of the grid location and the comparison to the RELAP5. Finally, the conclusions of 

this study are provided in Chapter 5.

3

2 TEST FACILITY AND MODEL DESCRIPTION

2.1 Description of the FLECHT-SEASET Experiment

The Full-Length Emergency Core Heat Transfer Separate Effects and Systems Effects Tests 

(FLECHT-SEASET) was constructed and conducted to identify the thermal hydraulic of forced and 

gravity reflooding in a 161-rod bundle without flow blockage [5]. The FLECHT-SEASET tests was a 

large and relatively well instrumented reflood experiments. This test facility was modified from the 

FLECHT facility to apply a new heater rod bundle of which dimensions were typical of the 

Westinghouse 17x17 fuel bundle. 

Figure 2-1 showed the FLECHT-SEASET test facility schematic. It comprised a cylindrical test section, 

a coolant accumulator, an entrained liquid separation tank, an external pipe downcomer for the gravity 

reflood tests, a steam boiler for back-pressure regulation, and the required piping and valves. The low

mass housing was designed to minimize the wall effects so that the rods one row or more away from 

the housing wall [6]. Three hosing windows were installed at 0.91, 1.83, and 2.74 m elevations. A 

cross section of the test bundle was shown in Figure 2-2. The bundle contained 177 heater rods which 

consisted of 161 heater rods (93 noninstrumented and 68 instrumented), 16 thimbles, and 12 steam 

probes [7]. The 177 heater rods were placed in a cylinder of 0.194 m diameter and the heated length 

was 3.66 m (12 ft). There were also 8 solid triangular fillers and 8 grids in the test bundle. The 

triangular fillers were welded to the grids to maintain the proper grid location and decreased the 

amount of flow area from 9.3%to 4.7%. 

The fuel rods were simulated with electrically heated rods with a Kanthal heater coil imbedded in 

boron nitride encased with stainless steel cladding. The outside diameter of heater rod is 9.5 mm 

(0.374 inch). The heater rod had the wall thickness of 0.64 mm (0.025 inch) and the heated length of 

3.66 m (12 ft) and the pitch of 12.6 mm (0.496 inch). The fuel bundle has 8 spacer grids which was the 

egg-crate type without the mixing vane. The grid locations were similar to a 17x17 PWR fuel 

assembly. The blockage ratio of the spacer grids was estimated to be 0.29. The axial power profile in 

the heater rod was considered as the cosine curve with a power peak-to-average ratio of 1.66 and the 

radial power profile was assumed as uniform. 

To determine the experimental conditions, the reflood phase following a LOCA transient was evaluated 

to start approximately 30 seconds after a hypothetical break. By referring the reflood transient for a 

standard Westinghouse 17x17 four-loop, the reference test conditions were followed: initial clad 

temperature (871 o

C), peak power (2.4 kW/m), upper plenum pressure (0.28 MPa), Injection rate with 

lower plenum initially full (25 mm/sec), coolant subcooling (78 o

C). 

For the reflood experiments, the test bundle was pre-heated to the desired pressure and temperature 

with dry steam, and then cooling water was delivered to the lower plenum of the bundle by a gascharged accumulator for the forced reflood tests to quench the rods, simulating the reflood process.
4

Figure 2-1 FLECHT-SEASET Test Facility
5

Figure 2-2 FLECHT-SEASET Test Bundle Cross Section
6

2.1.1 Spacer Grid Model in TRACE

There are the spacer grids in most fuel assembly. The functions of the spacer grid are to support fuel 

rods vertically and laterally, to maintain the space between rods and to enhance the flow mixing. In the 

thermal-hydraulic aspect, the spacer grid could provide flow obstacles in a core channel and then 

influence the heat transfer mechanism in a core. First, the spacer grid reduces the flow area and then 

the flow acceleration occurs at the location of a spacer grid. This could promote the local convective 

heat transfer. Second, the spacer grids can be quenched before the fuel rod during a reflood and it 

may increase the interfacial area between liquid and vapor. Third, the spacer grids can break the 

entrained droplets into smaller ones and this can increase the interfacial heat transfer to the vapor 

phase. 

Currently, there are 4 sub-models in TRACE, which are for the convective heat transfer enhancement, 

the pressure drop, the droplet breakup and the grid re-wetting, respectively. In this chapter, 4 submodels would be described shortly by referring the TRACE theory manual [3].

2.1.1.1 Convective heat transfer enhancement

The Yao, Hochreiter and Leech model [8] was applied in TRACE and it can be used in both egg-crate 

grids and mixing vane grids. This model consists of two parts; 1) the heat transfer enhancement due 

to the acceleration of the flow and the increased turbulence due to the spacer grid, and 2) the effects 

of mixing vanes: 

NNNN

NNNNoo

= �1 + 5.55ββ2ee

−0.13 xx

DDHH � ���������������

PPPPPPPP 1

�1 + aa2(tt tt∅)2ee

−0.034 xx

DDHH �

0.4

�������������������

PPPPPPPP 2

(1)

where NNNN is the local Nusselt number at the wall with the grid spacer, NNNNoo is the local Nusselt number 

at the wall without the spacer grid, ββ represents the spacer grid flow blockage area ratio as viewed 

from upstream, xx is the axial distance from the downstream edge of the spacer, and DDHH is the 

hydraulic diameter of the flow channel, aa is the mixing vane flow blockage area ratio when viewed 

from upstream, and ∅ is the angle of the vane with respect to the axial direction.

As shown in Eq. (1), this model depends on the spacer grid blockage ration, the mixing vane blockage 

ration, and the mixing vane angle. Also, in order to consider the increased enhancement effects for 

high void fraction, laminar flows, a laminar enhancement factor, F was introduced into this model as 

shown in TRACE theory manual [3]. 

2.1.1.2 Pressure losses

In order to consider the pressure drop due to the spacer grids, TRACE adopted the Yao, Loftus and 

Hochreiter [9] model. The loss coefficient for this model was improved by the Rehme [10] and the 40% 

increase of the loss factor was also applied to account for the sharp leading edge. 

∆ppgg gggg = KKgg gggg

ρρ

2

VV2, KKgg gggg = 1.4 CCVV(ββ + aa) 2 (2)

where, CCVV = �

196RR mm

−0.33,103 < RR < 104

41RR mm

−0.16 ,104 < RR < 105

6.5 , 105 < RR

.
7

2.1.1.3 Droplet breakup

In the dispersed droplets flow, the droplet can split apart by the spacer gird strap. The droplet breakup 

model in TRACE is based on the study which was conducted by Yao, Hochreiter and Cai [11]. The 

shattered small droplet ratio could be represented as a function of the droplet Weber number and the 

ration of the Sauter mean diameter to the initial diameter. If the Weber number is greater than 250, the 

droplets don’t have sufficient surface tension force to overcome the impact with the spacer grid straps 

and the droplet breakup is possible [3]. The droplet mass flow rate in the downstream cell is followed:

mṁ ss ssss = 0.6(ββ + aa)mṁ oo, mṁ ll = mṁ oo − mṁ ss ssss (3)

The diameters of the shattered droplets and the remained large droplets are given as:

DDss ss = 6.16WW −0.53DDoo, DD ll = DDoo (4)

where WW = ρρllDDddVVdd

2

σσdd

is the droplet Weber number. 

This model is subject to the completion of droplet field equation since droplets shattered from a large 

group are taken as source and sink terms in the downstream axial cell. However, the droplet breakup 

model is not currently activated in TRACE and it has been waiting the full implementation of the 

droplet field. 

2.1.1.4 Grid re-wetting

During reflood conditions, the spacer grid with no internal heat generation could fall below the 

minimum film boiling temperature before the fuel rods. Therefore, the liquid film formed on a spacer 

grid increases the local interfacial area for heat transfer and then it can result in significant desuperheating of the vapor. In TRACE, the modified radiation model of Paik, Hochreiter, Kelly, and 

Kohrt [12] was used after several simplifications. This model is about the heat transfer between fuel 

rods, a spacer grid, and the continuous vapor phase. The heat balance equation for the spacer grid is 

given as:

ρρssssCC ∂∂TTssss

∂∂ = WWPPggggggAAgggggg

�qq" − qqcc cccc " � (5)

The detailed description for this model would refer to TRACE manual [3]. According to this model, the 

spacer grid temperature was calculated by the spacer grid surface area is obtained as the spacer grid 

height times the grid wetted perimeter multiplied by two to account for both sides of the spacer grid 

straps. 

However, in current TRACE, the spacer grid re-wetting model is only implemented to calculate the 

transient spacer grid temperatures and add the spacer grid surface area to the wallFilmArea variable if 

the spacer grid is quenched. Further work for a critical film thickness will be required for determining 

the film Nusselt number from which the film interfacial heat transfer coefficient is calculated.

9 

3 EVALUATION OF THE SPACER GRID MODEL FOR 

FLECHT-SEASET REFLOOD TESTS 

In general, when the spacer grid is considered, the flow area is reduced and the convective heat 

transfer is promoted due to the flow acceleration and the turbulence increases. This is shown in both 

egg-crate style spacer grids and mixing vane grids. Also, the mixing vane increases the convective 

heat transfer additionally in the case of mixing vane grids. In this section, the effect of spacer grid 

model of TRACE was performed on FLECHT-SEASET (Full-Length Emergency Core Heat Transfer – 

Separate Effects And System Effect Test) reflood tests [7]. The tests comprised of forced and gravity 

reflood tests to simulate the nuclear fuel arrays of PWR, which was similar to Westinghouse 17x17 

assemblies. Especially, the tests were the important data sources to predict the characteristics of 

reflood phase for large-break (LB) loss-of-coolant accidents (LOCAs) since some tests were 

conducted at temperatures exceeding the requirement limit (2200 o

F). The test facility consisted of the 

cylindrical test section with a lower plenum, the heater rod bundle and the upper plenum, the cooling 

water injection system, the entrained water separation tank and the carry-over liquid collection tank as 

shown in Figure 2-1. This facility has a heated length of 3.66 m (12.0 ft) of typical pressurized water 

reactor (PWR) fuel assembly. The test section consisted of a 17x17 rod bundle with 161 heated 

electrical rods and 16 thimble rods. The 68 rods of the 161 heater rods were instrumented and the 

total power was about 850 KW. The bundle comprised of 8 spacer grids, 12 steam probes and 8 solid 

triangular fillers which were used to reduce the flow area near the housing wall. The electrical rods 

have a diameter of 9.5 mm and the wall thickness of 0.64 mm and are arranged in a 17x17 array with 

a 12.6 mm pitch. Eight forced reflood tests in the TRACE assessment manual [6] were selected to 

evaluate the effect of spacer grid model of TRACE.

3.1 TRACE Modeling

The FLECHT-SEASET facility was modeled by TRACE to simulate the forced reflood tests. The test 

facility was made as five components of TRACE to simplify the water injection system and the carryover liquid collection tank. The FLECHT-SEASET test section was modeled in Vessel and HTSTR 

components of TRACE as shown in Figure 3-1. TRACE modeling was almost same as that of the 

previous study [6] beside of some corrections. The Vessel was divided into 16 axial nodes; one node 

for lower plenum, fourteen nodes for heated section and one node for upper plenum and there were 

two cells between each gird. As in the TRACE assessment manual [6], the bottom of spacer grid was 

located in the bottom of corresponding cell. Two PIPE components were modeled as the lower- and 

the upper- plenum respectively. The 161 heated rods in 17x17 assemblies were modeled as a single 

heat structure (HTSTR Component 6) while the bundle housing wall was modeled as another heat 

structure (HTSTR Component 7). The heated length of heater rod was a 3.66 m with a cosine axial 

power profile as shown in Figure 3-2 and the radial power distribution was uniform. The rod power was 

designed to represent the decay heat prescribed by 10CFR part 50 Appendix K [13] from 30 sec 

following LBLOCA, which the transient power was 1.2 times to ANS 71 model. 

The heater rod was modeled with 7 radial cells and the material properties of the previous study [6] 

were also used in this calculation. The injected flow rates and temperatures was provided as a 

function of time by FILL component connecting to the bottom of the lower plenum. The BREAK 

component was used to set the pressure boundary at the top of the test section. 

For the spacer grid, the grid is the egg-crate type without the mixing vane. Eight spacer grids were 

installed along 3.6 m heated length in FLECHT-SEASET. The grid straps made by Inconel 718 alloy 

sheets which are 0.38 mm (0.015 in) thick and are 45.0 mm (1.75 in) in height. Each grid span has 

two equal distance nodes (~ 10 inch or 10.5 inch) as shown in Figure 3-1. Thus, the grid was located 
10

in the bottom of every other node. The first grid is located 0.57 m (22.25 in) above the test section, 

which is just under the heated section. Actually, the grid locations were similarto a 17x17 PWR fuel 

assembly. In order to model the spacer grid, the experimental design data [5, 7] and the general fuel 

data were considered to determine some parameters. The spacer grid for TRACE was modeled as 

follows;

*n: Grid- egg-crate style

*

* gridid 

101

* spbloc vnbloc phi wetperm

0.29 0.0 0.0 11.8

* height strthick spmatid

0.045 3.8.0E-4 10

Meaning of the grid parameters

- gridid : Grid number ID

- spbloc : Spacer grid flow blockage area ratio

- vnbloc : Mixing vane flow blockage area ratio

- phi : Mixing vane angle measured from parallel with the top of the spacer grid to the mixing 

vane

- wetperim : Spacer grid wetted perimeter

- height : Spacer grid axial height

- strthick : Grid strap thickness of modeled spacer grid (0.015 inch)

- spmatid : Grid material ID
11

Figure 3-1 TRACE Nodalization for FLECHT-SEASET Facility
12

Figure 3-2 Axial Power Profile of Heated Rod
13

3.2 Evaluation for Spacer Grid Model

As mentioned before, eight tests were chosen for this calculation as shown in Table 3-1. Tests were 

covering a range of flooding rate from 2.10 cm/sec to 15.50 cm/sec, subcooling temperature from 5 o

C 

to 79 o

C and upper plenum pressure from 0.13 MPa to 0.41 MPa. The initial rod power at the peak 

location is 2.3 KW/m (0.7 KW/ft) in all tests. The test conditions in Table 3-1 were nominal values. The 

actual values in tests changed with time and then the actual inlet flow rates, temperatures and the 

upper plenum pressures were used as input with a function of time. Also, the rod power was 

considered as a function of time that it was reduced with time like the decay heat formula in the 

Appendix K. 

As the general reflood test, the test bundle was pre-heated up to the predetermined pressure and 

temperature with dry steam. After that, the cooling water was injected to the lower plenum of the test 

bundle in order to quench the rods. 

The instrumentations of the FLECHT-SEASET Facility were very extensive, including 205 heater rod 

thermocouples, 12 differential pressure transmitters positioned 0.3048 m (1 ft.) apart along the axial 

direction of the heated section, 12 steam probes, and inlet and outlet flow meters. TRACE simulations 

in the assessment manual [6] showed the results for rod cladding temperatures, Vapor temperatures, 

heat transfer coefficients, quench profile, differential pressures and void fraction during a reflood test. 

In this study, the results for showing the effects of spacer grid well were given for tests as follows; 

Table 3-1 Test Matrix for TRACE Evaluation

Run No.

Flooding Rate Upper Plenum 

Pressure

Coolant 

Inlet Temp.

Coolant Inlet

Subcooling Temp.

cm/s (in/sec) MPa (psia) oC (oF) oC (oF)

1 31805 2.10 (0.81)

0.28 (40)

51 (124)

~79 

(143)

79 (143)

2 31504 2.40 (0.97) 51 (124) 79 (143)

3 31203 3.84 (1.51) 52 (126) 78 (141)

4 31302 7.65 (3.01) 52 (126) 78 (141)

5 31701 15.5 (6.10) 53 (127) 77 (140)

6 31108 7.90 (3.11) 0.13 (19) 33 (91) 74 (134)

7 32013 2.64 (1.04) 0.41 (60) 66 (150) 79 (143)

8 32114 2.5-3.1 (1.0-1.22) 0.28 (40) 125 (257) 5 (10)
14

1) Run No. 31805

Run No. 31805 was a test with a flooding rate of 2.1 cm/sec at 0.28 MPa and 79 oC inlet subcooling 

temperature as shown in Table 3-1. However, as described previously, the actual input conditions 

changed largely with time and showed the big oscillations. Therefore, the inlet flowrate and 

temperature were considered as time-dependent variables as shown in Figures 3.3 ~ 3.4. Most initial 

conditions in the TRACE assessment manual [6] were applied in this study.

The rod temperatures at various elevations are shown in Figure 3-5 through Figure 3-12. These 

figures showed well the process that the rod was heating up by the initiation of reflood, turning to 

reduce during the reflood, and finally quenching. In the calculation without the spacer grid model, the 

predicted peak temperatures agreed reasonably with the experimental data at low elevations (z ≤ 2.4 

m), but showed over predicted results at higher elevations (3.0 m ~ 3.3 m). The maximum peak clad 

temperature was shown at elevation z=1.98 m (6.5 ft). At low elevations (z ≤ 1.2 m), the quenching 

time showed a good agreement with the data, but under predicted at higher elevations. For low 

reflooding rates such as Run No. 31805, the dominant flow regime was a highly dispersed flow flimboiling region in which the heat transfer rates were very low. Therefore, the peak clad temperature 

usually occurred in this region and the maximum peak clad temperature at elevation z=1.98 m had the 

highest value as compared to following 4 tests. In the spacer grid model of TRACE, Only two models 

for the convective enhancement and pressure loss models were currently implemented to perform the 

rod temperature. The droplet breakup and grid rewet models were not fully implemented in TRACE. 

When the spacer grid model was applied, the effect of mixing vane was not considered since the eggcrate spacer grid was installed in the FLECHT-SEASET. Therefore, the convective Nusselt number 

was enhanced due to the flow acceleration and the turbulence increase for a spacer grid without the 

mixing vane. As would be expected, the lower rod temperatures and earlier rod quenches were 

predicted in the case with a spacer grid model. During a heat up period, the effect of the spacer grid 

model did not show at lower elevations (z ≤ 2.4 m), but the earlier rise of rod temperature was 

predicted at higher elevations. Those predictions for the spacer grid would come from the relatively 

high vapor temperature due to the promotion of heat transfer from lower elevations. The turnings of 

rod temperature to reduce were estimated earlier than those withoutthe spacer grid at all elevations 

since the spacer grid enhanced the increase of collapsed water level and heat transfer. The peak 

temperatures with the spacer grid model had the lower values at low elevations (z ≤ 3.0 m) and under 

predicted the experimental data in comparison with those without the spacer grid model, while the 

peak temperature with the spacer grid model had the higher value at z = 3.3 elevation as shown in 

Figure 3-11. This might be resulted from the earlier rise of rod temperature during a heat up region. 

The maximum peak clad temperature at z=1.98 m was reduced in case with the spacer grid model. 

From a previous study [14], the droplet break model among sub-models for the spacer grid would 

have the largest effect to the rod temperature. Since the droplet break model could influence largely 

the rod quenching downstream of grid, the full implementation of spacer grid model could predict the 

lower temperatures at high elevations. At elevation z=1.98 with the maximum peak cladding 

temperature, the decreasing temperature was ~ 57.0 K and the reduction of quenching time was ~ 32 

sec due to the spacer grid model. These values would seem to be the significant amount and were the 

largest values in tests of Table 3-1. This test have the lowest reflood rate in top 5 tests of Table 3-1and 

the effect of the spacer grid model was most dominant. 

As shown in Figure 3-13, the quench front without the spacer grid model showed a good agreement 

with data at all elevations. The steep rise at high elevations might be due to the de-entrainment of 

liquid from the upper part above the heated section. If the spacer grid model was applied, the quench 

front was increasing slightly faster than that without the spacer grid model as the elevation was higher. 
15

The differential pressure (DP) for entire 12 ft was plotted in Figure 3-14 and the differential pressure 

between 11 ft and 12 ft was shown in Figure 3-15. The growth of DP for entire test section was related 

to the collapsed liquid level and the DP increased gradually with the injected water. The DP for entire 

12 feet showed a reasonable prediction with the trend of data, but it showed the slight over-estimation 

of the experimental data. It would be considered to the behaviors of the rod temperature and the 

quench front. If the spacer grid model was applied, the DP for entire 12 feet was growing slightly faster 

than that without the spacer grid model as shown in Figure 3-14. As shown well in Figure 3-15, the DP 

between 11 ft and 12 ft was increased at elevation with the spacer grid and the pressure drop in the 

spacer grid model was predicted well.

The vapor temperatures at two elevations 1.8 m and 3.0 m were shown in Figures 3.16and 3.17, 

respectively. As described in the rod temperature, the vapor temperature at higher elevation 3.0 m was 

over predicted at the initial heat up period and it would result in the earlier increase of the rod 

temperature up to its turn-over time. The calculation with the spacer grid model showed the slightly 

higher vapor temperature during the heat up and the faster quenching due to the enhancement of 

heat transfer. 

The heat transfer coefficients (HTCs) at several elevations is in Figure 3-18 through Figure 3-20. 

Usually, as the rod was quenched, the heat transfer coefficient was increased suddenly. As shown in 

Figures, the HTC was reducing or stagnant during the rod heat up and the rod during the reflood 

phase cooled by the steam cooling and/or water droplets and the HTC increased gradually. TRACE 

seems to not predict the HTC decrease during the rod heat up beside of the result at elevation 3.0 m. 

The average HTC during the reflood predicted roughly the data, but the faster rise of HTC was 

predicted with the data due to the earlier turn-over of the rod temperature. TRACE predicted a sharp 

increase of HTC much earlier than experimental data since the rod quenching occurred too early. 

When the spacer grid model was applied, the steep increase of HTC was expedited since the earlier 

rod quench occurred due to the growth of convective heat transfer.

Figure 3-3 Liquid Inlet Flowrate – Run No.31805
16

Figure 3-4 Liquid Inlet Temperature – Run No.31805

Figure 3-5 Heater Rod Temperature at 0.6 m – Run No.31805
17

Figure 3-6 Heater Rod Temperature at 1.2 m – Run No.31805

Figure 3-7 Heater Rod Temperature at 1.8 m – Run No.31805
18

Figure 3-8 Heater Rod Temperature at 1.98 m – Run No.31805

Figure 3-9 Heater Rod Temperature at 2.4 m – Run No.31805
19

Figure 3-10 Heater Rod Temperature at 3.0 m – Run No.31805

Figure 3-11 Heater Rod Temperature at 3.3 m – Run No.31805
20

Figure 3-12 Heater Rod Temperature at 3.5 m – Run No.31805

Figure 3-13 Quench Front Profile – Run No.31805
21

Figure 3-14 Differential Pressure for Entire 12 ft – Run No.31805

Figure 3-15 Differential Pressure at 10~11 ft Elevation – Run No.31805
22

Figure 3-16 Vapor Temperature at 1.8 m – Run No.31805

Figure 3-17 Vapor Temperature at 3.0 m – Run No.31805
23

Figure 3-18 Heat Transfer Coefficient at 1.98 m – Run No.31805

Figure 3-19 Heat Transfer Coefficient at 2.4 m – Run No.31805
24

Figure 3-20 Heat Transfer Coefficient at 3.0 m – Run No.31805

3.2.1.1 Run No. 31504

Run No. 31504 was a test with a flooding rate of 2.4 cm/sec at 0.28 MPa and 79 oC inlet subcooling 

temperature as shown in Table 3-1. This was the same as Run No. 31805, except for an increase of a 

flooding rate. However, as described previously, the actual input conditions changed largely with time 

and showed the big oscillations. Therefore, the inlet flowrate and temperature were considered as 

time-dependent variables as shown in Figures 3.21~ 3.22. Most initial conditions in the TRACE 

assessment manual [6] were applied in this study.

Figure 3-23 through Figure 3-30 represent the rod temperatures at various elevations. These figures 

showed well the processes such as the heat up, turn-over of rod temperature during a reflood and a 

final quenching as in Run No. 31805. In the calculation without the spacer grid model, the peak 

temperatures were fairly predicted with the experimental data at low elevations (z ≤ 2.4 m), but were

over predicted at higher elevations (3.0 m ~ 3.3 m). The maximum peak clad temperature was shown 

at elevation z=1.98 m (6.5 ft) and decreased due to relatively higher reflood rate. The quenching times 

were reasonably agreed with the data beside of higher elevations (z > 3.3 m) and it showed the better 

results by the higher reflood rate compared to Run No. 31805. When the spacer model was applied, 

the rod temperatures were reduced and the quenching time was expedited at most elevations. The 

rod temperature and the quenching time showed the similar trend with Run No. 31805. During a 

heating up, the effect of the spacer grid model did not show at lower elevations (z ≤ 2.4 m), but the rod 

temperatures at higher elevations were increased faster than the case without the spacer grid model. 

The turn-over time of rod temperature were estimated earlier than those without the spacer grid at all 

elevations since the spacer grid enhanced the increase of collapsed water level. The peak 

temperatures with the spacer grid model had the lower values at low elevations (z ≤ 3.0 m) in 

comparison with those without the spacer grid model, while the peak temperature with the spacer grid 

model was similar to that at z = 3.3 elevation as shown in Figure 3-29. In the case with the spacer grid 

model, the decreasing temperature was ~ 10.8 K and the reduction of quenching time was ~ 9 sec 

due to the spacer grid model at elevation z=1.98with the maximum peak cladding temperature.
25

As shown in Figure 3-31, the quench front without the spacer grid model showed a very good 

agreement results at all elevations and this reflected well the quenching time of rod. With the spacer 

grid model, the quench front was increasing slightly faster than that without the spacer grid model as 

the elevation was higher. 

The differential pressure (DP) for entire 12 ft was shown in Figure 3-32and the differential pressure 

between 11 ft and 12 ft was illustrated in Figure 3-33. The rise of DP for entire test section was related 

to the collapsed liquid level and the DP increased gradually with the injected water. The DP for entire 

12 feet showed the large over-prediction with the experimental data. It appeared that TRACE 

predicted more water in the test section compared to the data. If the spacer grid model was applied, 

the faster increase of DP for entire 12 feet was predicted which is similar to Run No. 31805. As shown 

well in Figure 3-33, the DP between 11 ft and 12 ft rose at elevation with the spacer grid and the 

pressure drop was predicted well. The plots after 600 sec should be ignored because the test is 

finished at ~ 600 sec.

Figures 3.34 and 3.35represented the vapor temperatures at two elevations 1.8 m and 3.0 m, 

respectively. The vapor temperature at higher elevation 3.0 m was largely over predicted at the initial 

heat up period and it would result in the earlier increase of the rod temperature during the heating up 

and the earlier turn-over time. The calculation with the spacer grid model showed the higher vapor 

temperature during the heat up and the faster quenching at high elevation.

The heat transfer coefficients (HTCs) at several elevations is shown in Figure 3-36 through Figure 3-

38. As shown in Figures, the HTC was reducing or stagnant during the rod heat up, but TRACE did 

not predict well the experimental data for the heating up. However, the HTCs were fairly predicted 

during the reflood phase and the time of steep increase of HTC was agreed well with the experimental 

data, especially at elevation z ≥ 2.4. This would correspond with the behaviors of rod temperature. 

With the spacer grid model, the earlier sharp rise of HTC was predicted due to the earlier rod quench.

Figure 3-21 Liquid Inlet Flowrate – Run No.31504
26

Figure 3-22 Liquid Inlet Temperature – Run No.31504

Figure 3-23 Heater Rod Temperature at 0.6 m – Run No.31504
27

Figure 3-24 Heater Rod Temperature at 1.2 m – Run No.31504

Figure 3-25 Heater Rod Temperature at 1.8 m – Run No.31504
28

Figure 3-26 Heater Rod Temperature at 1.98 m – Run No.31504

Figure 3-27 Heater Rod Temperature at 2.4 m – Run No.31504
29

Figure 3-28 Heater Rod Temperature at 3.0 m – Run No.31504

Figure 3-29 Heater Rod Temperature at 3.3 m – Run No.31504
30

Figure 3-30 Heater Rod Temperature at 3.5 m – Run No.31504

Figure 3-31 Quench Front Profile – Run No.31504
31

Figure 3-32 Differential Pressure for Entire 12 ft – Run No.31504

Figure 3-33 Differential Pressure at 10~11 ft Elevation – Run No.31504
32

Figure 3-34 Vapor Temperature at 1.8 m – Run No.31504

Figure 3-35 Vapor Temperature at 3.0 m – Run No.31504
33

Figure 3-36 Heat Transfer Coefficient at 1.98 m – Run No.31504

Figure 3-37 Heat Transfer Coefficient at 2.4 m – Run No.31504
34

Figure 3-38 Heat Transfer Coefficient at 3.0 m – Run No.31504

3.2.1.2 Run No. 31203

Run No. 31203 was a test with a flooding rate of 3.84 cm/sec at 0.28 MPa and 78 oC inlet subcooling 

temperature as shown in Table 3-1. This was the same as Run No. 31805, except for an increase of a 

flooding rate. However, as described previously, the actual input conditions changed largely with time 

and showed the big oscillations. Therefore, the inlet flowrate and temperature were considered as 

time-dependent variables as shown in Figures 3.39~ 3.40. Most initial conditions in the TRACE 

assessment manual [6] were applied in this study.

The rod temperatures at various elevations were plotted in Figure 3-41 through Figure 3-48. These 

figures showed well the processes such as the heat up, turn-over of rod temperature during a reflood 

and a final quenching as in Run No. 31805. Without the spacer grid model, the peak temperatures 

were predicted well with the data at low elevations (z ≤ 2.4 m), but were over predicted at higher 

elevations (z = 3.0 m ~ 3.3 m). Especially, TRACE also over predicted the peak temperatures at z = 

3.5 m, It was a different result with previous tests. The maximum peak clad temperature was shown at 

elevation z=1.98 m (6.5 ft) and decreased as the reflood rate is higher. The turn-around times were 

reasonably predicted well in all elevations. The quenching times were fairly agreed with the data at 

lower elevations (z < 1.98 m). When the spacer model was applied, the rod temperatures were 

decreased and the final quenching time was shortened at all elevations, as expected. The trends of 

rod temperature and the quenching time were similar with Run No. 31805. During a heating up, the 

effect of the spacer grid model was not significant at lower elevations (z ≤ 3.0 m), but the earlier 

increase of the rod temperatures was predicted at higher elevations in the case with the spacer grid 

model. The turn-over time of rod temperature with the spacer grid model were expedited at all 

elevations since the spacer grid enhanced the convective heat transfer. The peak temperatures with 

the spacer grid model were lower than those without the spacer grid model at most elevations. In the 

case with the spacer grid model, the decreasing temperature was ~ 5.9 K and the reduction of 

quenching time was ~ 4 sec due to the spacer grid model at elevation z=1.98 with the maximum peak 

cladding temperature.
35

As shown in Figure 3-49, the quench front without the spacer grid model was predicted well for the 

lower 80% elevations. In this test, the data showed a top quench which was probably from the liquid 

de-entrainment above the active core, but TRACE did not predict this top quenching. The quench front 

with the spacer grid model was rising faster than that without the model as the elevation was higher.

The differential pressure (DP) for entire 12 ft was plotted in Figure 3-50 and the differential pressure 

between 11 ft and 12 ft was shown in Figure 3-51. The increase of DP for entire test section was 

correspond with the collapsed liquid level and the DP increased with the injected water. The DP for 

entire 12 feet showed the large over-prediction with the experimental data. TRACE predicted more 

water inventory in the test section compared to the data. If the spacer grid model was applied, the DP 

for entire 12 feet was increased at slightly earlier time. As shown well in Figure 3-51, the DP between 

11 ft and 12 ft rose at elevation with the spacer grid and the large pressure drop was predicted. 

Figures 3.52 and 3.53showed the vapor temperatures at two elevations 1.8 m and 3.0 m, 

respectively. The vapor temperature at higher elevation 3.0 m was over estimated at the initial heating 

up and it would result in the earlier increase of the rod temperature and the earlier turn-around time. In 

the case with the spacer grid model, the higher vapor temperature during the heat up and the faster 

quenching at high elevation were predicted.

Figure 3-54 through Figure 3-56 showed the heat transfer coefficients (HTCs) at several elevations. 

The HTC was reducing or stagnant during the rod heat up, but TRACE under predicted the 

experimental data for the heating up. During the reflood phase, the HTCs were over predicted and 

increased faster than the experimental data and the time of increase of HTC was expedited at 

elevations z ≥ 2.4. If the spacer grid model was applied, the earlier sharp rise of HTC was predicted 

according to the earlier quenching.

Figure 3-39 Liquid Inlet Flowrate – Run No.31203
36

Figure 3-40 Liquid Inlet Temperature – Run No.31203

Figure 3-41 Heater Rod Temperature at 0.6 m – Run No.31203
37

Figure 3-42 Heater Rod Temperature at 1.2 m – Run No.31203

Figure 3-43 Heater Rod Temperature at 1.8 m – Run No.31203
38

Figure 3-44 Heater Rod Temperature at 1.98 m – Run No.31203

Figure 3-45 Heater Rod Temperature at 2.4 m – Run No.31203
39

Figure 3-46 Heater Rod Temperature at 3.0 m – Run No.31203

Figure 3-47 Heater Rod Temperature at 3.3 m – Run No.31203
40

Figure 3-48 Heater Rod Temperature at 3.5 m – Run No.31203

Figure 3-49 Quench Front Profile – Run No.31203
41

Figure 3-50 Differential Pressure for Entire 12 ft – Run No.31203

Figure 3-51 Differential Pressure at 10~11 ft Elevation – Run No.31203
42

Figure 3-52 Vapor Temperature at 1.8 m – Run No.31203

Figure 3-53 Vapor Temperature at 3.0 m – Run No.31203
43

Figure 3-54 Heat Transfer Coefficient at 1.98 m – Run No.31203

Figure 3-55 Heat Transfer Coefficient at 2.4 m – Run No.31203
44

Figure 3-56 Heat Transfer Coefficient at 3.0 m – Run No.31203

3.2.1.3 Run No. 31302

Run No. 31302 was a test with a flooding rate of 7.65 cm/sec at 0.28 MPa and 78 oC inlet subcooling 

temperature as shown in Table 3-1. This was the same as Run No. 31805 except for increasing the 

reflood rate. However, as described previously, the actual input conditions varied largely with time and 

showed the big oscillations. So, the inlet flowrate and temperature were considered as timedependent variables as shown in Figures 3.57 ~ 3.58. Most initial conditions in the TRACE 

assessment manual [6] were applied in this study.

The rod temperatures at various elevations were shown in Figures 3.59 ~ 3.66. These figures showed 

fairly the reflood process that the rod was heating up by the initiation of reflood, turn-over of rod 

temperature and finally quenching. In the case without the spacer grid model, the predicted peak 

temperatures under estimated at elevation z = 0.6 m with low power density. This test had three times 

the reflood rate in Run No. 31504, that resulted in the low heat up at z = 0.6 m. However, the peak 

temperatures were over predicted with the data above elevation z = 1.8 m. Those predictions were 

reflected in the relatively high vapor temperature as shown in Figures 3.70 ~ 3.71. The maximum peak 

clad temperature was predicted at elevation z = 1.98 m (6.5 ft). At elevations (z ≤ 3.0 m), the 

quenching times were reduced as compared to the experimental data, while it were predicted at the 

later time at elevation z > 3.0 since TRACE did not predict the strong top-down quenching was 

observed in the experiment. In the case with a spacer grid model, the effect of the spacer grid model 

during a heat up period was not shown at all elevations, but the earlier turn-around of rod temperature 

was predicted since the spacer grid enhanced the increase of collapsed liquid level. The peak 

temperatures with the spacer grid model had the lower values at all elevations in comparison with 

those without the spacer grid model. The maximum peak clad temperature at z=1.98 m was reduced 

in case with the spacer grid model. At elevation z=1.98with the maximum peak cladding temperature, 

the decreasing temperature was ~ 2.4 K and the reduction of quenching time was ~ 4 sec due to the 

spacer grid model.
45

As shown in Figure 3-67, the quench front without the spacer grid model was predicted well with the 

data up to elevation z = 2.0 m. In this test, TRACE did not predict the top quenching at higher 

elevations which was from the liquid de-entrainment from the upper part above the active core. If the 

spacer grid model was applied, the quench front was increasing faster than that without the spacer 

grid model as the elevation was higher. 

The differential pressure (DP) for entire 12 ft was plotted in Figure 3-68 and the differential pressure 

between 11 ft and 12 ft was shown in Figure 3-69. The growth of DP for entire test section was 

correspond to the injected water. The DP for entire 12 ft showed a good prediction with the trend of 

data in spite of the difference of quenching behaviors at higher elevations (z ≥ 3.0 m). It would result 

from the top down quenching of test. If the spacer grid model was applied, the DP for entire 12 ft was 

growing a little faster than that without the spacer grid model as shown in Figure 3-68. As shown well 

in Figure 3-69, the DP between 11 ft and 12 ft was increased largely due to the spacer grid.

The vapor temperatures at two elevations 1.8 m and 3.0 m were shown in Figures 3.70and 3.71, 

respectively. As described in the rod temperature, the vapor temperatures at two elevations was over 

predicted during all reflood processes. The calculation with the spacer grid model had the slightly 

higher vapor temperature during the heat up and the faster quenching behaviors. 

The heat transfer coefficients (HTCs) at several elevations is in Figure 3-72 through Figure 3-74. As 

shown in Figures, the HTC was under predicted during the rod heat up and increased faster than the 

experimental data. TRACE predicted a sharp increase of HTC much earlier than experimental data 

since the rod quenching occurred too early. When the spacer grid model was applied, the steep rise of 

HTC was expedited because of the earlier quenching of rod. 
46

Figure 3-57 Liquid Inlet Flowrate – Run No.31302

Figure 3-58 Liquid Inlet Temperature – Run No.31302
47

Figure 3-59 Heater Rod Temperature at 0.6 m – Run No.31302

Figure 3-60 Heater Rod Temperature at 1.2 m – Run No.31302
48

Figure 3-61 Heater Rod Temperature at 1.8 m – Run No.31302

Figure 3-62 Heater Rod Temperature at 1.98 m – Run No.31302
49

Figure 3-63 Heater Rod Temperature at 2.4 m – Run No.31302

Figure 3-64 Heater Rod Temperature at 3.0 m – Run No.31302
50

Figure 3-65 Heater Rod Temperature at 3.3 m – Run No.31302

Figure 3-66 Heater Rod Temperature at 3.5 m – Run No.31302
51

Figure 3-67 Quench Front Profile – Run No.31302

Figure 3-68 Differential Pressure for Entire 12 ft – Run No.31302
52

Figure 3-69 Differential Pressure at 10~11 ft Elevation – Run No.31302

Figure 3-70 Vapor Temperature at 1.8 m – Run No.31302
53

Figure 3-71 Vapor Temperature at 3.0 m – Run No.31302

Figure 3-72 Heat Transfer Coefficient at 1.98 m – Run No.31302
54

Figure 3-73 Heat Transfer Coefficient at 2.4 m – Run No.31302

Figure 3-74 Heat Transfer Coefficient at 3.0 m – Run No.31302

3.2.1.4 Run No. 31701

Run No. 31701 was a test with a flooding rate of 15.5 cm/sec at 0.28MPa and 77 oC inlet subcooling 

temperature as shown in Table 3-1. This was the same conditions as Run No. 31805, except for 

increasing of the flooding rate. Compared to the previous 4 tests, this had the largest reflooding rate.
55

As described before, the actual input conditions changed largely with time and showed the big 

oscillations. Therefore, the inlet flowrate and temperature were considered as time-dependent 

variables as shown in Figures 3.75~ 3.76. Most initial conditions in the TRACE assessment manual 

[6] were applied in this study.

Figure 3-77 through Figure 3-84 represented the rod temperatures at various elevations. These 

figures also showed well the reflood processes such as the heat up, turnover of rod temperature and a 

final quenching. In the calculation without the spacer grid model, the peak temperatures were under 

predicted with the experimental data at low elevations (z ≤ 1.2 m) and were over predicted at higher 

elevations (z ≥ 1.98 m). Usually, for high flooding rates above 15 cm/sec (6 in/sec), the dominant flow 

regime would be an inverted annular regime in bundles and then the heat transfer was very high and 

the turn-over of cladding temperature occurred immediately. As shown in Figure 3-85, the earlier 

quenching from the middle elevation (z ~ 1.0 m) was observed in the experimental data as compared 

to TRACE results. This might come from the under predictions for the lumps of liquid from the lower 

elevations and the de-entrained liquid from the upper parts. It would result in the over prediction of rod 

temperature at higher elevations. Those predictions were also reflected as the relatively high vapor 

temperature in Figure 3-88and Figure 3-89. The maximum peak clad temperature was shown at 

elevation z=1.98 m (6.5 ft) and had the lowest value due to higher reflood rate as compared to above 

4 tests. The quenching times were over-estimated at most elevations due to the under prediction for 

the top-down quenching in test. The top-down quenching started to be observed from Run No. 31203 

and became more dominant as the reflood rate is higher. When the spacer model was applied, the rod 

temperatures were reduced and the quenching time was expedited at most elevations. However, 

because of the highest reflood rate, the differences of rod temperature and quenching time were the 

smallest as compared to 4 tests above. During a heating up, the effect of the spacer grid model was 

not shown at all elevations. The turn-over times of rod temperature were almost same as those 

without the spacer grid at all elevations due to the high reflood rate. The quenching time was slightly 

decreased because of the relatively faster rise of liquid level. In the case with the spacer grid model, 

the decreasing temperature was ~ 2.5 K and the reduction of quenching time was ~ 2 sec due to the 

spacer grid model at elevation z=1.98with the maximum peak cladding temperature.

The quench front profile was shown in Figure 3-85. The quench front without the spacer grid model 

was predicted well up to elevation z ~ 1.0 m. At elevations z ≥ 1.0 m, TRACE would under-predict the 

liquid chunks from lower part and/or the de-entrained liquid from upper part of the active core and this 

might result in the significantly delayed quench front in higher reflood rates. When the spacer grid 

model was used, the increase of quench front was almost identical up to elevation z = 2.0 m and the 

earlier rise of that was predicted at elevations z ≥ 2.0 m. 

The differential pressure (DP) for entire 12 ft was shown in Figure 3-86and the differential pressure 

between 11 ft and 12 ft was illustrated in Figure 3-87. The rise of DP for entire test section 

corresponded to the collapsed liquid level and the DP increased gradually with the injected water. The 

DP for entire 12 ft agreed fairly with the data until ~ 100 sec, but slightly under predicted after ~ 100 

sec. The DP between 11 ft and 12 ft under predicted for all times of test. It would result from the under 

prediction of the liquid chunks from upper parts above the active core. If the spacer grid model was 

applied, the faster increase of DP for entire 12 feet was predicted and the DP between 11 ft and 12 ft 

reflected well the larger pressure drop of the spacer grid model. 

Figures 3.88 and 3.89represented the vapor temperatures at two elevations 1.8 m and 3.0 m, 

respectively. As explained in the rod temperature, the vapor temperatures at two elevations was over 

predicted during all times of test. As shown in other tests, the spacer grid model resulted in the higher 

vapor temperature during the heat up and the faster quenching at high elevation, but the effect of the 

spacer grid model would be less, as the reflood rate is higher. 
56

The heat transfer coefficients (HTCs) at several elevations is shown in Figure 3-90 through Figure 3-

92. During the heating up, the reduction or stagnant for HTC was not observed in the data due to the 

high reflood rate and then the HTC was under predicted for this region. According to the over 

prediction of rod temperature at higher elevations (z ≥ 1.98 m), the under prediction of HTC was also 

shown continuously during the reflood phase. TRACE predicted a sharp increase of HTC at the 

delayed time as compared to experimental data due to the late rod quenching. In the case with the 

spacer grid model, the relatively higher HTC during the reflood phase and the earlier sharp increase of 

HTC were predicted. 

Figure 3-75 Liquid Inlet Flowrate – Run No.31701
57

Figure 3-76 Liquid Inlet Temperature – Run No.31701

Figure 3-77 Heater Rod Temperature at 0.6 m – Run No.31701
58

Figure 3-78 Heater Rod Temperature at 1.2 m – Run No.31701

Figure 3-79 Heater Rod Temperature at 1.8 m – Run No.31701
59

Figure 3-80 Heater Rod Temperature at 1.98 m – Run No.31701

Figure 3-81 Heater Rod Temperature at 2.4 m – Run No.31701
60

Figure 3-82 Heater Rod Temperature at 3.0 m – Run No.31701

Figure 3-83 Heater Rod Temperature at 3.3 m – Run No.31701
61

Figure 3-84 Heater Rod Temperature at 3.5 m – Run No.31701

Figure 3-85 Quench Front Profile – Run No.31701
62

Figure 3-86 Differential Pressure for Entire 12 ft – Run No.31701

Figure 3-87 Differential Pressure at 10~11 ft Elevation – Run No.31701
63

Figure 3-88 Vapor Temperature at 1.8 m – Run No.31701

Figure 3-89 Vapor Temperature at 3.0 m – Run No.31701
64

Figure 3-90 Heat Transfer Coefficient at 1.98 m – Run No.31701

Figure 3-91 Heat Transfer Coefficient at 2.4 m – Run No.31701
65

Figure 3-92 Heat Transfer Coefficient at 3.0 m – Run No.31701

3.2.1.5 Run No. 31108

Run No. 31108 was a test with a flooding rate of 7.90 cm/sec at 0.13 MPa and 74 oC inlet subcooling 

temperature as shown in Table 3-1. This was similar to Run No. 31302 except for decreasing the 

pressure for the upper plenum. However, as described previously, the actual input conditions varied 

largely with time and showed the big oscillations. So, the inlet flowrate and temperature were 

considered as time-dependent variables as shown in Figure 3-93 and 3.94, respectively. Most initial 

conditions in the TRACE assessment manual [6] were applied in this study.

The rod temperatures at various elevations were plotted in Figures 3.95 ~ 3.102. These figures 

showed fairly the reflood processes including the initiating heating up, the turn-over of rod temperature 

and finally quenching. When the spacer grid model was not applied, the predicted peak temperatures 

were reasonably agreed with the data at elevation z ≤ 1.8 m. The peak temperatures were over 

predicted with the data at elevations z ≥ 1.8 m. This over prediction of rod temperature corresponded 

to the relatively high vapor temperature as shown in Figures 3.106~ 3.107. The maximum peak clad 

temperature was showed at elevation z = 1.98 m (6.5 ft). At elevations (z ≤ 1.98 m), TRACE predicted 

fairly the quenching time, but the quenching times became shorten up to elevation z = 3.0 m. At 

elevation z = 3.0 m above, the quenching were predicted at the later time since the strong top-down 

quenching was observed in the experiment as shown in Figure 3-103. This test had the half of the 

upper plenum pressure of Run No. 31302 and the similar trends for the rod temperatures were shown 

with Run No. 31302. The rod temperatures at all elevations were slightly higher than those of Run No. 

31302 since the saturation temperature would be relatively low due to the lower pressure, but the 

difference of the rod temperatures would be not significant because of the high reflood rate. When the 

spacer grid model was applied, the differences of the rod temperatures was not shown well at all 

elevations during the heat-up phase, but the earlier turn-around of rod temperature was predicted 

since the spacer grid would enhance the heat transfer. The peak temperatures with the spacer grid 

model had the lower values at all elevations. The maximum peak clad temperature at z=1.98 m was 
66

decreased in case with the spacer grid model. At elevation z=1.98 with the maximum peak cladding 

temperature, the temperature reduction was ~ 16.3 K and the quenching time was decreased with ~ 8

sec due to the spacer grid model.

As shown in Figure 3-103, the quench front without the spacer grid model was predicted well with the 

data up to elevation z = 1.8 m. In this test, TRACE did not predict the top quenching at higher 

elevations (z ≥ 3.0 m) which was from the liquid de-entrainment from the upper part above the active 

core. The transition of quench front for the top quench could be identified in the experiment at around z 

= 3.0 m. If the spacer grid model was applied, the quench front was increasing faster than that without 

the spacer grid model as the elevation was higher. 

The differential pressure (DP) for entire 12 ft was plotted in Figure 3-104 and the differential pressure 

between 11 ft and 12 ft was shown in Figure 3-105. The growth of DP for entire test section was 

correspond to the collapsed water level. The DP for entire 12 ft was over predicted at all times of test 

despite of the top down quenching in test. TRACE predicted more water inventory in the test section. If 

the spacer grid model was applied, the DP for entire 12 ft was rising slightly earlier than that without 

the spacer grid model. As shown well in Figure 3-105, the DP between 11 ft and 12 ft was increasing 

due to the spacer grid.

The vapor temperatures at two elevations 1.8 m and 3.0 m were shown in Figures 3.106 and 3.107, 

respectively. As described in the rod temperature, the vapor temperatures at two elevations was over 

predicted during testing. The calculation with the spacer grid model had the slightly higher vapor 

temperature during the heat up and the quick quenching behaviors. 

The heat transfer coefficients (HTCs) at several elevations is in Figure 3-108through Figure 3-110. As 

shown in Figures, the HTC was under predicted during the rod heat up and increased faster than the 

experimental data. TRACE predicted a steep increase of HTC earlier than experimental data because 

of the earlier prediction of the rod quenching. When the spacer grid model was applied, the steep rise 

of HTC was expedited because of the earlier quenching of rod. 
67

Figure 3-93 Liquid Inlet Flowrate – Run No.31108

Figure 3-94 Liquid Inlet Temperature – Run No.31108
68

Figure 3-95 Heater Rod Temperature at 0.6 m – Run No.31108

Figure 3-96 Heater Rod Temperature at 1.2 m – Run No.31108
69

Figure 3-97 Heater Rod Temperature at 1.8 m – Run No.31108

Figure 3-98 Heater Rod Temperature at 1.98 m – Run No.31108
70

Figure 3-99 Heater Rod Temperature at 2.4 m – Run No.31108

Figure 3-100 Heater Rod Temperature at 3.0 m – Run No.31108
71

Figure 3-101 Heater Rod Temperature at 3.3 m – Run No.31108

Figure 3-102 Heater Rod Temperature at 3.5 m – Run No.31108
72

Figure 3-103 Quench Front Profile – Run No.31108

Figure 3-104 Differential Pressure for Entire 12 ft – Run No.31108
73

Figure 3-105 Differential Pressure at 10~11 ft Elevation – Run No.31108

Figure 3-106 Vapor Temperature at 1.8 m – Run No.31108
74

Figure 3-107 Vapor Temperature at 3.0 m – Run No.31108

Figure 3-108 Heat Transfer Coefficient at 1.98 m – Run No.31108
75

Figure 3-109 Heat Transfer Coefficient at 2.4 m – Run No.31108

Figure 3-110 Heat Transfer Coefficient at 3.0 m – Run No.31108

3.2.1.6 Run No. 32013

Run No. 32013 was a test with a flooding rate of 2.64 cm/sec at 0.41 MPa and 79 oC inlet subcooling 

temperature as shown in Table 3-1. This was the same conditions as Run No. 31504, except for one 
76

and a half times the pressure of the upper plenum. As described before, the actual input conditions 

changed largely with time and showed the big oscillations. Therefore, the inlet flowrate and 

temperature were considered as time-dependent variables as shown in Figures 3.111 ~ 3.112. Most 

initial conditions in the TRACE assessment manual [6] were applied in this study.

Figure 3-113 through Figure 3-120 plotted the rod temperatures at various elevations. These figures 

showed well the processes such as the heat up, turn-over of rod temperature during a reflood and a 

final quenching. In the case without the spacer grid model, the peak temperatures were reasonably 

predicted with the data at low elevations (z ≤ 2.4 m), but were over predicted at higher elevations (3.0 

m ~ 3.3 m). The maximum peak clad temperature was shown at elevation z=1.98 m (6.5 ft). The 

quenching times were fairly agreed with the data beside of higher elevations (z > 3.3 m). This test had 

1.5 times the upper plenum pressure of Run No. 31504 and showed a similar trend with Run No. 

31504. The lower rod temperatures and the earlier rod quench at all elevations were predicted as 

compared to Run No. 31504 since the saturation temperature would be relatively high due to the 

higher pressure. When the spacer model was applied, the rod temperatures were reduced and the 

quenching time was expedited at most elevations. During a heating up, the effect of the spacer grid 

model did not show at lower elevations (z ≤ 2.4 m), but the rod temperatures at higher elevations were 

increased faster than the case without the spacer grid model. The turn-over time of rod temperature 

were predicted earlier than those without the spacer grid at all elevations since the spacer grid 

enhanced the increase of collapsed water level. The peak temperatures with the spacer grid model 

had the lower values at low elevations (z ≤ 3.0 m) in comparison with those without the spacer grid 

model, while the peak temperature with the spacer grid model was a little higher than that at z = 3.3 

elevation as shown in Figure 3-119. This might be resulted from the earlier rise of rod temperature 

during a heat up region. In the case with the spacer grid model, the decreasing temperature was ~ 2.7

K and the reduction of quenching time was ~ 5 sec due to the spacer grid model at elevation z=1.98

with the maximum peak cladding temperature.

As shown in Figure 3-121, the quench front without the spacer grid model showed the slightly under 

prediction results except for the upper region at all elevations, but it was reasonable. With the spacer 

grid model, the quench front was increasing slightly faster than that without the spacer grid model as 

the elevation was higher. 

The differential pressure (DP) for entire 12 ft was shown in Figure 3-122 and the differential pressure 

between 11 ft and 12 ft was illustrated in Figure 3-123. The rise of DP for entire test section was 

related to the collapsed liquid level and the DP increased gradually with the injected water. The DP for 

entire 12 ft agreed fairly with the data and the DP between 11 ft and 12 ft was slightly over predicted 

for all times of test. If the spacer grid model was applied, the faster increase of DP for entire 12 feet 

was predicted which is similar to Run No. 31504. As shown well in Figure 124, the DP between 11 ft 

and 12 ft rose at elevation with the spacer grid and the pressure drop was predicted well.

Figures 3.124 and 3.125 represented the vapor temperatures at two elevations 1.8 m and 3.0 m, 

respectively. The vapor temperature at higher elevation 3.0 m was predicted well at the initial heat up 

period as compared to Run No. 31504 and it would be related to the lower rod temperature and the 

earlier quench time. The calculation with the spacer grid model showed the higher vapor temperature 

during the heat up and the faster quenching at high elevation. 

The heat transfer coefficients (HTCs) at several elevations is shown in Figure 3-126through Figure 3-

128. As shown in Figures, the HTC was reducing or stagnant during the rod heat up, but TRACE did 

not predict well the experimental data for the heating up. However, the HTCs were fairly predicted 

during the reflood phase and the time of steep increase of HTC was agreed reasonably with the 

experimental data, especially at elevation z = 2.4. This would correspond with the behaviors of rod 
77

temperature. With the spacer grid model, the earlier sharp rise of HTC was predicted due to the earlier 

rod quench.

Figure 3-111 Liquid Inlet Flowrate – Run No.32013

Figure 3-112 Liquid Inlet Temperature – Run No.32013
78

Figure 3-113 Heater Rod Temperature at 0.6 m – Run No.32013

Figure 3-114 Heater Rod Temperature at 1.2 m – Run No.32013
79

Figure 3-115 Heater Rod Temperature at 1.8 m – Run No.32013

Figure 3-116 Heater Rod Temperature at 1.98 m – Run No.32013
80

Figure 3-117 Heater Rod Temperature at 2.4 m – Run No.32013

Figure 3-118 Heater Rod Temperature at 3.0 m – Run No.32013
81

Figure 3-119 Heater Rod Temperature at 3.3 m – Run No.32013

Figure 3-120 Heater Rod Temperature at 3.5 m – Run No.32013
82

Figure 3-121 Quench Front Profile – Run No.32013

Figure 3-122 Differential Pressure for Entire 12 ft – Run No.32013
83

Figure 3-123 Differential Pressure at 10~11 ft Elevation – Run No.32013

Figure 3-124 Vapor Temperature at 1.8 m – Run No.32013
84

Figure 3-125 Vapor Temperature at 3.0 m – Run No.32013

Figure 3-126 Heat Transfer Coefficient at 1.98 m – Run No.31701
85

Figure 3-127 Heat Transfer Coefficient at 2.4 m – Run No.31701

Figure 3-128 Heat Transfer Coefficient at 3.0 m – Run No.31701

3.2.1.7 Run No. 32114

Run No. 32114 was a test with a flooding rate of 2.5 ~ 3.1 cm/sec at 0.28 MPa and 5 oC inlet 

subcooling temperature as shown in Table 3-1. This was similar to Run No. 31504, except for a large 
86

decrease of a subcooling degree. There was only test which used the low subcooling rate in Table 3-1. 

As described previously, the actual input conditions changed largely with time and showed the big 

oscillations. Therefore, the inlet flowrate and temperature were considered as time-dependent 

variables as shown in Figures 3.129 ~ 3.130. Most initial conditions in the TRACE assessment manual 

[6] were applied in this study.

The rod temperatures at various elevations were shown in Figures 3.131 ~ 3.138. These figures 

showed well the processes such as the heat up, turn-over of rod temperature during a reflood and a 

final quenching as in other tests. In the calculation without the spacer grid model, the peak 

temperatures were fairly predicted with the experimental data up to middle elevations (z ≤ 1.98 m) and 

the calculated temperatures were located within the data spread range. However, the rod 

temperatures were over predicted at higher elevations (2.4 m ~ 3.3 m). The maximum peak clad 

temperature was shown at elevation z=1.98 m (6.5 ft) and increased largely due to a relatively higher 

inlet temperature as compared to Run No. 31504. Those predictions could be also induced well in 

Figures 3.142 ~ 3.143 with the relatively high vapor temperature. The quenching times were 

reasonably agreed with the data up to middle elevations (z ≤ 1.8 m), but they were over predicted at 

higher elevations (z ≥ 1.98 m). The quenching times were delayed due to the higher inlet temperature 

than Run No. 31805. When the spacer model was used, the rod temperatures were reduced and the 

quenching time was expedited at most elevations. During a heating up, the effect of the spacer grid 

model was not shown at all times of test. The turn-over time of rod temperature were estimated earlier 

than those without the spacer grid at all elevations since the spacer grid model enhanced the heat 

transfer. In the case with the spacer grid model, the decreasing temperature was ~ 20.2 K and the 

reduction of quenching time was ~ 27 sec due to the spacer grid model at elevation z=1.98 with the 

maximum peak cladding temperature.

As shown in Figure 3-139, the quench front without the spacer grid model showed a very good 

agreement results up to elevation z ~ 1.5 m, but it started deviating from the data above this elevation 

and over predicted due to the delayed prediction of rod quenching. With the spacer grid model, the 

quench front was increasing slightly faster than that without the spacer grid model as the elevation 

was higher. 

The differential pressure (DP) for entire 12 ft was plotted in Figure 3-140 and the differential pressure 

between 11 ft and 12 ft was shown in Figure 3-141. The DP for entire test section was increased 

gradually with the injected water. The DP for entire 12 ft showed the over-prediction with the 

experimental data and TRACE predicted more water inventory in the test section compared to the 

data. The large oscillations of the DP, as in vapor temperatures, might be estimated due to the 

oscillation of upper plenum pressures. The slightly faster increase of DP for entire 12 feet was 

predicted as the spacer grid model was applied. As shown well in Figure 3-141, the DP between 11 ft 

and 12 ft was increased at elevation with the spacer grid and the pressure drop was predicted well. 

The vapor temperatures were shown at two elevations 1.8 m and 3.0 m in Figures 3.142 and 3.143, 

respectively. The vapor temperature at higher elevation 3.0 m was largely over predicted during all 

times of test period and it would be related to the higher the rod temperature and the earlier turn-over 

time. The calculation with the spacer grid model showed well the faster quenching at high elevation.

The heat transfer coefficients (HTCs) at several elevations is illustrated in Figure 3-144through Figure 

3-146. As shown in Figures, TRACE under predicted the experimental data for the heating up, but the 

HTCs were fairly predicted during the reflood phase. The time of steep increase of HTC was delayed 

as the late prediction of rod quench in comparison with the experimental data at elevations z ≥ 1.98. 

When the spacer grid model was applied, the earlier sharp rise of HTC was predicted due to the 

earlier rod quench.
87

Figure 3-129 Liquid Inlet Flowrate – Run No.32114

Figure 3-130 Liquid Inlet Temperature – Run No.32114
88

Figure 3-131 Heater Rod Temperature at 0.6 m – Run No.32114

Figure 3-132 Heater Rod Temperature at 1.2 m – Run No.32114
89

Figure 3-133 Heater Rod Temperature at 1.8 m – Run No.32114

Figure 3-134 Heater Rod Temperature at 1.98 m – Run No.32114
90

Figure 3-135 Heater Rod Temperature at 2.4 m – Run No.32114

Figure 3-136 Heater Rod Temperature at 3.0 m – Run No.32114
91

Figure 3-137 Heater Rod Temperature at 3.3 m – Run No.32114

Figure 3-138 Heater Rod Temperature at 3.5 m – Run No.32114
92

Figure 3-139 Quench Front Profile – Run No.32114

Figure 3-140 Differential Pressure for Entire 12 ft – Run No.32114
93

Figure 3-141 Differential Pressure at 10~11 ft Elevation – Run No.32114

Figure 3-142 Vapor Temperature at 1.8 m – Run No.32114
94

Figure 3-143 Vapor Temperature at 3.0 m – Run No.32114

Figure 3-144 Heat Transfer Coefficient at 1.98 m – Run No.32114
95

Figure 3-145 Heat Transfer Coefficient at 2.4 m – Run No.32114

Figure 3-146 Heat Transfer Coefficient at 3.0 m – Run No.32114
96

3.2.1.8 Summary

In general, the spacer grid would enhance the convective heat transfer due to the flow acceleration 

and the turbulence increase downstream of grid. Currently, among 4 sub-models, the droplet breakup 

model is not activated and the grid re-wetting model is not fully implemented in TRACE. Therefore, the 

enhancement of convective heat transfer and the pressure drop due to the spacer grid could be only 

evaluated in this study. 

As would be expected, the rod temperatures were decreased and the rods were quenched at the 

earlier time in most tests as the spacer grid model was used. From upper 5 tests of Table 3-2 that the 

reflood rate was only changed, it could be identified that the lower of the peak rod temperature and the 

earlier quenching time were predicted as the reflood rate was increased. For high reflood rate as Run 

No. 31701, the inlet flow is bigger than the quenching rate of rods and the heat transfer became very 

high. In tests with high reflood rate, the change of peak temperature due to the spacer grid was not 

large, which resulted from the short heat up period and the faster increase of liquid level by the high 

reflooding rate. 

When the test pressure was lower, the higher rod temperature and the later rod quench were 

predicted since the liquid approached faster to the relatively lower saturation temperature. The use of 

spacer grid model predicted relatively larger differences for rod temperature, not for quenching time at 

lower pressure condition. Also, as the subcooling degree was higher, the decrease amount of 

quenching time due to the spacer grid was reduced since the high subcooling degree could enhance 

the heat transfer rate. 

In all tests, Run No. 31805 with the lowest reflood rate was affected most significantly for the peak 

temperature and the quenching time by using the spacer grid model as shown in Table 3-2 and 

Figures 3.147 ~ 3.148.
97

Table 3-2 Peak Temperature and Quenching Time at Elevation z = 1.98

Run No.

Peak Temperature (K)

(at time)

Quenching Time (sec)

W/O grid With grid ∆ Temp W/O grid With grid ∆ Time

31805

1440.9

(113 sec)

1383.9

(65 sec)

57.0 379 347 32

31504

1371.4

(80 sec)

1360.6

(77 sec)

10.8 330 321 9

31203

1270.6

(57 sec)

1264.7

(37 sec)

5.9 252 248 4

31302

1196.4

(24 sec)

1194.0

(19 sec)

2.4 156 152 4

31701

1153.5

(7 sec)

1151.0

(7 sec)

2.5 85 83 2

31108

1202.8

(50 sec)

1186.5

(25 sec)

16.3 200 192 8

32013

1359.2

(60 sec)

1356.5

(58 sec)

2.7 273 268 5

32114

1406.6

(99 sec)

1386.4

(85 sec)

20.2 461 434 27
98

Figure 3-147 Variation of Peak Rod Temperature at Elevation z = 1.98

Figure 3-148 Variation of Quenching Time at Elevation z = 1.98
99

4 SENSITIVITY ANALYSIS

4.1 Effect of the Grid Location

From the input manual [2], the spacer grid input for the core region in the PIPE, CHAN, or VESSEL 

component would be internally considered in the core fuel rod HTSTR models (i.e., enhanced wall 

heat transfer due to the spacer grid) and in the fluid solution (i.e., additional additive flow loss due to 

the spacer grid). The HTSTR component could have only one spacer grid located within a given 

coarse mesh axial level, and then the renoding of a fuel HTSTR component might be required if the 

core region included a number of spacer grids. 

Therefore, the selection of the proper grid location would be very significant in modeling the nodes for 

separate effect tests and actual plant analyses. Eight spacer grids were installed along 3.6 m heated 

length in FLECHT-SEASET. In chapter 2, each grid was modeled to locate in the bottom of every 

other node, and the first grid was located at the bottom of second node, which was just under the 

heated section. In this study, in orderto perform the sensitivity study for the grid location, the locations 

of spacer grids would change to the top of every other node as shown in Figure 4-1. The first grid was 

located at the top of first node of test section, and then the lengths of first and last nodes were 

changed. 

- Case-1 : Each grid was located in the bottom of every other node 

- Case-2 : Each grid was located in the top of every other node

Among 8 tests in Table 3-1, Run No. 31805 and 31302 were selected for the sensitivity calculations for 

the grid location in low and high reflood test, respectively. 

4.1.1 Run No. 31805

Run No. 31805 was a test with a flooding rate of 2.1 cm/sec at 0.28 MPa and 79 oC inlet subcooling 

temperature. This is a test with the lowest reflood rate in Table 3-1, and it could be identified that the 

effect of the spacer grid model was most dominant as described in previous chapter. 

When the locations of spacer grids were varied (Case-1 Case-2), the rod temperatures at various 

elevations were shown in Figure 4-2 through Figure 4-5. The effect of the grid locations would not be 

shown at lower elevation (z < 1.2 m) with low power density. However, the rod temperature of Case-2 

had higher values at elevation z = 1.98 with the maximum peak cladding temperature as compared to 

Case-1. At elevation z=1.98, the temperature reduction for Case-2 was ~ 22.9 K, and the quenching 

time was decreased with ~ 13 sec in comparison to the case without the spacer grid model as shown 

in Table 4-1. It would be more reasonable to predict the experimental data and showed more 

conservative results. In TRACE, the heat transfer enhancement effects by the spacer grid were 

integrated over the downstream axial cells for 50 hydraulic diameters to consider the exponential 

decay of the enhancement downstream of a spacer grid [3]. The spacer grid enhancement effects for 

a given cell as well as design factors are multiplied by the original heat transfer coefficient predicted by 

TRACE for the given cell. Therefore, the HTSTR component had the effect of only one spacer grid 

located within a given coarse mesh. When the spacer grid was located at the top of node, the original 

heat transfer coefficient could be evaluated by the heat flux upstream of a spacer grid. The low heat 

flux upstream of a spacer grid woulddecrease the enhancement of heat transfer, and then it resulted 

in the relatively high rod temperature up to the elevation with the peak power. The delayed quenching 

was predicted due to high rod temperatures for Case-2. However, the difference for the rod 

temperatures was reduced, and the quenching time was just a little changed for the grid locations at 

higher elevations (z > 2.4 m). The effect of the grid location could depend on the power shape, the 
100

reflood rate, and the node modeling, and then the detailed sensitivity studies was required for various 

tests. 

As shown in Figure 4-6, the quench front for Case-2 was increased rather late, but its difference was 

not big. Because the number of spacer grids was not changed, the pressure losses due to the spacer 

grids would be unchanged. The relatively high rod temperature could delay the rise of the quench 

front. 

The heat transfer coefficients (HTCs) at elevation z = 1.98 m is presented in Figure 4-7. The average 

HTC for Case-2 during the reflood was under-predicted in comparison with Case-1. Those predictions 

were reflected in the relatively high rod temperatures. The steep increase of HTC for Case-2 was 

delayed since the rod was quenched late due to the high rod temperatures.

Table 4-1 Peak Temperature and Quenching Time at Elevation z = 1.98 for Run No. 31805

Case

Peak Temperature (K)

(at time)

Quenching Time (sec)

W/O grid With grid ∆ Temp W/O grid With grid ∆ Time

Case 1

1440.9

(113 sec)

1383.9

(65 sec)

57.0

379

347 32

Case 2

1418.0

(85 sec)

22.9 366 13
101

• • Case-1 Case-2

Figure 4-1 Node Change for Sensitivity of Grid Location
102

Figure 4-2 Heater Rod Temperature at 1.2 m – Run No.31805

Figure 4-3 Heater Rod Temperature at 1.98 m – Run No.31805
103

Figure 4-4 Heater Rod Temperature at 2.4 m – Run No.31805

Figure 4-5 Heater Rod Temperature at 3.3 m – No.31805
104

Figure 4-6 Quench Front Profile – Run No.31805

Figure 4-7 Heat Transfer Coefficient at 1.98 m – Run No.31805
105

4.1.2 Run No. 31302

Run No. 31302 was a test with a flooding rate of 7.65 cm/sec at 0.28 MPa and 78 oC inlet subcooling 

temperature as shown in Table 3-1. This was the same as Run No. 31805 except that it had three 

times over the reflood rate. 

The rod temperatures at various elevations were shown in Figure 4-8 through Figure 4-11. The rod 

temperatures for Case-2 were similar to those for Case-1 at elevation z < 1.98 m (6.5 ft). The effect of 

the grid locations would be weakened by the high reflood rate which did not largely change the original 

heat transfer coefficient with time. However, the rod temperatures for Case-2 were predicted higher 

than those for Case-1 at elevation z = 2.4 m. This reflected the behavior of HTC as shown in Figure 4-

13. The effect of the grid locations was not showed well for the rod temperature and the quenching 

time at elevation z = 3.3 m with low power density. At elevation z=1.98 with the maximum peak 

cladding temperature, the temperature reduction for Case-2 was ~ 7.8 K, and the quenching time was 

decreased with ~ 5 sec in comparison to the case without the spacer grid model as shown in Table 4-

2.

As shown in Figure 4-12, in Case-2 that the spacer grid was at the top of node, the quench front 

increased with almost same speed as Case-1. Because of the high reflood rate, the effect of the grid 

locations was not predicted for this test. The heat transfer coefficients (HTCs) at elevation z = 2.4 m 

was plotted in Figure 4-13. As shown in Figure, the HTC for Case-2 was under-predicted during the 

reflood phase as compared to Case-1. This might result from the use of the low original heat transfer 

coefficient since the spacer grid was modeled at the top of node. The steep increase of HTC for two 

cases occurred almost simultaneously since the rod quenching time was the same due to the high 

reflood rate. 

Table 4-2 Peak Temperature and Quenching Time at Elevation z = 1.98 for Run No. 31302

Case

Peak Temperature (K)

(at time)

Quenching Time (sec)

W/O grid With grid ∆ Temp W/O grid With grid ∆ Time

Case 1

1196.4

(24 sec)

1194.0

(19 sec)

2.4

156

152 4

Case 2

1188.6

(18 sec)

7.8 151 5
106

Figure 4-8 Heater Rod Temperature at 1.2 m – Run No.31302

Figure 4-9 Heater Rod Temperature at 1.98 m – Run No.31302
107

Figure 4-10 Heater Rod Temperature at 2.4 m – Run No.31302

Figure 4-11 Heater Rod Temperature at 3.3 m – No.31302
108

Figure 4-12 Quench Front Profile – Run No.31302

Figure 4-13 Heat Transfer Coefficient at 2.4 m – Run No.31302
109

4.2 Comparison with Spacer Grid Model of RELAP5

Currently, the RELAP5 [15] was updated as version 3.3jy by implementing the KNF (KEPCO Nuclear 

Fuel Co.) reflood model and the spacer grid model [16]. These models were theoretically based on the 

research was conducted by T.S. Choi [14]. Especially, the spacer grid model could be divided into 

three sub-models: Single-phase heat transfer enhancement, Grid rewet, and Droplet breakup. 

The spacer grid model of RELAP5 is somewhat different from that of TRACE. For the convective heat 

transfer enhancement, the enhancement due to the acceleration of the flow (Part 1 of Eq. (1)) was 

only considered in RELAP5 that it has been used also in COBRA-TF [17]. The convective heat 

transfer enhancement could be under-estimated in tests with mixing vanes since Part 2 of Eq. (1) 

could be large enough to influence to downstream near to spacer grid. For fuel bundles with typical 

mixing vane, the enhancement for mixing vanes could be below ~ 20% of the enhancement due to the 

flow acceleration up to ~ 0.4 m downstream of the spacer grid. However, the effect for mixing vanes 

might be not shown in FLECHT-SEASET since the spacer grid without mixing vanes was installed in 

it. The laminar enhancement factor, F was used in TRACE, but it was not in RELAP5. 

For the grid re-wetting model, the similar heat balance equation to Eq. (5) of TRACE was also 

implemented in RELAP5, but there were some differences in the detailed modelling. The radiation 

heat flux from the rods to the grid was obtained by using an electrical circuit analogy [3] in TRACE, but 

it was calculated explicitly in RELAP5 [14]. The correlation for the rewetting temperature in RELAP5 

was also different from that in TRACE. The rewetting temperature in RELAP5 was selected as the 

maximum value between the homogeneous nucleation temperature and two other minimum film 

boiling temperatures, while it in TRACE was determined by the minimum film boiling temperature. 

For the droplet breakup model, the suggested model by Yao, Hochreiter, and Cai [11] was used in 

TRACE, but the KAIST model [18] was used in RELAP5 since it could cover a wider range of the 

droplet Weber number, including Yao, Hochreiter, and Cai’s data. As the droplet Weber number was 

larger, the smaller shattered droplets occurred. In the upper region where the droplet dispersed flow 

regime was long maintained in which the droplet velocity and the number of entrained droplet were 

large, the droplet Weber number was large, and then the vapor temperature could be reduced 

because of the higher interfacial heat transfer between the droplets and the vapor phase. This could 

decrease the rod temperature and expedite the quenching time at that region. 

Currently, the droplet breakup model and the grid re-wetting model were not fully implemented in 

TRACE. Therefore, it was very difficult to compare directly to the effect of the spacer grid model 

between RELAP5 and TRACE. In this study, the results by RELAP5 could provide insights into the 

implementation and the modification associated with the spacer grid model. 

The RELAP5 nodalization for the FLECHT-SEASET was shown in Figure 4-14 and was almost similar 

to the previous study [19]. The test section including a heater with 3.66 m length was divided into 20 

nodes, and the upper and lower time-dependent volumes were modeled to define the fluid conditions, 

which represented the upper and lower plenums, respectively. Seven spacer grids were considered 

except for a first grid below the heater. Some initial conditions could be modified to use the same 

conditions as TRACE calculations in Chapter 3. Therefore, the actual inlet flow rate, temperature, and 

the upper plenum pressure were used as input with a function of time. The rod power was considered 

as a function of time. 

Among 8 tests in Table 3-1, Run No. 31805 and 31302 were selected as low and high reflood test for 

the assessment for the spacer grid model of RELAP5, respectively.
110

Figure 4-14 RELAP5 Nodalization for the FLECHT-SEASET Test

4.2.1.1 Effect of the new reflood model in RELAP5

RELAP5 has been broadly used in licensing LBLOCA analyses such as Westinghouse, OPR1000, 

and APR1400 plants in Korea. The KNF reflood model [14] was developed to improve the potential 

problems of RELAP5 Mod3.3 that the peak cladding temperature were under-predicted and the rod 

quenching occurred too early in FLECHT-SEASET tests. Therefore, the dry wall selection logic, the 

droplet size and inverted slug size, the post-dryout interfacial, film boiling heat transfer, and transition 

boiling models, etc. were modified in RELAP5. The effect of the new reflood model for Run. No 31805 

was shown in Figure 4-15 through Figure 4-17. The developmental code version of KNF 

(r5m33p03rev0-F6.exe) was used in this calculation. As mentioned in previous study [14], it could be 

identified that the rod temperatures were increased and the rod quenches were delayed, especially at 

high elevations. It was confirmed that the KNF reflood model could show the more improved results in 

Run. No 31805 although the rod quenching time was over-predicted in elevation z = 3.0 m.

At that time, the spacer grid model was implemented together with this reflood model. It was modeled 

that the KNF reflood model could be invoked without the spacer grid input, but the spacer grid input 

could not be activated without the KNF reflood model. Figure 4-18through Figure 4-20plotted the 

relations between the KNF reflood model and the spacer grid input. As shown in Figures, if the KNF 

reflood model was not used, the calculation results were not changed with or without the spacer grid 

input. It could be found that the KNF reflood model should be required to activate the spacer grid input. 
111

Figure 4-15 Heater Rod Temperature at 1.2 m – Run No.31805
112

Figure 4-16 Heater Rod Temperature at 1.98 m – Run No.31805

Figure 4-17 Heater Rod Temperature at 3.0 m – Run No.31805
113

Figure 4-18 Heater Rod Temperature at 1.2 m – Run No.31805

Figure 4-19 Heater Rod Temperature at 1.98 m – Run No.31805
114

Figure 4-20 Heater Rod Temperature at 3.0 m – Run No.31805

4.2.1.2 Effect of the spacer grid model in RELAP5

In order to identify the effect of the spacer grid model of RELAP5 for Run No.31805, the rod 

temperatures at various elevations were shown in Figures 4.21 ~ 4.26. The typical reflood process 

was shown well that the rod was heating up in the initiation of reflood, turning to reduce during the 

reflood, and finally quenching. In the RELAP5 calculation without the spacer grid model, the predicted 

peak temperature along elevation agreed reasonably with the experimental data. Especially, the peak 

rod temperature at higher elevations (z ≥ 3.0 m) had better prediction results as compared to those in 

TRACE. The maximum peak rod temperature was also shown at elevation z=1.98 m (6.5 ft). 

However, the quenching time was over-predicted even at the low elevation. This would show well the 

characteristics of new KNF reflood model such as the increase of rod temperature, the delay of 

quenching time as described in Figure 4-15 through Figure 4-17. When the spacer grid model was 

applied, the effect of mixing vane was not considered since the egg-crate grid was installed in the 

FLECHT-SEASET. As shown Figures 4.21 ~ 4.26, the effect of the spacer grid model did not show 

significantly up to elevation z = 2.4 m, but the peak temperatures with the spacer grid model had the 

lower values at higher elevations (z ≥ 3.0 m) and were under-predicted the experimental data in 

comparison with those without the spacer grid model. The earlier rod quenches were also predicted in 

the case with a spacer grid model. 

For the effect of the spacer grid model, it is difficult to find the meaningful comparison results between 

TRACE and RELAP5 in a quantitative perspective since they were different in the nodalization, the 

numerical scheme, and the applied thermal-hydraulic models. Therefore, several potential causes for 

differences of two codes could be considered in a qualitative point of view. 

When the results of RELAP5were compared with those of TRACE, the effect of the spacer grid model 

was more significantly shown in TRACE although the convective heat transfer enhancement was only 

considered in TRACE. The rod temperatures of TRACE started to reduce due to the spacer grid even 
115

at elevation z = 1.2 m as shown in Figure 3-6. However, the rod temperatures of RELAP5 did not 

changed at elevation z ≤ 1.98 m except for the earlier quenching at elevation z=1.98 m. As the 

elevation was higher over z = 1.98 m, the effect of the spacer grid in RELAP5 became larger as 

shown in Figure 4-27. 

The axial temperature profile at 150 sec was shown in Figure 4-28 to identify the effect of sub-models 

in RELAP5. The droplet breakup model gave the biggest effect on the rod temperatures among three 

models taking into account the effect of spacer grids. The effect of the droplet breakup model was 

more significant at the higher elevation since the droplet velocity and the number of entrained droplet 

could be larger at the higher elevation. The rod temperature was locally decreased because of the 

increase of convective heat transfer immediately downstream of a spacer grid (z ~ 2.1 m), but the 

convective heat transfer enhancement did not have a significant effect on the rod temperatures during 

the entire transient as shown in Figure 4-28. This was the biggest difference between TRACE and 

RELAP5. For the spacer grid heat transfer enhancement, the effect of the convective heat transfer in 

TRACE was much larger than that in RELAP5. 

This might be come from the modeling characteristics in TRACE and the use of the laminar 

enhancement factor. As explained above, the convective heat transfer enhancement effects in TRACE 

were integrated over the downstream axial cells for 50 hydraulic diameters to consider the exponential 

decay of the enhancement downstream of a spacer grid. In this test, 50 hydraulic diameters is about 

0.5 m, and this effect could be considered for longer downstream cells in comparison with that of 

RELAP5. The distance from the grid location to the cell center was only considered in RELAP5 with 

an assumption being far enough apart between the spacer grids. 

Secondly, the laminar enhancement factor was implemented in TRACE to take account for the 

additional enhancement effects that were observed for high void fraction, laminar flows. This factor can 

vary from 1.0 to 1.75 with Reynolds number, and the convective heat transfer enhancement of TRACE 

could be varied as shown Figure 4-29. Therefore, these differences between two codes could result in 

the more significant enhancement for the convective heat transfer in TRACE. As shown in Figure 3-6

~ 3.12, the effect of the convective heat transfer enhancement in TRACE could be shown in the entire 

heated rod, but its effect in RELAP5 could be limited immediately downstream of a spacer grid. 

Actually, the convective heat transfer enhancement had a significant effect immediately downstream of 

a spacer grid, and its effect may be evaluated too much in TRACE. The detailed further studies would 

need to be performed to evaluate the effect of sub-models. The convective heat transfer enhancement 

ought to be also modified since the spacer grid could influence upstream convective heat transfer.
116

Figure 4-21 Heater Rod Temperature at 1.2 m – Run No.31805

Figure 4-22 Heater Rod Temperature at 1.98 m – Run No.31805
117

Figure 4-23 Heater Rod Temperature at 2.4 m – Run No.31805

Figure 4-24 Heater Rod Temperature at 3.0 m – Run No.31805
118

Figure 4-25 Heater Rod Temperature at 3.3 m – Run No.31805

Figure 4-26 Heater Rod Temperature at 3.5 m – Run No.31805
119

Figure 4-27 Axial Rod Temperature Profile at 150 sec – Run No.31805

Figure 4-28 Axial Rod Temperature Profile for Sub-models at 150 sec – Run No.31805
120

Figure 4-29 Variation of Heat Transfer Enhancement for Laminar Enhancement Factor

Figure 4-30 through Figure 4-35 represent the rod temperatures at various elevations for Run 

No.31302 with relatively high flooding rate. In the RELAP5 calculation without the spacer grid model, 

the rod temperatures along elevation were predicted well the experimental data. Especially, the peak 

rod temperature at higher elevations (z ≥ 2.4 m) had better prediction results as compared to those in 

TRACE. The maximum peak rod temperature was also shown at elevation z=1.98 m (6.5 ft). The 

stepwise behaviors of the rod temperature couldbe showed at elevations z ≥ 2.4 m due to the use of a 

coarse node and the relatively high reflood rate. The quenching time also showed a good agreement 

with the experimental data except for earlier quenching at some elevations. When the spacer grid 

model was applied, the rod temperatures did not changed significantly up to elevation z = 1.98 m. The 

peak temperatures with the spacer grid model had the lower values at higher elevations (z ≥ 3.0 m), 

but they were not largely deviated from the experimental data in comparison with those without the 

spacer grid model. The rod quenches were also occurred at earlier time as a result of the heat transfer 

enhancement due to a spacer grid.

As compared to TRACE results in Figures 3.59 ~ 3.66, the effect of the spacer grid model in TRACE 

was more significantly shown, and it was also similar to the Run No.31805. As described above, these 

results may be come from the modeling characteristics in TRACE and the use of the laminar 

enhancement factor. The convective heat transfer enhancement may be excessively large if the 

laminar enhancement factor will be large. 
121

Figure 4-30 Heater Rod Temperature at 1.2 m – Run No.31302

Figure 4-31 Heater Rod Temperature at 1.98 m – Run No.31302
122

Figure 4-32 Heater Rod Temperature at 2.4 m – Run No.31302

Figure 4-33 Heater Rod Temperature at 3.0 m – Run No.31302
123

Figure 4-34 Heater Rod Temperature at 3.3 m – Run No.31302

Figure 4-35 Heater Rod Temperature at 3.5 m – Run No.31302
124

Based on the theoretical study [14] and the developmental version of KNF, the KNF reflood model and 

the spacer grid model were initially implemented to RELAP5/MOD3.3 patch4 Version 3.3jy in October 

2014, and they has been modeled in the latest developmental Version 3.3km (February 2016). 

According to RELAP5 input manual, the KNF reflood model could be used by Option 40 of Card 1 and 

invoked regardless of the KNF grid input (43000000 cards). However, the spacer grid model could be 

applied only when the option 40 was used in Card 1. This could be identified in the calculations of 

Figures 4.18 ~ 4.20. 

In this study, the calculations for various FLECHT-SEASET tests were performed by the latest 

distributed RELAP5 versions (Version 3.3jz & 3.3kl), and two codes actually produced same results for 

this tests. The used input decks were the same as the calculations for a developmental version of 

KNF. The rod temperatures at various elevations for Run No.31805 were shown in Figures 4.36 ~ 

4.38. When the spacer grid model was applied, the peak rod temperature had the higher values up to 

elevation z = 1.98 m. The peak rod temperature especially at elevation z = 1.98 m was largely 

increased due to the spacer grid. It is not the expected result. When the spacer grid was applied, the 

flow area is reduced and the convective heat transfer is promoted due to the flow acceleration and the 

turbulence increases and then the rod temperature would be usually decreased and the rod quenches 

were expedited. However, this decrease of the rod temperature was not shown in Figures 4.36 ~ 4.38. 

The results was contrary to those in Figures 4.21 ~ 4.24 using the same input decks. 

Figure 4-39 through 4.41 showed the results for the use of KNF reflood model (Option 40) and spacer 

grid input (43000000 cards). The results with a spacer grid and no Option 40 (Green line) were 

completely in accord with those with a spacer grid and Option 40 (Cyan line). It was also an unusual 

thing that the effect of spacer grid model was significantly shown in Figures 4.39~ 4.41 even though 

Option 40 was not used in Card 1. For example, in the case with a spacer grid model (Green line), the 

growth of rod temperature and the delay of quenching time were predicted as compared with the case 

without it (Red line). It was also contrary to what we expected. 

Consequently, as these results were compared to those in Figures 4.15 ~ 4.26, the current RELAP5 

version (Version 3.3jz ~ 3.3kl) including the KNF reflood and spacer grid models may have some 

troubles to implement these models. Therefore, the detailed errors should be found and corrected in 

the further study. 
125

Figure 4-36 Heater Rod Temperature at 1.2 m – Run No.31805

Figure 4-37 Heater Rod Temperature at 1.98 m – Run No.31805
126

Figure 4-38 Heater Rod Temperature at 3.0 m – Run No.31805

Figure 4-39 Heater Rod Temperature at 1.2 m – Run No.31805
127

Figure 4-40 Heater Rod Temperature at 1.98 m – Run No.31805

Figure 4-41 Heater Rod Temperature at 3.0 m – Run No.31805

129

5 CONCLUSIONS

The Full-Length Emergency Core Heat Transfer Separate Effects and Systems Effects Tests 

(FLECHT-SEASET) was conducted to identify the thermal hydraulic phenomena of forced and gravity 

reflooding in a 161-rod bundle without flow blockage. In this study, the spacer grid model of TRACE 

was evaluated for the FLECHT-SEASET test by the TRACE V5.0 patch4. The FLECHT-SEASET test 

section was modeled in the VESSEL component of TRACE. The test section was divided into 16 axial 

nodes; one node for lower plenum, fourteen nodes for heated section and one node for upper plenum 

and there were two cells between each gird. The 161 heated rods in 17x17 assemblies were modeled 

as a single HTSTR component. The injected flow rates and temperatures was provided as a function 

of time by FILL component connecting to the bottom of the lower plenum. The BREAK component 

was used to set the pressure boundary at the top of the test section. The egg-crate grids without the 

mixing vane were located in the bottom of every other node. The grid parameters were determined by 

the experimental design data [5,7] and the general fuel data. 

Eight tests were chosen in this study and tests were covering a range of flooding rate from 2.10 

cm/sec to 15.50 cm/sec, subcooling temperature from 5 o

C to 79 oC and upper plenum pressure from 

0.13 MPa to 0.41 MPa. The initial rod power at the peak location is 2.3 KW/m (0.7 KW/ft) in all tests. 

As would be expected, the rod temperatures were decreased and the rods were quenched at the 

earlier time in most tests if the spacer grid model was used. In addition, as the reflood rate was 

increased, the lowerof the peak rod temperature and the earlier quenching time were predicted. In 

tests with high reflood rate, the change of peak temperature due to the spacer grid was not large, 

which resulted from the short heat up period and the faster increase of liquid level by the high 

reflooding rate. As the test pressure was lower, the higher rod temperature and the later rod quench 

were predicted since the liquid approached faster to the relatively lower saturation temperature. The 

use of spacer grid model showed relatively bigger differences for rod temperature, not for quenching 

time at lower pressure condition. As the subcooling degree was higher, the decrease amount of 

quenching time due to the spacer grid was reduced since the high subcooling degree could enhance 

the heat transfer rate.

The sensitivity studies were performed to identify the effect of the grid locations and the difference 

from the spacer grid model of RELAP5. When the locations of spacer grids were changed into the top 

of every other node, the rod temperature had higher values and the rod quenching was delayed at 

elevation with a high power as compared to the case which the spacer grids were located in the 

bottom of the node. It would be more reasonable to predict the experimental data and showed more 

conservative results. The RELAP5 code currently implemented the KNF reflood model and the spacer 

grid model. This reflood model considered three sub-models such as single-phase heat transfer 

enhancement, grid rewet, and droplet breakup. However, the current RELAP5 version (Version 3.3jz ~ 

3.3kl) may have some error to implement the KNF reflood & spacer grid model. It may be because 

there are some troubles in the use of KNF reflood model (Option 40) and spacer grid input (43000000 

cards). Therefore, the developmental version of KNF (r5m33p03rev0-F6.exe) was used to identify the 

effect of sub-models and then the droplet breakup model gave the biggest effect on the rod 

temperatures among three models. The effect of the droplet breakup model was more significant at 

the higher elevation since the droplet velocity and the number of entrained droplet could be larger at 

the higher elevation. When the RELAP5 results were compared with the TRACE results, the effect of 

the spacer grid model in TRACE was more significantly shown even though the single-phase heat 

transfer enhancement was only implemented in TRACE. These results may be come from the 

modeling characteristics and the use of the laminar enhancement factor in TRACE. The convective 

heat transfer enhancement may be excessively large if the laminar enhancement factor will be large.
130

In conclusion, the effect of the spacer grid model in TRACE was shown well to simulate the FLECHTSEASET reflood heat transfer experiments. However, there are some limitations to quantitatively 

predict the effect of the droplet breakup and the grid rewetting models. From the comparison with the 

RELAP5, it may be found that the current RELAP5 version had some errors to implement the spacer 

grid model, and the effect of the spacer grid of TRACE could be over-estimated for the rod 

temperature behaviors. Therefore, the current RELAP5 needs to be modified to correct their errors for 

the spacer grid model and the TRACE code should be improved to implement the droplet breakup 

and the grid rewetting models in the future study. 
131

6 REFERENCES

[1] USNRC, “Improvement of RELAP5/MOD3.3 Reflood Model Based on the 

Assessments against FLECHT-SEASET Tests”, NUREG/IA-0251, 2011.

[2] USNRC, “TRACE V5.840 User’s manual, Volume 1: Input Specification”, 2014. 

[3] USNRC, “TRACE V5.840 Theory Manual, Field Equations, Solution Methods, and 

Physical Models”, 2013.

[4] RBHT IA reportssss............

[5] USNRC, “PWR FLECHT SEASET Unblocked Bundle, Forced and Gravity Reflood 

Task Data Evaluation and Analysis Report”, NUREG/CR-2256, 1982.

[6] USNRC, “TRACE V5.0 Assessment Manual, Appendix B: Separate Effects Tests”, 

2007.

[7] USNRC, “PWR FLECHT SEASET Unblocked Bundle, Forced and Gravity Reflood 

Task Data Report”, NUREG/CR-1532, 1980.

[8] Yao, S. C., L. E. Hochreiter, and W. J. Leech., "Heat-Transfer Augmentation in 

Rod Bundles Near Grid Spacers", Journal of Heat Transfer, 104, 76-81, 1982.

[9] Yao, Shi-Chune, M. Loftus, and L. Hochreiter., "Hydraulics of Blocked Rod 

Bundles", Nuclear Technology, 65, 444-453, 1984.

[10] Rehme, Klaus, J. Marek, and K. Maubach. "Heat Transfer and Pressure Drop 

Performance of Rod Bundles Arranged in Square Arrays", International Journal of 

Heat and Mass Transfer, 16, 2215-2228, 1973.

[11] Yao, S. C., L. E. Hochreiter, and K. Y. Cai. "Dynamics of Droplets Impacting on 

Thin Heated Strips," 23rd National Heat Transfer Conference, American Society of 

Mechanical Engineers, Heat Transfer Division. New York, 1985.

[12] USNRC, “Analysis of FLECHT-SEASET 163-Rod Blocked Bundle Data Using 

COBRA-TF”, NUREG/CR-4166, 1985.

[13] USNRC, “NRC Regulations (10 CFR) Appendix K to Part 50-ECCS Evaluation 

Models”, 2015.

[14] T.S. Choi, “Development of an Improved Reflood Model for RELAP5 and SPACE”, 

PhD Thesis of KAIST, 2013.

[15] USNRC, “RELAP5/MOD3.3 Code Manual Volume II: Appendix A Input 

Requirements”, 2010.

[16] D. Barber., “RELAP5 Status and User Problem Report”, Fall 2014 CAMP Meeting, 

Information Systems laboratories (ISL), Inc., 2014.

[17] Thurgood, M.J., Kelly, J.M., Guidotti, T.E., Kohrt, R.J., and Crowell, K.R., 

“COBRA/TRAC-A Thermal Hydraulic Code for Transient Analysis of Nuclear 

Reactor Vessels and Primary Coolant Systems”, NUREG/CR-3046, 1983. 

[18] Park, B.H., Kim, Y.S., Jeon, B.K., Kim, H.S., Choi, J.Y., and NO, H.C., “Models 

Evaluation of Reflood Heat Transfer by Spacer Grids in the Safety Analysis Code”, 

KAERI/CM-1592/1211. 2012.

[19] USNRC, “RELAP5/MOD3.3 Code Manual Volume III: Developmental Assessment 

Problems”, 2010.

Evaluation of TRACE Spacer Grid Model with FLECHT-SEASET Reflood 

Test

NUREG/IA-0481

Byung-Gil HUH, Ae-Ju CHEONG, Kyung-Won LEE

Korea Institute of Nuclear Safety

62 Gwahak-ro, Yuseong-gu 

Daejeon, 34142, Korea

K.Tien, NRC Project Manager

Division of Systems Analysis

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, D.C. 20555-0001

The effects of spacer grid model of TRACE V5.0 patch 4 were assessed for the Full-Length Emergency Core Heat 

Transfer Separate Effects and Systems Effects Tests (FLECHT-SEASET). The main parameters of the spacer grid were 

defined by the experimental design data and eight egg-crate grids were modeled in the VESSEL component of TRACE. 

The calculations for eight tests of FLECHT-SEASET revealed that when the spacer grid model was used, the rod 

temperatures decreased and the rods were quenched at an earlier time in most other tests. In addition, as the reflood rate 

increased, the lower peak rod temperature and the earlier quenching time were predicted. When the test pressure was 

lower, the higher rod temperature and the later rod quenching were predicted since the liquid approached a relatively lower 

saturation temperature faster. When the subcooling degree was higher, the reduced degree of quenching time due to the 

spacer grid was further decreased because the higher subcooling degree enhanced the heat transfer rate. Sensitivity 

studies were performed to identify the effect of the grid locations and the difference from the spacer grid model of RELAP5. 

In this study, the effect of the spacer grid model in TRACE is shown well to simulate the FLECHT-SEASET reflood heat 

transfer tests. However, since the droplet breakup and the grid rewetting models were not fully implemented yet, there 

were some limitations in quantitatively predicting their effects. The comparison with the RELAP5 revealed that the current 

RELAP5 version had some errors in implementing the spacer grid model, and the effect of the spacer grid of TRACE could 

have been over-estimated for the rod temperature behaviors as compared with RELAP5. 

Full-Length Emergency Core Heat Transfer Separate Effects and 

Systems Effects Tests (FLECHT-SEASET) 

Mixing Vane Spacer Grid

Egg-crate Spacer Grid

TRACE Version 5.0 Patch Release 4

Reflood Heat Transfer

Droplet Breakup

Spacer Grid Rewetting

August 2017

Technical



NUREG/IA-0481 Evaluation of TRACE Spacer Grid Model with FLECHT-SEASET Reflood Test August 2017