NUREG-1752

Assessment of Modernization and 

Integration of BWR 

Components and Spatial 

Kinetics in the TRAC-M, 

Version 3690, Code 

U.S. Nuclear Regulatory Commission 

Office of Nuclear Regulatory Research -oeoft REQ&4 

Washington, DC 20555-0001 ,A
NUREG-1752 

Assessment of Modernization and 

Integration of BWR 

Components and Spatial 

Kinetics in the TRAC-M, 

Version 3690, Code 

Manuscript Completed: September 2001 

Date Published: December 2001 

Principal Author: F. Odar (NRC)* 

Contributing Authors: 

J. Uhle, S. Lu (NRC)* 

J. F. Dearing (LANL)** 

B. Aktas, R. Shumway, S. Lucas (ISL)*** 

T. Downar (Purdue University)**** 

*Division of Systems Analysis and Regulatory Effectiveness 

Office of Nuclear Regulatory Research 

U.S. Nuclear Regulatory Commission 

Washington, DC 20555-0001 

**Los Alamos National Laboratory 

Los Alamos, NM 87545 

***ISL, Inc. 

11140 Rockville Pike 

Rockville, MD 20852 

****Purdue University 

1290 Nuclear Engineering Building 

W. Lafayette, IN 47907-1290
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ABSTRACT

This report presents an assessment of three different activities which led to the creation of 

TRAC-M, Version 3690. The first is the modernization of the old Transient Reactor Analysis 

Code for pressurized-water reactors (TRAC-P), the second is the integration of boiling water 

reactor (BWR) components from the TRAC-B code and the third is the integration of a spatial 

kinetics capability from the PARCS Code. The assessment shows that all three aspects of the 

modernization and integration have been successful.

Page iii
CONTENTS

Abstract ..................... 

Executive Summary ............ 

Foreword .................... 

Acknowledgments ............. 

Acronyms ....................

Page 

. .. . .. . . . . . .. . .. . .. . . . . . . . . .. . .. . .. . .iii 

. .. . xxxv ...............

................................. xxxvii 

................................. xxxviii 

.............. o ...................... XXX'X

1. Introduction ......................................................... 1-1

Description of the Consolidated TRAC Release and History of its 

D evelopm ent ......................................................... 2-1 

Analytical Test Problem s ............................................. 3-1 

Steady-State Conduction Problems (HCOND#; # = 1,2,3) ...................... 3-3 

1-D Radial Conduction .................................................. 3-3 

2-D Radial and Axial Conduction ......................................... 3-5 

1-D Drain and Fill Test Problem (DRAIN) ................................. 3-11 

U-Tube Test Problem (UTUBE) .......................................... 3-16 

C onclusions .......................................................... 3-20 

Page v

2.

3. 

3.1 

3.1.1 

3.1.2 

3.2 

3.3 

3.4
4. PWR Developmental Assessment ................................ 4-1 

4.1 Countercurrent Flow Limitation ........................................... 4-1

4.1.1 Bankoff Tests ........................................ 

4.1.1.1 Test Facility ....................................... 

4.1.1.2 TRAC M odel ..................................... 

4.1.1.3 Comparison of Predicted and Measured Results .......... 

4.1.1.4 Conclusions ....................................... 

4.2 Condensation ........................................ 

4.2.1 Akim oto Tests ........................................ 

4.2.1.1 Condensation Experiments ........................... 

4.2.1.2 TRAC M odel ..................................... 

4.2.1.3 Comparison of Predicted and Measured Results .......... 

4.2.1.4 Conclusions ....................................... 

4.3 Critical Flow ......................................... 

4.3.1 M arviken Tests ....................................... 

4.3.1.1 Facility Description ................................. 

4.3.1.2 TRAC M odel ..................................... 

4.3.1.3 Comparison of Predicted and Measured Results for Test 4.. 

4.3.1.4 Comparison of Predicted and Measured Results for Test 13 . 

4.3.1.5 Comparison of Predicted and Measured Results for Test 20. 

4.3.1.6 Comparison of Predicted and Measured Results for Test 22. 

4.3.1.7 Comparison of Predicted and Measured Results for Test 24. 

4.3.1.8 Conclusions .......................................

4-1 

................. 4-1 

................. 4-2 

................. 4-3 

................. 4-4 

................ 4-13

4-13 

4-13 

4-14 

4-14 

4-16 

4-32 

4-32 

4-32 

4-32 

4-33 

4-33 

4-34 

4-34 

4-35 

4-35

Page vi
4.3.2 Edwards Blowdown Test ............................................... 

4.3.2.1 Facility Description ................................................. 

4.3.2.2 TRA C M odel ..................................................... 

4.3.2.3 Comparison of Predicted and Measured Results .......................... 

4.3.2.4 Conclusions .......................................................

4.4 Reflood ................................... 

4.4.1 Lehigh Rod Bundle Test 02/24/85-20 ............ 

4.4.1.1 Facility Description ....................... 

4.4.1.2 Test Procedure and Description ............. 

4.4.1.3 TRAC Model ........................... 

4.4.1.4 Comparison of Predicted and Measured Results

4.4.1.5 Conclusions.

.......................... 4-63

4-63 

4-64 

4-65 

4-65 

4-66

........................... 4-68

4.4.2 FLECHT-SEASET Test 31504 ................................. 

4.4.2.1 Facility Description ....................................... 

4.4.2.2 Test Procedure ........................................... 

4.4.2.3 TRAC M odel ........................................... 

4.4.2.4 Comparison of Predictions with Experimental Data ............. 

4.4.2.5 Discussion of TRAC Models ............................... 

4.4.2.6 Conclusions ............................................. 

4.4.3 Cylindrical Core Test Facility (CCTF), Core I Test Cl-5 (Run 14) ..... 

4.4.3.1 Facility Description ....................................... 

4.4.3.2 Test Procedure ........................................... 

4.4.3.3 TRAC M odel ........................................... 

4.4.3.4 Comparison of TRAC-M(F77) and TRAC-M(F90) Calculations ... 

4.4.3.5 Conclusions .............................................

.......... 4-69 

.......... 4-69 

4-69 

.......... 4-70 

.......... 4-71 

.......... 4-73 

.......... 4-75 

.......... 4-76 

.......... 4-76 

.......... 4-77 

.......... 4-78 

.......... 4-79 

.......... 4-79

Page vii

4-36 

4-36 

4-36 

4-36 

4-37
4.4.4 Cylindrical Core Test Facility (CCTF), Core-II Run 54 ........................ 4-80 

4.4.4.1 Facility D escription ................................................. 4-80 

4.4.4.2 Test Procedure ..................................................... 4-82 

4.4.4.3 TRAC Input M odel ................................................. 4-84 

4.4.4.4 Comparison of TRAC-M(F77) and TRAC-M(F90) Predictions .............. 4-84 

4.4.4.5 Conclusions ....................................................... 4-84 

4.4.5 Slab Core Test Facility (SCTF), Run 719 ................................... 4-85 

4.4.5.1 Facility Description ................................................. 4-85 

4.4.5.2 Test Procedure ..................................................... 4-86 

4.4.5.3 TRAC Input M odel ................................................. 4-86 

4.4.5.4 Comparison of TRAC-M(F77) and TRAC-M(F90) Calculations ............. 4-88 

4.4.5.5 Conclusions ....................................................... 4-89 

4.5 Flow Patterns in Cold Legs During ECC Injection .......................... 4-180 

4.5.1 Upper-Plenum Test Facility (UPTF), Cold Leg Flow Test 8, Phase B, Part 1 ...... 4-180 

4.5.2 Test Procedure .................................................... 4-180 

4.5.3 TRAC M odel .................................................... 4-181 

4.5.4 Comparison of TRAC-M(F77) and TRAC-M(F90) Predictions .............. 4-183 

4.5.5 Conclusions ...................................................... 4-183 

4.6 ECC Bypass and Downcomer and Lower Plenum Refill ...................... 4-201 

4.6.1 Upper-Plenum Test Facility (UPTF), Downcomer Test 6, Run 133 .............. 4-201 

4.6.2 Test Procedure .................................................... 4-201 

4.6.3 TRAC M odel .................................................... 4-202 

4.6.4 Comparison of TRAC-M(F77) and TRAC-M(F90) Calculations ............ 4-203 

4.6.5 Conclusions ...................................................... 4-203

Page viii
4.7 Steam Cooling ....................................................... 4-226 

4.7.1 FLECHT-SEASET Test 32753 .......................................... 4-226 

4.7.2 TRAC Input Model for FLECHT-SEASET Steam Cooling Run 32753 .......... 4-226 

4.7.3 Comparison of TRAC-M(F77) and TRAC-M(F90) Calculations with 

Experimental Data from FLECHT-SEASET Run 32753 ...................... 4-226 

4.7.4 Conclusions ......................................................... 4-227 

4.8 Loss-of-Fluid Test (LOFT), Test L2-6 (Large-Break LOCA) .................. 4-231 

4.8.1 Facility Description ................................................... 4-231 

4.8.2 Test Procedure ....................................................... 4-232 

4.8.3 TRAC M odel ....................................................... 4-232 

4.8.4 Comparison of Calculated and Measured Results ........................... 4-233 

4.8.5 Conclusions ......................................................... 4-234 

4.9 Loss-of-Fluid Test (LOFT), Test L6-1 (Loss of Steam Load) .................. 4-250 

4.9.1 Facility Description ................................................... 4-250 

4.9.2 Test Procedure ....................................................... 4-251 

4.9.3 TRAC-M M odel ..................................................... 4-251 

4.9.4 Comparison of Predicted and Measured Results ............................ 4-253 

4.9.5 Conclusions ......................................................... 4-254 

5. Assessment of Integration of BWR Components ...................... 5-1 

5.1 Heater (HEATR) Component ............................................. 5-1 

5.1.1 R equirem ents ......................................................... 5-3 

5.1.2 Verification Testing and Assessment ....................................... 5-4

Page ix
5.1.3 C onclusion ........................................................... 5-8

5.2 Jet Pump (JETP) Component ......................................

R equirem ents ........................................................ 

Verification Testing and Assessment ...................................... 

C onclusions ..........................................................

5.3 Fuel Channel (CHAN) Component ........... 

5.3.1 Requirem ents ............................ 

5.3.2 Verification Testing and Assessment .......... 

5.3.3 Conclusions ..............................

............................ 5-39

5-39 

5-40 

5-48

5.4 Level Tracking (LTRCK) Model ......................................... 5-70 

5.4.1 R equirem ents ........................................................ 5-70 

5.4.2 Verification Testing and Assessment ...................................... 5-80 

5.4.2.1 Verification Tests for Requirements LTRCK 1.1, 2.1, and 2.2 ................ 5-80 

5.4.2.2 Verification Tests and Their Assessment for Requirement LTRCK 1.2 ......... 5-84 

5.4.2.3 Verification Tests and Their Assessment for Requirement LTRCK 2.3 ......... 5-85 

5.4.2.4 Verification Tests and Their Assessment for Requirement LTRCK 2.4 ......... 5-86 

5.4.2.5 Verification Tests and Their Assessment for Requirement LTRCK 3.0 ......... 5-88 

5.4.3 Conclusions .......................................................... 5-88 

5.5 BW R Control System ................................................. 5-118 

5.5.1 R equirem ents ....................................................... 5-120 

5.5.2 Verification Testing and Assessment ..................................... 5-126 

5.5.3 C onclusions ......................................................... 5-132

Page x

5.2.1 

5.2.2 

5.2.3

..... 5-20

5-20 

5-22 

5-28
5.6 Single Junction (SJC) Component ....................................... 5-151 

5.6.1 Requirem ents ....................................................... 5-152 

5.6.2 Verification Testing and Assessment ..................................... 5-153 

5.6.2.1 End-Junction Replacement Tests and Their Results ....................... 5-153 

5.6.2.2 Composite TEE Tests and Their Results ............................... 5-154 

5.6.2.3 Link to VESSEL, Static Check (Input Files tinm3pl-b and tinm3pl-s) ........ 5-155 

5.6.2.4 Link to VESSEL, Simulated BWR CHAN Component .................... 5-155 

5.6.2.5 SJACS PUMP Component (Input Files tinm5pl-b and tinm5pl-s) ............ 5-156 

5.6.2.6 SJACS VALVE Component ......................................... 5-156 

5.6.2.7 Test tfpipe2 M odification........................................... 5-157 

5.6.3 Conclusions ......................................................... 5-158 

5.7 Turbine (TURB) Component ........................................... 5-178 

5.7.1 Requirem ents ....................................................... 5-178 

5.7.2 Verification Testing and Assessment ..................................... 5-183 

5.7.3 Conclusions ......................................................... 5-188 

6. Assessment of Integration of Spatial Kinetics ......................... 6-1 

6.1 Description of Linkage .................................................. 6-1 

6.1.1 Variable Transfer ....................................................... 6-1 

6.1.2 General Interface Structure and Data Transfer ................................ 6-2 

6.1.3 Design of the Mapping Function .......................................... 6-4 

6.1.4 Calculational Control ................................................... 6-6 

6.1.5 Tim e Step Control ...................................................... 6-7 

6.1.6 Programming Standard for Interface Routines ................................ 6-7 

6.1.7 Interface Sum m ary ..................................................... 6-8

Page xi
6.2 R equirem ents ........................................................ 6-12

6.2.1 

6.2.2 

6.2.3 

6.2.4 

6.2.5 

6.2.6 

6.2.7

Description of Test Cases ............................................... 

Permutation Submatrices in Coordinate Storage Format ....................... 

Input Vectors After Permutation .......................................... 

Output Vectors After Permutation ........................................ 

Initial Control Buffers .................................................. 

Time-Dependent Control Buffers ......................................... 

Test Plan Verification and Acceptance Criteria ..............................

6.3 Verification Testing and Assessment ........ 

6.4 Conclusions ............................ 

7. Plant Transient Calculations ........... 

7.1 AP600 Small-Break LOCA Calculations .....

7.1.1 

7.1.2 

7.1.3 

7.1.4 

7.1.5

AP600 System Description ......................................... 

Accident Scenario Investigated ..................................... 

TRAC-M M odeling .............................................. 

TRAC-M Calculations ............................................ 

C onclusions .....................................................

....... 7-1 

....... 7-1 

....... 7-2 

....... 7-3 

....... 7-3

7.2 TRAC-M Spatial Kinetics Functionality: BWR Application ........ ............. 7-7

7.2.1 Benchmark Model Specifications .......................................... 

7.2.1.1 Core Geometry and Fuel Assembly Specifications ......................... 

7.2.1.2 PARCS Neutronics M odeling ..........................................

Page xii

6-15 

6-16 

6-20 

6-23 

6-26 

6-28 

6-31 

6-34 

6-44 

7-1

7-1

7-7 

7-7 

7-8
7.2.1.3 TRAC-M Thermal-Hydraulic Model and Coupling of Neutronics to Thermal

H ydraulics ......................................................... 7-8 

7.2.2 Peach Bottom Benchmark Results with TRAC-M / PARCS ..................... 7-8 

7.2.2.1 PBTT2 Steady-State Initialization Results ................................ 7-8 

7.2.2.2 PBTT2 Transient Results with TRAC-M / PARCS ......................... 7-9 

8. C onclusions .......................................................... 8-1

Page xiii
FIGURES

Figure 3.1.1 

Figure 3.1.2 

Figure 3.2.1 

Figure 3.2.2 

Figure 3.2.3 

Figure 3.2.4 

Figure 3.3.1 

Figure 3.3.2 

Figure 3.3.3 

Figure 4.1.1 

Figure 4.1.2 

Figure 4.1.3 

Figure 4.1.4 

Figure 4.1.5 

Figure 4.1.6 

Figure 4.1.7 

Figure 4.1.8 

Figure 4.2.1 

Figure 4.2.2 

Figure 4.2.3 

Figure 4.2.4 

Figure 4.2.5

Boundary Conditions of a 1 -D Heat-Conduction Problem in the Radial 

D irection .................................................... 3-10 

The 2-D Radial and Axial Heat-Conduction Problem .................. 3-10 

TRAC Model of the DRAIN Test Problem .......................... 3-12 

Comparison of Void Fraction Predictions in DRAIN and FILL Test 

Problem ..................................................... 3-13 

Comparison of Mass Flow Predictions in DRAIN and FILL Test Problem. 3-14 

Comparison of Pressure Predictions in DRAIN and FILL Test Problem ... 3-15 

TRAC Model of the UTUBE Test Problem .......................... 3-17 

Comparison of Gas Volume Fraction Oscillations in the UTUBE Test Problem, 

DT(m ax)=0.05s ............................................... 3-18 

Comparison of Gas Volume Fraction Oscillations in the UTULBE Test Problem, 

D T(m ax)=0.01s ...... I ........................................ 3-19 

Sketch of the Perforated Plate Air-Water Countercurrent Flow System Used by 

Bankoff et al ................................................... 4-5 

TRAC Model for the Test Apparatus Used by Bankoff et al .............. 4-6 

Comparison of Calculated Liquid Mass Flows at the Perforated Plate in the 

Bankoff CCFL Test ............................................. 4-7 

Comparison of Calculated Air Mass Flows at the Perforated Plate in the Bankoff 

CCFL Test .................................................... 4-8 

Liquid Mass Flow at the Perforated Plate in the Bankoff CCFL Test ....... 4-9 

Air Mass Flow at the Perforated Platein the Bankoff CCFL Test ......... 4-10 

Liquid and Air Dimensionless Fluxes at the Perforated Plate in the Bankoff 

CCFL Test ................................................... 4-11 

Comparison of TRAC Results with Bankoff Correlation ............... 4-12 

Sketch of the Experimental Apparatus of Akimoto et al ................ 4-18 

TRAC Input M odel ............................................ 4-19 

Noding Cell Lengths of the Test Section ............................ 4-20 

Comparison of Calculated and Measured Liquid Film Temperatures in the 

Akimoto Stratified Mist Flow Test ................................ 4-21 

Comparison of Predictions of Pressure Variations at Injection Location in the 

Akimoto Oscillatory Plug Test .................................... 4-22

Page xv
Figure 4.2.6 

Figure 4.2.7 

Figure 4.2.8 

Figure 4.2.9 

Figure 4.2.10 

Figure 4.2.11 

Figure 4.2.12 

Figure 4.2.13 

Figure 4.2.14 

Figure 4.3.1 

Figure 4.3.2 

Figure 4.3.3 

Figure 4.3.4 

Figure 4.3.5 

Figure 4.3.6 

Figure 4.3.7 

Figure 4.3.8 

Figure 4.3.9 

Figure 4.3.10 

Figure 4.3.11

Comparison of Predictions of Pressure Variations at Injection Location in the 

Akimoto Oscillatory Plug Test .................................... 4-23 

Comparison of Predictions of Pressure Variations at Injection Location in the 

Akimoto Oscillatory Plug Test .................................... 4-24 

Comparison of Vapor Fraction Predictions at Injection Location in the Akimoto 

Oscillatory Plug Test ........................................... 4-25 

ECC Injection and Steam Flow in the Akimoto Transient Test ........... 4-26 

Comparison of Pressure Oscillations at the Injection Point in the Akimoto 

Transient Test (0 - 140s) ........................................ 4-27 

Comparison of Pressure Oscillations at the Injection Point in the Akimoto 

Transient Test (0 - 40s) ......................................... 4-28 

Comparison of Pressure Oscillations at the Injection Point in the Akimoto 

Transient Test (50 - 90s) ........................................ 4-29 

Comparison of Pressure Oscillations at the Injection Point in the Akimoto 

Transient Test (100 - 140s) ...................................... 4-30 

Transition from Stable Wavy to Oscillatory Flow in the Akimoto Transient 

T est ......................................................... 4-3 1 

M arviken Pressure Vessel ....................................... 4-41 

Marviken Discharge Pipe, Test Nozzle, and Rupture Disk Assembly ...... 4-42 

Pressure-Vessel and Discharge-Pipe Noding ......................... 4-43 

Nozzle and Rupture-Disk Assembly Noding for Marviken Tests ......... 4-44 

Comparison of Calculated and Measured Break Flows for Marviken 

Test 4 ....................................................... 4-45 

Comparison of Calculated and Measured Pressures in the Upper Vessel for 

M arviken Test 4 ............................................... 4-46 

Comparison of Calculated and Measured Discharge Line Pressures for Marviken 

4 Test ....................................................... 4-47 

Comparison of Calculated and Measured Lower Vessel Fluid Temperatures for 

M arviken Test 4 ............................................... 4-48 

Comparison of Calculated and Measured Break Flows for Marviken 

Test 13 ...................................................... 4-49 

Comparison of Calculated and Measured Upper Vessel Pressures for Marviken 

Test 13 ...................................................... 4-50 

Comparison of Calculated and Measured Lower Vessel Temperatures for 

M arviken Test 13 .............................................. 4-51

Page xvi
Figure 4.3.12 

Figure 4.3.13 

Figure 4.3.14 

Figure 4.3.15 

Figure 4.3.16 

Figure 4.3.17 

Figure 4.3.18 

Figure 4.3.19 

Figure 4.3.20 

Figure 4.3.21 

Figure 4.3.22 

Figure 4.3.23 

Figure 4.4.1 

Figure 4.4.2 

Figure 4.4.3 

Figure 4.4.4 

Figure 4.4.5 

Figure 4.4.6 

Figure 4.4.7 

Figure 4.4.8 

Figure 4.4.9

Comparison of Calculated and Measured Break Flows for Marviken 

Test 20 ...................................................... 4-52 

Comparison of Calculated and Measured Pressures in the Upper Vessel for 

M arviken Test 20 .............................................. 4-53 

Comparison of Calculated and Measured Break Flows for Marviken 

Test 22 ...................................................... 4-54 

Comparison of Calculated and Measured Pressures in the Upper Vessel for 

M arviken Test 22 .............................................. 4-55 

Comparison of Calculated and Measured Break Flows for Marviken 

Test 24 ...................................................... 4-56 

Comparison of Calculated and Measured Pressures in the Upper Vessel for 

Marviken Test 24 4-57 

Edwards Blowdown Test Apparatus ............................... 4-58 

TRAC Model from the Edwards Blowdown Test ..................... 4-58 

Comparison of Calculated and Measured Pressures at GS-1 for the Edwards 

Blowdown Test ............................................... 4-59 

Comparison of Calculated and Measured Pressures at GS-7 for the Edwards 

Blowdown Test ............................................... 4-60 

Comparison of Calculated and Measured Vapor Fractions at GS-5 for the 

Edwards Blowdown Test ........................................ 4-61 

Comparison of Calculated and Measured Liquid Temperatures at GS-5 for the 

Edwards Blowdown Test ........................................ 4-62 

Two-Phase Heat Transfer Test Loop .............................. 4-100 

Cross-Sectional View of the Lehigh Test Bundle .................... 4-101 

Lehigh Rod Bundle Test-Section Schematic ........................ 4-102 

Lehigh Rod Bundle Facility Input Model Schematic ................. 4-103 

Comparison of Calculated and Measured Clad Temperatures (Z=0.15m) for 

Lehigh Reflood Test 02/24/85-20 ................................ 4-104 

Comparison of Calculated and Measured Clad Temperatures (Z=0.30m) for 

Lehigh Reflood Test 02/24/85-20 ................................ 4-105 

Comparison of Calculated and Measured Clad Temperatures (Z=0.45m) for 

Lehigh Reflood Test 02/24/85-20 ................................ 4-106 

Comparison of Calculated and Measured Clad Vapor Temperatures (Z=0.60m) 

for Lehigh Reflood Test 02/24/85-20 ............................. 4-107 

Comparison of Calculated and Measured Clad & Vapor Temperatures 

(Z=0.965m) for Lehigh Reflood Test 02/24/85-20 ................... 4-108

Page xvii
Figure 4.4.10 

Figure 4.4.11 

Figure 4.4.12 

Figure 4.4.13 

Figure 4.4.14 

Figure 4.4.15 

Figure 4.4.16 

Figure 4.4.17 

Figure 4.4.18 

Figure 4.4.19 

Figure 4.4.20 

Figure 4.4.21 

Figure 4.4.22 

Figure 4.4.23 

Figure 4.4.24 

Figure 4.4.25 

Figure 4.4.26 

Figure 4.4.27 

Figure 4.4.28 

Figure 4.4.29 

Figure 4.4.30 

Figure 4.4.31 

Figure 4.4.32 

Figure 4.4.33 

Figure 4.4.34 

Figure 4.4.35 

Figure 4.4.36

FLECHT-SEASET System Configuration ......................... 4-109 

FLECHT-SEASET 161 Rod Test Bundle .......................... 4-110 

FLECHT-SEASET Axial Power Shape ............................ 4-111 

Decay Power Curve (ANS + 20%) ............................... 4-112 

TRAC-M Input Model Schematic for FLECHT-SEASET Run 31504 .... 4-113 

Initial Axial Temperature Variation ............................... 4-114 

Initial Axial Steam Temperature Variation ......................... 4-115 

Comparison of Calculated and Experimental Clad Temperatures 

@ z= 1.22m .................................................. 4-116 

Comparison of Calculated and Experimental Clad Temperatures 

@ z=1.98m .................................................. 4-117 

Comparison of Calculated and Experimental Clad Temperatures 

@ z -2.59m .................................................. 4-118 

Comparison of Calculated and Experimental Clad Temperatures 

@ z=3.35m .................................................. 4-119 

Calculated and Measured Quench Front Propagation ................. 4-120 

Experimental Differential Pressure Measurements Along the Test Section 4-121 

Calculation of Axial Void Fraction Distribution by TRAC-M(F77) ...... 4-122 

Comparison of Differential Pressures (Taps 1 - 2) ................... 4-123 

Comparison of Differential Pressures (Taps 2 - 3) ................... 4-124 

Comparison of Differential Pressures (Taps 3 - 4) ................... 4-125 

Comparison of Differential Pressures (Taps 4 - 5) ................... 4-126 

Comparison of Differential Pressures (Taps 5 - 6) ................... 4-127 

Comparison of Differential Pressures (Taps 6 - 7) ................... 4-128 

Steam Temperatures (Z=1.83 - 1.94m) ............................ 4-129 

Steam Temperatures (Z=2.82 - 3.03m) ............................ 4-130 

Steam Temperatures (Z=3.35 - 3.5 1 m) ............................ 4-131 

Comparison of Test Section Pressures ............................. 4-132 

Comparison of Liquid Carryouts ................................. 4-133 

Comparison of Steam Carryouts ................................. 4-134 

TRAC-M(F77) Clad Temp. Predictions with and without Choking 

@ z= l.22m .................................................. 4-135

Page xviii
Figure 4.4.37 

Figure 4.4.38 

Figure 4.4.3 9 

Figure 4.4.40 

Figure 4.4.41 

Figure 4.4.42 

Figure 4.4.43 

Figure 4.4.44 

Figure 4.4.45 

Figure 4.4.46 

Figure 4.4.47 

Figure 4.4.48 

Figure 4.4.49 

Figure 4.4.50 

Figure 4.4.51 

Figure 4.4.52 

Figure 4.4.53 

Figure 4.4.54 

Figure 4.4.55 

Figure 4.4.56 

Figure 4.4.57 

Figure 4.4.58 

Figure 4.4.59 

Figure 4.4.60

TRAC-M(F77) Clad Temp. Predictions with and without Choking 

@ z=1.98m .................................................. 4-136 

TRAC-M(F77) Clad Temp. Predictions with and without Choking 

@ z=2.59m .................................................. 4-137 

TRAC-M(F77) Clad Temp. Predictions with and without Choking 

@ z=3.35m .................................................. 4-138 

TRAC-M Reflood Flow Regime Map ............................. 4-139 

Calculated Relative Velocities at Cell #10 .......................... 4-140 

Calculated Relative Velocities at Cell # 17 .......................... 4-141 

Isometric View of CCTF ....................................... 4-142 

Top View of Primary-Loop Piping Layout ......................... 4-143 

Cross-Section of the CCTF Pressure Vessel Through the Core-I Vessel Core and 

Primary-Loop Nozzles ......................................... 4-144 

CCTF Input Model Schematic ................................... 4-145 

TRAC Model of the CCTF Run 14 Power Profile and Instrument 

Locations ................................................... 4-146 

CCTF Run #14, Reflood Test - Comparison of TRAC-M(F77) and TRAC

M(F90) Calculations (Z=0.38m) ................................. 4-147 

CCTF Run #14, Reflood Test - Comparison of TRAC-M(F77) and TRAC

M(F90) Calculations (Z=l.105m) ................................ 4-148 

CCTF Run #14, Reflood Test - Comparison of TRAC-M(F77) and TRAC

M(F90) Calculations (Z=l.83m) ................................. 4-149 

CCTF Test Run #14, Reflood Test - Comparison of TRAC-M(F77) and TRAC

M(F90) Calculations (Z=2.44m) ................................. 4-150 

CCTF Run #14, Reflood Test - Comparison of TRAC-M(F77) and TRAC

M(F90) Calculations (Z=3.05m) ................................. 4-151 

Isometric View of CCTF ....................................... 4-152 

CCTF Pressure Vessel ......................................... 4-153 

Cross-Sections of the CCTF Pressure Vessel ....................... 4-154 

CCTF Heater-Rod Axial Power Profile ............................ 4-155 

Top View of Primary-Loop Piping Layout ......................... 4-156 

CCTF Pressure Vessel Noding: 2-Theta Model ...................... 4-157 

CCTF Intact-Loop Noding ...................................... 4-158 

CCTF Broken-Loop Noding .................................... 4-158

Page xix
Figure 4.4.61 

Figure 4.4.62 

Figure 4.4.63 

Figure 4.4.64 

Figure 4.4.65 

Figure 4.4.66 

Figure 4.4.67 

Figure 4.4.68 

Figure 4.4.69 

Figure 4.4.70 

Figure 4.4.71 

Figure 4.4.72 

Figure 4.4.73 

Figure 4.4.74 

Figure 4.4.75 

Figure 4.4.76 

Figure 4.4.77 

Figure 4.4.78 

Figure 4.4.79 

Figure 4.4.80 

Figure 4.4.81

CCTF, Core II, Run 54 - Comparison of Hot Rod Clad Temperature Predictions 

at Z= 2.48m .................................................. 4-159 

CCTF, Core II, Run 54 - Comparison of Hot Rod Clad Temperature Predictions 

at Z=3.115m ....... ..................................... 4-160 

CCTF, Core II, Run 54- Comparison of Hot Rod Clad Temperature Predictions 

at Z -3.93m .................................................. 4-161 

CCTF, Core II, Run 54 - Comparison of Hot Rod Clad Temperature Predictions 

at Z=4.54m .................................................. 4-162 

CCTF, Core II, Run 54 - Comparison of Hot Rod Clad Temperature Predictions 

at Z=5.15m .................................................. 4-163 

Overview of the Slab Core Test Facility ........................... 4-164 

Vertical Cross-Section View of SCTF Pressure Vessel ................ 4-165 

Axial Power Distribution of Heater Rods .......................... 4-166 

Pressure Vessel Noding Diagram ................................. 4-167 

TRAC Heater-Rod Noding, Thermocouple Locations, and Modeled 

Axial Pow er Shape ............................................ 4-168 

Hot-Leg and S/W-Separator Noding Diagram ...................... 4-169 

Intact Cold-Leg Noding Diagram ................................ 4-169 

Broken Cold-Leg Noding from Pressure Vessel to Containment Tank 1 .. 4-170 

Broken Cold-Leg Noding from S/W Separator to Containment Tank 2... 4-170 

SCTF, Run 719 - Comparison of Clad Temperature Calculations, Bundle 2 

(Z=1.829m ) ................................................. 4-171 

SCTF, Run 719 - Comparison of Clad Temperature Calculations, Bundle 2 

(Z=2.239m ) ................................................. 4-172 

SCTF, Run 719 - Comparison of Clad Temperature Calculations, Bundle 2 

(Z=2.669m ) ................................................. 4-173 

SCTF, Run 719 - Comparison of Clad Temperature Calculations, Bundle 2 

(Z=3.099m ) ................................................. 4-174 

SCTF, Run 719 - Comparison of Clad Temperature Calculations, Bundle 2 

(Z=3.454m ) ................................................. 4-175 

SCTF, Run 719 - Comparison of Clad Temperature Calculations, Bundle 2 

(Z=4.049m ) ................................................. 4-176 

SCTF, Run 719 - Comparison of Clad Temperature Calculations, Bundle 2 

(Z=4.479m ) ................................................. 4-177

Page xx
Figure 4.4.82 

Figure 4.4.83 

Figure 4.5.1 

Figure 4.5.2 

Figure 4.5.3 

Figure 4.5.4 

Figure 4.5.5 

Figure 4.5.6 

Figure 4.5.7 

Figure 4.5.8 

Figure 4.5.9 

Figure 4.5.10 

Figure 4.5.11 

Figure 4.5.12 

Figure 4.5.13 

Figure 4.5.14 

Figure 4.5.15 

Figure 4.5.16 

Figure 4.5.17 

Figure 4.5.18 

Figure 4.5.19 

Figure 4.5.20 

Figure 4.6.1 

Figure 4.6.2 

Figure 4.6.3 

Figure 4.6.4 

Figure 4.6.5

SCTF, Run 719 - Comparison of Clad Temperature Calculations, Bundle 2 

(Z=4.909m ) ................................................. 4-178 

SCTF, Run 719 - Comparison of Clad Temperature Calculations, Bundle 2 

(Z=5.292m ) ................................................. 4-179 

UPTF Flow Diagram .......................................... 4-185 

UPTF Primary System ......................................... 4-186 

Major Dimensions of the UPTF Primary System .................... 4-187 

Plan View of UPTF Test Vessel .................................. 4-188 

System Configuration for UPTF Test 8, Part 1 ...................... 4-189 

UPTF8B Vessel Model Elevation View ............................ 4-190 

UPTF8B Vessel Model Plan View ................................ 4-191 

UPTF Loop-1 Noding Diagram .................................. 4-191 

UPTF Loop-2 Noding Diagram .................................. 4-192 

UPTF Loop-3 Noding Diagram .................................. 4-192 

UPTF Broken-Loop-4 Cold-Leg Noding Diagram ................... 4-193 

UPTF Broken-Loop-4 Hot-Leg Noding Diagram .................... 4-193 

UPTF Water Drainage System Noding Diagram ..................... 4-194 

UPTF Core-Simulator-Injection Noding Diagram ................... 4-194 

UPTF, Test 8B, Part 1 - Vessel Downcomer Pressure (Cell 2, 1, 12) ..... 4-195 

UPTF, Test 8B, Part 1 - Vessel Upper Plenum Pressure (Cell 1, 1, 12) ... 4-196 

UPTF, Test 8B, Part 1 - Liquid Height in Vessel (Signal Var. 106) ...... 4-197 

UPTF, Test 8B, Part 1 - Liquid Temp. Cl. Leg Loop 2, Downstream ECC 

Injection, (Comp. 24, Cell 18) ................................... 4-198 

UPTF, Test 8B, Part 1 - Steam Flow in Loop 3 (Loop Seal, Comp. 34, 

C ell 5) ..................................................... 4-199 

UPTF, Test 8B, Part 1 - Steam Flow in Loop 2 (Loop Seal, Comp. 24, 

C ell 5) ..................................................... 4-200 

UPTF Flow Diagram .......................................... 4-205 

Overview of UPTF Primary System .............................. 4-206 

Major Dimensions of the UPTF Primary System .................... 4-207 

Plan View of UPTF Test Vessel .................................. 4-208 

System Configuration for UPTF Test 6 ............................ 4-209

Page xxi
Figure 4.6.6 

Figure 4.6.7 

Figure 4.6.8 

Figure 4.6.9 

Figure 4.6.10 

Figure 4.6.11 

Figure 4.6.12 

Figure 4.6.13 

Figure 4.6.14 

Figure 4.6.15 

Figure 4.6.16 

Figure 4.6.17 

Figure 4.6.18 

Figure 4.6.19 

Figure 4.6.20 

Figure 4.6.21 

Figure 4.7.1 

Figure 4.7.2 

Figure 4.8.1 

Figure 4.8.2 

Figure 4.8.3 

Figure 4.8.4 

Figure 4.8.5 

Figure 4.8.6 

Figure 4.8.7 

Figure 4.8.8 

Figure 4.8.9

Vessel Model Elevation View for uIPTF Test 6 ...................... 4-210 

Vessel Model Plan View for UPTF Test 6 .......................... 4-211 

UPTF Loop-1 Noding Diagram .................................. 4-212 

UPTF Loop-2 Noding Diagram .................................. 4-213 

UPTF Loop-3 Noding Diagram .................................. 4-214 

UPTF Loop-4 Broken-Hot-Leg Noding Diagram .................... 4-215 

UPTF Loop-4 Broken-Cold-Leg Noding Diagram ................... 4-216 

UPTF Water Drainage System Noding Diagram ..................... 4-217 

UPTF Core-Simulator-Injection Noding Diagram ................... 4-218 

UPTF, Test 6, Run 133 - Comparison of Downcomer Pressures at Mid

E levation ................................................... 4-219 

UPTE, Test 6, Run 133 - Downcomer Liquid Temperatures at Mid

Elevation ................................................... 4-220 

UPTF, Test 6, Run 133 - Downcomer Steam Temperature at Mid

Elevation ................................................... 4-221 

UPTF, Test 6, Run 133 - Downcomer Void Fraction at Bottom Elevation. 4-222 

UPTF Test 6, Run 133 - Downcomer Void Fraction at Mid-Elevation.... 4-223 

UtPTF, Test 6, Run 133 - Downcomer Void Fraction at the Top Elevation. 4-224 

UPTF, Test 6, Run 133 - Comparison of Vessel Mass Calculations ...... 4-225 

Experimental Data, Clad Temperatures at Z=1.52m .................. 4-229 

Comparison of Predicted and Measured Clad Temperatures ............ 4-230 

Overview of the LOFT Facility .................................. 4-237 

Piping Schematic of the LOFT Facility ............................ 4-238 

Intact-Loop Steam Generator in the LOFT Facility .................. 4-239 

LOFT L2-6 Reactor Vessel Noding Diagram ....................... 4-240 

LOFT L2-6 Intact-Loop Noding Diagram .......................... 4-241 

LOFT L2-6 Broken-Loop Noding Diagram ........................ 4-242 

LOFT L2-6 Steam-Generator Noding Diagram ..................... 4-243 

Comparison of Calculated and Measured Pressurizer Pressures ......... 4-244 

Comparison of Calculated and Measured Broken-Loop Hot Leg Mass 

Flow R ates .................................................. 4-245

Page xxii
Figure 4.8.10 

Figure 4.8.11 

Figure 4.8.12 

Figure 4.8.13 

Figure 4.9.1 

Figure 4.9.2 

Figure 4.9.3 

Figure 4.9.4 

Figure 4.9.5 

Figure 4.9.6 

Figure 4.9.7 

Figure 4.9.8 

Figure 4.9.9 

Figure 4.9.10 

Figure 4.9.11 

Figure 4.9.12 

Figure 5.1.1 

Figure 5.1.2 

Figure 5.1.3 

Figure 5.1.4 

Figure 5.1.5 

Figure 5.1.6 

Figure 5.1.7 

Figure 5.1.8 

Figure 5.1.9 

Figure 5.1.10

Comparison of Calculated and Measured Broken-Loop Cold Leg Mass 

Flow Rates .................................................. 4-246 

Comparison of Calculated and Measured Intact Loop Hot Leg Mass Flow 

R ates ....................................................... 4-247 

Comparison of Calculated and Measured Intact Loop Cold Leg Mass 

Flow Rate ................................................... 4-248 

Comparison of Calculated and Measured Rod Clad Temperatures ....... 4-249 

Intact-Loop Steam Generator in the LOFT Facility .................. 4-259 

TRAC Model of the LOFT Facility ............................... 4-260 

TRAC Model for the LOFT Facility Intact-Loop Steam Generator ...... 4-261 

3-D Vessel, Fueled-Core Rod, and Structure Rod Components Modeling the 

LOFT Facility Vessel .......................................... 4-262 

Comparison of Core Powers .................................... 4-263 

Comparison of Vessel Upper-Plenum Pressures ..................... 4-264 

Comparison of Flow Rates in the Hot Leg of the Loop on the Steam Generator 

Side ....................................................... 4-265 

Comparison of Steam Generator Primary Side Inlet Liquid Temperatures. 4-266 

Comparison of Steam Generator Primary Side Outlet Liquid 

Temperatures ................................................ 4-267 

Comparison of Steam Flow Rates ................................ 4-268 

Comparison of Steam Line Control Valve Flow Area Restrictions ....... 4-269 

Comparison of Steam Dome Pressures ............................ 4-270 

Typical Two-Zone Feedwater Heater (condensing and subcooling zones).. 5-10 

Sketch of a Simple Feedwater Heater .............................. 5-11 

Heat Balance and Flows for the First and Second Point Heaters of River Bend 

Station ...................................................... 5-12 

Temperature Path Length Diagram for a Two-Zone Closed Heater ....... 5-12 

Schematic of a Two-Zone Closed Feedwater Heater ................... 5-13 

Comparison of Shell and Drain Cooler Pressures ..................... 5-13 

Shell Liquid Level ............................................. 5-14 

Cell Edge Void Fraction Defined at the Drain Cooler Inlet .............. 5-14 

Timestep Size Used by TRAC-B and TRAC-M ...................... 5-15 

Drain Valve Flow Area Driven by Equation 5.1.1 ..................... 5-15

Page xxiii
Figure 5.1.11 

Figure 5.1.12 

Figure 5.1.13 

Figure 5.1.14 

Figure 5.1.15 

Figure 5.1.16 

Figure 5.1.17 

Figure 5.1.18 

Figure 5.2.1 

Figure 5.2.2 

Figure 5.2.3 

Figure 5.2.4 

Figure 5.2.5 

Figure 5.2.6 

Figure 5.2.7 

Figure 5.2.8 

Figure 5.2.9 

Figure 5.2.10 

Figure 5.2.11 

Figure 5.2.12 

Figure 5.2.13 

Figure 5.3.1 

Figure 5.3.2 

Figure 5.3.3 

Figure 5.3.4 

Figure 5.3.5 

Figure 5.3.6 

Figure 5.3.7 

Figure 5.3.8 

Figure 5.3.9

Shell Condensate Level ......................................... 5-16 

Condensate Level (no condensate flow from the HP heater) ............ 5-16 

Drain Flow Rate (no condensate flow from the HP heater) ............. 5-17 

Drain Cooler Approach (no condensate flow from the HP heater) ........ 5-17 

Feedwater Temperature Rise (no condensate flow from the HP heater) .... 5-18 

D rain Flow Rate ............................................... 5-18 

Drain Cooler Approach ......................................... 5-19 

Feedwater Temperature Rise ..................................... 5-19 

Schematic of a JETP Component ................................. 5-32 

Comparison of Discharge and Drive Line Pressures ................... 5-32 

Bernoulli Losses and Gains Across Diffuser and Nozzle ............... 5-33 

Mixing of Suction and Drive Line Flows ........................... 5-33 

Jet Pump Flow Regimes ........................................ 5-34 

INEL Tests with Positive Drive Flow .............................. 5-35 

INEL Tests with Negative Drive Flow ............................. 5-35 

Comparison of TRAC-B and TRAC-M vs. the INEL Data ............. 5-36 

Comparison of TRAC-B and TRAC-M vs. the INEL Data ............. 5-36 

Pressures of Single vs. Lumped Jet Pump(s) ......................... 5-37 

Velocities of Single vs. Lumped Jet Pump(s) ........................ 5-37 

Mass Flow Rate of Single vs. Lumped Jet Pump(s) ................... 5-38 

M-N Ratio for Full-Scale Jet Pumps ............. ................. 5-38 

Boiling-Water Reactor Schematic ................................. 5-54 

Hydraulic and Heat Transfer Test Nodalization Diagram ............... 5-55 

Results of Requirement 1 Test .................................... 5-56 

CHAN Cell Leaking to a VESSEL Cell ............................ 5-57 

Requirem ent 3 (1-D Test) ....................................... 5-57 

Requirement 3 (1-D Results) ..................................... 5-58 

3-D Requirement 3 Nodalization Diagram .......................... 5-59 

3-D Results for Requirement 3 ................................... 5-59 

One Cell CHAN with FILL and BREAK ........................... 5-60

Page xxiv
Figure 5.3.10 

Figure 5.3.11 

Figure 5.3.12 

Figure 5.3.13 

Figure 5.3.14 

Figure 5.3.15 

Figure 5.3.16 

Figure 5.3.17 

Figure 5.3.18 

Figure 5.3.19 

Figure 5.3.20 

Figure 5.4.1 

Figure 5.4.2 

Figure 5.4.3 

Figure 5.4.4 

Figure 5.4.5 

Figure 5.4.6 

Figure 5.4.7 

Figure 5.4.8 

Figure 5.4.9 

Figure 5.4.10 

Figure 5.4.11 

Figure 5.4.12 

Figure 5.4.13 

Figure 5.4.14

Effect of Multiple Rods (Requirement 4) on Rod Heat Up .............. 5-60 

Vessel and Canister Wall Temperatures with 1 CHAN Compared to 10 .... 5-61 

TRAC-B Results for Requirement 6 ............................... 5-62 

TRAC-M Results for Requirement 6 ............................... 5-63 

Geometry of GOTA Radiation Test Section ......................... 5-64 

GOTA Radiation Test 27 Results .................................. 5-64 

Effect of Canister Quenching on Rod Radiation (Requirement 7.1) ....... 5-65 

GOTA Radiation Test 27 Results with 5 Rod Groups .................. 5-66 

Demonstration of Radiation Cut-Off (Requirement 7.6) ................ 5-67 

Demonstration of Energy Absorption by Steam (Requirement 7.7) ....... 5-68 

Effect of Water on the Radiation to the Canister Wall (Requirement 7.7) .. 5-69 

Flow Regime Dilemma ......................................... 5-93 

Two-Phase Level in Vertical Flows ................................ 5-93 

1-D and 3-D Component Connection with Two-Phase Level ............ 5-94 

Flux Terms Above and Below a Two-Phase Level .................... 5-94 

Single-Pipe Fill and Drain Test Problem ............................ 5-95 

Pipe Fill & Drain Test, Water Level Height, TRAC-BF1 with Level 

Tracking ..................................................... 5-95 

Pipe Fill & Drain Test, Water Level Height, TRAC-M with Level 

Tracking ..................................................... 5-96 

Pipe Fill & Drain Test, Pressure in Cell #1, TRAC-BFI without Level 

Tracking ..................................................... 5-96 

Pipe Fill & Drain Test, Pressure in Cell #1, TRAC-BFI with Level 

Tracking ..................................................... 5-97 

Pipe Fill & Drain Test, Pressure in Cell #1, TRAC-M without Level 

Tracking ..................................................... 5-97 

Pipe Fill & Drain Test, Pressure in Cell #1, TRAC-M with Level 

Tracking ..................................................... 5-98 

Valve Fill & Drain Test, Pressure in Cell #1, TRAC-M with Level 

Tracking ..................................................... 5-98 

Prizer Fill & Drain Test, Pressure in Cell #1, TRAC-M with Level 

Tracking ..................................................... 5-99 

Pump Fill & Drain Test, Pressure in Cell#I, TRAC-M with Level 

Tracking ..................................................... 5-99

Page xxv
Figure 5.4.15 

Figure 5.4.16 

Figure 5.4.17 

Figure 5.4.18 

Figure 5.4.19 

Figure 5.4.20 

Figure 5.4.21 

Figure 5.4.22 

Figure 5.4.23 

Figure 5.4.24 

Figure 5.4.25 

Figure 5.4.26 

Figure 5.4.27 

Figure 5.4.28 

Figure 5.4.29 

Figure 5.4.30 

Figure 5.4.31 

Figure 5.4.32 

Figure 5.4.33 

Figure 5.4.34 

Figure 5.4.35

Intentionally Left Blank ........................................ 5-100 

TEE Component Fill and Drain Test Problem Nodalization ............ 5-100 

TEE Fill & Drain Test, Main Tube Cell #1 Pressure, TRAC-M with Level 

Tracking .................................................... 5-101 

TEE Fill & Drain Test, Side Ann Bottom Cell Pressure, TRAC-M with Level 

Tracking .................................................... 5-101 

Two-Pipe Fill & Drain Test, Bottom Cell Pressure, TRAC-M with Level 

Tracking .................................................... 5-102 

Single-Ring Vessel Fill & Drain Test, Water Level Height TRAC-B with Level 

Tracking .................................................... 5-102 

Single-Ring Vessel Fill & Drain Test, Water Level Height, TRAC-M with Level 

Tracking .................................................... 5-103 

Single-Ring Vessel Fill & Drain Test, Bottom Cell Pressure, TRAC-B with Level 

Tracking .................................................... 5-104 

Single-Ring Vessel Fill & Drain Test, Bottom Cell Pressure, TRAC-B without 

Level Tracking ............................................... 5-104 

Single-Ring Vessel Fill & Drain Test, Bottom Cell Pressure, TRAC-M with 

Level Tracking ............................................... 5-105 

Single-Ring Vessel Fill & Drain Test, Bottom Cell Pressure, TRAC-M without 

Level Tracking ............................................... 5-105 

Multiple-Ring/Azimuthal Vessel Component Fill and Drain Test Problem 5-106 

Two-Ring Vessel Fill & Drain Test, Water Level Height, TRAC-M with Level 

Tracking .................................................... 5-106 

Two-Vessel Fill & Drain Test, Water Level Height, TRAC-M with Level 

Tracking .................................................... 5-107 

Inverted Void Fraction Profile Level Tracking Test .................. 5-108 

Inverted Void Profile Test, Water Level Height, TRAC-M with Level 

Tracking .................................................... 5-108 

Inverted Void Profile in Vessel Test, Water Level Height, TRAC-M with Level 

Tracking .................................................... 5-109 

3-D to 1-D Above-Level Suction Test ............................. 5-110 

Vessel Above-Level Suction Test, Void Fraction .................... 5-110 

3-D to 1 -D Below-Level Suction Test .............................. 5-111 

Vessel Below-Level Suction Test, Void Fraction in Level 6 of the Vessel and 

Cell #1 of the Pipe ............................................ 5-112

Page xxvi
Figure 5.4.36 

Figure 5.4.37 

Figure 5.4.38 

Figure 5.4.39 

Figure 5.4.40 

Figure 5.4.41 

•Figure 5.4.42 

Figure 5.4.43 

Figure 5.4.44 

Figure 5.4.45 

Figure 5.4.46 

Figure 5.5.1 

Figure 5.5.2 

Figure 5.5.3 

Figure 5.5.4 

Figure 5.5.5 

Figure 5.5.6 

Figure 5.5.7 

Figure 5.5.8 

Figure 5.5.9 

Figure 5.5.10 

Figure 5.5.11 

Figure 5.5.12 

Figure 5.5.13 

Figure 5.6.1 

Figure 5.6.2

Single-Pipe Cold Water Flooding Test, Steam Flow Rate TRAC-B without Level 

Tracking .................................................... 5-112 

Single-Pipe Cold Water Flooding Test, Steam Flow Rate, TRAC-B with Level 

Tracking .................................................... 5-113 

Single-Pipe Cold Water Flooding Test, Steam Flow Rate, TRAC-M with Level 

Tracking .................................................... 5-113 

Single-Ring Vessel Cold Water Flooding Test, Steam Flow Rate, TRAC-M 

without Level Tracking ........................................ 5-114 

Single-Ring Vessel Cold Water Flooding Test, Steam Flow Rate, TRAC-M with 

Level Tracking ............................................... 5-114 

Complete PIPE Fill and Drain Transient ........................... 5-115 

First Half of the Transient ...................................... 5-115 

Second Half of the Transient .................................... 5-116 

Complete Single-Ring Vessel Fill and Drain Transient ................ 5-116 

First Half of the Single-Ring Vessel Fill and Drain Transient ........... 5-117 

Second Half of the Single-Ring Vessel Fill and Drain Transient ........ 5-117 

TRAC-M and Control System Data Passing Flow Diagram ............ 5-140 

TRAC-BF1 Control System Data Passing Flow Diagram .............. 5-140 

Test Problem Control Block Diagram ............................. 5-141 

Implicit Loop Solution Test Problem .............................. 5-142 

TRAC-M and TRAC-BF1 Implicit Loop Solution Scheme ............ 5-142 

TRAC-M Control System Time Step Size Sensitivity Test ............. 5-143 

Lumped Flow Controller Test Problem ............................ 5-144 

Lumped Flow Controller Test Results ............................. 5-145 

Lumped Level Controller Test Problem ........................... 5-146 

Lumped Level Controller Test Results ............................ 5-147 

Lumped Pressure Controller Test Problem ......................... 5-148 

Lumped Pressure Controller Test Results .......................... 5-149 

Mixture Enthalpy and Collapsed Water Level Test Problem ............ 5-150 

Representative Component Network .............................. 5-159 

Noding Diagram for End Junction Replacement Test Cases - Parametric 1 

(B ase Case) ................................................. 5-159

Page xxvii
Figure 5.6.3 

Figure 5.6.4 

Figure 5.6.5 

Figure 5.6.6 

Figure 5.6.7 

Figure 5.6.8 

Figure 5.6.9 

Figure 5.6.10 

Figure 5.6.11 

Figure 5.6.12 

Figure 5.6.13 

Figure 5.6.14 

Figure 5.6.15 

Figure 5.6.16 

Figure 5.6.17 

Figure 5.6.18 

Figure 5.6.19 

Figure 5.6.20 

Figure 5.6.21 

Figure 5.6.22 

Figure 5.6.23 

Figure 5.6.24

Noding Diagram for End-Junction Replacement Test Cases - Parametric 2 

(Reverse Orientation) .......................................... 5-160 

Noding Diagram for End-Junction Replacement Test Cases - Parametric 3 

(Non-Zero Gravity) ........................................... 5-160 

Noding Diagrams for Composite TEE Test Cases - Parametric 1 

(B ase C ase) ................................................. 5-161 

Noding Diagrams for Composite TEE Test Cases - Parametric 2 (In Flow at 

F ills) ....................................................... 5-162 

Noding Diagrams for Composite TEE Test Cases - Parametric 3 (Reversed SJC 

O rientation) ................................................. 5-163 

Noding Diagrams for Composite TEE Test Cases - Parametric 4 (Non-Zero 

Gravity Term s) ............................................... 5-1⁄464 

Noding Diagrams for Composite TEE Test Cases - Parametric 5 (Non-Zero 

Elevation Term s) ............................................. 5-165 

Noding Diagrams for Composite TEE Test Cases with Single-Cell Main Leg 

Param etric 7 ................................................. 5-166 

Noding Diagrams for Link-to-VESSEL Test Cases - Parametric 1 

(B ase Case) ................................................. 5-167 

Noding Diagrams for Link-to-VESSEL Test Cases with BWR Channel 

Simulation - Parametric 1 (Static Check) .......................... 5-168 

Noding Diagrams for Link-to-VESSEL Test Cases with BWR Channel 

Simulation - Parametric 2 (Cold Leg In-flow) ....................... 5-169 

Relative Difference in Leakage Flows for the SJC and Non-SJC Models . 5-170 

Noding Diagrams for Pump Test Cases - Parametric 1 (Base Case) ...... 5-170 

Relative Difference in Pump Flows for the SJC and Non-SJC Models ... 5-171 

Noding Diagrams for VALVE Test Cases .......................... 5-171 

Relative Difference in Valve Flows for the SJC and Non-SJC Models .... 5-172 

Noding Diagrams for VALVE Test Cases Next to BREAK - Parametric 2 

(VALVE Next to BREAK) ...................................... 5-172 

Relative Difference in Valve Flows for the SJC and Non-SJC Models .... 5-173 

Noding Diagrams for TFPIPE2 Test Cases - Parametric 1 (Base) ....... 5-174 

Noding Diagrams for TFPIPE2 Test Cases - Parametric 2 (Reversed 

O rientation) ................................................. 5-175 

Noding Diagrams for PLENUM Test Cases - Parametric 1 (Base Case) .. 5-176 

Relative Difference in Side-Leg Flows for the SJC and Non-SJC Models. 5-177

Page xxviii
Figure 5.7.1 

Figure 5.7.2 

Figure 5.7.3 

Figure 5.7.4 

Figure 5.7.5 

Figure 5.7.6 

Figure 5.7.7 

Figure 5.7.8 

Figure 5.7.9 

Figure 5.7.10 

Figure 5.7.11 

Figure 5.7.12 

Figure 5.7.13 

Figure 5.7.14 

Figure 5.7.15 

Figure 6.1.1 

Figure 6.1.2 

Figure 6.1.3 

Figure 6.1.4 

Figure 6.1.5 

Figure 6.1.6 

Figure 6.1.7 

Figure 6.2.1 

Figure 6.2.2 

Figure 7.1.1 

Figure 7.1.2 

Figure 7.1.3 

Figure 7.2.1 

Figure 7.2.2 

Figure 7.2.3

Turbine Nodalization .......................................... 5-191 

Entropy-Enthalpy Process Diagram ............................... 5-191 

Simple Turbine Test Model ..................................... 5-192 

Pressure Comparison for TURB1 ................................ 5-192 

Mass Conservation Test (Semi-Implicit) ........................... 5-193 

Mass Conservation Test with Courant Violating Numerics ............. 5-193 

TRAC-B and TRAC-M Nozzle Velocities ......................... 5-194 

Pressure Response in Second Cell of Turbine ....................... 5-194 

Pressure Response for TRAC-B and TRAC-M with 0.1, 1.0, and 5.0 Second 

Tim e Steps .................................................. 5-195 

Turbine System M odel ......................................... 5-196 

Comparison of TRAC-M and TRAC-B Liquid Drain Velocities ........ 5-197 

Comparison of Drain Models for Low Flow ........................ 5-198 

System Pressure Comparison .................................... 5-198 

System Velocity Comparisons ................................... 5-199 

System Vapor Temperature Comparison ........................... 5-199 

Diagram of Interface Implementation ............................... 6-9 

Structure of General Interface Module .............................. 6-9 

Auxiliary Data Structure ........................................ 6-10 

Cartesian Thermal-Hydraulic to Neutronic Mapping .................. 6-10 

Cylindrical to Cartesian Mapping ................................. 6-10 

Coupling with PVM ............................................ 6-11 

Example of Kinetics Controlling Time Step ......................... 6-11 

Problem Nodalization for Base Test Case ........................... 6-33 

Thermal-Hydraulic Nodalization for Annulus Test Case ................ 6-33 

Schematic of the AP600 System ................................... 7-4 

Comparison of Pressurizer Pressure Calculations ...................... 7-5 

Comparison of Break Flow Rate Calculations ........................ 7-6 

Reactor Core Cross-Sectional View ............................... 7-16 

Two-Dimensional Assembly Type Map (Half Core) ................... 7-17 

Mapping of Fuel Assemblies to CHAN Components .................. 7-18

Page xxix
Figure 7.2.4 

Figure 7.2.5 

Figure 7.2.6 

Figure 7.2.7 

Figure 7.2.8 

Figure 7.2.9 

Figure 7.2.10

Steady-State Convergence During Initialization of TRAC-M Model ...... 

Axial Power Distribution for the PBTT2 Initial Conditions ............. 

Steady-State Radial Power Distribution of the PBTT2 ................. 

Upper Plenum Pressure During PBTT2 ............................. 

PBTT2 Core Average Power and Void Fraction Predicted by TRAC-M .. 

Comparison of Measured and TRAC-M Predicted Core Reactivity ....... 

Comparison of Measured and TRAC-M Predicted Core Average Power...

Page xxx

7-19 

7-20 

7-21 

7-22 

7-22 

7-23 

7-23
TABLES

Table 3.1.1 

Table 3.1.2 

Table 3.1.3 

Table 3.1.4 

Table 4.2.1 

Table 4.3.1 

Table 4.3.2 

Table 4.4.1 

Table 4.4.2 

Table 4.4.3 

Table 4.4.4 

Table 4.4.5 

Table 4.4.6 

Table 4.4.7 

Table 4.4.8 

Table 4.4.9 

Table 4.4.10 

Table 4.5.1 

Table 4.6.1 

Table 4.7.1 

Table 4.8.1 

Table 4.8.2 

Table 4.9.1

Comparison Between TRAC Codes and Analytical Solutions for a 1-D Heat

Conduction Problem in the Radial Direction at Steady State (TRAC Input Model 

HCOND2) ............................................ ....... 3-6 

Comparison Between TRAC Codes and Analytical Solutions Along a Center 

Line of a Rod in a 2-D Heat-Conduction Problem While at Steady State with 

Fine Mesh (TRAC Input Model HCOND3) .......................... 3-7 

Comparison Between TRAC Codes and Analytical Solutions Along a Center 

Line of a Rod in a 2-D Heat-Conduction Problem While at Steady State with 

Fixed Noding (TRAC Input Model HCOND1) ........................ 3-8 

Comparison Between TRAC Codes and Analytical Solutions Along a Radius at 

Midplane (Elevation 100 mm) of a Rod in a 2-D Heat-Conduction Problem at 

Steady State (TRAC Input Model HCOND1) ......................... 3-9 

Data Comparison in the Junction Cell of the Test Section .............. 4-17 

Marviken Tests and TRAC Noding ................................ 4-39 

Comparison of Measured Mass Fluxes and the Mass Fluxes Calculated by 

TRAC-M(F77) and Different Critical Flow Models ................... 4-40 

Predicted and Measured Quench Times ............................. 4-91 

Predicted and Measured Quench Velocities .......................... 4-91 

Conditions and Key Results for Run 31504 ......................... 4-92 

Flow Regime Transition Criteria .................................. 4-92 

Operating Parameters for CCTF Run 14 ............................ 4-93 

CCTF Component Scaled Dimensions ............................. 4-94 

Component Elevations of the CCTF ............................... 4-96 

Initial Test Conditions .......................................... 4-97 

Chronology of Testing Events for CCTF Run 54 ..................... 4-98 

Summary of Test Conditions for SCTF Run 719 (S3-15) ............... 4-99 

Transient Conditions .......................................... 4-184 

Initial Test Conditions ......................................... 4-204 

Initial Conditions for FLECHT-SEASET Run #32753 ................ 4-228 

LOFT Test L2-6 Initial Conditions and Initial Test Data .............. 4-236 

LOFT Test L2-6 Sequence of Events .............................. 4-236 

Trip Set Points for Experiment L6-1 .............................. 4-256

Page xxxi

x -
Table 4.9.2 

Table 4.9.3 

Table 5.1.1 

Table 5.2.1 

Table 5.2.2 

Table 5.2.3 

Table 5.2.4 

Table 5.2.5 

Table 5.3.1 

Table 5.3.2 

Table 5.3.3 

Table 5.3.4 

Table 5.3.5 

Table 5.3.6 

Table 5.3.7 

Table 5.4.1 

Table 5.4.2 

Table 5.4.3 

Table 5.5.1 

Table 5.5.2 

Table 5.5.3 

Table 5.5.4 

Table 5.5.5 

Table 5.5.6 

Table 5.7.1 

Table 5.7.2 

Table 5.7.3 

Table 5.7.4 

Table 6.2.1

LOFT Experiment L6-1 Initial Conditions and TRAC-Evaluated Steady-State 

Calculation Results ........................................... 4-257 

LOFT Experiment L6-1 Sequence of Events ....................... 4-258 

Typical Feedwater Heater Boundary Conditions ....................... 5-9 

The INEL Tests with Positive Drive Flow ........................... 5-29 

Boundary Conditions for Negative Flow Tests ....................... 5-29 

The INEL Tests with Negative Drive Flow .......................... 5-30 

Boundary Conditions for Positive Flow Tests ........................ 5-30 

Full-Scale Jet Pump Dimensions .................................. 5-31 

TRAC-B CHAN Models That Will Not Be Ported .................... 5-49 

CHAN Requirem ents ........................................... 5-50 

CHAN Radiation Requirements .................................. 5-51 

TRAC-B View Factors for GOTA Test with 5 Rod Groups ............. 5-52 

TRAC-M View Factors for GOTA Test with 5 Rod Groups ............. 5-52 

TRAC-B Beam Lengths for GOTA Test with 5 Rod Groups ............ 5-52 

TRAC-M Beam Lengths for GOTA Test with 5 Rod Groups ............ 5-53 

LTRCK Nomenclature .......................................... 5-90 

Requirements for Two-Phase Level Tracking Model .................. 5-91 

Test Problem Summary ......................................... 5-92 

Control System Consolidation Software Requirement Specifications .... 5-134 

Control System Test Problem List ................................ 5-135 

TRAC-BF1 Control Block Sorting Results ......................... 5-136 

TRAC-M Control System Sorting Results ......................... 5-137 

TRAC-M Control Block Sorting Results with 1 CsSort=2 ............. 5-138 

Test Problem 9 Signal Variable Definitions and Results ............... 5-139 

Nomenclature Used in Equations Related to TURB Component Design.. 5-189 

TURB Requirements .......................................... 5-190 

Energy Error in TRAC-M ...................................... 5-190 

Nozzle Velocity Comparison .................................... 5-190 

Test Case Matrix for General Interface ............................. 6-32

Page xxxii
Table 6.3.1 

Table 6.3.2 

Table 6.3.3 

Table 6.3.4 

Table 6.3.5 

Table 6.3.6 

Table 6.3.7 

Table 7.2.1 

Table 7.2.2 

Table 7.2.3 

Table 7.2.4 

Table 7.2.5

Matrix of Test Results for Case 1 ................................. 6-34 

Matrix of Test Results for Case 2 ................................. 6-37 

Matrix of Test Results for Case 3 ................................. 6-39 

Matrix of Test Results for Case 4 ................................. 6-40 

Matrix of Test Results for Case 5 ................................. 6-41 

Matrix of Test Results for Case 6 ................................. 6-42 

Matrix of Test Results for Case 7 ................................. 6-43 

PB2 Fuel Assembly Data ........................................ 7-11 

PB2 Reference Design Information ................................ 7-12 

Reactor Core Layout ........................................... 7-13 

Steady-State Parameters ........................................ 7-14 

Comparison of T/H Steady-State Parameters for PBTT2 ............... 7-15

Page xxxiii
EXECUTIVE SUMMARY

This report documents assessment findings, which demonstrate that the pressurized-water 

reactor (PWR) version of the Transient Reactor Analysis Code (TRAC-P) has successfully been 

modernized. This assessment was performed by comparing the code predictions that were 

produced using TRAC-P(Version 5.50), which is generically designated as TRAC-M(F77) 

because it was written in FORTRAN 77, and the modernized TRAC-M(Version 3690), which is 

generically designated as TRAC-M(F90) because it is written in FORTRAN 90. Specifically, the 

results reflect the predictions obtained for a series of test cases, which are presented in Chapter 3, 

"Analytical Test Problems," Chapter 4, "PWR Developmental Assessment," and Chapter 7, 

"Plant Transient Calculations," of this report. 

The test matrix for "Analytical Test Problems" contains 42 cases, all of which are relatively 

simple and small problems designed to exercise various parts of the code, with an estimated 

coverage of about 80%. (Chapter 3 discusses five representative test cases. The results of other 

test cases are available to interested parties.) These test cases are exercised as each version of the 

code is created, to ensure that new errors have not been introduced as a result of the 

modernization. The resulting code predictions are then compared, and the comparisons show 

either "Excellent" or "Reasonable" agreement. 

The test matrix for "PWR Developmental Assessment" contains 20 test cases that also 

exercise various parts of the code. These cases are more complicated than the "Analytical Test 

Problems," in that they test the coding and modeling of two-phase flow phenomena in PWR plant 

system transients. Again, comparisons between the code predictions produced using TRAC

M(F77) and TRAC-M(F90) show either "Excellent" or "Reasonable" agreement. Most of these 

test cases also provide comparison with the test data. Since the modernization did not change the 

physical models used in the code, agreements between the code predictions and the test data 

remain the same as those presented in the report prepared by Los Alamos National Laboratory, 

LA-UR-99-6480, "TRAC-M: FORTRAN 77, Version 5.5, Developmental Assessment Manual, 

Volume I: Nonproprietary Assessment Sections," dated December 1999. Like the foregoing 

report, this report finds that the reflood model needs improvement. It should also be noted that 

the test data obtained in the 2-D/3-D program are not presented in this report, since the data are 

restricted for exclusive use by 2-D/3-D program participants. When the participants agree to 

release these test data, this report will be revised accordingly. 

Finally, a limited number of plant transients are calculated using both TRAC-M(F77) and 

TRAC-M(F90). Plant transients are usually long-running; therefore, it is not possible to include a 

large number of transients in a report that is prepared in relatively a short time. Comparisons 

between the resulting code predictions for the transients presented show "Excellent" or 

"Reasonable" agreement. 

The consolidated TRAC-M(F90) integrates many boiling water reactor (BWR) components, 

as well as one component that is designed to provide a capability similar to the Reactor Excursion 

and Leak Analysis Program (RELAP). In addition, TRAC-M(F90) integrates some BWR

specific models. Specifically, TRAC-M(F90) integrates the following components, models and 

modules:

Page xxxv
Components:

* Channel (CHAN) component "* Containment (CONTAN) component 

"• Feedwater heater (HEATR) component 

* Jet pump (JETP) component "* Separator (SEPT) component 

"* Single-junction (SJACS) component 

"* Turbine (TRBN) component 

Models: 

"* BWR control system (CNSYS) model 

"* Level tracking (LTRCK) model 

Modules: 

. Power module 

0 Spatial kinetics module 

The assessment of the integration of the CHAN, HEATR, JETP, SJACS, and TRBN 

components, as well as the CONSYS and LTRCK models and the 3-D Spatial Kinetics module is 

reported in Chapters 5 and 6 of this report. Verification tests regarding the integration show that 

these five components, two models, and one module are correctly integrated. The only exception is that the tests revealed an error in the TRBN component, which prevents verification of the 

control system operation for the turbine. This error will be corrected in a future version of the 

code, and the assessment will be repeated using the corrected version. Documentation of the 

SEPT and CONTAN components is not complete; therefore, their assessments are not discussed 

in this report. The modernization of the power module is complete, and was incorporated into 

TRAC-M(F90) as part of the BWR consolidation. In addition, heat structure nodes and hydraulic cells were recently aligned, and work was done to ensure that the BWR components use this 

enhanced heat structure. This work is complete, but there was insufficient time to reflect it in this 

report. As a result, the assessment of the enhanced heat structure will be included in a future 

report. 

This report also discusses the Peach Bottom Turbine Trip transient, which was used to 

exercise the BWR components, spatial kinetics, and the BWR control system. Although the resulting calculations are preliminary, the results of these calculations and a comparison with the 

test data indicate the successful integration of the spatial kinetics module from the Purdue 

Advanced Reactor Core Simulator (PARCS) code to TRAC-M.

Page xxxvi
FOREWORD

To advance its current capability, the United States Nuclear Regulatory Commission (NRC) is 

consolidating the capabilities of its suite of thermal-hydraulic system analysis codes into one code 

to save resources needed for maintenance and development of four different codes. The base 

code for the consolidation is the Transient Reactor Analysis Code for pressurized water reactors, 

TRAC-P. The architecture of the code has been completely revamped to a modular structure 

using the Fortran90 language. This code is now referred to as TRAC-M to reflect the modernized 

architecture. The functionality of the predecessor codes are being incorporated into TRAC-M. 

To date, the functionality of the Transient Reactor Analysis Code for boiling water reactors, 

TRAC-B, and RAMONA have been incorporated and the consolidation of the Reactor Excursion 

and Leak Analysis Program, RELAP5, is in progress. Once completed, TRAC-M will have the 

ability to read both RELAP5 and TRAC-B input decks as well as legacy TRAC-P decks. A 

developmental assessment will then be performed to select a set of physical models that will 

allow TRAC-M to simulate the applications of the predecessor code with equivalent fidelity. 

The resulting consolidated code, TRAC-M, will enhance the effectiveness and efficiency of 

the NRC thermal-hydraulic program, since effort will be focused on only one code. Therefore, 

user needs will be accommodated and improvements, aimed at increasing the code's realism, will 

be made more efficiently. Analysis capabilities will also improve, as NRC and the user 

community can focus their attention on one code, thereby developing the collective expertise far 

more efficiently than is possible when four codes are utilized. 

This report demonstrates that the architectural modifications made to TRAC-P to produce 

TRAC-M have been successful and have not modified the answers. During the development, 

some code errors in TRAC-P were identified and fixed in TRAC-M. The resulting variation in 

code results are described and reported in this document. In addition, other analyses presented 

demonstrate that the functionality of RAMONA and TRAC-B capabilities have been successfully 

consolidated. 

Thomas L. King, Director 

Division of Systems Analysis and Regulatory Effectiveness

Page xxxvii
ACKNOWLEDGMENTS

This project, undertaken under the leadership of Dr. Farouk Eltawila, began in 1996 upon 

approval of the "Thermal-Hydraulics 5-Year Research Plan." That plan called for the 

consolidation of TRAC-P and TRAC-B, as well as the inclusion of a spatial kinetics capability 

similar to that in the RAMONA code. This large-scale project required close coordination 

between the staff of the U.S. Nuclear Regulatory Commission (NRC) and its contractors. Some 

staff and management thought that the undertaking was very ambitious. 

This report demonstrates that the project has been successful. The authors appreciate the 

support given by Dr. Eltawila throughout the conduct of this project. Without his support (despite 

budget cuts), and without many countless hours of hard work by the contractors and NRC staff 

who are coauthors to this report, this project would not have been successful. The authors also 

acknowledge contributions by Brent E. Boyack, James F. Lime, David A. Pimentel, Jay W. Spore, 

and James L. Steiner from Los Alamos National Laboratory; Weidong Wang from the NRC; 

Glen Mortensen from ISL, Inc.; and Douglas Barber from Purdue University. 

The authors also thank Ms. Colleen Amoruso of ISL, Inc., for her word processing, editing, 

and expertise in preparing the report in FrameMaker format. She has spent many hours (including 

some weekends) to keep up with the schedule for timely preparation of this report. The authors 

appreciate the fine manuscript that she prepared.

Page xxxviii
ACRONYMS

I-D one-dimensional 

2-D two-dimensional 

3-D three-dimensional 

ACC accumulator 

ADS automatic depressurization system 

A/E architect/engineer 

ANS American Nuclear Society 

B&W Babcock & Wilcox 

BE best-estimate 

BN boron nitride 

BOCREC bottom of core recovery 

BWR boiling-water reactor 

CCFL countercurrent flow limitation 

CCTF Cylindrical Core Test Facility 

CDAS Computer Data Acquisition System 

CHF critical heat flux 

CLV Courant Limit Violating 

CMT core makeup tank 

CPU central processing unit 

CSAU code scaleability, applicability, and uncertainty 

DCA drain cooler approach 

DINL double integrator with XOUT limited 

DM1D Detailed Model One-Dimensional 

ECC emergency core coolant 

ECCS emergency core cooling system

Page XXXiX
EM evaluation model 

F77 FORTRAN 77 

F90 FORTRAN 90 

FA fuel assembly 

FLECHT Full-Length Emergency Core Heat Transfer 

GI General Interface 

HEM Homogeneous Equilibrium Model 

HPIS high-pressure injection system 

HSTR heat structure 

HTC heat-transfer coefficient 

IAF inverted annular flow 

IB intermediate-break 

IET integral effect test 

IFDC interfacial-drag coefficient 

1/0 input/output 

INEL Idaho National Engineering Laboratory 

INT integrator 

INTM integrator with mode control 

IRWST in-containment refueling water storage tank 

JAERI Japan Atomic Energy Research Institute 

JCELL junction cell 

KWU Kraftwerk Union Aktiengesellschaft 

LAG first order lag 

LANL Los Alamos National Laboratory 

LB large-break 

LBLOCA large-break loss-of-coolant accident 

LDLY logic delay

Page x I
L/D length-to-diameter 

LINT limited integrator 

LLAG lead-lag transfer function 

LMTD logarithmic mean temperature difference 

LOCA loss-of-coolant accident 

LOFT loss-of-fluid test 

LOSP loss-of-offsite power 

LPCI low-pressure coolant injection 

LPIS low-pressure injection system 

LRB Lehigh University rod bundle 

LWR light-water reactor 

MATLAB Matrix Laboratory 

MgO magnesium oxide 

MSFCV main steam flow control valve 

MSJACS multiple-side junction attribute component set 

NDM Neutronics Data Map 

NEA Nuclear Energy Agency 

NRC U.S. Nuclear Regulatory Commission 

OECD Organization for Economic Cooperation and Development 

PARCS Purdue Advanced Reactor Core Simulator 

PBL pressure balance line 

PBTT2 Peach Bottom Turbine Trip 2 

PCT peak cladding temperature 

PI proportional plus integral 

PID proportional plus integral plus derivative 

PORV pressure-operated relief valve 

PRHR passive residual heat removal

Page x l i
PVM parallel virtual machine 

PWR pressurized-water reactor 

QA quality assurance 

QATP Quality Assurance Test Plan 

RELAP5 Reactor Excursion and Leak Analysis Program 

S/W steam/water 

SB small-break 

SBLOCA small-break loss-of-coolant accident 

SCTF Slab Core Test Facility 

SEASET Separate-Effects and System-Effects Test 

SET separate effects test 

SETS stability-enhancing two-step 

SJACS single-junction attribute component set 

SJC single-junction component 

SOTF second order transfer function 

TDMR TRAC Data Mapping Routine 

T/H thermal hydraulic 

TF two-fluid 

THDM Thermal-Hydraulic Data Map 

TRAC Transient Reactor Analysis Code 

TRAC-B Generic designation for the Transient Reactor Analysis Code for 

boiling-water reactors 

TRAC-M Modernized Transient Reactor Analysis Code 

TRAC-P Generic designation for the Transient Reactor Analysis Code for 

pressurized-water reactors 

TSV turbine stop valve 

TT turbine trip 

TTD terminal temperature difference

Page x I i i
UCSP upper-core support plate 

UPTF Upper-Plenum Test Facility 

USAR Upated Safety Analysis Report 

XTV X-TRAC-View Code

Page x l i i i
1. Introduction

The first internally-released version (Version 3690) of the consolidated Transient Reactor 

Analysis Code (TRAC-M) was produced by modernizing TRAC-P, the generic designation for 

the pressurized-water reactor (PWR) version of the code. As part of the modernization, the PWR 

components in TRAC-P were consolidated and integrated with the boiling-water reactor (BWR) 

components from the TRAC-B code, the generic designation for the BWR version of the code. In 

addition, the modernized version integrates the three-dimensional (3-D) spatial kinetics capability 

from the Purdue Advanced Reactor Core Simulator (PARCS) code. The TRAC modernization 

also included changing the programming language from FORTRAN 77 to FORTRAN 90, and 

reprogramming the code to create new modules. The modernization did not involve changing the 

existing physical models, and did not add any new models for thermal-hydraulics or neutronics. 

(Chapter 2, "Description of the Consolidated Code," describes the modernization and integration 

activities in greater detail.) 

The purposes of this assessment are to demonstrate that a) the modernization has been 

successful, and b) BWR components and spatial kinetics have been successfully integrated with 

the PWR components of the TRAC code. The success of the modernization is demonstrated by 

comparing the results of calculations produced using TRAC-P(Version 5.5.0), which is 

generically designated as TRAC-M(F77) because it was written in FORTRAN 77, and the 

released version (Version 3690) of the modernized and consolidated code, which is generically 

designated as TRAC-M(F90) because it was written in FORTRAN 90. Specifically, the test plan 

is designed to use two assessment test matrices and some plant transients to test phenomena of 

increasing complexity, as well as plant transient calculations involving very large input decks and 

a good number of plant components. The first matrix is the Standard Assessment Test Matrix 

presented in Ref. 1.1, and the second is the PWR Developmental Assessment Test Matrix 

presented in Ref. 1.2. For plant calculations, some of the existing plant input decks are used. 

The Standard Assessment Test Matrix contains 42 short, mostly analytical, test cases, all of 

which represent relatively simple phenomena designed to exercise various parts of the code, with 

an estimated coverage of about 80%. This matrix was used to test each internal (not released) 

version of the consolidated code, as the modernization progressed. This testing was used to 

ensure that new errors have not been introduced, and that the scope of the modernization for that 

version has been successful. The criterion for success was that differences between the test 

results for two sequential versions would be "small." These differences occurred because of 

truncation errors, compiler differences, changes in numerics, and error corrections. Acceptance 

was based on the subjective judgment of the authors. 

This report uses the same acceptance criterion in comparing the released version of the 

consolidated code, TRAC-M(F90) and TRAC-M(F77), when the Standard Assessment Test 

Matrix is exercised. Graphics comparing the results from these two codes will be presented to 

support the authors' conclusions that the differences in these results are "small." Chapter 3, 

"Analytical Test Problems," contains the results of these analyses. 

The PWR Developmental Assessment Test Matrix, which contains the test problems used to 

assess TRAC-M(F77) (Ref. 1.2), and the companion document containing proprietary data 

obtained in the two- and three-dimensional (2-D/3-D) program, increase the test coverage to 

encompass more than the 80% of the code that is exercised using the Standard Assessment Test

Page 1-1
Matrix. In addition, the PWR Developmental Assessment Test Matrix uses long-running test 

problems that increase the complexity of the phenomena. These problems are not used in the 

Standard Assessment Test Matrix, and they also include experimental test data from facilities. 

Since the physical models have not been changed, differences between the predictions obtained 

using TRAC-M(F77) and TRAC-M(F90) are expected to be "small." If so, the agreement between the calculations will be deemed "Excellent." It is also expected that comparison of code 

predictions with test data will lead to conclusions similar to those reached in Ref 1.2 and the 

companion proprietary document. If corrections of errors or bugs involving thermal-hydraulic 

models lead to large differences between the two sets of code predictions, such that the 

conclusions reached in Ref. 1.2 are changed, an assessment of changes in physical models should 

be performed. The need for this type of assessment will be noted; but, performing this type of 

assessment is outside the scope of this report. If corrections of errors or bugs lead to some 

differences that are not "small," but earlier conclusions regarding the assessment of models 

simulating physical phenomena are not changed, the differences will be deemed "Reasonable," 

and the modernization and integration for that particular application will be deemed acceptable. 

Graphics comparing the two sets of code results together with the experimental data, when 

available, will be presented to support the conclusion that differences between two code 

predictions are "Excellent" or "Reasonable." Chapter 4, "PWR Developmental Assessment," 

contains these analyses. Results show that agreements are "Excellent" or "Reasonable." 

PWR plant transient calculations are performed using two codes, and the results of two sets of 

predictions are compared with each other. Differences between predictions should be "Excellent" 

or "Reasonable." If they are not, the reasons for these differences will be understood and noted. 

Graphics comparing the two sets of predictions will be presented to demonstrate that any differences are generally "small" or "reasonable." Chapter 7, "Plant Transient Calculations," 

contains the results of these analyses. 

As stated earlier, the consolidated TRAC version contains BWR components that have been 

integrated from the TRAC-B code. The assessment demonstrates that this integration has been 

performed correctly. The NRC's Quality Assurance (QA) procedures (Ref. 1.3) has been used as 

a guidance. As required by the QA procedures, the requirements for functionality of each 

component were defined. Then, in order to demonstrate that these requirements were met, test 

cases were designed. The consolidated TRAC version and TRAC-B were then run for these test 

cases, and two sets of predictions were obtained. Demonstration of a successful integration of a 

component requires that differences between the two code predictions are "small." These 

differences occur because of truncation errors, compiler differences, changes in numerics, and 

error corrections. If the differences are "small," the agreement is deemed "Excellent." Graphics comparing the two sets of predictions will be presented to demonstrate that any differences are 

generally "small," and the agreements are "Excellent." Chapter 5, "Assessment of Integration of 

BWR Components," describes the requirements, test cases, and results of these analyses. 

The consolidated TRAC version also contains a spatial neutronics capability that was 

integrated from the PARCS code. At each time step, fuel and thermal-hydraulic data are 

transferred from TRAC to the spatial kinetics module, and the power data are transferred from the 

spatial kinetics module to the TRAC code. Successful integration of the spatial kinetics capability requires correct implementation of a new data mapping function. Test cases have been designed, and successful implementation of the mapping function has been demonstrated. Chapter 6,

Page 1-2
"Assessment of Integration of Spatial Kinetics," describes the test cases and results of these 

analyses. Chapter 7, "Plant Transient Calculations," contains preliminary results of calculations 

for the Peach Bottom Turbine Trip 2 transient, which was used to exercise the BWR components, 

spatial kinetics, and BWR control system of the modernized TRAC release. The results show 

successful integration of spatial kinetics. 

It should be noted that requirements, tests and test results presented in Chapters 5 and 6 have 

not been independently reviewed as required by Ref. 1.3. Reviews have been conducted by the 

contractor management, by the author of this report and by the NRC staff while incorporating 

code changes. These reviews were not formal and they are not a substitute for a detailed and 

independent review following software quality assurance procedures as defined in Ref. 1.3. 

However, it appears that the quality of code development work has been maintained during the 

code development. Hence, conclusions in Chapters 5 and 6 assume that requirements, tests and 

test results would not change if an independent formal review as defined in Ref. 1.3 is conducted. 

REFERENCES 

1.1 Steinke, R.G., "A Description of the Test Problems in the TRAC-P Standard Test Matrix," 

Los Alamos National Laboratory, LA-UR-96-1475, May 1996. 

1.2 Boyack, B.E., J.F. Lime, D.A. Pimentel, J.W. Spore, and J.L. Steiner, NUREG/CR-6730, 

"TRAC-M/F77, Version 5.5, Developmental Assessment Manual, Volume I: Assessments, 

Volume II: Appendices," NRC: Washington, D.C. August 2001. 

1.3 Odar, F., NUIREG-1737, "Software Quality Assurance Procedures for NRC Thermal

Hydraulic Codes," NRC: Washington, D.C. December 2000.

Page 1-3
2. Description of the Consolidated TRAC Release and History of its 

Development 

The modernized, consolidated, and internally-released TRAC-M(F90), Version 3690, is an 

advanced, best-estimate (BE) computer program for calculating the transient reactor behavior of a 

PWR or BWR. As such, TRAC-M(F90) incorporates a four-component (liquid water, liquid 

solute, water vapor, and noncondensable gas), two-fluid (liquid and gas) model of the thermal

hydraulic (T/H) processes that are involved in such transients. The complexity of the T/H 

modeling requires many additional models and correlations with logic imposed into a coherent 

description of a phenomenon. 

TRAC-M(F90), Version 3.0, (Ref. 2.1), is one of the latest in a series of TRAC releases. The 

manuals for TRAC-M(F90), Version 3690 and later versions are under preparation. The 

development history represented by earlier versions begins with a very fundamental and 

important improvement in analyzing the behavior of light-water reactors (LWRs), namely, the 

decision to separately track the liquid and vapor fields in the reactor coolant system. These earlier 

versions required the greatly improved computer systems that were becoming available during the 

1970s. They also pushed advances in numerical techniques to solve the complex equation set, 

and to permit the large number of nodes that are required in the BE analysis of some transients of 

interest. 

A preliminary TRAC version for PWRs consisting of only one-dimensional (l-D) 

components was completed in December 1976. Although this version was neither publicly 

released nor formally documented, it was used in developing the first publicly released version, 

TRAC-P1, which was completed in December 1977 (Ref. 2.2), and it formed the basis for the 1-D 

loop component modules. 

The TRAC-P 1 program was primarily designed for use in analyzing large-break (LB) loss- of

coolant accidents (LOCAs) in PWRs. Because of its versatility, however, TRAC-P1 could be 

directly applied to many analyses, ranging from blowdowns in simple pipes to integral LOCA 

tests in multiloop facilities. A refined version, TRAC-P1A, was released to the National Energy 

Software Center in March 1979 (Ref. 2.3). Although it still treated the same class of problems, 

TRAC-P1A was more efficient than TRAC-PI and incorporated improved hydrodynamic and 

heat transfer models. It was also easier to implement on various computers. A subsequent 

version, known as TRAC-PD2 (Ref. 2.4), contained improvements in reflood, heat transfer 

models, and numerical solution methods. Although it was primarily a LBLOCA code, TRAC

PD2 was also successfully applied to small-break (SB) problems and the Three Mile Island 

accident. 

The 1984 release, known as TRAC-PF1 (Ref. 2.5), was designed to improve the ability of 

TRAC-PD2 to handle SBLOCAs and other transients. In addition to all of the major 

improvements of TRAC-PD2, TRAC-PF1 used a two-fluid model with stability-enhancing two

step (SETS) numerics (Ref. 2.6) in the 1-D components. This two-fluid model, in conjunction 

with a stratified flow regime, modeled countercurrent flow better than the drift flux model that 

was used in previous versions, and the two-step numerics allowed large time steps for slow 

transients. Also, the l-D core component permitted calculations with reduced dimensionality, 

although the 3-D vessel option was retained. In addition, noncondensable gas field was added to

Page 2-1
the 1-D and 3-D hydrodynamics, and significant improvements were made to the trip logic and 

the input. 

The development ofTRAC-PFl/MOD1 (Refs. 2.7 and 2.8) maintained the models that were 

needed to apply the code to LBLOCAs, and added or modified models as necessary to enhance the application of the code to SBLOCAs and operational transients. In particular, many user 

convenience features were added or enhanced to promote the application of the code to transients 

involving more complex control of the nuclear plant. Specifically, TRAC-PF1/MODI contained 

generalized, reactivity feedback models (subject to point kinetics assumptions), generalized trip and control system modeling, and necessary components to model the rest of the plant. The code 

was also applicable to most transients for which large asymmetries in the power generation do not 

exist, for which the 1 -D fluid modeling in the pipe is valid, and for which thermal stratification in the liquid in the 1-D components is not important (the 3-D vessel component can model thermal 

stratification in a coarse manner). In addition, the code maintained the capability to run in either a 

1 -D or a mixed 1 -D/3-D mode, with SETS numerics in both the 1 -D and 3-D components. 

TRAC-PFI/MOD2, Version 5.5, is the latest released version of the code for PWRs. This 

code, with some minor corrections, is used is this report and referred to as TRAC-M(F77). This 

code differs from the last formal release, Version 5.4.15, in three major features. First, standard 

FORTRAN 77 (F77) has been implemented throughout the code with a commensurate increase in 

portability and maintainability. Second, the platform-dependent binary file, named TRCGRF, has 

been replaced by the XTVGR.b and XTVGR.t files, which can be processed by the X-TRAC

View code (XTV) visualization and plotting tool (Ref. 2.9). Third, the code includes a new option (newrfd=3) for the reflood model (Refs. 2.10 and 2.11). TRAC-M(F77), Version 5.5, was used to 

perform the developmental assessment calculations in Ref 1.2. 

In 1996, the NRC decided to consolidate the functionality of its suite of codes (TRAC-P, TRAC-B, RAMONA, and RELAP5) into a single, consolidated code that is capable of 

performing the applications of the four codes. The goal was to minimize the dilution of resources 

and the fragmentation of the knowledge base that occurs with the existence of four separate codes. 

TRAC-PF 1/MOD2, Version 5.4.25, a version that preceded Version 5.5, was selected as the basis 

for the consolidation because its architecture was consistent with the ultimate design goal of 

modularity. It had a 3-D hydraulic modeling capability, and it was a better target for installation 

of special-purpose BWR component models that were developed for the TRAC-B code series. Before beginning the consolidation, TRAC-PF1/MOD2, Version 5.4.25, was converted to 

FORTRAN 90 (F90). The container array and integer pointers that were used in Version 5.4.25 as 

a form of home spun dynamic memory allocation, as well as common blocks to provide communication of the global data, made the coding difficult to decipher. The restrictions in F77 

that drove the original TRAC-P data structures were eliminated with the introduction of derived 

types, allocatable arrays, pointers, and modules in F90. The code was denoted TRAC-M to 

signify the modernization of its architecture, which enhanced readability, portability, extensibility, 

and maintainability. 

To recover the functionality of the RAMONA code, BWR stability, and 3-D kinetics, a semi

implicit numerics scheme was enabled in TRAC-M in place of SETS in cases where the 

preservation of property gradients is important. 3-D kinetics capability was recovered by utilizing 

a parallel virtual machine (PVM) to tightly couple the T/H model to a 3-D neutronics package 

(PARCS) running outside of TRAC-M.

Page 2-2
To recover the functionality of TRAC-B, which included transients and LOCAs in BWRs, the 

modeling philosophy used in TRAC-B was duplicated in TRAC-M to construct the components 

that are necessary to model a BWR. The work performed to enhance modularity, and in turn, 

extensibility in TRAC-M during the modernization activities facilitated the TRAC-B 

consolidation. The object-oriented programming concept of inheritance was used to consolidate 

the TRAC-B components with minimal maintenance points and code complexity. The TRAC-M 

and TRAC-B codes share the same generic components, such as PIPE, TEE, VESSEL, BREAK, 

and FILL. The BWR components, including the jet pump, feedwater heater, turbine, 

containment, and BWR fuel channel, were built from these generic components, and only 

component-specific terms were added to the finite volume equations in order to model the BWR 

components. 

REFERENCES 

2.1 Spore, J.W., et al., NUREG/CR-6724, "TRAC-M/Fortran 90 (Version 3.0) Theory Manual," 

NRC: Washington, D.C. July 2001. 

2.2 Safety Code Development Group, "TRAC-P 1: An Advanced, Best-Estimate Computer 

Program for PWR LOCA Analysis," Los Alamos Scientific Laboratory Report LA-7279

MS, NUREG/CR-0063, May 1978. 

2.3 Safety Code Development Group, "TRAC-P1A: An Advanced, Best-Estimate Computer 

Program for PWR LOCA Analysis," Los Alamos Scientific Laboratory Report LA-7777

MS, NUREG/CR-0665, May 1979. 

2.4 Safety Code Development Group, "TRAC-PD2: An Advanced, Best-Estimate Computer 

Program for Pressurized-Water Reactor Loss-of-Coolant Accident Analysis," Los Alamos 

Scientific Laboratory Report LA-8709-MS, NUREG/CR-2054, April 1981. 

2.5 "Computer Program for Pressurized-Water Reactor Analysis," Los Alamos National 

Laboratory Report LA-9944-MS, NUREG/CR-3567, February 1984. 

2.6 Mahaffy, J.H., "A Stability-Enhancing Two-Step Method for Fluid Flow Calculations," 

J. Computational Physics 46, 329-341, 1982. 

2.7 Safety Code Development Group, "TRAC-PFl/MODI Correlations and Models," Los 

Alamos National Laboratory Report LA-11208-MS, NUREG/CR-5069, December 1988. 

2.8 Safety Code Development Group, "TRAC-PFl/MODI: An Advanced, Best-Estimate 

Computer Program for Pressurized-Water Reactor Thermal-Hydraulic Analysis," Los 

Alamos National Laboratory Report LA-10157-MS, NUREG/CR-3858, July 1986. 

2.9 Dearing, J.F., and R. Johns, "XTV User's Guide, Release 2.3c," Los Alamos National 

Laboratory Report LA-UR-01-921, May 1998.

Page 2-3
2.10 Nelson, R.A., Jr., D.A. Pimentel, S. Jolly-Woodruff, and J.W. Spore, "Reflood Completion 

Report, Volume I, A Phenomenological Thermal-Hydraulic Model of Hot Rod Bundles 

Experiencing Simultaneous Bottom and Top Quenching and an Optimization Methodology 

for Closure Development," Los Alamos National Laboratory Report LA-UR-98-3043, 

April 1998. 

2.11 Boyack, B.E., J.F. Lime, D.A. Pimentel, J.W. Spore, and T.D. Knight, "Reflood Completion Report, Volume II, Developmental Assessment of a New Reflood Model for the TRAC-M/ 

F77 Code," Los Alamos National Laboratory Report LA-UR-98-3043, April 1998.

Page 2-4
3. Analytical Test Problems

This chapter describes comparisons between the predictions obtained using TRAC-M(F77) 

and TRAC-M(F90) for analytical test problems selected from the "Standard Assessment Test 

Matrix." The 42 primarily simple, fast-running test problem-calculation cases in this matrix 

exercise approximately 80% of the code, and are used when a new version of the code is created 

in order to ensure that errors are not introduced by the new updates. 

The matrix consists of two problem series including (1) a short-running series of 10 test 

problems with 34 problem-calculation cases, and (2) a long-running series of 5 test problems with 

8 problem-calculation cases (Ref. 1.1). The test problem names and their number of calculation 

cases (shown in brackets) are as follows: 

Short-Running Series Long-Running Series 

1. DRAIN [1] 11. BANKOFF [1] 

2. HCOND#; #=1,2,3,4,5 [5] 12. LEHIGH [1] 

3. MARVIK [1] 13. LOFTL26 [2] 

4. MS#&; # - C,P,R,S,T,V; & = 2,L,V; #& = VLR [19] 14. UPTF8B [3] 

5. ROD2 [1] 15. UTUBE3D [1] 

6. TFPIPE2 [1] 

7. TIN#; # = DWN,UP [2] 

8. UTUBE [1] 

9. W4LOOP [2] 

10. ZIONPWR [1] 

Each version of the code has been exercised using all 42 problem-calculation cases. 

Comparisons of the results obtained using TRAC-M(F77) and TRAC-M(F90) show "Excellent" 

to "Reasonable" agreement between the two codes. The following sections present calculations 

for a sample of five short-running test problems. The files for all other cases are available for 

review by interested parties. 

In Section 3.1, both codes are compared using steady-state conduction solutions for 1 -D radial 

and 2-D radial and axial geometries. Analytical steady-state conduction solutions have been 

developed in literature with fixed heat transfer coefficients and fixed fluid conditions.

Page 3-1
Comparisons are made with the TRAC predicted temperature distributions for the same boundary 

conditions. 

In Section 3.2, both codes are compared using a drain and fill problem. This analytical test 

problem demonstrates the TRAC capability to accurately calculate mixture levels and level 

crossings at cell boundaries without significant numerical instabilities. 

In Section 3.3, both codes are compared using a manometer oscillation problem with a known 

analytical frequency. This test problem indicates that both codes accurately calculate the gravity 

head and oscillation in this U-tube geometry.

Page 3-2
3.1 Steady-State Conduction Problems (HCOND#; # = 1,2,3) 

These steady-state conduction problems were provided by the Japan Atomic Energy Research 

Institute (JAERI) as part of the checkout for their fully implicit axial conduction solution. As 

discussed in the following subsections, these test problems provide an analytical check to verify 

that the TRAC conduction solution is accurate. The results for a 1-D radial steady-state 

conduction solution and a 2-D radial and axial steady-state conduction solution are compared to 

analytical solutions. 

3.1.1 1-D Radial Conduction 

The 1 -D radial steady-state conduction was determined for a generic nuclear fuel rod, with the 

dimensions shown in Fig. 3.1.1. For this test problem, the inner material is representative of fuel 

that is surrounded by a gap that is, in turn, surrounded by cladding. The material from r = 0 to 

r = r1 is a uniform source of 1000 W. From r = r1 to r = r2 , a gap occurs with a conductance of 

1000 Wm' 2

K-1. From r = r 2 .to r = r3 , the thermal conductivity is 13.8 Wm-IK- 1

. The outside 

fluid temperature is 300 K, and the rod surface heat transfer coefficient is 2836 Wm' 2K-1. 

The governing differential equation is ld 

rdr dr 

where T is the temperature, q"' is the heat generation rate per unit volume, k is the thermal 

condictivity and r is the radial coordinate. 

The boundary conditions are 

dT 0 r 0, 

dr 

where there is no heat conduction across the centerline of the fuel rod, and 

(-kdd7 h- T - Tf) = Q/(2irr3L) d 3 

where h3 is the surface heat transfer coefficient, T3 and Tf are rod surface and outside fluid 

temperatures respectively, r3 is the distance between the rod surface and rod centerline, L is the 

height of the rod and Q is the heat generation rate within the rod. 

Q 2 r7 q = 1000W 

at steady-state, all of the energy produced must appear as surface heat flux.

Page 3-3
q1 = Q/(2rrrL), q 2 = Q/(2rir 2L) and q 3 = Q/(27rr 3L).

The solution for this problem is 

T= T3 - qlrlln(r/r3 )/k 3 for r 2 < r < r 3 

2 2 T= Ti +ql(r -r )(2k rl) for r<r1 , 

where 

T3 = Q/(27tr 3Lh3) + Tf, 

T2 = T3- qIrln(r2 /r 3 )/k 3 

T1 = T2 + ql1h2, 

where h2 is the gap conductance, 

Q = 1000 W, 

L= 0.1 m, 

r, =6.35 mm, 

r2 6.426 mm, 

r3= 7.239 mmm, 

k, 2 Wm-IK- 1

, 

k3= 13.8 Wm-nK 1 , 

T= 300 K, 

h2 1000 Wm-2K 1 , and 

h3 2836 Wm-2K 1 . 

A TRAC model using 18 nodes in the radial direction (input model HCOND2) was used to 

represent the geometry of the analytical test problem described above. Calculations were 

performed using TRAC-M(F77) and TRAC-M(F90), and the results are compared with the 

analytical model in Table 3.1.1. 

The results given in Table 3.1.1 indicate that the finite-difference solutions of the heat 

conduction equation are accurate in the radial direction. The surface temperature is exact to four 

digits of accuracy, which is the number of digits that TRAC edits. The results in Table 3.1.1 also 

indicate a TRAC error of 0.17% in the centerline temperature, and this error is associated with an

Page 3-4
error in the temperature solution across the gap, where the steepest temperature profile occurs. 

Consequently, inaccuracies in the finite-difference approximation would tend to occur where the 

temperature profile is steep. Given that the agreement between TRAC-M(F77) and TRAC

M(F90) calculations is "Excellent," the results confirm that the radial heat conduction solution in 

TRAC is accurate and adequate for LWR safety applications, and that conversion from TRAC

M(F77) to TRAC-M(F90) has been successful. 

3.1.2 2-D Radial and Axial Conduction 

The analytical solution for the 2-D radial and axial heat conduction problem was developed 

for a solid rod with a constant and uniform heat source, with a constant heat transfer coefficient in 

the radial direction, and adiabatic boundary conditions at both ends of the rod (as shown in Fig. 

3.1.2). The axial variation in the temperature profile is obtained by having one fluid temperature 

boundary condition on the lower half of the rod, and another fluid temperature boundary 

condition on the upper half of the rod. 

This problem is solved in Ref. 3.1, and the solution is not repeated here. The analytical 

solution for this problem is, however, compared with the TRAC solution in Table 3.1.2 for rod 

temperatures along the centerline of the rod. As presented in Table 3.1.2, the results obtained 

using TRAC-M(F77) and TRAC-M(F90) with input model HCOND3 indicate that 2-D 

conduction solutions in both codes are accurate. Again, some inaccuracies appear in the region 

where the temperature profile is steep; however, the error percentage is quite small (0.05%). The 

results given in Table 3.1.2 were obtained with the fine mesh option on. That option adds axial 

levels in regions where the temperature profile is steep, in an attempt to reduce the finite 

difference error in that region. 

Table 3.1.3 presents a comparison between the results obtained using TRAC-M(F77) and 

TRAC-M(F90) and the analytical solution using 21 fixed axial levels (input model HCOND1). 

This comparison demonstrates that the fine mesh option improves the accuracy of the finite 

difference approximations relative to the original differential equation. The maximum error for 

the fixed node case is 0.75%, whereas the maximum error for the fine mesh case is 0.05%. 

Agreement between the TRAC-M(F77) and TRAC-M(F90) predictions is "Excellent." 

In Table 3.1.4, the radial solutions at z = 100 mm for the 21 fixed axial heat-conduction model 

(HCOND 1) are compared with the analytical solution at the same location. Table 3.1.4 shows 

that the maximum error of 0.018% along that radius occurs at the outer edge. The results indicate 

that the 2-D heat conduction solution obtained using both codes is very accurate in both the axial 

and radial directions. Nevertheless, the accuracy of the axial solution is significantly improved 

with the fine mesh option turned on. The results also confirm that conversion of the TRAC

M(F77) code to the TRAC-M(F90) code has been successfully performed.

Page 3-5
Table 3.1.1 Comparison Between TRAC Codes and Analytical Solutions 

for a 1-D Heat-Conduction Problem in the Radial Direction at Steady State 

(TRAC Input Model HCOND2)

Node 

Location (mm) 

0.000 

1.830 

2.590 

3.175 

3.670 

4.100 

4.490 

4.850 

5.185 

5.500 

5.800 

6.080 

6.350 

6.426 

6.670 

6.840 

7.040 

7.239

Analytical 

Solution (K) 

1039.8 

1006.7 

973.6 

940.3 

906.9 

873.9 

840.9 

807.7 

774.5 

741.3 

707.8 

675.0 

641.9 

391.3 

387.0 

384.1 

380.7 

377.5

TRAC-M(F77) 

Solution (K) 

1038.0 

1005.0 

972.1 

938.8 

905.4 

872.4 

839.4 

806.2 

773.0 

739.8 

706.3 

673.5 

640.4 

391.3 

387.0 

384.1 

380.7 

377.5

TRAC-M(F90) 

Solution (K) 

1038.0 

1005.0 

972.1 

938.8 

905.4 

872.4 

839.4 

806.2 

773.0 

739.8 

706.3 

673.5 

640.4 

391.3 

387.0 

384.1 

380.7 

377.5

Page 3-6

Error (K) 

-1.8 

-1.7 

-1.5 

-1.5 

-1.5 

-1.5 

-1.5 

-1.5 

-1.5 

-1.5 

-1.5 

-1.5 

-1.5 

0.0 

0.00 

0.0 

0.0 

0.0
Table 3.1.2 Comparison Between TRAC Codes and Analytical Solutions 

Along a Center Line of a Rod in a 2-D Heat-Conduction Problem 

While at Steady State with Fine Mesh 

(TRAC Input Model HCOND3)

Node 

Elevation (mm) 

0 

10 

20 

30 

40 

50 

60 

70 

80 

90 

100 

110 

120 

130 

140 

150 

160 

170 

180 

190 

200

Analytical 

Solution (K) 

658.1 

658.1 

658.1 

658.1 

658.1 

658.1 

658.1 

658.1 

658.2 

662.7 

758.1 

853.9 

857.5 

858.1 

858.1 

858.1 

858.1 

858.1 

858.1 

858.1 

858.1

TRAC-M(F77) 

Solution (K) 

658.1 

658.1 

658.1 

658.1 

658.1 

658.1 

658.1 

658.1 

658.2 

662.7 

758.5 

853.6 

857.9 

858.1 

858.1 

858.1 

858.1 

858.1 

858.1 

858.1 

858.1

TRAC-M(F90) 

Solution(K) 

658.1 

658.1 

658.1 

658.1 

658.1 

658.1 

658.1 

658.1 

658.1 

662.6 

758.5 

853.6 

857.9 

858.1 

858.1 

858.1 

858.1 

858.1 

858.1 

858.1 

858.1

Page 3-7

Error (K) 

0.0 

0.0 

0.0 

0.0 

0.0 

0.0 

0.0 

0.0 

0.0 

0.0 

+0.4 

-0.3 

+0.4 

0.0 

0.0 

0.0 

0.0 

0.0 

0.0 

0.0 

0.0
Table 3.1.3 Comparison Between TRAC Codes and Analytical Solutions 

Along a Center Line of a Rod in a 2-D Heat-Conduction Problem 

While at Steady State with Fixed Noding 

(TRAC Input Model HCOND1)

Node 

Elevation (mm) 

0 

10 

20 

30 

40 

50 

60 

70 

80 

90 

100 

110 

120 

130 

140 

150 

160 

170 

180 

190 

200

Analytical 

Solution (K) 

658.1 

658.1 

658.1 

658.1 

658.1 

658.1 

658.1 

658.1 

658.2 

662.7 

758.1 

853.9 

857.5 

858.1 

858.1 

858.1 

858.1 

858.1 

858.1 

858.1 

858.1

TRAC-M(F77) 

Solution (K) 

658.1 

658.1 

658.1 

658.1 

658.1 

658.1 

658.1 

658.2 

658.8 

667.7 

758.1 

848.5 

857.4 

858.0 

858.1 

858.1 

858.1 

858.1 

858.1 

858.1 

858.1

TRAC-M(F90) 

Solution (K) 

658.1 

658.1 

658.1 

658.1 

658.1 

658.1 

658.1 

658.2 

658.8 

667.7 

758.1 

848.5 

857.4 

858.0 

858.1 

858.1 

858.1 

858.1 

858.1 

858.1 

858.1

Page 3-8

Error (K) 

0.0 

0.0 

0.0 

0.0 

0.0 

0.0 

0.0 

+0.1 

+0.6 

+5.0 

0.0 

-5.4 

-0.1 

-0.1 

0.0 

0.0 

0.0 

0.0 

0.0 

0.0 

0.0
Table 3.1.4 Comparison Between TRAC Codes and Analytical Solutions 

Along a Radius at Midplane (Elevation 100 mm) 

of a Rod in a 2-D Heat-Conduction Problem at Steady State 

(TRAC Input Model HCOND1)

Node 

Location (mm) 

0.0 

0.5 

1.0 

1.5 

2.0 

2.5 

3.0 

3.5 

4.0 

4.5 

5.0

Analytical 

Solution (K) 

758.1 

756.1 

750.1 

740.2 

726.3 

708.4 

686.5 

660.6 

630.8 

596.9 

559.1

TRAC-M(F77) 

Solution (K) 

758.1 

756.1 

750.1 

740.2 

726.3 

708.4 

686.5 

660.6 

630.8 

597.0 

559.2

TRAC-M(F90) 

Solution K 

758.1 

756.1 

750.1 

740.2 

726.3 

708.4 

686.5 

660.6 

630.8 

597.0 

559.2

Page 3-9

Error (K) 

0.0 

0.0 

0.0 

0.0 

0.0 

0.0 

0.0 

0.0 

0.0 

+0.1 

+0.1
Figure 3.1.1 Boundary Conditions of a I-D Heat-Conduction Problem in the Radial Direction 

U2!. 

L/21 r 

U2 

Figure 3.1.2 The 2-D Radial and Axial Heat-Conduction Problem

Page 3-10
3.2 1-D Drain and Fill Test Problem (DRAIN)

The drain test problem consists of a vertical PIPE component that is partially filled with water, 

a FILL component, and a BREAK component, as shown in Fig. 3.2.1. The BREAK component 

provides a constant pressure boundary condition (1.Oe+05 Pa) for the top of the pipe. The FILL 

component slowly drains water from the pipe, and then refills the pipe to its original level. The 

purpose of this test problem is to verify that (1) TRAC versions do not calculate numerical 

instabilities as the cells in a PIPE component slowly drain or fill, and (2) TRAC codes accurately 

calculate the gravity head. 

Figs. 3.2.2 and 3.2.3 show the void fraction and mass flow rate within the pipe as the pipe 

drains and fills. In particular, these plots show that no numerical problems appear as the level 

crosses the cell boundaries. In Fig. 3.2.4, the pressure in cell 1 is plotted for the drain and fill 

transient, and no significant pressure oscillation occurs as the pipe drains and fills. 

The PIPE component is made up of 15 cells, with each cell being 1.0 m long. With the PIPE 

component initially full of 10 m of water, the cell center pressure of the first cell should be the 

1.0e+05 Pa plus 9.5 m of water or, 

p = p(break) + rho*g*h 

= 1.0e+05 + 997.4*9.8*9.5 = 1.929e+05 Pa. 

TRAC calculates the initial cell 1 pressure before the drain starts as 1.92e+05 Pa, and the final 

cell 1 pressure at the end of the fill transient as 1.93e+05 Pa. The minimum level during this 

calculation is 4 m of water above the cell 1 center. Given the static head above cell 1, the pressure 

would be 

p = p(break) + rho*g*h 

= 1.0e+05 + 997.4*9.8*4.0 = 1.391e+05 Pa. 

TRAC calculates a minimum pressure in cell I of 1.387e+05 Pa. 

From these calculations and Figs. 3.2.3 and 3.2.4, it is apparent that both TRAC-M(F77) and 

TRAC-M(F90) accurately calculate the gravity head as the cells slowly drain and fill. It is also 

apparent that the codes encounter no significant numerical problems when a liquid level crosses a 

cell boundary. Agreement between TRAC-M(F77) and TRAC-M(F90) predictions is also 

"Excellent."

Page 3-11
Initial 11 

Liq uid ....... ....... Level 10 

9 PIE 

8 700 

7 

6 

5 

2 

Figure 3.2.1 TRAC Model of the DRAIN Test Problem

Page 3-12
Drain and Fill Test Problem 

Comparison of TRAC Codes

50 100 150 200 250 

Time (s) 

Figure 3.2.2 Comparison of Void Fraction Predictions in DRAIN and FILL Test Problem

(

1 

0.8 

c0.6 

"0 

0 

> 0.4 

0.2 

0

0 300

r' t
Drain and Fill Test Problem

Comparison of TRAC Codes

1

--- -TRAC-M(F77), Cell #4 Edge 

uuTRAC-M(F77), Cell #5 Edge 

STRAC-M(F77), Cell #6 Edge k- iTRAC-M(F77), Cell #7 Edge *--A TRAC-M(F77), Cell #8 Edge 

-TTRAC-M(F77), Cell #9 Edge 

S---+ TRAC-M(F77), Cell #10 Edge 

- - 0 TRAC-M(F90), rmvm-700004 

0- - e TRAC-M(F90), rmvm-700005 

+ - E TRAC-M(F90), rmvm-700006 

+. - A5 TRAC-M(F90), rmvm-700007 

* - A TRAC-M(F90), rmvm-700007 

- V TRAC-M(F90), rmvm-700009 

D- -• TRAC-M(F90), rmvm-700010

50 100 

Time (s)

150

Figure 3.2.3 Comparison of Mass Flow "-,dictions in DRAIN and FiLL, Test Problem

12.5 

7.5

(b 

4•

0 

uI 

C,) 

0 

F2.5

-2.5

-7.5 

-12.5

0

200
/

Drain and Fill Test Problem

Comparison of TRAC Codes

- r -___

50 100 150 

Time (s)

Figure 3.2.4 Comparison of Pressure Predictions in DRAIN and FILL Test Problem

C'

2e+05 

1.8e+05 

1.6e+05 

1.4e+05 

1.2e+05 

le+05

(L 

CI.

0 200
3.3 U-Tube Test Problem (UTUBE) 

This test problem (illustrated in Fig. 3.3.1) consists of a single pipe representing a U-tube that 

is -6.8 m high. Both legs of the U-tube are connected to constant-pressure BREAK components. 

Under the initial conditions, the liquid level in one leg is 0.3 m higher than the liquid level in the 

other leg. This results in a U-tube oscillation that has a period (Ref. 3.2) of 

"- = 27L/2g = 2.954s 

Fig. 3.3.2 shows that the periods calculated by TRAC-M(F77) and TRAC-M(F90) for the U

tube oscillation are almost identical. The period is -3.1 s. In addition, it is apparent that the gas 

volume fraction oscillation in cell 6 is the same as the gas volume fraction oscillation in cell 10, 

with the exception that the oscillations in cells 6 and 10 are out of phase. That is, when the cell 6 

gas volume fraction peaks high, the cell 10 gas volume fraction peaks low, and conversely. 

However, because of wall friction in the pipe, the amplitude of the gas volume fraction oscillation 

should decay to zero. However, the gas volume fraction oscillation does not completely dampen 

out in either calculation because of the time and spatial averaging used to calculate the interfacial 

shear between cells 6 and 7 and between cells 9 and 10. For relatively large time-steps, the time 

averaging model for the interfacial shear tends to move closer to the new time interfacial shear. 

Some time averaging of the interfacial shear is required for stability. The oscillations can be 

reduced through finer noding, reduced maximum time-step size, or increased wall drag. Fig. 

3.3.3 shows that the oscillations dampen out considerably when the maximum time-step size is 

reduced from 0.05 s to 0.01 s. Overall, the agreement between TRAC-M(F77) and TRAC

M(F90) predictions is "Excellent."

Page 3-16
inwea 

LUqu 

LOW

BREAK mm BREAK 701 U702 

1 15 

2 14 

3 13 

PIPE 

700 --- --

4 12 

id- 6 10. 

7,9 

J# 8

aid

Figure 3.3.1 TRAC Model of the UTUBE Test Problem 

Page 3-17
U -Tube Test Problem

Comparison of TRAC Codes. DT(max)=0.05s r I

10 20 30 40

Time (s)

Figure 3.3.2 Comparison of Gas Volume Fraction Oscillations in the UTUBE Test Problem, 

DT(max)=0.05s

0.3

00

C: 

0 

U

E 

(.9 CU, 0

0.2 

0.1

0

0 50
\

U - Tube Test Problem

Comparison of TRAC Codes, DT(max)=0.01 s

0.3

0.2 

0.1

0 LL 

0 10 20 30 40 

Time (s) 

Figure 3.3.3 Comparison of Gas Volume Fraction Oscillations in the UTUBE Test Problem, 

DT(max)=0.01s

(

C 

0 

U

E 

(.)

50
3.4 Conclusions

Analytical test problem comparisons show that the TRAC finite-difference solution to the heat 

conduction equations in rod geometries is accurate. The hydro models are also accurate and 

stable during the filling and draining of cells. The conversion from TRAC-M(F77) to TRAC

M(F90) has been successful for these applications and other cases in the Standard Assessment 

Test matrix. The files for all other cases are available for review by interested parties. 

REFERENCES 

3.1 Carslaw, H.S., and J.C. Jaeger, Conduction of Heat in Solids, Clarendon Press, Oxford, 1959. 

3.2 Streeter, V.L., and E.B. Wylie, Fluid Mechanics, McGraw-Hill Book Company, New York, 

1975.

Page 3-20
4. PWR Developmental Assessment

This chapter presents results of a developmental assessment of TRAC-M(F77), Version 5.50 

and TRAC-M(F90), Version 3690 using the separate effects test (SET) and integral effects test 

(lET) data for PWRs. The test matrix has been used in Ref. 1.2 in the developmental assessment 

of TRAC-M(F77). Since no physics modeling has been changed between these two codes, it is 

expected that calculations obtained from exercising these two versions should be close to each 

other. There will be some differences because of some error corrections in TRAC-M(F90). As 

long as these error corrections do not change conclusions on the capabilities of the codes to 

predict the separate effects phenomena and integral effects test results, conversion of TRAC

M(F77) to TRAC-M(F90) will deem acceptable. As the results of calculations show, the 

comparisons are "Excellent" or "Reasonable." 

4.1 Countercurrent Flow Limitation 

The countercurrent flow limitation (CCFL) is defined when the downward flowing liquid in a 

countercurrent flow is about to change its direction to flow upward. This condition can be 

attained if the upward steam flow rate is increased, while the liquid is flowing downward. 

Depending upon the geometry of the equipment used, the occurrence of the CCFL varies. In 

reactor applications, CCFL can occur at flow area restrictions when liquid flows downward 

through rising gas or vapor. For example, in the upper tie-plate region of a PWR during reflood, 

the upward flow of steam can prevent or limit the downflow of liquid, thereby causing the rods to 

cool less than they otherwise would. 

In TRAC-M, a special model allows the user to invoke characteristic CCFL correlations at 

specific locations. Rather than try to develop a new mechanistic CCFL model that could predict 

the complex behaviors in a variety of geometrical configurations, TRAC enables the user to apply 

a CCFL correlation for a specific geometry and apply it at a specific location. 

4.1.1 Bankoff Tests 

This section assesses the CCFL model against the data of Bankoff et al. (Ref. 4.1.1). 

4.1.1.1 Test Facility 

As illustrated in Fig. 4.1.1, the experimental apparatus used by Bankoff et al. was a vertical 

channel with a flow area of 31 cm2. The air-water experiments were used to determine CCFL 

independent of the effects of phase change. The air was introduced below a perforated plate, 

while the water was introduced to the upper-plenum chamber and overflowed to drain. This test 

was used to assess the capability of TRAC to predict CCFL with a 15-hole geometry 

corresponding to a small segment of the upper tie plate of a PWR. 

Pressure and temperature measurements were available at various places, and air and water 

flow rates were measured using rotameters. The test procedure used in the air-water tests 

consisted of establishing the water inlet flow rate, increasing the air flow rate, and determining the 

water delivery rate at each step until a point of essentially zero downward delivery was finally 

obtained. That point is defined as the CCFL.

Page 4-1
4.1.1.2 TRAC Model

Fig. 4.1.2 illustrates the TRAC CCFL model. The test section is modeled with a 1 -D VESSEL component with 11 axial cells. The rectangular perforated plate is located at the interface between the fifth and sixth cells, and the flow area of all 15 holes is specified at this interface. The hydraulic diameter of 10.5 mm is the same as that of the hole diameter. The water injection PIPE component is connected to the middle of the water pool that initially existed in cells 6 through 9. The height of this water pool is controlled by the drain tube, which is modeled 

by a PIPE component, and horizontally connected to cell 9. The other side of the drain tube is connected to an atmospheric BREAK component. Downward flowing water is collected at the bottom of the vessel, which is connected to a zero FILL boundary condition. The volume of the bottom level of the VESSEL component is sufficient for water flowing through the tie plate to accumulate. The air flows through an air outlet pipe to a BREAK component. 

The TRAC CCFL model allows the user to input the characteristic flooding curve parameters for a specific geometry applied at a particular location. Typically, these parameters have been developed from experimental data for the geometry of interest, or for hardware of at least similar dimensions. Bankoff et al. have shown that the data correlate well with the relationship 

HI /2 + MH12 =C, (4.1.1.1) 

where 

Hg is the dimensionless gas flux, 

H, is the dimensionless liquid delivery, 

C is the abscissa intercept, and 

M is the slope. 

The Bankoff relationship is sufficiently general that either Wallis scaling for diameter dependence, Kutaleladze scaling for surface-tension dependence, or a combination of the two can be implemented. The scaling is performed by defining a variable length scale in the 

determination of the dimensionless flux: 

Hk = J Pk 2 (4.1.1.2) 

kkgcwApi 

w =HD1-B L , (4.1.1.3) 

1

Page 4-2
where

k refers to the phase (gas or liquid), 

] is the superficial velocity, 

HD is the hydraulic diameter, 

gc is the gravitational constant, 

a is the surface tension, 

p is the density, 

Ap is the difference between the phasic densities, and 

B is a factor between 0 and 1. 

The correlation reverts to the Wallis scaling for B = 0, and to the Kutaleladze scaling for 

B = 1. For B between 0 and 1, the user can input the scaling proposed by Bankoffet al., which 

can be used for tie plate geometry even when no experimental data are available. 

4.1.1.3 Comparison of Predicted and Measured Results 

Parametric calculations have been performed by varying both TRAC-M(F77) and TRAC

M(F90). Figs. 4.1.3 and 4.1.4 show the liquid and air mass flow rates calculated by two codes. 

The agreement between predictions is "Excellent," and the frequency of oscillations of water 

downflow and air upflow agree with each other every well. The amplitudes also agree quite well. 

From these two figures, we conclude that the conversion from TRAC-M(F77) to TRAC-M(F90) 

has been successful for this type of application. 

The rest of the analysis has been performed using the results obtained from TRAC-M(F77). 

Smoothing functions are applied to the mass flows as shown in Figs. 4.1.5 and 4.1.6, and the 

smoothed mass flows are then used to calculate the dimensionless fluxes H]/ 2 and H1/ 2 , which 

are shown in Fig. 4.1.7. 

Fig. 4.1.8 shows the TRAC-M(F77) results for the flow of air and water at I atm, along with 

the correlation of the Bankoff data. The TRAC input parameters for the Bankoff data are B = 

0.884, M = 1, and C = 1.92. As demonstrated, the TRAC results agree quite well with the 

experimental data of Bankoff. Thus, it is expected that when the TRAC CCFL option is used at a 

given location with appropriate input parameters for the flooding curve, the liquid downflow will 

be predicted reasonably well for plates that are similar to those used in the Bankoff experiment (as 

well as other types of plates). Since liquid and air mass flow rates calculated by TRAC-M(F90) 

are in excellent agreement with those calculated by TRAC-M(F77), we expect similar agreement 

with the experimental data of Bankoff when the TRAC-M(F90) code is used.

Page 4-3
4.1.1.4 Conclusions

The TRAC CCFL model allows the user to input the characteristic flooding curve parameters for a specific geometry applied at a particular location. When used to predict the air-water CCFL 

data obtained by Bankoff et al., this model gave reasonable results in predicting the CCFL. The 

assessment is limited to air-water mixtures at atmospheric conditions, and the current code CCFL 

model for steam-water mixtures is not assessed. The assessment will be extended to cover steam

water mixtures in a future program. 

The conversion from TRAC-M(F77) to TRAC-M(F90) has been successful for this type of 

application. 

REFERENCES 

4.1.1 Bankoff, S.G., R.S. Tankin, M.C. Yuen, and C.L. Hseih, "Countercurrent Flow of Air/ 

Water and Steam/Water Through a Horizontal Perforated Plant," Int. J. Heat Mass 

Transfer 24(8), 1381-1385, 1981.

Page 4-4
Steam

mm

Water 

rH

Plate

Rotameters

Water

71.5 mm

00000 

00000 

00000

-- I42.9 mm

D=1 0.5 mm

Rotameters

Air 

Steam

Beam scale 

Figure 4.1.1 Sketch of the Perforated Plate Air-Water Countercurrent Flow System 

Used by Bankoff et al.

Page 4-5
24100 

1.700 

1.397 10Pipe C7 Atrn. Break C4 

1.387 

1.259-0- cWaterlnj. Pipe 06W 

1.130 0.07 o• 

0.980 a15 Holes Rectangular 

Hole 

Plate (0.0715x0.0429 mxm ID=10.5 mm 0.600 4 

0.400 3 i Air Inj. Ppe C5 KJ Air inj. Fill C2 

0.2 2 

0.000 -Vessel 

4 I r . o de fed 

Cll 

Figure 4.1.2 TRAC Model for the Test Apparatus Used by Bankoff et al.

Page 4-6
(

Bankoff CCFL Test

Comparison of Calculated Liquid Mass Flows at the Perforated Plate

0 10 20 30 40 

Time (s) 

Figure 4.1.3 Comparison of Calculated Liquid Mass Flows at the Perforated Plate 

in the Bankoff CCFL Test

,I

0.1 

0

0' ýO 

"to4

0) CM 

0 

CO, 

Cu 

0*

-0.1 

-0.2

-0.3 

-0.4

50 -0.4
Bankoff CCFL Test

Comparison of Calculated Air Mass Flows at the Perforated Plate I I -r - I-- 

10 20 30 40

Time (s)

Figure 4.1.4 Comparison of Calculated Air Mass Flows at the Perforated Plate 

in the Bankoff CCFL Test

(

0.04

0.03

00

CY) 

0L 0.02

0.01 

0

0

50
(

10 20 30 40 

Time (s) 

Figure 4.1.5 Liquid Mass Flow at the Perforated Plate in the Bankoff CCFL Test

Bankoff CCFL Test 

Liquid Mass Flow at the Perforated Plate

0.1 

0

(

0 LL 

C/, 

C/, 0co

It

-0.1 

-0.2

-0.3 

-0.4

0 50
Bankoff CCFL Test 

Air Mass Flow at the Perforated Plate

I T I -I

10 20 30 40 50

Time (s)

Figure 4.1.6 Air Mass Flow at the Perforated Plate in the Bankoff CCFL Test

(

0.04 

0.03

0 

co 

co 

cO

cIQ 

,0 0.02

0.01 

0 

0
225

2 ........ APOR. 

1.75 

o 0.75 

W'3 0.50

0.25o 

1.0 

• -0.25 

0 5 10 15 20 25 30 35 40 45 

TIME (s) 

Figure 4.1.7 Liquid and Air Dimensionless Fluxes at the Perforated Plate 

in the Bankoff CCFL Test

Page 4-11

50
2 1 rI 1 I I I I 

0 TRAC 

1.75 -- -- BANKOFF CORRELATION. C=192 

b~ 

1.50 b 

>< 

r00 

0 %° 

LLJ 

z 

0. 

-J I * 100".. 

z •. 0.75-, 00 "" 

CY 0.50 

0.25 

01h 

0 0.25 0.50 0.75 1 125 1.50 1.75 

SQRTDUODNSONLESS UQUID FLUX) 

Figure 4.1.8 Comparison of TRAC Results with Bankoff Correlation

Page 4-12

2
4.2 Condensation

When a thermal-hydraulic (T/H) analysis of PWRs is performed to calculate the consequences 

of a small-break loss-of-coolant accident (SBLOCA) or a large-break loss-of-coolant accident 

(LBLOCA), direct contact occurs between subcooled liquid and steam after the emergency 

injection system is turned on. This contact causes a violent pressure transient and a phenomenon 

similar to a water hammer. Consequently, the prediction of the correct interfacial condensation 

rate for a variety of flow regimes is very important for LOCA analyses. The condensation model 

used in TRAC-M is founded upon the Chen-Mayinger condensation correlation. The details of 

the model with its limitations and applicability range are discussed in the TRAC-M/FORTRAN 

90 (Version 3.0) Theory Manual (Ref. 2.1). The assessment of the model against small-scale, 

separate effect data will be discussed in the following subsections. 

4.2.1 Akimoto Tests 

The data and correlations for this assessment are taken from Akimoto et al. (Refs. 4.2.1 and 

4.2.2), and the same test section was used in all experiments. The experimental facility is 

illustrated in Fig. 4.2.1. The test section was horizontal, 3.35 m long, and 0.02 x 0.05 m in 

rectangular cross-section. The steam produced by a steam generator was injected at one end of 

the test section after it was metered. The subcooled water was injected 1.5 m downstream of the 

steam entrance, with an injection tee along the entire width of the test section. The test section 

was transparent, which allowed flow visualizations. The two-phase fluid was emptied to an exit 

tank that was exhausted to the atmosphere. 

4.2.1.1 Condensation Experiments 

Akimoto et al. observed distinct oscillatory and stable wavy flow regimes. The water was 

injected after the steam flow rate was established. When the steam flow rate was small relative to 

the liquid injection rate, the downstream of the test section was plugged by water while the 

upstream of the section was filled with steam. The liquid plug traveled back and forth from the 

liquid injection point to the channel exit. Oscillations were caused by changing condensation 

rates and pressures, and their frequencies and amplitudes changed with the steam injection rate. 

When the liquid injection point was fully exposed to the steam, rapid condensation led the local 

pressure to drop below that of the channel exit. This forced the liquid plug to move toward the 

channel entrance. When the plug moved upstream of the liquid injection point, this movement 

insulated the injected liquid from the steam. The decreasing condensation rate caused the local 

pressure to increase until the plug moved toward the channel exit, and a new oscillation began. 

When the steam flow rate was sufficiently large relative to the liquid injection rate, the liquid 

plug was swept away from the test section, and a stratified flow regime with mist flow on top and 

stable wavy liquid film flow at the bottom was established in the test section. 

Akimoto et al. (Refs. 4.2.1 and 4.2.2) reported the following simple relation for the boundary 

between the oscillatory and stable flow regimes: 

M, = 0-84Mcm,, (4.2.1.1) 

and

Page 4-13
mcmax- Cpi{Tsat(Pexit)-Ti} (4.2.1.2) 

hig 

where 

M, is the mass flux of the injected steam, 

Mti is the mass flux of the injected liquid through the steam flow channel, 

CPI is the specific heat of the injected liquid, 

hig is the injected liquid's latent heat of vaporization, 

Tsat(Pexil) is the saturation temperature at atmospheric pressure, 

Ti, is the injected liquid's temperature, and 

MI, max is defined by Akimoto et al. as the condensability of the injected water. 

"the condensability is a measure of the maximum mass flux of the steam that the injected water can condense before reaching thermal equilibrium, assuming a complete mixing of water 

with steam (Ref 4.2.2). 

4.2.1.2 TRAC Model 

The TRAC input model for this facility is shown in Figs. 4.2.2 and 4.2.3. Two TEE 

components were used to simulate the horizontal test section and the exit tank. The rectangular cross-section is considered to be a circular geometry, with a hydraulic diameter of 0.286 m and 

actual flow area of 0.001 m2. The node sizes are selected in such a way that the length-to diameter (L/D) ratio is similar to that used in full plant analyses. 

4.2.1.3 Comparison of Measured and Predicted Results 

The following subsections discuss the assessment results and the analysis of the stratified mist 

flow, oscillatory plug flow, and transition from stable to oscillatory flows. Both TRAC-M(F77) 

and TRAC-M(F90) were run to predict the test data. 

4.2.1.3.1 Stratified-Mist Flow 

When the steam flow rate is high (M/Mc, max > 0.84), Akimoto et al. observed that the 

injected liquid was quickly atomized into droplets immediately downstream of the injection nozzle. The mixing process in the injection region was described as very rapid and complicated. Some coalesced droplets eventually reattached to the wall, and formed a wavy liquid film at the 

bottom of the test section. 

The liquid film temperature is determined by the condensation rate of steam, and was 

measured as a function of distance from the injection nozzle. For the purpose of this assessment, 

we chose the test results given in Fig. 4 of Ref. 4.2.1, with the following test conditions:

Page 4-14
steam mass flux

injected liquid mass flux =330 kgn-2s-1 

injected liquid temperature =323 K 

The predicted and measured liquid film temperatures are plotted in Fig. 4.2.4 as a function of 

the distance from the injection nozzle. The predicted temperatures correspond to t = 150 s 

because they remain constant. The liquid entering the test section is subcooled to 323 K. Shortly 

downstream (0.1 m), experimental data indicate that the liquid is heated up to 361.6 K, whereas, 

the calculated liquid temperature at this location is 358.0 K (indicating a disagreement of only 

3.6 K between calculated and measured data). The predicted and measured values at 0.2 m 

downstream of the injection point are almost the same (366.4 and 365.7 K, respectively). At 0.3 

m downstream of the injection point, the predicted and measured values (369.6 and 368.8 K) also 

agree very well. The correct prediction of the liquid film temperature indicates that the TRAC 

condensation model (including the heat transfer coefficient and the interfacial area) for the mist 

flow adequately predicts the rate of condensation and is expected to predict a reasonably correct 

rate of condensation for the stratified-mist flow regime encountered in LOCA situations. Since 

both code predictions are essentially the same, the conversion of TRAC-M(F77) to TRAC

M(F90) for this application is successful. 

4.2.1.3.2 Oscillatory Plug Flow 

Unstable flow with regular oscillations occurred when the liquid injection rate was increased. 

Akimoto et al. classified the oscillatory flow regime into three subregimes, identified as Type-A, 

Type-B, and Type-C. Type-A flow included very well-defined oscillatory behavior. The test 

reported in Fig. 5 of Ref. 4.2.2, performed with 20 kgm' 2 s-I steam and 600 kgm'2s-1 water mass 

fluxes, exhibited well-defined oscillatory behavior. This test was selected for assessment. Type

B flow included not so well-defined oscillatory behavior, while Type-C flow exhibited only small 

interface fluctuations and did not exhibit any periodicity. 

The predicted pressure trace at the injection location shown in Fig. 4.2.5 indicates that the 

pressure oscillated between 36.8 to 335.8 kPa. Table 4.2.1 lists the measured and calculated 

frequencies and amplitudes of the pressure. Notably, the amplitudes of the predicted pressure 

oscillation are higher than those measured. Nonetheless, a count of the number of oscillations 

between 15 and 40 s yielded a predicted oscillation frequency of 1.73 s-1, which agrees 

reasonably well with the measured data of 1.67 s-1. (See Fig. 4.2.6.) 

As seen in Figs. 4.2.7 and 4.2.8, a pressure increase near the liquid injection point (cell 4) 

corresponds to both a decrease in the void fraction in cell 4 and an increase in the void fraction in 

cell 13, clearly indicating that a plug of water is moving toward the steam injection point. When 

the plug moves upstream of the injection point, the condensation at the liquid-steam interface 

continues, heating up the liquid and eventually reducing the condensation rate. The pressure then 

continues to rise, and causes the plug movement to reverse, thereby exposing the injection point 

to the steam. When the injection point is exposed to the inflowing steam, a rapid condensation

Page 4-15

=40 kgm-2sql
occurs and the pressure drops below that of the exit tank. The plug of water then starts to move 

toward the injection point again, and a new oscillation follows the same history. 

Despite the complexity of the phenomenon described above, the frequencies and minimum 

pressures predicted by both codes agree reasonably well with the measured data. The 

overprediction of the maximum pressures can be explained if the condensation rate is 

underpredicted when the liquid steam interface passes the liquid injection point, where the 

condensation rate should reach a maximum. Inadequate modeling of the pressure losses can be 

another reason for the overprediction of the maximum pressure. Further investigation is needed 

to correctly identify the cause. 

4.2.1.3.3 Stable-to-Oscillatory Flow Transition 

This section assesses the ability of TRAC to predict oscillatory condensation, and compares the results with the transition correlation of Akimoto et al. (Eq. 4.2.1.1). In order to assess this 

capability, a test was designed in which the liquid flow rate for a given steam flow rate in the 

stable flow region is increased until the flow starts to show some oscillations. It is assumed that 

an oscillatory flow occurs when the amplitude of the pressure oscillations is >±15%. 

Fig. 4.2.9 shows how the steam and ECC injection flow rates were varied during this test, while Fig. 4.2.10 shows the results of oscillations calculated by TRAC-M(F77) and TRAC

M(F90). Figs. 4.2.11 through 4.2.13 contain the pressure oscillation traces for different portions of the transient. The predicted results and experimental observations are presented in Fig. 4.2.14, which plots the steam mass flux as a function of the condensability, given by Eq. 4.2.1.1. 

When the ratio Mv/Mc max becomes smaller than the experimental value of 0.84 at a given 

liquid flow rate, the liquid plugs the channel because the steam provided is not sufficient to 

overcome the condensate energy of the injected liquid. As shown in Fig. 4.2.14, the experimental 

constant of 0.84 in Eq. 4.2.1.1 is reasonably predicted for the stable-to-oscillatory flow transition 

for lower mass fluxes. However, for the higher steam flux, both versions of TRAC overpredict the condensability. The prediction of condensation rate depends on predictions of the 

condensation heat transfer coefficient (HTC) and the interfacial area. Unfortunately, no 

experimental data are available from Akimoto's tests to determine which parameter is 

overpredicted. 

4.2.1.4 Conclusions 

In both stable and mist-flow regimes, the TRAC condensation model yielded good predictions of the liquid temperature along the test section measured by Akimoto et al. The results indicated 

that the rate of condensation in the mist-flow regime was predicted reasonably well; therefore, it 

is expected that the model will also give reasonable results for a similar range of operating 

parameters in the mist-flow regime (i.e., lower steam mass flux) when it is used in full-plant 

simulations. 

The model also predicted the frequency of oscillations when the plug flow is considered. The 

motion of the plug in the predictions was similar to that described by Akimoto et al. and the 

mimimum pressure agreed with the data; however, the predicted pressure amplitude disagreed with the measured data, and the maximum peak pressure was overpredicted. The overpredicted

Page 4-16
maximum pressure could be attributable to the overprediction of the condensation rate when the 

steam-water interface was upstream of the injection point. In addition, the transition from stable 

to oscillatory flow was reasonably predicted for a steam mass flux at or below 40 kgm-2s'. For a 

higher steam flux, the condensability was overpredicted. 

Since predictions obtained using both TRAC-M(F77) and TRAC-M(F90) are essentially the 

same, the conversion of TRAC-M(F77) to TRAC-M(F90) is successful for this application. 

REFERENCES 

4.2.1 Akimoto, H., Y. Tanaka, T. Kozwa, A. Inoue, and S. Aoki, "Oscillatory Flows Induced 

by Direct-Contact Condensation of Flowing Steam with Injected Water," J. Nucl. Sci. 

Tech. 22(4), 269-283, 1985.

4.2.2 Akimoto, H., Y. Kozawa, A. Inoue, and S. Aoki, "Analysis of Direct Contact Conden

sation of Flowing Steam onto Injected Water with Multifluid Model of Two-Phase 

Flow," J. Nuci. Sci. Tech. 20 (12), 1006-1022, 1983.

Table 4.2.1 Data Comparison in the Junction Cell of the Test Section 

Type Experiment TRAC-M(F77) TRAC-M(F90) 

Oscillation Frequency (s- 1) 1.67 1.58 1.65 

Maximum Pressure (kPa) 140 335 400 

Minimum Pressure (kPa) 40 37 37

Page 4-17
Water

Figure 4.2.1 Sketch of the Experimental Apparatus of Akimoto et al.

Page 4-18
Water 

lr ection

Steam 

Injection

1 2 3 415 6 7 8 9 10 11 12 13 

1.5 m 2.15 m

Figure 4.2.2 TRAC Input Model

Page 4-19
Water 

Injection

5 6 7 8 9 10 

0.05 0.05 0.15 0.125 0.2 0.2

11 

0.4

12 

0.6

I I I I I I 1 1 ] 

0 0.05 0.1 0.2 0.36 0.5 0.7 1.0 1. 

Distance from Water Injection Nozzle to Cell Center

Figure 4.2.3 Noding Cell Lengths of the Test Section

Page 4-20

Cell Lengths

13 

0.35

5 1.98

I I i I B P I I I I I I I

I
( Akimoto Mist Flow Test

Comparison of Calculated and Measured Liquid Film Temperatures

0.5 1 1.5 2 

Distance from ECC Injection, (m) 

Figure 4.2.4 Comparison of Calculated and Measured Liquid Film Temperatures 

in the Akimoto Stratified-Mist Flow Test

380

360 

340

CL 

E 

a)

320

0

t
4e+05 

3e+05 

ci) 

S2e+05 

le+05

Akimoto Oscillatory Plug Test 

Comparison of Predictions of Pressure Variations at Injection Location

10 20 30

Time (s)

Figure 4.2.5 Comparison of Predictions of Pressure Variations at Injection Location 

in the Akimoto Oscillatory Plug Test

crq

0

0 40
4e+05 

3e+05 

S2e+05 

1le+05 

0

Akimoto Oscillatory Plug Test 

Comparison of Predictions of Pressure Variations at Injection Location

15 20 25 30 35 

Time (s) 

Figure 4.2.6 Comparison of Predictions of Pressure Variations at Injection Location 

in the Akimoto Oscillatory Plug Test

/

40

!
Akimoto Oscillatory Plug Test 

Comparison of Predictions of Pressure Variations at Injection Location

25 26 27 28 29

Time (s) 

Figure 4.2.7 Comparison of Predictions of Pressure Variations at Injection Location 

In the Akimoto Oscillatory Plue Test

4e+05 

3e+05

GrO

a

cI) 

C,, 

C,, 

0.

2e+05

le+05 

0

30
Akimoto Oscillatory Plug Test 

Comparison of Vapor Fraction Predictions at Injection Location

25 26 27 28 29 

Time (s) 

Figure 4.2.8 Comparison of Vapor Fraction Predictions at Injection Location 

In the Akimoto Oscillatory Plug Test

(

1.2 

1 

0.8

,-Q 

C-o

0 

"C5 

-0.6 

0 

CL 

co

0.4 

0.2 

0

30
Akimoto Transient Test 

ECC Injection and Steam Flow

50 100 

Time (s)

Figure 4.2.9 ECC Injection and Steam Flow in the Akimoto Transient Test

2 

1.5

,.j

C,

cc) 

0 

Co CA)

1

0.5 

0

0 150
(

50 100 

Time (s) 

Figure 4.2.10 Comparison of Pressure Oscillations at the Injection Point 

in the Akimoto Transient (0 - 140s)

Akimoto Transient Test (0 - 140s) 

Comparison of Pressure Oscillations at the Injection Point

8e+05 

6e+05

(

03 

L-. 

0.

4e+05

2e+05 

0

0 150
Akimoto Transient Test ( 0 - 40s) 

Comparison of Pressure Oscillations at the Injection Point

10 20 30

Time (s) 

Figure 4.2.11 Comparison of Pressure Oscillations at the Injection Point 

in the Akimoto Transient Test (0 - 40s)

1.6e+05 

1.4e+05 

1.2e+05 

le+05

0I 

SD 

Ci)

80000

60000

0 40

t
Akimoto Transient Test ( 50- 90s) 

Comparison of Pressure Oscillations at the Injection Point

60 70 80 

Time (s) 

Figure 4.2.12 Comparison of Pressure Oscillations at the Injection Point 

in the Akimoto Transient Test (50 - 90s)

K

6e+05 

4e+05

K

cJQ 

(b 

1.

a-, 

a.

2e+05 

0

50 90
6e+05 

4e+05

Akimoto Transient Test ( 100 - 140s) 

Comparison of Pressure Oscillations at the Injection Point

2e+05 

0 

100 110 120 130

Time (s) 

Figure 4.2.13 Comparison of Pressure Oscillations at the Injection Point 

in the Akimoto Transient Test (100 - 140s)

CU) 

ca) 

a.

0

140
(

Akimoto Transient Test 

Transition from Stable Wavy to Oscillatory Flow 

100 , , , 

Akimoto Formula Eq. 4.2.1.1 

* TRAC-M(F77) 

Il TRAC-M(F90) 

80 

E.' 

"- 60 

0 

M 40 

E 

Co, 

20 

0• I , I, 

0 20 40 60 80 100 

Condensability, Mc,max (kg/m**2*s)

Figure 4.2.14 Transition from Stable Wavy to Oscillatory Flow in the Akimoto Transient Test

/r 

\
4.3 Critical Flow

The TRAC critical flow model is described in Ref. 2.1. The 1-D critical flow model was 

programmed in TRAC-PFl to evaluate choked flow conditions with a coarse spatial mesh. 

Without the critical flow model, a fine mesh is required to accurately evaluate choking from the 

hydrodynamic equations. However, evaluating choking with a fine mesh is accurate only for 

smooth flow area changes between upstream mesh cells. The TRAC critical flow model 

improves computational efficiency by allowing a coarse mesh, and it also allows abrupt upstream 

flow area changes to be modeled while accurately predicting choked flow conditions. 

The critical flow model comprises three separate models, including a single-phase subcooled

liquid choked flow model, a two-phase liquid and gas choked flow model, and a single-phase gas, choked flow model. The single-phase subcooled liquid choked flow model is a modified form of 

the Burnell model (Ref. 4.3.1) for an upstream gas volume fraction <0.01, and it incorporates a 

nucleation delay model developed by Jones (Ref. 4.3.2). The two-phase liquid and gas choked 

flow model is an extension of the model developed by Ransom and Trapp (Ref. 4.3.3), on the 

basis of first principles and a characteristic analysis approach for an upstream gas volume fraction 

between 0.1 and 0.999. This model assumes thermal equilibrium between the gas and liquid 

phases. Because there is a discontinuity in the sound speed during the transition from liquid to 

two-phase flow, linear interpolation is used between the subcooled liquid and two-phase models 

for an upstream gas volume fraction between 0.01 and 0.1. The single-phase gas choked flow 

model assumes ideal gas isentropic expansion for an upstream gas volume fraction >0.999. 

4.3.1 Marviken Tests 

The Marviken facility, located in Sweden, was originally designed to be part of a nuclear plant and, thus, is a full scale facility. The Marviken blowdown tests provide data for assessing the 

ability of computer codes to predict subcooled and saturated critical flow in large-diameter pipes. 

4.3.1.1 Facility Description 

The Marviken test facility consists of a pressure vessel and a discharge pipe, which contains a 

test nozzle with the minimum flow area, a rupture disk assembly, and a ball valve that is used to 

isolate the vessel after testing. The pressure vessel and discharge pipe are shown in Figs. 4.3.1 

and 4.3.2. The vessel includes part of the core superstructure and three gratings to limit vortex 

formation. Pressure and temperature transducers are located in the vessel and discharge pipe, as 

shown in Figs. 4.3.1 and 4.3.2. The break flow is measured with a pitot tube, and is also 

calculated from differential pressure measurements. The data uncertainties are 90 kPa for the 

pressure data, 2 K for the temperatures, 10% and 15% for subcooled and saturated break flows 

determined by the pitot-static method, and 15% and 20% for subcooled and saturated break flows 

derived from differential pressure. 

4.3.1.2 TRAC Model 

Five critical flow experiments (Marviken Tests 4, 13, 20, 22, and 24) from the Marviken 

blowdown test facility are selected for assessment calculations. 

The TRAC model of the Marviken facility has four components. Two PIPE components are 

used to represent the vessel and discharge pipe, a BREAK component provides the pressure

Page 4-32
boundary condition downstream of the break, and a zero-velocity FILL closes the top of the 

vessel. The vessel, discharge pipe, and break nozzle noding for Test 4 are typical of the input 

models, and are shown in Figs. 4.3.3 and 4.3.4. The internally set default subcooled and saturated 

break flow multipliers of 1.0 were used for all calculations. 

The initial and boundary conditions were varied in the remaining Tests 13, 20, 22, and 24. 

Table 4.3.1 summarizes the differences in the Marviken tests and the TRAC models. Both 

TRAC-M(F77) and TRAC-M(F90) were run to predict test data. 

4.3.1.3 Comparison of Predictions and Measurements for Test 4 

The calculated pressures, fluid temperatures, and break flow are in reasonable agreement with 

the measured data from Test 4. Calculated values are within the data uncertainties for most of the 

50 s test duration. The values calculated by TRAC-M(F90) closely agree with the results that 

were previously calculated by TRAC-M(F77). 

The calculated break flow is compared with the break flow measurements from the pitot tube 

and the vessel differential pressure in Fig. 4.3.5. It should be noted, however, that the break flow 

derived from the vessel differential pressure is not valid for the first -5 s because of momentum 

effects on the measured pressures. As a result, the pitot tube measurements should be used for 

comparison purposes. The TRAC break flow underpredicts the subcooled critical flow for the 

first 19 s of the test, but was in reasonable agreement with the test data during the saturated 

blowdown after that time. The underprediction of the break flow results when the flow conditions 

at the nozzle become saturated even though the upstream incoming flow is subcooled. 

The vessel pressure comparison is shown in Fig. 4.3.6. Except for the initial pressure dip in 

the test data caused by delayed nucleation, the calculated pressure is just under the lower 

uncertainty limit until -27 s, and then is within the uncertainty limit for the rest of the test. A 

similar comparison was obtained for the discharge line pressure shown in Fig. 4.3.7. Again, the 

calculated pressure is very close to the lower uncertainty of the measured pressure for most of the 

time. Both the calculated and measured fluid temperatures are very close to saturation for most of 

the time. 

The fluid temperature comparison in the lower vessel shown in Fig. 4.3.8 is very similar to 

comparisons in the middle of the vessel and the discharge pipe. A temperature increase occurs in 

the first 20 s as warmer fluid from the upper vessel drains into the lower vessel. The calculated 

temperature increase is smoother because of the averaging of the liquid temperature within a 

hydraulic cell. The fluid temperature comparison in the upper vessel is very similar to the 

pressure comparison because both the measured and calculated fluid temperatures are very close 

to saturation. 

4.3.1.4 Comparison of Predictions and Measurements for Test 13 

The Test 13 nozzle was shorter, and had a smaller throat diameter than the nozzle in Test 4. 

Test 13 also had more initial subcooling at the nozzle inlet. The lower velocities and pressure 

drops upstream of the nozzle, along with the higher subcooling, were sufficient to prevent vapor 

formation from reducing the subcooled break flow in the calculation for Test 13.

Page 4-33
Fig. 4.3.9 compares the calculated and measured break flows in the first 75 s. When the break 

flow is subcooled, the calculated flow is generally within the uncertainty of the test data. Voiding 

at the nozzle exit remains below 0.01, which allows delayed nucleation in the subcooled critical 

flow model to take effect. As warmer liquid from the vessel enters the discharge line, the 

calculated break flow is reduced by vapor formation, and the break flow is underpredicted for the 

remainder of the test. As suggested in Ref. 4.3.1, the nonequilibrium effects in Test 13 were more 

pronounced during the saturated flow because of the short nozzle length. In addition, there is a 

large underprediction of the flow because the two-phase TRAC critical flow model assumes 

thermal equilibrium between phases. 

The upper vessel pressure comparison for Test 13 is shown in Fig. 4.3.10. Except for the 

initial dip in pressure caused by the delayed nucleation in the test, the calculated pressure is within 

the uncertainty of the test data. Similar comparisons were obtained at lower vessel elevations and 

in the discharge pipe. 

The fluid temperature predictions in the upper vessel are within the uncertainty of the test 

data, with the exception of the initial dip that was caused by delayed nucleation in the test. When 

the fluid reaches saturation, the measured and calculated fluid temperatures closely follow 

saturation. As in Test 4, the calculated lower vessel fluid temperature for Test 13 increased to 

saturation more smoothly than the data, as shown in Fig. 4.3.11. Calculations by both TRAC

M(F77) and TRAC-M(F90) are essentially the same. 

4.3.1.5 Comparison of Predictions and Measurements for Test 20 

The break nozzle in Test 20 was shorter than in Test 4, but it had the same throat diameter. 

Test 20 also had less initial subcooling in the bottom of the vessel. With the reduced initial 

subcooling, the subcooled break flow was again underpredicted by both codes during the first 7 s 

of the test, as shown in Fig. 4.3.12. However, the saturated break flow predictions obtained using both codes were within the uncertainty of the test data for the remainder of the test after 7 s. 

Fig. 4.3.13 shows the upper-vessel pressure comparison for Test 20. The pressure calculated 

by both codes is within the measurement uncertainty, with the exception of the initial pressure dip that was caused by the delayed nucleation in the test. Calculated pressures at lower vessel 

elevations and in the discharge pipe compared similarly with the measured data from Test 20. 

The fluid temperature comparisons for Test 20 were much like those of the other Marviken tests, in that (1) the predicted temperatures in the vessel are generally within the data uncertainty except for the initial dip, (2) the discharge line temperatures increase to saturation more smoothly in 

calculations and are within the uncertainty afterward, and (3) calculations produced by TRAC

M(F90) are essentially the same as those produced by TRAC-M(F77). 

4.3.1.6 Comparison of Predictions and Measurements for Test 22 

Test 22 had the same break nozzle as Test 20, which was shorter than the Test 4 nozzle but had 

the same throat diameter as the Test 4 nozzle. Test 22 had the maximum initial subcooling, both 

in the bottom of the vessel and in the discharge pipe. The duration of the subcooled break flow 

was prolonged until -30 s in Test 22 (Fig. 4.3.14) because of the high initial subcooling. Fig. 4.3.14 shows that both codes underpredict portions of the subcooled break flow, but are 

consistently within the data uncertainty during the saturated blowdown after 30 s.

Page 4-34
The vessel pressure comparison for Test 22 is shown in Fig. 4.3.15. The initial dip in the 

pressure is not predicted because the codes do not have delayed nucleation in the vapor generation 

model. After the initial pressure dip, pressures are generally underpredicted for the remainder of 

the subcooled blowdown. However, during the saturated blowdown after 30 s, the break flow and 

nonequilibrium effects are reduced, and the predicted pressures are within the data uncertainty. 

Pressures calculated by both codes at lower vessel elevations and in the discharge pipe compared 

similarly with the measured data from Test 22. Fluid temperature comparisons for Test 22 are 

very much like those of the other Marviken tests. 

4.3.1.7 Comparison of Predictions and Measurements for Test 24 

Test 24 had the shortest break nozzle, but its throat diameter was the same as the Test 4 nozzle. 

Test 24 also had the same initial subcooling in the bottom of the vessel as Test 4, and greater 

subcooling in the discharge pipe. The short nozzle increases the acceleration and magnitude of 

nonequilibrium effects upstream of the throat, and the codes do not correctly calculate the 

nonequilibrium effects. As a result, both codes underpredict the subcooled break flow to a greater 

degree than in the other tests, as shown in Fig. 4.3.16. The calculated break flows are within the 

data uncertainty during the saturated blowdown after 30 s, but the vessel emptied -15 s late in the 

calculation because of the underprediction of the subcooled break flow. 

Fig. 4.3.17 shows the upper vessel pressure comparison for Test 24. Again, both codes fail to 

predict the initial pressure dip, and underpredict the pressures during much of the subcooled 

blowdown. The pressure comparisons are closer after the transition to saturated break flow at 

-30 s. Similar pressure comparisons were obtained in the lower vessel and discharge pipe. Fluid 

temperature comparisons for Test 24 are very much like those of the other Marviken tests. 

43.1.8 Conclusions 

Calculations were performed using TRAC-M(F90) and TRAC-M(F77) codes, and the results 

were compared with test data to assess the TRAC critical flow model. Results of the comparisons 

for the Marviken tests with larger nozzle diameters generally showed that both codes correctly 

predict the saturated critical flow, but sometimes predict an early transition from subcooled to 

saturated critical flow. For Marviken Test 13, which has a smaller nozzle diameter, both codes 

underpredict the saturated critical flow. The early transition to the saturated critical flow is a 

result of the vapor generation model, which does not model delayed nucleation. The 

depressurization and break flow rates are both adversely affected by the early transition in the 

calculations. The comparisons are best for the tests with higher initial subcooling, longer break 

nozzle inlet sections, and larger nozzle diameters. 

Comparisons of code predictions with the Marviken tests indicate that whereas both codes 

generally calculate saturated break flow correctly, the absence of nucleation delay in the vapor 

generation model over most of the two-phase flow regime can cause differences in the break flow 

and depressurization rate during the subcooled break flow. The differences seem to correspond to 

nonequilibrium conditions just upstream of the break, and result in a brief period when both the 

break flow and pressure are underpredicted. For the short-nozzle and small-diameter tests, 

Marviken Test 13 and Test 24, nonequilibrium effects become dominant. In these tests, during the 

period where the measured flow is highly underpredicted by both codes, the measured mass-flux 

generally agrees with that predicted by the Henry-Fauske Critical-Flow Model (Ref. 4.3.4). This

Page 4-35
can be seen in Table 4.3.2, which compares the TRAC calculated and measured critical mass 

fluxes at selected test times to the mass fluxes calculated by three different critical flow models, including (1) Homogeneous Equilibrium Model (HEM), (2) Moody Model, and (3) Henry-Fauske Model. For Tests 13 and 24, the measured mass fluxes agree more closely to those predicted by the Henry-Fauske model. It is recommended that a nonequilibrium critical-flow model, such as 

the Henry-Fauske Model, be implemented into the code as a user option for modeling the critical 

flow for small-diameter nozzles, orifices, and short tubes. 

The magnitude of the initial pressure dip in the Marviken tests, which is not calculated, seemed to depend on the rate of depressurization more than any other single parameter. Comparisons are generally better for the tests with slower depressurization rates. It is clear that 

additional modeling of nonequilibrium effects is needed. Plausible solutions might be to extend the Jones nucleation delay model to higher void fractions and/or to develop an improved thermal 

nonequilibrium model for the two-phase flow. 

All calculations performed using TRAC-M(F90) are essentially the same as those performed using TRAC-M(F77), demonstrating that the conversion of TRAC-M(F77) from FORTRAN 77 

to FORTRAN 90 has been successful for this type of application. 

4.3.2 Edwards Blowdown Test 

The Edwards blowdown test facility (Ref. 4.3.5) is a separate-effects facility that was built in 

England in the late 1960s to study depressurization phenomena associated with subcooled water 

that is initially stagnant. This facility is much smaller than any component of a full-scale reactor 

system. 

4.3.2.1 Facility Description 

The Edwards blowdown test apparatus (shown in Fig. 4.3.18) consists of a horizontal pipe that 

is sealed at one end, an orifice at the pipe exit, and a glass rupture disk. The pipe, which is 

initially filled with subcooled liquid, is electrically heated and insulated with asbestos. Pressures 

are measured at four locations along the pipe, and the void fraction and fluid temperature are 

measured at a point near the middle of the pipe. The break flow rate and pipe wall temperatures 

are not measured. 

4.3.2.2 TRAC Model 

The TRAC-M model of the Edwards blowdown test facility has three components, as shown 

in Fig. 4.3.19. A PIPE component represents the blowdown pipe, a zero-velocity FILL terminates 

the closed end of the blowdown pipe, and a BREAK component provides the pressure boundary 

condition downstream of the break. 

4.3.2.3 Comparison of Predicted and Measured Results 

Pressures, fluid temperature, and void fractions in the middle of the blowdown pipe are 

calculated using both TRAC-M(F77) and TRAC-M(F90). The predictions reasonably agree with 

the measured data from the test. Although the calculated parameters are sometimes outside the 

data uncertainty, the major trends in the data were predicted.

Page 4-36
The pressure comparisons at the closed end and the break end of the blowdown pipe are 

shown in Figs. 4.3.20 and 4.3.21, respectively. Measurement uncertainties were not given in the 

original publication for the Edwards blowdown test; however, an uncertainty of -0.3 MPa has 

been suggested in Ref. 4.3.3. Figs. 4.3.20 and 4.3.21 show that the initial subcooled 

depressurization lasted -0.1 s in both the test and the calculation. However, the predictions are 

frequently outside the data uncertainty during the saturated blowdown after 0.01 s. The pressure 

comparisons for the other two pressure measurements, in the midsection of the pipe, are very 

similar to those shown in Figs. 4.3.20 and 4.3.21. 

The fluid temperature measured near the center of the pipe is compared with the 

corresponding calculated temperature in Fig. 4.3.22. After -0.1 s, the liquid and vapor 

temperatures in the calculation are equal to the saturation temperature. Thus, the fluid 

temperature comparison is very similar to the pressure comparisons shown in Figs. 4.3.20 and 

4.3.21 after this time. The void fraction was measured at the same location as the fluid 

temperature, and the void fraction comparison is shown in Fig. 4.3.23. The overall trend in the 

void fraction is captured in the calculations; however, there are significant differences between 

the measured and calculated void fractions in the first half of the test when the flows and 

depressurization rate are higher. Also, the oscillations in the measured void fraction from 0.15 to 

0.25 s do not occur in the calculations, 

4.3.2.4 Conclusions 

The parameters calculated by both codes for the Edwards blowdown test followed the same 

trends as the test data, and correctly predicted the timing of major events. However, there are 

periods when the calculated parameters noticeably differ from the measured values, and the 

oscillations in the measured void fraction are not captured in the calculations. Sparse data from 

the facility and the lack of data uncertainty information make it difficult to determine which 

models in the code contribute most to the differences between the test results and predictions, or 

whether the differences are attributable to incorrect initial conditions in the input model. 

The Edwards blowdown test comparison shows that although the major trends and timing of 

events in the test data are captured in calculations, the calculated parameters do not always match 

the values measured during the test. Also, the oscillations in the measured pressure upstream of 

the break and in the void fraction in the middle of the pipe are not completely captured in the 

calculations. Uncertainties in the initial temperature distribution and the lack of break flow and 

pipe wall temperature measurements make it difficult to attribute differences to a specific model 

in the code. 

Calculations performed using TRAC-M(F77) and TRAC-M(F90) are reasonably close, but 

they are not identical. TRAC-M(F90) contains some updates that may affect the critical flow 

calculations. Consequently, it is necessary to incorporate the same updates in TRAC-M(F77) to 

make precise comparisons. At this point, we can say that the predictions are reasonably close, and 

that the conversion from TRAC-M(F77) to TRAC-M(F90) has been successful for this type of 

application.

Page 4-37
REFERENCES

4.3.1 Ransom, V.H., R.K. Wagner, J.A. Trapp, K.E. Carlson, D.N. Kaiser, H. Kuo, H. Crow, R.A. Nelson, and S.W. James, "RELAP/MOD1 Code Manual, Volume 1: System Models and Numerical Methods," Idaho National Engineering Laboratory Report EGG-2070, NUREG/CR-1826, March 1982. 

4.3.2 Jones, O.C., "Flashing Inception in Flowing Liquids," Brookhaven National 

Laboratory report BNL-NUREG-51211, 1980. 

4.3.3 Ransom, V.H., and J.A. Trapp, "The RELAP5 Choked Flow Model and Application to 

a Large-Scale Flow Test," in ANS/ASME/NRClnternational Topical Meeting on Nuclear Reactor Thermal-Hydraulics, pp. 799-819, Saratoga Springs, New York, 1980. 

4.3.4 Henry, R.E., and H.K. Fauske, "The Two-Phase Critical Flow of One-Component Mixtures in Nozzles, Orifices, and Short Tubes," Journal of Heat Transfer, Transactions, 93, May 1971. 

4.3.5 Edwards, A.R., and T.F. O'Brien, "Studies of Phenomena Connected with the 

Depressurization of Water Reactors," Journal of British Nuclear Energy Society, 9, 125-135, April 1970.

Page 4-38
Table 4.3.1 Marviken Tests and TRAC Noding

Test 4 Test 13 Test 20 Test 22 Test 24 

Nozzle straight section 1.500 0.590 0.731 0.731 0.166 

length (m) 

Nozzle diameter (m) 0.509 0.200 0.500 0.500 0.500 

Length-to-diameter ratio 2.95 2.95 1.46 1.46 0.33 

Initial pressure (MPa) 4.94 5.09 4.987 4.93 4.96 

Initial subcooling near 37 31 7 52 33 

vessel bottom (K) 

Initial subcooling at nozzle 63 95 77 95 83 

inlet (K) 

Initial liquid level above 17.59 17.52 16.65 19.64 19.88 

vessel bottom (m) 

Number of cells in PIPE 15 15 38 38 15 

representing vessel 

Number of cells in PIPE 17 13 14 14 17 

representing lower vessel 

and discharge pipe

Page 4-39
Table 4.3.2 Comparison of Measured Mass Fluxes and the Mass Fluxes Calculated by 

TRAC-M(F77) and Different Critical Flow Models 

Marviken Test No. 4 13 20 22 24 

Test Time (s) 41 100 40 40 15 

Measured Mass Flux (pitot) 15971 34364 18640 14860 40881 

(kg/(s-m2

)a (Ap) 16428 34841 21279 17045 41171 

TRAC-M(F77) Calculations 

Exit Pressure (MPa) 3.281 3.963 3.676 2.394 2.855 

Exit Void Fraction 0.2941 0.0608 0.3321 0.2581 0.0448 

Mass Flux (kg/(s-m2) 18466 23130 21448 15854 30805 

Mass Flux (kg/(s-m2

) 

determined from: 

HEM 14782 17679 15851 11910 14022 

Moody Model 26635 29558 27978 22982 25233 

Henry-Fauske Model 27668 35641 28391 24285 30549 

a The first value given is the mass flux determined from measured pitot tube data. The second value is the mass flux 

determined from differential vessel pressure measurements.

Page 4-40
24.55 m 

1.5 

21.515 m 

11.4 m 

3.4 m

1.4 m 

0.0 m

P103 

T401

*- 5.22 m --

(23.130 m)

(20.543 m)

VORTEX 

MITIGATOR 

MODERATOR TANK

T402 (0.740 m) _P106 (0.525 m) 

--14 0.752 m 

Figure 4.3.1 Marviken Pressure Vessel

Page 4-41

.... ......... ... . /
0.740 m0.0 m 

-2.790 m 

-4.568 min

-5.568 m

T402 

P 106 

4--- DISCHARGE 

0.752 m

(0.740 

'0.525

BALL VALVE

0.780 m

0.752 m

P109 

T405

TEST NOZZLE

: . - RUPTURE DISKS 

Figure 4.3.2 Marviken Discharge Pipe, Test Nozzle, and Rupture Disk Assembly

Page 4-42

m) 

m)

PIPE

-4ý.868 m)
FILL 

1

PIPE 

Note: For Tests 20 and 

22, this PIPE component 

is noded into 38 cells 

PIPE 

Noding in this section 

vanes with test See 

detailed noding diagrams

Q Zero-Row Boundary

C ) Pressure Vessel 

G Discharge Section

BREAK 0 (D Pressure Boundary 

Figure 4.3.3 Pressure-Vessel and Discharge-Pipe Noding

Page 4-43
14 

(a) Test 4

6 

11 

(b) Test 13

Note: Not to scale

6

12 

1

(c) Tests 20 and 22

6 

15 

16 

1

(d) Test 24

Figure 4.3.4 Nozzle and Rupture-Disk Assembly Noding for Marviken Tests

Page 4-44
\

20 40 

Time (s)

Figure 4.3.5 Comparison of Calculated and Measured Break Flows for Marviken Test 4

15000

(

"(1 CO) 

E 

0

10000 

5000

0

0 60

Marviken Test 4 

Comparison of Calculated and Measured Break Flows
Marviken Test 4

Comparison of Calculated and Measured Pressures in Upper Vessel

10 20 30 40

Time (s) 

Figure 4.3.6 Comparison of Calculated and Measured Pressures in the Upper Vessel 

for Marviken Test 4

5e+06 

4.5e+06 

4e+06

L. 

L.

3.5e+06 

3e+06

0 50

K
(

Marviken 4 Test 

Comparison of Calculated and Measured Discharge Line Pressures

20 40 

Time (s) 

Figure 4.3.7 Comparison of Calculated and Measured Discharge Line Pressures for 

Marviken 4 Test

(

6e+06 

5e+06

(

0

CU (. 

CD I

Co

4e+06 

3e+06

2e+06 

1 e+06

0 60
Marviken Test 4

Comparison of Calculated and Measured Lower Vessel Fluid Temperatures 

540 

0-4 TRAC-M(F77) 

530 53 X..... X Exp. Temperature Data, T402 113.- -.E TRAC-M(F901 tln-nn0nn4 /

I .I

0 20 40 60

Time (s)

Figure 4.3.8 Comparison of Calculated and Measured Lower Vessel Fluid Temperatures for 

Marviken Test 4

520

E 

CD 

.'a

510 

500 

490

480 

470 

460

i
(

Marviken Test 13

Comparison of Calculated and Measured Break Flows

50 100 

Time (s)

Figure 4.3.9 Comparison of Calculated and Measured Break Flows for Marviken Test 13

/

3000

o0 

OQ 0 

(I 

CO 

E 

I--

2000 

1000

0

0 150
Marviken Test 13

Comparison of Calculated and Measured Upper Vessel Pressures

0 50 100

Time (s) 

Figure 4.3.10 Comparison of Calculated and Measured Upper Vessel Pressures 

for Marviken Test 13

5.5e+06 

5e+06

(L 

0

0A 4.5e+06 

4e+06 

3.5e+06

150

t l
Marviken Test 13 

Comparison of Calculated and Measured Lower Vessel Temperatures

50 100 

Time (s) 

Figure 4.3.11 Comparison of Calculated and Measured Lower Vessel Temperatures 

for Marviken Test 13

/

540 

520

,ia

I

CT 

.L E 

Cir

500 

480

460 

440

0 150

C
Marviken Test 20 

Comparison of Calculated and Measured Break Flows

Time (s)

Figure 4.3.12 Comparison of Calculated and Measured Break Flows for Marviken Test 20

10000 

8000

a.Q 0 

CO, 

W E 

I-F

6000 

4000

2000 

0
( (

Marviken Test 20

6e+06 

5.5e+06 

5e+06 i 

4.5e+06 

4e+06 

3.5e+06 

3e+06 

2.5e+06 

2e+06 

1.5e+06 

1 e+06 

0

Comparison of Calculated and Measured Pressures in Upper Vessel

20 40 

Time (s) 

Figure 4.3.13 Comparison of Calculated and Measured Pressures in the Upper Vessel for 

Marviken Test 20

OQ 

0 

(Ji

0

ci) 

1... 

(I) 

U, ci) 

0

60
Marviken Test 22 

Comparison of Calculated and Measured Break Flows

Time (s)

Figure 4.3.14 Comparison of Calculated and Measured Break Flows for Marviken Test 22

K

15000

cia 

JI

0) 

Ch "Cn 

CU) 

E 

3§ 

4-0 0

10000 

5000

0

t
Marviken Test 22

Comparison of Calculated and Measured Pressures in Upper Vassel

20 40 60

Time (s) 

Figure 4.3.15 Comparison of Calculated and Measured Pressures in Upper Vessel 

for Marviken Test 22

(

5e+06 

4e+06

rlQ

ci) 

0.. 

a3e+06

2e+06 

le+06

0
Marviken Test 24 

Comparison of Calculated and Measured Break Flows

Time (s)

Figure 4.3.16 Comparison of Calculated and Measured Break Flows for Marviken Test 24

15000

-P, 

(A. 

LCh

0 

E 

0.

10000 

5000

0

i
Marviken Test 24 

Comparison of Calculated and Measured Pressures in Upper Vessel

20 40 60

Time (s) 

Figure 4.3.17 Comparison of Calculated and Measured Pressures in the Upper Vessel for 

Marviken Test 24

5e+06 

4e+06

(

cial

CO) 

CO) 

0~

3e+06

2e+06 

1 e+06

0 80
4.096 m 

GS-7 GS-6 GS-5 G5-4 0S-3 GS-2CS- m .PJPE G z J C -8A EAK m PIPE_____________ BREAK END

-Dimension 

A 

B 

C 

D E 

F 

G 

H

m 

0.158 

0.835 

0.911 

0.555 

0.555 

0.835 

0.079

VACUUM VALVE

Figure 4.3.18 Edwards Blowdown Test Apparatus

GS-7 GS-6 GS-5 GS-4 GS-3 

113 181 1261

GS-2 GS-1 

1341 6172:

PIPE 

(0

BREAK 

0

Figure 4.3.19 TRAC Model from the Edwards Blowdown Test

Page 4-58

0.073 

CONTROL 

VALVE 

HYDRAUL I 

PUMP 

WATER 

SUPPLY 

TANK

C

FILL 

0
Edwards Blowdown Test

Comparison of Calculated and Measured Pressures at GS-1

0.2 0.4 

Time (s) 

Figure 4.3.20 Comparison of Calculated and Measured Pressures at GS-1 

for the Edwards Blowdown Test

(/

8e+06 

6e+06 

4e+06 rQQ

a. 

ao 

I.-

2e+06

0 

0 0.6
Edwards Blowdown Test 

Comparison of Calculated and Measured Pressures at GS-7

0.2 0.4 

Time (s) 

Figure 4.3.21 Comparison of Calculated and Measured Pressures at GS-7 

for the Edwards Blowdown Test

8e+06 

6e+06

tI

ca 

() 1.

4e+06

2e+06 

0

0 0.6
(

0.2 0.4 

Time (s) 

Figure 4.3.22 Comparison of Calculated and Measured Vapor Fractions at GS-5 

for the Edwards Blowdown Test

( 

Edwards Blowdown Test 

Comparison of Calculated and Measured Vapor Fractions at GS-5

1 

0.8

(

GO 

Cb

S0.6 

0 

IJ... U-' 

0 

Cz 

> 0.4

0.2 

0

0 0.6
Edwards Blowdown Test 

Comparison of Calculated and Measured Liq. Temps. at GS-5

0 0.2 0.4

Time (s)

Figure 4.3.23 Comparison of Calculated and Measured Liquid Temperatures at GS-5 

for the Edwards Blowdown Test

550 

500

CD 

E 

a) 

_0 

.5

450

400 

350

0.6

t
4.4 Reflood

This section presents the results of an assessment of the capabilities of two codes, namely 

TRAC-M(F77), Version 5.5, and TRAC-M(F90), Version 3670, to predict reflood phenomenon 

and provides some conclusions on the conversion of TRAC-M(F77) to TRAC-M(F90). The 

assessment of the predictive capability of reflood is based on some of the tests performed in the 

Lehigh University test facility, the FLECHT-SEASET test facility, the Cylindrical Core Test 

Facility (CCTF) and the Slab Core Test Facility (SCTF). An assessment with Run 31504 from the 

FLECHT-SEASET test facility shows that the predictions of the reflood phenomenon by the 

TRAC codes are not accurate, but will yield conservative results although agreements in other 

tests are better. One reason for better agreement is that some of the TRAC reflood modeling 

parameters are based on test results of Run 14 performed at CCTF (Ref. 2.1). 

The Office of Nuclear Regulatory Research has a research program in place at Pennsylvania 

State Univeristy (PSU) to improve the agency's capability to model reflood phenomena using its 

system analysis codes. The findings in this report provide guidance on how future modeling of 

reflood should proceed. Cases in this report exhibit the overall trend that TRAC-M(F90) 

overpredicts peak clad temperature (PCT), even though some predictions are more in reasonable 

agreement in some runs made at the CCTF, SCTF, and LOFT (IET) facilities than in Run 31504 in 

the FLECHT-SEASET facility. The TRAC-P code series has been used by NRC to perform 

LOCA analysis because of this conservative trend. However, as the agency begins to risk-inform 

its regulations, new designs are submitted and licensees request further uprates. NRC would be 

better equipped to audit licensee submittals with a less conservative, more accurate reflood 

model. In order to model reflood accurately for all designs, conditions, and test facilities, the 

proper physics must be modeled mechanistically. This assessment provides a guidance for the 

development of more accurate models. 

4.4.1 Lehigh Rod Bundle Test 02/24/85-20 

The Lehigh University rod bundle (LRB) test facility was designed, built, and operated by 

Lehigh University to simulate post-critical heat flux (CHF) phenomena under nonequilibrium 

conditions in a rod bundle at low mass fluxes, low to moderate vapor qualities, and near 

atmospheric pressures. A total of 467 data points were generated from 144 experimental runs, 

which were executed with the following range of conditions: 

Coolant mass flux 0.1 to 26 kgm-2s-I 

Inlet temperature/quality 40'C subcooled to 0.4 

Pressure 105 to 120 kPa 

Heat Flux 5 to 43 kWmn2 

Post-CHF experiments were conducted in three different modes, including (1) steady state 

experiments with a fixed CHF location, (2) reflood experiments with slowly advancing CHF 

locations (propagating quench front), and (3) boiloff experiments with retreating CHIF locations. 

The test used for this assessment, LRB Run 02/24/85-20, was one in a series of advancing quench 

front tests.

Page 4-63
4.4.1.1 Facility Description 

The LRB test facility was designed and built for convective boiling and heat transfer studies 

using steam/water mixtures. The purpose of this research was to obtain post-CHF heat transfer 

data with nonequilibrium vapor temperature measurements at two axial positions and several 

radial positions downstream from a fixed or advancing CHF point. A detailed facility description is provided in Ref. 4.4.1, and much of the descriptive information that follows was taken directly 

from that document. A sketch of the facility is provided in Fig. 4.4.1. 

Water was drawn from a surge tank, passed through filters to remove solid impurities, through 

a flow metering station and pumped to a vertical boiler that produced the desired flow quality to 

the test section. After passing through the test section, the fluid passed to a tank where the water 

and steam were separated. A regulating valve on the steam outlet was varied to produce the 

desired test section pressure. The two-phase fluid was then condensed and cooled by a water

cooled condenser, and returned to the surge tank. 

The test section consisted of a heated shroud containing a 3 x 3 rod bundle assembly that was 

representative of a PWR rod bundle. A cross-section of the bundle is shown in Fig. 4.4.2. The 

dimensions of the rods and the shroud were selected to provide equal hydraulic diameters in all 

subchannels. The shroud, made of 2-mm thick Inconel 625 alloy, was heated by radiation from a 

tubular furnace with three independently controlled zones. The shroud heat flux was to be equal 

to the test rod heat flux, and 12 chromel alumel thermocouples were brazed onto the outside 

surface of the shroud at different axial locations. The test section had a length of 1.220 m, and 

incorporated four ports for vapor temperature and pressure probes. A rod spacer was located 762 

mm from the test section inlet, as shown in Fig. 4.4.3, and the inlet contained a strainer to provide 

uniform flow across the bundle cross-section. 

The test rods (9.5-mm O.D.) were heated internally by electric resistance ribbons that were 

embedded in boron nitride. The rod clad was made of stainless steel. A hot patch was also 

available to condition the thermal state of the inlet two-phase flow, and to stabilize the quench 

front in tests for which a fixed CHF point was desired. The hot patch technique was used to 

produce dryout at the test section inlet. The bottom hot patches consisted of an outer copper block 

surrounding the shroud and nine short hot rods, each of which acted as an individual hot patch for 

one of the nine rods in the test bundle. The hot rods were similar to test rods in construction, but 

provided higher heat fluxes and were individually temperature controlled. 

The test section had 108 thermocouples on the test rods in the 3 x 3 bundle assembly. Each of 

the 9 test rods was equipped with 12 internal thermocouples at different axial elevations. There 

were also 2 thermocouples in the aspirated vapor probes to measure the vapor superheat 

temperature. Because the clearance between test bundle rods was only 3 umm, the vapor probes 

were of a small size (2.13-mm-diameter outer tube) to minimize disruption of the flow as it passed 

the probe. The probe was also traversable in the cross-sectional plane to enable the measurement 

of lateral variations in vapor superheat. 

Absolute pressure in the test bundle was measured, as well as the pressure drop between the 

four taps on the shroud walls. Flow rates were measured with variable area flow meters. Power 

fed to all components of the test loop was measured by power transducers. The loop was 

designed and instrumented to collect the desired data on flow conditions.

Page 4-64
4.4.1.2 Test Procedure and Description

With single-phase water flow in the test section, the furnace power was turned on. It took 

about 1 to 2 hours to reach the steady-state furnace operating temperature, because of its large 

thermal mass. When the steady-state operating temperature was reached, the single-phase water 

flow was diverted from the test section to the bypass line by closing the test section inlet and 

outlet, and opening the bypass valves. At the same time, the pre-test cooling fluid (steam) was 

introduced to the test section. (Pre-test cooling was necessary to keep the shroud temperatures 

within limits, because of the radiation from the furnace, while the hot patches and the test section 

were preheated.) The steam flow rate was manually adjusted to maintain the shroud temperatures 

and test section pressure in the ranges of 300-400'C and 1-10 psig, respectively. The hot-patch 

was heated to approximately 5000C, while the test rods and hot rods were heated to approximately 

6000C. 

During this pre-heating period, the boiler power was adjusted to obtain the desired test section 

inlet quality, while the flow was passing through the bypass line. Finally, when all temperatures 

were high enough to initiate the test, the two-phase fluid was diverted to the test section. 

At the time of flow diversion, all instrumentation was activated and recorded until the quench 

front passed the second vapor probe. Each experimental run was stopped when all of the 

thermocouples on the test section were quenched. 

4.4.1.3 TRAC Model 

The TRAC model included only the test section of the Lehigh University two-phase flow 

facility and LRB Run 02/24/85-20, because the T/H conditions in the other components of the 

two-phase loop were not of interest. Also, it was possible to specify the necessary boundary 

condition time histories using experimental data. This approach permitted a more direct 

assessment of the core T/H models that are used in TRAC. As shown in the schematic in 

Fig. 4.4.4, the model consisted of 7 components, including 13 fluid cells. For reference, the 

entrance to the component 1 model in Fig. 4.4.4 corresponds to the bottom of the test section in 

the schematic presented in Fig. 4.4.3 (i.e., the interface between the copper block/hot patch and 

the test section). 

A 1-D VESSEL model (Component 1) was used to simulate the heated test section, which was 

divided into 11 fluid cells with node sizes between 0.1 and 0.11--m (3.9 to 4.3-in.) to facilitate 

detailed modeling of the grid spacer and the locations of vapor probes. The superheated vapor 

temperatures were measured at locations corresponding to the middle of the sixth and ninth cells 

in this model. A grid spacer that was 0.79 cm (2 in.) long was located in cell 8. The constant 

power-generation option was used to simulate the electrically heated test rods. The radiation 

heated square shroud was modeled with a cylindrical slab having the same perimeter as the 

shroud. The energy transfer from the radiation furnace was modeled by specifying a proper heat 

transfer coefficient and a sink temperature at the outside of the cylindrical slab. 

The axial noding of-0.1m used in the LRB Run 02/24/85-20 model is finer than the noding 

used in standard full-plant models. Core nodes in an AP600 model, for example, are -0.6 m high. 

The fine nodalization is used to simulate the phenomena occurring in this test and to accurately 

predict experimental data.

Page 4-65
To model the unheated part of the test section (hot patch), a PIPE (Component 2) was 

connected to the lower end of the vessel. This extension PIPE consisted of only one fluid cell, 

and is considered to be very well insulated against heat losses to the environment. At the inlet end 

of this PIPE component, a FILL (Component 3) was used to model the constant mass injection to 

the test section. The upper end of the test section was open to a plenum. A BREAK (Component 5) was connected to the upper end via an exit PIPE to specify the boundary conditions at this 

location. 

The boundary conditions input to the model were the test section inlet mass flux, void 

fraction, inlet temperature (FILL Component 3), and the outlet pressure (BREAK Component 5). The initial wall temperatures were also specified for the ROD and SLAB components. 

4.4.1.4 Comparison of Predicted and Measured Results 

Lehigh Rod Bundle Test 02/24/85-20 was one of a series of advancing quench front tests 

conducted in the LRB facility. Specifically, this test was a subcooled reflood experiment, with a 

slowly moving quench front. The inlet mass flux was 25.13 kg/m2s, and the inlet coolant 

temperature of 351.9 K was subcooled 21.1 K relative to the saturation temperature of 373 K at 

101.3 kPa. Calculations were performed using both TRAC-M(F77) and TRAC-M(F90). 

The prediction of the clad cooling rate throughout the core strongly depends upon the 

prediction of the correct void fraction distribution, which is primarily determined by the core 

reflood interfacial drag coefficient (IFDC) model. The IFDC model was developed from the 

inverted annular flow (IAF) regime map suggested by Ishii and his coworkers (Ref. 4.4.2), in 

which the IAF regimes along the core are determined from a capillary number defined at the 

quench front location. In addition to the criteria suggested by Ishii, TRAC also forces IAF 

regimes to occur within certain void fraction ranges (Ref. 2.1). In some IAF regimes, the core 

reflood IFDC model uses weighting functions, as well as empirical constants, to make transitions 

between them.. (The empirical constants are used because the experimental pressure drop or void 

fraction data for each IAF regime are not available in the literature to assess each IFDC model. 

The available information is usually the pressure drop over a certain length of the core that may 

experience several IAF regimes, depending upon the distance between the pressure taps.) The 

pressure drop data of Cylindrical Core Test Facility (CCTF) Run 14 (Ref. 4.4.3) was used to 

determine the empirical constants. 

The TRAC model included the single grid spacer in the Lehigh rod bundle facility, which is 

located 0.762 m above the test section inlet (the same axial position as in the test facility). 

In Figs. 4.4.5 through 4.4.7, the predicted and measured heater rod wall temperatures at 

elevations 0.15, 0.3, and 0.45 m above the core inlet are plotted against the transient time. The 

predicted and measured heater rod wall and vapor temperatures at an elevation of 60 cm are 

shown in Fig. 4.4.8. This is the elevation of the first vapor temperature probe (see Figs. 4.4.3 and 

4.4.4). The predictions of wall temperatures tend to increase until -70 s, whereas the measured 

data show an immediate decrease at a very small rate. The wall temperature is overpredicted by -80 K, while the vapor temperature at this location is underpredicted by -200 to 300 K. Note that 

the measured vapor temperatures are the maximum possible values and are obtained by drawing a 

tangent to the peak values of the recorded trace.

Page 4-66
At the axial level of 96 cm, the vapor temperature is also underpredicted as shown in 

Fig. 4.4.9. After 65 s, the predicted vapor temperature comes close to the saturation temperature 

of the fluid, well below the measured vapor temperature, which remains between 800 and 900 K. 

The predicted wall temperatures are less than the measured temperatures, indicating that the 

predicted cooling rate is higher than in the test at this elevation. Whereas the wall temperature is 

overpredicted at the 60 cm level, it is underpredicted at the 96 cm level. The change from 

undercooling at the 60 cm level to overcooling at the 96 cm level is caused by the TRAC grid

spacer model implemented at the 76.2 cm level. 

We have used the measured and predicted quench times at three levels to calculate an average 

quench front velocity over the time intervals (Tables 4.4.1 and 4.4.2). The quench-front velocity 

is not correctly predicted. The main controlling mechanism in determining the rate of quench 

front propagation is the axial distance dependent transition model. From this assessment, it seems 

that the empirical constants used in the transition boiling correlation should be reviewed. 

As a result of this assessment, it is obvious that the TRAC grid spacer model is seriously 

flawed. The grid spacer model for interfacial heat transfer is activated when the code finds that a 

grid spacer is located within a hydro cell, the hydro cell is above the transition boiling regime for 

bottom-up reflood and the vapor temperature in this hydro cell is above the saturation 

temperature. The TRAC grid spacer model for interfacial heat transfer is: 

higAi = CsHD/(vol(Tv-Tsat)) 

where 

hig = interfacial heat transfer coefficient from the vapor phase to the interface 

Ai = interfacial area 

Cs = correlation constant = 106 

HD = hydraulic diameter 

vol = volume of the hydro cell 

Tv = vapor temperature 

Tsat = saturation temperature 

When a grid spacer is modeled in the input deck, and the conditions governing its initiation 

are satisfied, the grid spacer model for interfacial heat transfer is actuated. 

A constant interfacial heat transfer rate from the vapor phase to the interface continues until 

the vapor temperature reaches approximately the saturation temperature. The TRAC grid spacer 

model for interfacial heat transfer in the TRAC reflood model is ad hoc; it was developed on the 

basis of the Lehigh test rod surface temperatures, which showed significant levels of cooling in 

the vicinity of the grid spacer. 

The absence of the grid spacer model can most easily be observed in the vapor superheat 

comparisons beyond the grid spacer location and in all the axial cladding temperature

Page 4-67
comparisons. As discussed above, the grid spacer model, when activated, reduces the vapor 

temperature to near saturation (Figs. 4.4.8 and 4.4.9). Calculations have also been made without 

the grid spacer model using TRAC-M (F77), Ref. 1.2. Without the grid spacer model, the 

predicted superheat temperature, although fluctuating, is near the measured value. With the grid 

spacer model, the vapor superheat varies between values near measured to saturation for the first 

70 s with the saturation temperatures resulting from the grid spacer model. In addition, the wall 

temperature is significantly higher when the grid spacer model is absent. The impact of the grid 

spacer model on the wall temperatures is more graphically illustrated in the series of wall axial 

temperature profiles with and without the grid spacer model, Ref. 1.2. Without the grid spacer model, the predicted wall temperatures remain much higher than measured. 

We have concluded that the TRAC grid spacer model is severely flawed; and therefore, it 

should not be used. However, the results of LRB Test 02/24/85-20 clearly show that the grid 

spacer is a highly significant component in reflood cooling, and a good grid spacer model is 

clearly needed. The impact of the grid spacer varies from facility to facility. However, the 

assessment shows that the grid spacers have a first order impact on the measured cladding 

temperature. We emphasize that the grid spacer model, used in these calculations does not model 

the key physical processes associated with a grid spacer. 

4.4.1.5 Conclusions 

Overall, the agreement between the predicted and measured parameters, with or without the 

grid spacer model, is minimal or insufficient for both codes. 

Early in the transient, the code predicted a rapid advancement of the quench front. However, 

the predicted rate of quench front advancement slowed as the transient progressed, and the 

predicted and measured axial locations of the quench front were identical by 149 s. Subsequently, 

the predicted position of the axial quench front fell behind the measured position. 

With the grid spacer model activated, the vapor temperatures were underpredicted throughout the transient. Further, the predicted vapor superheats were highly variable, with short periods of 

high superheat temperatures approaching the measured superheated vapor temperatures, interspersed with intervals during which the steam is at saturation temperatures. This behavior 

was directly related to the TRAC grid spacer model. 

The TRAC grid spacer model lacks an appropriate physical basis and, therefore, its use is not 

recommended. The Lehigh data clearly shows enhanced wall cooling caused by the bundle grid 

spacer. Therefore, it is clear that a well-founded grid spacer model should be developed for 

TRAC. This assessment also shows that the predictions produced by TRAC-M (F77) and TRAC

M (F90) are very close to each other. We therefore conclude that the conversion from TRAC-M 

(F77) to TRAC-M(F90) is successful for this type of application.

Page 4-68
4.4.2 FLECHT-SEASET Test 31504

This section presents assessment studies of TRAC-M using the data from the 161-rod 

unblocked test section at the Full-Length Emergency Core Heat Transfer/Separate-Effects and 

System-Effects Test (FLECHT-SEASET) facility. The FLECHT-SEASET facility simulates only 

1-D reflood phenomena, but it allows visual observations through view ports located at several 

elevations of the bundle. Since the hydraulics and heat transfer are closely coupled, the 

interactions of such system parameters as flooding rates, pressure, subcooling, initial cladding 

temperature, heat transfer, and mass effluent fraction have been studied. Among these 

parameters, the flooding rate was found to be the most influential. Different flooding rates give 

rise to different flow regimes which, in turn, determine the cladding temperature rise and the 

quenching of the core. A detailed assessment of the code capabilities is presented in Ref. 4.4.4. 

4.4.2.1 Facility Description 

The FLECHT-SEASET test facility was developed by modifying the existing FLECHT test 

facility. A detailed description of the facility is presented in Ref. 4.4.5. 

A cross-section of the test bundle with the heater rod instrumentation groups is shown in 

Fig. 4.4.10. The bundle comprised 161 heater rods (93 uninstrumented and 68 instrumented), 

4 instrumented thimbles, 12 steam probes, 8 solid triangular fillers, and 8 grids. The triangular 

fillers were welded to the grids to maintain the proper grid location. The fillers also reduced the 

amount of excess flow area from 9.3 to 4.7 percent. 

Locations of electrically heated rods, thimbles, and steam probe tubes are identified by alpha 

numeric coordinates indicated in Fig. 4.4.11. Each heater rod is identified by a circle. Some 

heater rods are instrumented by thermocouples welded inside its cladding, and such instrumented 

rods are identified by a group number shown in the circle. These group numbers are associated 

with the locations of thermocouples. 

The heater rods were designed to provide an axial cosine shaped power profile. As a run 

started, a computer provided decay power to the rods. Figs. 4.4.12 and 4.4.13 show the axial 

power shape and variation of decay power with time, respectively. 

Steam probe tubes and thimbles are identified by double circles in Fig. 4.4.10. Steam probe 

tubes contain steam probes designed to measure superheated steam temperatures. These steam 

probes were located in the bundle, as well as the test section outlet pipe, and the bundle steam 

probes were located in thimble tubes. The probes were designed to separate moisture from the 

high-temperature steam, and to aspirate the steam across the thermocouple into an ice bucket. 

4.4.2.2 Test Procedure 

The following is a general procedure used to establish the initial test conditions and to 

perform a typical FLECHT-SEASET unblocked bundle reflood test. The accumulator was filled 

with water, and heated to the desired coolant temperature of 53 °C (127 OF) nominal. The boiler 

was turned on and brought up to the nominal gage pressure of 0.62 MPa (75 psig). The carryover 

vessel, entrainment separator, separator drain tank, test section upper plenum, and test section 

outlet piping (located before the entrainment separator) were heated while empty to slightly above 

the saturation temperature corresponding to the test run pressure. The exhaust line between the

Page 4-69
separator and the exhaust orifice was heated to 260'C (500 'F) nominal, and the test section lower 

plenum was heated to the temperature of the coolant in the accumulator. The above component heating was accomplished using clamp-on strip heaters. A Computer Data Acquisition System (CDAS) was used to provide controls for component action and to measure test data during a test 

procedure. 

The test section, carryover vessel, and exhaust line components were pressurized to the 

desired system pressure of 0.14 to 0.41 MIPa (20 to 60 psia) by valving the boiler into the system and setting an exhaust line air-operated control valve to the desired pressure. The coolant in the 

accumulator was pressurized to 2.76 MPa (400 psia). Water was then injected into the test section 

lower plenum until it reached the beginning of the heated length of the bundle heater rods. 

Coolant was circulated and drained to ensure that the water in the lower plenum and injection line 

were at the specified temperature before the run. Power was then applied to the test bundle, and 

the rods were allowed to heat up. When the temperature in any two designated bundle 

thermocouples reached a preset value between 260'C and 871 'C (500'F to 1600 'F), the 

computer automatically initiated flooding and controlled power decay. Solenoid valves, in 

conjunction with a hydraulic control valve, controlled coolant injection into the test section. 

During an experimental run, the boiler was valved out of the system, and the pressure was 

maintained by the control valve located in the exhaust line. Liquid effluent leaving the test 

section was separated in the upper plenum and collected in a carryover tank. A baffle assembly in 

the upper plenum was used to improve liquid carryout separation and to minimize liquid entrainment in the exhaust vapor. An entrainment separator (located in the exhaust line) was used 

to separate any remaining liquid carryout from the vapor. Dry steam flow leaving the separator 

was measured by an orifice before exhausting to the atmosphere. In order to ensure single-phase 

flow measurement, the piping upstream of the orifice was heated to a temperature well above the 

saturation temperature. After all of the designated heater rods were quenched, as indicated by the 

rod thermocouples, power to the heater rods and coolant injection to the test section were 

terminated. The entire system was depressurized by opening a control valve, and the CDAS was 

deactivated. Water stored in all components was then drained and weighed. 

4.4.2.3 TRAC Model 

Figure 4.4.14 shows a schematic of the TRAC input model that was used to model the forced 

flow reflood in FLECHT-SEASET Run 31504. The input deck contains one 1-D VESSEL 

component (Component 5) to model the test section and plenums, one FILL component 

(Component 1) where the flow rate is defined, one BREAK component (Component 4) where the 

system pressure is defined, and two PIPE components (Components 2 and 3) connecting the FILL 

and BREAK components, respectively. The fuel (heater) rods are modeled using a ROD 

component (Component 6), and the test section housing wall is modeled using another ROD 

component (Component 7). 

A 1-D VESSEL component is used because the reflood phenomenon that occurs in the 

FLECHT-SEASET test section is one-dimensional. The component contains 18 cells. The lower 

plenum is modeled with cells 1 and 2, the rod bundle occupies cells 3 through 17, and the upper 

plenum is modeled with cell 18. The 161 heater rods are modeled as a single ROD component. The input deck does not model grid spacers, since the effects of grid spacers on T/H phenomena

Page 4-70
are not correctly modeled in TRAC. Coolant injection was applied to the lower plenum through 

the PIPE and FILL components, and the flow exits from the BREAK component. 

The axial power shape modeling using 18 axial cells accurately duplicated the axial power 

shape constructed in heater rods (including all steps indicating different power levels). This was 

done in order to accurately compare the temperature predictions with the measured values. This 

is much finer nodalization than used for standard reactor fuel rod modeling. Normally, fuel rods 

are modeled using three or four axial elevations. However, such coarse nodalization introduces 

an additional uncertainty in calculations that should be accounted for. 

4.4.2.4 Comparisons of Predictions with Experimental Data 

Calculations were performed using both TRAC-M(F77) and TRAC-M(F90). Both 

calculations used the newrfd=l option, which specifies the bottom-up quenching reflood model. 

FLECHT-SEASET Run 31504 is one of a series of forced reflood tests that was conducted in 

the facility. The initial test conditions and key results are presented in Table 4.4.3. 

The input deck models the initial conditions reasonably well. Figs. 4.4.15 and 4.4.16 compare 

the calculated initial heater rod cladding and steam temperatures along the test section with 

experimental data, respectively. The calculated temperatures agree very well with the 

experimental data. 

The uncertainty of experimental measurements was on the order of 100 K in a two-phase flow. 

Figs. 4.4.15 and 4.4.16 illustrate the initial experimental clad and steam temperatures, as well as 

the calculations produced by TRAC-M(F77). The experimental data shown in Fig. 4.4.15 include 

all of the thermocouples, including those that gave erroneous readings. One reading seems to be 

an outlier, and can be eliminated. The data spread seems to be larger than 100 K (on the order of 

150 K), probably as a result of natural circulation of the steam before the testing was initiated. 

Such natural circulation would occur because of heat losses from the test section wall, and 

nonuniformity of the axial power shape. This is also substantiated in Fig. 4.4.16, which depicts 

the initial axial variation of steam temperatures. Calculations were performed using a 1-D model, 

which did not model the natural circulation. Considering this limitation, the agreement between 

the temperature calculations and the measured data is "Excellent" for the initial conditions. 

This test was initiated by applying the decay power and forcing the specified inlet flow, which 

simulated the emergency core cooling system (ECCS) injection into the test section full of steam 

when the temperature in any rod at any location reached the specified level. As the inlet flow 

progressed, water slugs and droplets provided precursory cooling to the rods. As the flow 

progressed, the rods quenched. Comparisons between the calculated clad temperatures and the 

experimental data are illustrated at different elevations of the test section in Figs. 4.4.17 through 

4.4.20. Clad temperatures were calculated using both TRAC-M(F77) and TRAC-M(F90). Only 

bounding experimental temperature curves are shown. At t=O.O, the clad temperatures are at their 

initial values. 

Temperatures initially increase because the precursory cooling is not sufficient to immediately 

cool down the rods. As the time progresses, the quench front propagates, precursory cooling 

increases, and both decay power and rod power decrease. At a chosen elevation, clad 

temperatures first increase until they attain their maximum values, and then turn around as

Page 4-71
precursory cooling increases, and finally drop to the saturation value as the rod quenches. Fig. 4.4.21 illustrates the propagation of the quench front. Quench times predicted by TRAC

M(F77) during the first 200 s of the transient are slightly delayed, but they meet the acceptance criterion of a 120-s spread. (See Ref. 4.4.4.) After 200 s but before 500 s, the predictions indicate 

substantially earlier quench, and do not meet the acceptance criterion. Between 500 and 620 s, quenching occurs within the acceptance criterion. Quench times predicted by TRAC-M(F90) are 

always delayed. Except for a short time in the beginning of the transient (about the first 150 s of 

the transient), the delay does not meet the acceptance criterion (Ref. 4.4.4). 

Figs. 4.4.17 through 4.4.20 illustrate the clad temperatures predicted by both codes, as well as 

the bounding experimental data for various elevations of the test section. The bounding experimental data traces bound the spread of the experimental data, which is 100 K. Since the 

tests were run with grid spacers, and grid spacers are not modeled in the codes, the codes should 

predict hotter temperatures (up to 100 K), which is the expected bias (Ref. 4.4.4). Figs. 4.4.17 

through 4.4.20 show that the maximum clad temperatures predicted by TRAC-M(F77) meet the 

acceptance criterion of 100 K ± 100 K for all elevations (Ref. 4.4.5). Those predicted by TRAC

M (F90) meet the acceptance criterion for low elevations (below -2.00 m), but not at higher 

elevations. 

Fig. 4.4.22 illustrates experimental differential pressure data taken along the test section. 

These data were taken at 1-foot increments, starting from the 0-foot reference and ending at a 

height of 10 feet. Note that the data indicate orderly and monotonic progression of a sharp quench front (suggested by rapid increases of pressure drops at lower elevations). The sharpness of the 

quench front disappears at higher elevations; however, progression of the quench front is still 

orderly and monotonic (i.e., the height of the quench front increases with increasing time). 

Fig. 4.4.23 illustrates calculations of vapor fraction along axial distance with TRAC-M(F77). Although the calculations show predictions of quench fronts at different times, they indicate that 

progression is not orderly and monotonic. Calculations and animation of the calculated results 

with the XTV code show that the quench front and vapor fractions along the test section 

continuously oscillate, indicating that the code does not correctly predict the phenomenon. 

Figs. 4.4.24 through 4.4.29 present comparisons between the experimental data and the code 

predictions of differential pressure drops. Both codes predict large-amplitude, high-frequency 

oscillations, while the measurements show relatively smooth conditions. The predictions of 

large-amplitude, high-frequency oscillations are not considered realistic. Qualitatively, there is 

"Insufficient" agreement between the calculations and the test data. 

Figs. 4.4.30 through 4.4.32 present comparisons between the experimental data and the 

calculated steam temperatures at various elevations. There are very few steam temperature 

measurements; hence, it is not possible to estimate their uncertainties. However, the comparisons indicate that the calculated values differ by several hundred degrees from the experimental data. 

Hence, the agreement is "Insufficient." 

These comparisons indicate that both codes are not correctly predicting the reflood 

phenomena. Although peak clad temperature predictions produced using TRAC-M(F77) are 

excellent, the rest of the clad temperature predictions minimally agree with the data, and there is 

insufficient agreement between the test data and the code predictions of differential pressures and

Page 4-72
steam temperatures. One can conclude that the predictions of peak clad temperatures are 

fortuitous. 

A big surprise in the code calculations was the very large differences between the TRAC

M(F77) and TRAC-M(F90) predictions of almost any parameter. Examination of the output files 

indicates that TRAC-M(F77) was predicting choked flow conditions at the exit of the test section. 

There were some corrections in the algorithm of the choked flow model in TRAC-M(F90); 

therefore, TRAC-M(F90) was not predicting choked flow conditions. 

Fig. 4.4.33 demonstrates the effects of choked flow predictions. The choking of the flow 

produces high pressures in the test section, as predicted by TRAC-M(F77) while the test section 

pressures predicted by TRAC-M(F90) (without choked flow) remain almost constant at 40 psia, 

which was the pressure level at which the experiments were conducted. 

Choking should not have been predicted, since this phenomenon does not occur during 

reflood. It is clear that the TRAC-M(F77) predictions are completely incorrect, and the code is 

predicting different phenomena. Prediction of choking implies high two-phase velocities. It is 

expected that predictions by both codes would indicate substantially more liquid carryout than in 

testing. This would lead to higher clad temperature predictions, since less liquid would contact 

the surface of the rods and less cooling would occur. Figs. 4.4.34 and 4.4.35 show comparisons 

between the experimental data and the liquid and steam carryouts (integrated liquid and steam 

outflows) predicted by both codes. 

To evaluate the effect of choking on the predictions, the choking model was artificially turned 

off during a sensitivity study performed using TRAC-M(F77). Figs. 4.4.36 through 4.4.39 

compare the calculations produced using TRAC-M(F77) with and without the choking model, 

and those produced using TRAC-M(F90) with the experimental data. As expected, the 

predictions produced using TRAC-M(F77) without the choking model are very close to those 

produced using TRAC-M(F90). Some differences are expected as a result of the bug fixes in 

TRAC-M(F90) and the results seem to be "close enough." This comparison shows that the 

conversion to TRAC-M(F90) is successful for this type of application. 

4.4.2.5 Discussion of TRAC Models 

Excessive liquid carryout, which has been predicted in many reflood tests (Ref. 1.2), is an 

indication that the interfacial area used by the code is larger than in tests. This raises the question 

of how well the flow regime maps have been modeled in the code. 

Fig. 4.4.40 shows the flow regime map used in the TRAC code (which depicts "inverted 

annular flow") during reflood when the inlet conditions are subcooled or low-quality. Figs. 4.4.41 

and 4.4.42 show the relative velocities calculated using TRAC-M(F77) and TRAC-M(F90) at 

Cell 10 (at about the middle of the test section) and Cell 17 (almost at the end of the test section). 

Note that these relative velocities are very high. High steam velocities may tear the liquid away 

from the liquid core and also carry large slugs of liquid out of the test section without permitting 

the liquid to contact heated walls for a long time. Consequently, these high relative velocities 

explain why the liquid carryout from the test section is higher than the measured data and why the 

calculated clad temperatures, in general, are higher than the experimental values.

Page 4-73
The TRAC flow regime maps are predicated on the work performed by Ishii and his 

coworkers (Ref. 4.4.2), which indicates that these flow regime maps exist at relatively low 

relative velocities. For example, Ref. 4.4.2 states that if the relative velocity is increased to 

2.0 m/s, smooth and rough wavy sections are nonexistent; however, the relative velocities shown 

in Figs. 4.4.41 and 4.4.42 are much higher than 2.0 m/s (up to 20.0 m/s). This would lead to the 

conclusion that, at these predicted velocities, the flow regime maps used by TRAC-M(F77) or 

TRAC-M(F90) cannot be consistent with the flow regimes that actually occur in tests. 

According to Ref. 4.4.2, the inverted annular flow regime has four different regions with 

distinct characteristics, 1) smooth, 2) rough-wavy, 3) agitated flow, and 4) dispersed flow. 

Fig. 4.4.40 depicts these regions, as well as the transition flow region (i.e., the transition boiling region between the CHF point and the T(min)). The tests in Ref. 4.4.2 were performed under 

conditions that did not simulate the transition flow region. 

Ref. 4.4.2 defines the boundaries of four different regions on the basis of the Capillary number 

with restrictions on their applicability. The regions of applicability are predicated on the Weber 

number and vapor fraction, and the Weber number is founded on the relative velocity. Examination of the documentation on the TRAC code (Ref. 2.1 and Ref. 4.4.6) indicates that 

although the same Ishii correlations based on the Capillary number have been used, the regions of 

applicability are not the same, because they are based on the vapor fraction alone and not on the 

Weber number. 

Table 4.4.4 shows the Capillary number criteria and the Weber number restrictions in 

Ref 4.4.2 and TRAC code application in Ref. 4.4.6 and Ref 2.1, where in Ref, 4.4.2, Z is the 

average break-up length measured from the inlet of the test section, D is the liquid jet hole 

diameter, Ca is the Capillary number defined using fluid jet properties (viscosity, velocity, and 

surface tension) at the inlet (Ca = g, V11/a), We is the Weber number defined using fluid jet and 

gas properties at the inlet (We = p( Vg - V) 2D/la), and ai is the void fraction at the inlet of the 

test section. In Ref. 4.4.2, for a constant inlet jet flow and a constant inlet gas flow, which is 

injected from a nozzle at the periphery of the test section, all breakout points (boundaries of flow 

regimes) are determined by the relationships in the second and third columns. Experiments were 

run at the steady state. Flow regimes vary axially but the boundaries remain constant with time. 

The void fraction at the inlet of the section is determined by the test section geometry alone, and 

does not change with changing jet liquid velocity or gas injection velocity. Hence, for a given test 

section, the breakout points are determined only by the liquid velocities (i.e., Capillary numbers, 

provided that the gas velocities are within the ranges determined by the correlations with Weber 

numbers). Note that both liquid and gas velocities are quantities measured at the inlet, not local 

quantities. It is surprising that the breakout points do not vary with the amount of power or power profile applied to the test section. This can be true only if the amount of power and power profile did not change during the tests. Hence, power or power profile did not enter into the correlations. 

However, heating and power profile can be a major factor in reactor applications. 

The TRAC modeling is different, in that the Capillary number is calculated using fluid 

properties at the CHF point, and the criteria for the applicability of Ishii correlations (i.e., breakout points) are founded on local void fractions. Details concerning the calculations of the 

local void fractions, presented in Ref, 2.1, do not include any restriction on the basis of Weber 

numbers calculated using relative velocities.

Page 4-74
It is clear that the flow regime maps modeled by Ishii in Ref. 4.4.2, have not been correctly 

implemented in TRAC. On the other hand, Ishii correlations may have an important shortcoming 

in that breakout points are independent of power input to the test section, and relative velocities 

are predicated on quantities measured at the inlet. It is clear that new experiments should be 

performed, and a new reflood model should be developed. Determination of the type of reflood 

modeling in the TRAC code and the type of experiments to be performed for accurate reflood 

modeling is outside of the scope of the present assessment. 

4.4.2.6 Conclusions 

The top-level requirement in the development of a T/H code is that the code should correctly 

model the physics of the phenomena. This assessment shows that TRAC does not correctly 

model the flow regime maps during the reflood and, consequently, it cannot correctly model the 

interfacial areas between phases, and cannot correctly predict vapor fractions, phasic velocities, 

and temperatures. Almost none of the important parameters (such as clad temperatures, steam 

temperatures, differential pressures, and liquid and steam carryouts) are consistently predicted 

with acceptable accuracy throughout the range. The predictions in ranges where clad 

temperatures are predicted within accuracy limits, are accidental. Also, the predicted clad 

temperatures are higher than the experimental data because predictions of liquid carryout by the 

codes during reflood have consistently been high (Ref. 1.2). Hence, these predictions are 

conservative. 

Comparison between TRAC-M(F77) and TRAC-M(F90) code calculations is favorable. The 

sensitivity study performed with TRAC-M(F77) without the choking model indicates that the 

calculations agree reasonably well with the calculations performed with TRAC-M(F90). The 

observed differences are probably attributable to many error fixes in TRAC-M(F90). We believe 

that the conversion from TRAC-M(F77) to TRAC-M(F90) has been successful for this 

application.

Page 4-75
4.4.3 Cylindrical Core Test Facility (CCTF), Core I Test C1-5 (Run 14) 

The Cylindrical Core Test Facility (CCTF) was designed, built, and operated by the Japan Atomic Energy Research Institute (JAERI) to provide information on thermal-hydraulic behavior 

during the refill and reflood phases of a LOCA in a PWR. The CCTF was used to run two test 

series, which were designated as CCTF Core I (CCTF-I) and CCTF Core II (CCTF-II) tests. The 

blowdown period was not simulated for either series, and the facility configuration was slightly 

modified for the second series. 

The CCTF tests were conducted with the following objectives: 

"* Demonstrate ECC behavior during refill and reflood phases. 

"* Verify reflood analysis codes. 

"• Collect information to improve the T/H models in such analysis codes as (a) multidimensional 

core thermal-hydrodynamics including the radial power distribution effect, failback effect, 

and spatial oscillatory behavior; (b) flow behavior in the upper plenum and hot legs; (c) 

behavior of accumulated water at the bottom of the upper plenum, including possible 

countercurrent flow and sputtering effects; (d) hydrodynamic behavior of the injected ECC 

water and water passing through the steam generator; (e) multidimensional T/H behavior in 

the hot annular downcomer; and (f) overall oscillatory behavior in the system. 

4.4.3.1 Facility Description 

The CCTF is a full-height, experimental facility designed to model a four-loop, 11 00-MWe 

PWR. Figs. 4.4.43 and 4.4.44 show a schematic of the CCTF facility and a top view of the 

primary system piping, respectively. Fig. 4.4.45 shows a top view of the pressure vessel, and the 

configuration of bundle assemblies. Additional descriptive information regarding the facility can 

be found in Refs. 4.4.7 and 4.4.8. All pertinent PWR components have been modeled. Vessel and 

primary-loop components are full height and their dimensions are volume-scaled in a ratio of 1/ 

21.4, as compared with a 11 00-MWe PWR. The pressure vessel downcomer annulus is larger than the width that results from an exact 1/21.4 scaling because the scaled core bypass area is 

included as part of the downcomer area. The upper plenum internals are reduced in radial 

dimension by the ratio of 8/15. Depending on facility configuration and operation, tests in the 

CCTF have been variously defined as either separate effects tests (SETs) or integral effect tests 

(IETs) (Refs. 4.4.9 and 4.4.10, respectively). 

The CCTF vessel contains a core, an annular downcomer, and upper and lower plena. The 

CCTF core contains thirty-two 8 x 8 bundles, each containing 57 electrically heated rods and 7 

nonheated rods (a total of 1824 heated rods and 224 unheated rods). Bundles are assembled to 

form different powered zones, as shown in Fig. 4.4.45. In CCTF-I, each bundle includes rods 

with three different power densities, resulting in local peaking factors. The heated rods are held in 

place by grids, as in a PWR. Each heated rod consists of a nichrome heating element, magnesium oxide and boron nitride insulators, and an Inconel sheath. The heated length and outer diameter 

of the heater rods are 3.66 m and 10.7 mm, respectively. The average axial power profile is 

simulated by a 17-step chopped cosine profile, with an axial peaking factor of 1.492. The 

unheated rods simulate the guide and instrument thimble tubes in PWR fuel assemblies.

Page 4-76
The primary loops consist of three intact loops and a broken loop. Each loop consists of hot 

leg and cold leg piping, a steam generator simulator, and a pump simulator. The 200% cold leg 

break is simulated in the broken loop. The two ends of the broken loop are connected to two 

containment tanks through blowdown valves. The steam generator simulators are of the U-tube 

and shell type. Each steam generator simulates two loops that are housed in a single shell 

assembly having two compartments, one simulator for each loop in a single compartment. The 

primary coolant passes through the tube side, and the secondary coolant is stagnant in the shell 

side. The steam generator tubes are 25% shorter than those in a commercial reactor. 

The ECCS consists of two accumulators and a low-pressure coolant injection (LPCI) system. 

The instrumentation consists of detectors, signal conditioners, signal processors, recording 

equipment, and display equipment. Temperature measurements are made with alumel chromel 

thermocouples having an expected error of 0.5%. Pressure and differential pressure 

measurements are made with differential pressure cells, with an expected error of 0.3%. In the 

pressure vessel, liquid level measurements are made in two ways. Differential pressure 

measurements are used to determine the collapsed liquid level. The liquid level is also directly 

measured using conductivity or optical detectors. ECC water mass flow rate measurements are 

made with electromagnetic flow meters, with an expected error of 0.5%. Mass flow rate 

measurements in the outlet nozzles of the steam generators are made with pitot tubes. Mass flow 

rate measurements from the downcomer into the core are made with four drag disks. The steam 

mass flow rate from containment tank II is measured with a venturi tube. Two-phase flow 

measurements in the primary loops are made at spool pieces. A spool piece is installed in each 

hot and cold leg. Each spool piece consists of a three-beam gamma densitometer, a turbine 

flowmeter, three drag transducers attached to a drag screen, fluid and metal temperature 

thermocouples, an absolute pressure transducer, and a differential pressure transducer across the 

drag screen. The fluid densities, velocities, mass flow rates, and void fraction can be obtained 

using data processing software. Veloeimeters were installed in the lower plenum to measure the 

core flooding rate. 

4.4.3.2 Test Procedure 

CCTF Test C1-5 (Run 14), hereafter referred to as CCTF Run 14, was the CCTF Core-I 

facility base case (Ref. 4.4.7). CCTF Run 14 does not correspond to either a best-estimate (BE) 

or evaluation model (EM) condition (Ref. 4.4.11). The power level for this test is 7% greater than 

the value used in the current EM licensing analyses, and 35% higher than a realistic BE case. 

ECC flows are 25% lower than the scaled ECC flow that is expected in a PWR. Pump simulator 

orifices are sized to provide an equivalent locked rotor pump resistance. The 0.2-MPa 

containment pressure chosen for this test is lower than that used in most licensing calculations. 

In preparation for the test, the accumulator tank, LPCI tank, saturated water tank, and 

secondary sides of the steam generator simulators were filled with purified water. The 

instruments were checked for their zero points and sensitivities. 

After these preparatory operations, the primary system was heated with the preheaters to its 

specified temperatures (downcomer wall, core internals, and primary-loop piping wall), and was 

pressurized to the specified pressure using steam. The water in the accumulator tank was 

electrically heated to its specified temperature, and pressurized with nitrogen gas to provide 

sufficient head to drive the required injection flow. The water in the LPCI tank was also heated to

Page 4-77
its specified temperature, and was circulated through the delivery lines to preheat the lines to the 

same temperature as the water. The water in the secondary side of each steam generator simulator 

was also heated and pressurized to the specified temperature and pressure. The water in the 

saturated water tank was heated to approximately the saturation temperature of the expected 

primary pressure. 

After establishing the initial test conditions, the electric power for preheating was turned off 

and the lower plenum was filled with water to the specified level (0.9 m) from the saturated water tank. When the initial test conditions were stabilized at allowable tolerances, electric power was 

applied to the heater rods in the core and the data recording was started. The temperature increases in the rods were monitored by a computer. When a specified initial cladding temperature was reached, direct injection of the accumulator water into the lower plenum was 

initiated. The system pressure was maintained at the specified initial pressure throughout the test 

by controlling the outlet valve of containment tank II. The decay power input to the rods was programmed to begin when the water reached the bottom of the heated region of the core. The 

specified initial cladding temperature of the heater rods was predetermined by interpolation between the cladding temperature after preheating and the cladding temperature assumed for the 

time of bottom of core recovery (BOCREC). When the water reached the specified level below 

the core, the injection port for the accujmulator water was switched from the lower plenum to the three intact cold-leg ECC ports. The accumulator (ACC) injection flow rate was then reduced 

during the cold-leg injection. At a specified time after BOCREC, the valves in the accumulator 

lines and the LPCI circulation lines were closed, and the valves in the LPCI injection lines were 

opened. These actions transferred the ECC injection from the ACC mode (high flow rate) to the 

LPCI mode (lower flow rate). The steam and entrained water flowed via the broken and intact 

loops to the containment tanks. The steam was then vented to the atmosphere to maintain a 

constant pressure in the containment tanks. The test was terminated after all of the thermocouples 

on the surface of the heater rods indicated quenching of the rods. 

CCTF Run 14 proceeded as described in the previous paragraph. Direct injection of ACC 

water into the previously filled lower plenum began after power was applied to the core, and after 

the cladding temperature at a specified location reached 5060C. When the lower plenum filled 

with water and the level reached the bottom of the heated core, the core power followed a 

programmed transient to simulate decay power. The injection port was changed from the lower 

plenum to the three intact cold-leg ECC ports 14 s after the start of ACC injection. As specified, 22 s after the initiation of ACC injection, ECC injection was transferred from the ACC injection mode (high flow rate) to the LPCI mode (lower flow rate). Operating parameters for CCTF 

Run 14 are presented in Table 4.4.5. 

4.4.3.3 TRAC Model 

Several TRAC input models have been developed for simulation of CCTF Run 14. These include (1) a finely noded, 3-D vessel model with 16 axial levels, 4 radial sectors, and 4 azimuthal 

sectors; (2) a coarsely noded, 3-D vessel model with 16 axial levels, 2 radial sectors, and 2 

azimuthal sectors; and (3) a finely noded 1 -D vessel model with 24 axial levels, 1 radial sector, and 1 azimuthal sector. Of these three versions of the CCTF Run 14 input model, only the latter 

has been updated for TRAC-M(F77). This model has also been used for assessment of an earlier

Page 4-78
code version by JAERI (Ref. 4.4.12). The noding schematic for the input model is presented in 

Fig. 4.4.46. The power profile is illustrated in Fig. 4.4.47. 

The CCTF Run 14 1-D TRAC input model used in this assessment directly models only the 

core and upper plenum of the CCTF-I facility. The boundaries of the model have been reduced by 

artificially limiting the model to a limited region of the larger test facility (i.e., not modeling the 

downcomer and cold legs), thereby permitting a more direct assessment of the core reflood T/H 

modeling. The added degrees of freedom associated with the downcomer and cold legs have also 

been eliminated. However, this approach requires application of appropriate boundary conditions 

if the simplified model is to be acceptable. Toward this end, the core inlet boundary conditions 

are specified by a FILL component with time-dependent, core inlet mass flow and fluid 

temperature. The core outlet boundary conditions are specified by a single BREAK component, 

with the time-dependent pressure specified. The total power supplied to the heater rods is also 

provided. 

The CCTF core and upper plenum are modeled with 24 axial levels, and the model included 

grids. The fine axial noding used in the CCTF Run 14 model (-0.2-m high axial levels) is coarser 

than the noding used in standard full-plant models. (For example, core nodes in an AP600 model 

are -0).6-m high.) 

The inlet FILL (Component 1) is connected via the PIPE (Component 2) to axial cell 1. The 

BREAK (Component 5) is connected to axial cell 23 via the PIPE (Component 4). The core 

extends axially from level 2 through 19. 

4.4.3.4 Comparison of TRAC-M(F77) and TRAC-M(F90) Calculations 

Calculations were performed using both TRAC-M(F77) and TRAC-M(F90). Figs. 4.4.48 

through 4.4.52 compare the clad temperatures calculated at different elevations in the core. In 

these figures, experimental data are not shown, since the data were obtained under the 2-D/3-D 

program, and are not publicly available. When an agreement with other partners is reached to 

release these data, this report will be revised to include the experimental data and to illustrate the 

comparisons between the data and the code predictions.* 

Figs. 4.4.48 through 4.4.52 show comparisons of code predictions, which demonstrate 

"Excellent" to "Reasonable" agreement. There are some differences between the predictions in 

some areas; however, these differences are attributable to error corrections and bug fixes in the 

TRAC-M(F90) code, and we can conclude that the conversion from TRAC-M(F77) has been 

successful for this application. 

4.4.3.5 Conclusions 

The conversion from TRAC-M(F77) has been successful for this type of application. 

* 2-D/3-D partners agreed to the release of data on CCTF-14, CCTF-54, SCTF-719, UPFF 6, and UPTF 8b 

in NRC developmental assessment reports. These data will be included in future reports.

Page 4-79
4.4.4 Cylindrical Core Test Facility (CCTF), Core-Il Run 54 

The CCTF is an experimental test facility designed to model a full-height core section and 

four primary loops with components of a PWR. The facility was completed on March 10, 1979, 

and 22 tests were performed with CCTF Core-I by April 1981. For the CCTF Core-II test series, 

the first core was replaced by the new second core on November 1981, and six tests, including 

two acceptance tests and two shakedown tests, were performed with CCTF Core-I by April 1982. 

Although there were some differences between the core designs in the CCTF Core-I and CCTF 

Core-II programs, the special purposes of the CCTF Core-II program were to investigate the 

effects of an alternative ECCS (such as the combined and downcomer injections), and to extend 

the experimental range of the CCTF Core-I test series. 

Core-II Run 54 was successfully conducted on March 30, 1982. The objectives of the test 

were to check the functions of the modified CCTF Core-II facility, to confirm the similarities 

between Core-I and Core-H, and to study the effect of the lower power supplied into the core. 

The test was conducted under the same initial and boundary conditions as in the base case of the 

CCTF Core-II test series (Run 62), which simulated EM conditions, except for the lower power. 

The ECC flow conditions were the same as those of the EM tests in CCTF Core-I. 

4.4.4.1 Facility Description 

Fig. 4.4.53 shows an isometric view of the facility, which was designed and operated by JAERI to provide information regarding fluid behavior in the core, downcomer, and upper 

plenum, including steam and water carryover (steam binding) and integral system (steam 

generator and pump simulator) effects during the refill and reflood phases of a hypothetical 

LOCA in a PWR. The central part of the test facility is a nonnuclear core that consists of 1824 

electrically heated rods and 224 nonheated rods arranged in a cylindrical array. The core is 

housed in a test vessel that includes a downcomer, lower plenum, upper plenum, and core region. The core design is based on 8 x 8 rod assemblies that model the typical 15 x 15 fuel assembly of a 

PWR. Volumetric scaling is based on core flow area scaling. 

The scaled dimensions of the components are given in Table 4.4.6, and elevations of the 

components are given in Table 4.4.7. 

4.4.4.1.1 Pressure Vessel and Internals 

The pressure vessel is a cylindrical type, as shown in Fig. 4.4.54. The height is the same as 

the reference reactor pressure vessel. The radial direction is scaled down in proportion to the core 

flow area scaling (that is, 1/21.44). The upper ring was newly installed in the Core-Il facility for 

the upper plenum ECC water injection nozzles (called core flooding nozzles). Core flooding nozzles were not used in the Run 54 test. The vent valves and core flooding nozzles are intended 

to simulate a Babcock & Wilcox (B&W) type PWR. This facility includes the following features: 

The cross-section of the pressure vessel is shown in Fig. 4.4.55. The core consists of thirty

two 8 x 8 electrically heated rod bundles, arranged in a cylindrical configuration, to simulate 

Westinghouse 15 x 15 fuel assemblies.

Page 4-80
" The downcomer is an annulus of a 61.5-mm gap. In determining the gap size, the flow area of 

the core baffle flow area is included in the downcomer region. Thus, the core baffle flow area 

is included in the downcomer simulation, and is not simulated separately in this vessel. 

" The vessel wall is constructed of carbon steel that is clad with a stainless steel lining. The 

wall is 90 mm thick to simulate the stored energy as closely as is reasonably possible during 

ECC water injection. 

" The design of upper plenum internals is based on that of the Westinghouse 17 x 17 fuel 

assemblies. The internals consist of 10 control rod guide tubes, 10 support columns, and 12 

open holes. Flow resistance baffles are inserted into the guide tubes. The radius of each 

internal is scaled down by a factor of 8/15 from that of an actual reactor. 

" The end box and the upper-core support plate (UCSP) are installed between the core and 

upper plenum. The end box tie plate is 10 mm thick, and perforated with round flow holes. 

Plugging devices are installed in the Core-fl facility to simulate the flow resistance. The 

UCSP is a 60-mm thick perforated plate. The geometry of the perforation is analogous to that 

of an actual reactor. 

4.4.4.1.2 Heater Rod Assembly 

The heater rod assembly that simulates the fuel assembly consists of thirty-two 8 x 8 array rod 

bundles. Each bundle consists of 57 electrically heated rods and 7 unheated rods. The core is 

usually subdivided into three regions to achieve the desired radial power distribution. The local 

peaking factor of heated rods in a bundle is unity; that is, all heated rods in a bundle have the same 

power density in the Core-II facility. 

A heater rod consists of a nichrome heating element, magnesium oxide (MgO) and boron 

nitrite (BN) insulators, and an Inconel-600 sheath. BN is used only for the central part of the 

heated region, and MgO is used for the other part. The heated length and outer diameter of the 

heater rods (3.66 m and 10.7 mm, respectively) are identical to the corresponding dimensions of 

actual PWR fuel rods. The sheath-wall thickness is 1.0 mm, which is thicker than the actual fuel 

cladding because of the requirements for thermocouple installation. The heating element is a 

helical coil with a varying pitch to generate a 17-step, chopped cosine axial power profile, as 

shown in Fig. 4.4.56. Nonheated rods are either stainless steel pipes or solid bars of 13.8-mm 

outside diameter. All of the pipes are used to install such instruments as steam superheat probes 

and thermocouples, and to carry the assembly loads. 

The heater rods and nonheated rods are held in a radial position by grid spacers that are 

located at six elevations along the axial length, as shown in Fig. 4.4.56. A grid spacer is a lattice 

structure composed of stainless steel plates. The top and bottom edges of the stainless steel plates 

are sharpened in CCTF Core-IL. The rod pitch is 14.3 mm, which is same as that of the reference 

PWR. 

The heater rods penetrate through the bottom plate to facilitate lead out of the power cables 

from the bottom of the vessel. The outer diameter of the rods in the lower plenum is reduced to 

8.6 mm. Three-phase electric current is used to heat the heater rods, and the electrical neutral 

point is at the top of the rods (where they are interconnected to each other).

Page 4-81
4.4.4.1.3 Primary Loops and the Emergency Core Cooling System (ECCS) 

The primary loop consists of three intact loops and a broken loop. Each loop consists of hot

and cold-leg piping, a steam-generator simulator, and a pump simulator. The cold-leg break is 

simulated for the broken loop. The broken cold leg is connected to two containment tanks 

through blowdown valves. The primary loop arrangement is shown in Fig. 4.4.57. 

The inner diameter of the piping is scaled down in proportion to the core flow area scaling. The length of each piping section is almost the same as the corresponding sections of the 

reference PWR. 

The steam generator simulators are the U-tube and shell type. The tube length is -5 m shorter 

than in the reference PWR, and the vertical height of the steam generator simulators is -5 m lower 

than in the reference PWR. The primary coolant passes through the tube side, and the secondary coolant is stagnant in the shell side. The steam generator simulators of two loops are housed in a 

single shell assembly that has two sets of separated inlet and outlet headers for the two loops. The 

wall thickness of the U-tube is 2.9 mm (instead of 1.27 mm for the reference PWR) because of a 

higher pressure difference between the primary and secondary sides in the steam-generator simulator. The pump simulator consists of the casing and vane simulators and an orifice plate. 

The resistance of each looop is simulated with the orifice plate, and has a hole with a diameter 

of 95 mm and a thickness of 10 mm. 

The ECCS consists of an ACC and an LPCI. The injection points are at each cold leg and at 

the lower plenum. The upper plenum and downcomer injection system are available for further 

alternative ECCS tests. 

4.4.4.1.4 Instrumentation 

The instrumentation is divided into two groups. One group consists of JAERI-supplied instruments that measure the temperatures, absolute pressure, differential pressures, water levels, and flow rates. Thermocouples measure the temperatures of the rod surface, fluid, and structure. 

The absolute pressures are measured in the upper and lower plena, steam generator plena, and 

containment tanks. The differential pressure measurements are carried out at many locations that 

almost completely cover the entire system. The liquid levels are measured in the ECC water 

supply tanks and containment tank 1. The flow meters measure the ECC water flow rates, and 

flow rates in the downcomer and loop seal piping vent line from containment tank 2 to the 

atmosphere are measured with the drag disk flow meter, Pitot tubes, and venturi tube, respectively. The total number of JAERI-supplied instruments are 1316 channels, and the signals 

from these instruments are recorded on a magnetic tape. 

The other instrumentation group consists of NRC-supplied advanced instruments totaling 536 

channels for the two-phase flow measurement. 

4.4.4.2 Test Procedure 

In preparation for the test, the ACC tank, LPCI tank, saturated water tank, and secondary sides 

of the steam generator simulators were filled with water purified with ion exchange resin. The 

instruments were checked for their zero points and sensitivity after all of the components and 

instruments were inspected for mechanical and electrical leakage.

Page 4-82
After the preparatory operations, the primary system was heated to its specified temperatures 

(downcomer wall, 468 K; core internals, 423 K; and primary loop piping wall, 406 K) with the 

preheaters, and pressurized to 0.2 MPa by substituting steam for nitrogen gas in the system. The 

water in the ACC tank was electrically heated to 310 K, and pressurized with nitrogen gas to 

provide sufficient head to drive the required injection flow. The water in the LPCI tank was also 

heated to 310 K, and was circulated through the circulation line (including the LPCI line) to 

preheat the line to the same temperature as the water. The water in the saturated water tank was 

heated close to the saturation temperature (394 K) of the expected primary system pressure 

(0.2 MPa). The water in the secondary side of each steam-generator simulator was heated to 

539 K, and and pressurized to 5.3 Mpa. 

After the initial conditions of the test were established, the electric power for preheating was 

turned off, and the lower plenum was filled to 0.86 m directly from the saturated water tank. 

Electric power was applied to the heater rods in the core, and the data recording was started when 

the water level in the lower plenum reached the specified level and other initial conditions of the 

test stabilized at the allowable tolerance. A computer was used to monitor the temperature 

increases of the rods. When a specified initial clad temperature (1003 K) was reached, direct 

injection (0.104 m3s-1

) of the ACC water into the lower plenum was initiated. The system 

pressure was maintained at 0.2 MPa throughout the test by controlling the outlet valve of 

containment tank 2. The decay of power input to the rods was programmed to begin when the 

water reached the bottom of the heated region of the core. The specified initial clad temperature 

(995 K) of the heater rods for initiation of coolant injection was predetermined by interpolation 

between the clad temperature (394 K) after preheating and the clad temperature (1073 K) 

assumed for the time of core bottom recovery. The specified power decay was obtained by 

normalizing the 1.0 X ANS standard + U2 38 - capture decay curve at 30 s after shutdown. 

When the assumed water level reached the specified level (0.5 m from the bottom of the 

heated region of the core), the injection port was changed from the lower plenum to the three 

intact cold-leg ECC ports. This water level was assumed to be the level at which considerable 

steam generation occurs in the core to minimize oscillatory behavior caused by the condensation 

at the ECC ports. The ACC injection flow rate was then reduced to 0.088 m3s"1 in the cold-leg 

injection period. At a specified time (16.5 s) after the time of core bottom recovery, the valves in 

the ACC lines and LPCI circulation lines were closed, and the valves in LPCI injection lines were 

opened. These actions transferred the ECC injection from ACC injection mode to LPCI mode. A 

specified LPCI flow rate (0.0116 m3s"') was maintained constantly until the ECC injection was 

turned off. 

The generated steam and the entrained water flowed via broken and intact loops to the 

containment tanks. The steam was then vented to the atmosphere to maintain a constant pressure 

in the containment tanks. After all thermocouples on the surface of the heater rods indicated 

quenching of the rods, the power supply to the heater rods decreased linearly. The linear power 

decay was performed to study any particular reflood phenomena under very low power supply. 

The linear power decay was initiated at 690.5 s, and the power was turned off at 898 s. After the 

ECC injection was turned off, the recording system was stopped, thus terminating the test.

Page 4-83
4.4.4.3 TRAC Input Model

The input model was a straightforward conversion of an earlier TRAC model that was used 

for the analysis of CCTF Run 54, which was conducted as part of the 2-D/3-D program. This 

input model used the intermediate-node model of the CCTF reactor vessel, as shown in 

Fig. 4.4.58. The vessel nodalization includes 2 azimuthal zones, 4 radial rings, and 16 axial 

levels. 

The three CCTF intact loops are combined and modeled as a single loop. The intact loop components are shown in Fig. 4.4.59, and Fig. 4.4.60 shows the noding for the broken loop 

components. 

Initial thermal and hydraulic parameter values are generally consistent with the initial values 

listed in Table 4.4.8. In most instances, initial pressures and temperatures are at the nominal 

values given in the table. The axial and radial distribution of the power in the core agree with the 

distribution given in the Run 54 data report. Table 4.4.9 presents the chronology of testing events. 

The seven time-dependent boundary conditions used for these calculations included the total 

core power, lower plenum injection velocity and ECC temperature, cold leg injection velocity and 

ECC temperature, and both containment tank pressures. The inputs were tabulated from 

instrument records. 

4.4.4.4 Comparison of TRAC-M(F77) and TRAC-M(F90) Predictions 

Calculations are performed using both TRAC-M(F77) and TRAC-M(F90), and the results are 

compared in Figs. 4.4.61 through 4.4.65. Since CCTF data are restricted to 2-D/3-D program 

participants, this report does not contain any graphics of CCTF test data.* These figures compare code to code results. The figures show clad temperature predictions of the hot rod at various 

elevations. Agreements between code calculations are considered "Excellent" to "Reasonable" 

for this test. Predictions of peak clad temperatures are almost the same. There are some 

differences in quenching times. Considering all error corrections and changes that have been 

implemented in TRAC-M(F90), the agreement between predictions is deemed to be 

"Reasonable." 

4.4.4.5 Conclusions 

The conversion from TRAC-M(F77) to TRAC-M(F90) has been successful for this type of 

application. 

* See footnote on page 4-79.

Page 4-84
4.4.5 Slab Core Test Facility (SCTF), Run 719

The Slab Core Test Facility (SCTF) is designed to investigate the 2-D thermal-hydraulic 

phenomena that occur in a PWR during the end of the blowdown, refill, and reflood phases of a 

postulated LBLOCA. The SCTF is a full-height facility; most major vessel and primary loop 

dimensions are volume scaled in a ratio near 1/21 relative to a four-loop PWR with lower plenum 

and cold-leg ECC injection. A description of the SCTF Core-III facility is provided in 

Section 4.4.5.1. Fig. 4.4.66 shows an overview of the SCTF Core-Ir system. A major feature of 

the SCTF is that it contains a full-height and full-radius heated core with a width of one tube 

bundle. The current configuration of the core is designated as Core-HI, with the upper plenum 

and upper head representative of a German PWR. The facility is operated by JAERI. 

4.4.5.1 Facility Description 

Fig. 4.4.67 shows a vertical cross-sectional view of the SCTF pressure vessel. The vessel 

contains a downcomer, lower plenum, electrically heated core simulator, core baffle region, upper 

plenum, and upper head. The core simulator models a sector of the core from the centerline to the 

outer wall of the vessel. The name of the facility, "slab core," is descriptive of the core simulator 

arrangement. A slab-like sector of the vessel from the core centerline to the vessel outer wall 

contains eight heater rod bundles arranged in a row, the core baffle region, and the downcomer. 

Whereas the design of the Core-II upper plenum was based on that of a Westinghouse plant with 

17 x 17 array fuel assemblies, the design of the Core-HI upper plenum is based on that of a 

German PWR. 

Each heater rod bundle contains 236 electrically heated rods and 20 unheated rods. A heater 

rod consists of a Nichrome heater element wound in a helix about a ceramic core rod (magnesium 

hydroxide), and a Nichrofer 7216 (equivalent to Inconel 600) sheath, with boron nitride (central 

region) or magnesium hydroxide (upper and lower regions) packed between the heater element 

and the sheath. The pitch of the heater element winding is varied to produce a 17-step, chopped

cosine axial power profile, as shown in Fig. 4.4.68. 

Simplified models of the primary coolant loops are provided in the facility. The facility is 

equipped with a hot leg that is equivalent to four hot legs in a PWR. The hot leg cross-sectional 

shape is not typical of a PWR, in that it is tall and narrow rather than circular in shape. The hot 

leg connects the vessel upper plenum to a steam/water (S/W) separator. The S/W separator 

simulates single-phase steam flow downstream of a steam generator in a PWR, and also facilitates 

the measurement of liquid carryover from the upper plenum, through the hot-leg simulator, and 

into the S/W separator. The S/W separator inlet plenum models a steam generator inlet plenum to 

simulate the flow characteristics of carryover water into the U-tubes. The intact cold leg in the 

SCTF is equivalent to three intact loops in a PWR. A pump simulator and a loop seal are 

provided in the intact cold leg simulator. Two broken loop components are provided, and each of 

these is connected to a separate containment tank through a break valve. The loop resistance is 

adjusted with orifice plates that can be placed in the intact cold leg, the S/W separator side and 

pressure vessel side broken cold legs, and the pump simulator in the intact loop. 

The coolant flows delivered by the ECCS are also modeled. These simulate flows from the 

ACC, an LPCI system, and a combined injection system. ECC water injection ports are at the 

cold leg, the hot leg, through the upper head into the top of the upper plenum, into the upper

Page 4-85
plenum at the UCSP, into the downcomer, and into the lower plenum. The specific injection ports 

are selected according to the objectives of the given test. 

4.4.5.2 Test Procedure 

SCTF Run 719 (Test S3-15) was an integral, lower plenum injection test that simulated the 

condition of a US/Japan-type PWR during the reflood phase of a 200% cold leg LBLOCA. The 

test objective of Run 719 was to investigate the effect of radial power distribution shape and ratio 

on 2-D heat transfer behavior. The radial power distribution was inclined, and the peak-to

average power ratios were 1.36, 1.20, 1.20, 1.00, 0.91, 0.86, 0.81, and 0.76 for bundles 1 through 8, respectively. Before test initiation, the facility was filled with saturated steam at 0.2 MPa. The lower plenum was filled with saturated water (393 K) to a height of 1.2 m. The core then began to 

heat in near adiabatic conditions. The core power was tripped and ECC initiated when a preselected number of rod cladding temperatures exceeded 1134 K. The maximum cladding temperature reached was 1183 K; this temperature was measured at the midplane of bundle 1 (the bundle receiving the highest power). The entire core was quenched at 574.5 s. The core T/H behavior was strongly 2-D in Run 719. A single recirculating flow was established with upflow through the high-power bundles and downflow through the low-power bundles. A summary of 

the test conditions for this test is presented in Table 4.4.10. 

4.4.5.3 TRAC Input Model 

The pressure vessel noding is shown in Fig. 4.4.69. The heated core section is noded into six 

axial levels. Fig. 4.4.70 shows the heater rod noding, along with the thermocouple locations and the modeled axial power shape. The hot leg and the S/W separator are shown in Fig. 4.4.71; the 

intact cold leg noding is shown in Fig. 4.4.72; the broken cold leg noding from the pressure vessel 

to containment tank 1 is shown in Fig. 4.4.73; and the broken cold leg noding from the S/W 

separator to containment tank 2 is shown in Fig. 4.4.74. The primary features of each element in 

the TRAC input model are discussed in the following sections. 

4.4.5.3.1 Pressure Vessel 

The pressure vessel is modeled with the TRAC VESSEL component in 2-D Cartesian 

coordinates with 12 sectors in the horizontal direction and 15 levels in the vertical direction, as shown in Fig. 4.4.69. All of the major elements in the experimental facility are modeled in the 

input. The lower plenum is modeled in sectors 1 through 9 and levels 1 through 3. The lower 

core support plate is modeled at the top of level 3. The lower plenum is noded into three axial 

levels to capture the geometry below the tube bundles, and to model the opening between the lower plenum and downcomer. The electrically heated core is modeled in sectors 1 through 8 and 

axial levels 4 through 9. Each of the eight sectors represents a heater rod bundle. The core baffle occupies sectors 9 and 10 and the same axial levels as the core. The upper plenum is modeled in 

sectors 1 through 9 and axial levels 10 through 14. The upper tie plate, which contains the 

minimum flow area between the core and upper plenum, is located in level 10, and the upper core 

support plate is located in level 11. The support columns connecting the upper head and the upper plenum are modeled with PIPE components in sectors 1, 3, 5, 6, and 7 to provide a flow path between the upper plenum and the upper head. The upper head is modeled in sectors 1 through 12 

at level 15. The downcomer is modeled in sectors 11 through 12 and axial levels 2 through 14.

Page 4-86
The intact cold leg and broken cold leg connections to the pressure vessel are in level 12 of 

sectors 11 and 12, respectively. The hot leg connections to the upper plenum are at the outer 

interface of sector 10 in level 13. The lower plenum ECC injection point is located in sector 10 at 

level 1. 

4.4.5.3.2 Heater Rods 

Heater rods in the test facility were fabricated to provide a 17-step, chopped-cosine 

distribution for the axial power. The TRAC fuel rod model, shown in Fig. 4.4.70, has only six 

axial sectors. This presented several challenges in modeling the electrically heated core as 

described in earlier SCTF Run 179 assessment calculations. To model the axial power 

distribution, the areas under the power distribution of the heater rods in the test facility and that of 

the TRAC model were set equal to preserve the total power. Also, the TRAC six-level axial 

power shape had to be adjusted to match the pre-ECC injection axial cladding temperatures of the 

test. Several parametric runs were made to determine the axial power shape and peak power so 

that the axial distribution of calculated cladding temperatures compared well with those provided 

by JAERI. 

A second problem encountered with the earlier SCTF assessment calculations involved 

specifying the correct heater rod material properties. SCTF Core-I studies showed that the heater

rod material properties had a large effect on cladding temperature. A parametric study was 

conducted by varying the density and specific heat of magnesium hydroxide. To verify the input, 

the cladding temperature was calculated for the adiabatic heatup conditions of Run 719, and then 

compared with the initial and boundary condition data provided by JAERI. For case 1, the values 

for density and specific heat provided by the heater rod manufacturer were used (density = 2500 

kg/m3, with specific heat defined from a temperature-dependent table). For case 2, the values for 

density and specific heat selected for use in the TRAC Core-I and Core-II models were used 

(density = 2800 kg/m3 , with specific heat defined from a temperature-dependent table increased 

by 2800/2500 from the manufacturer's recommended values). For case 3, intermediate values 

were used (density = 2625 kg/m3). For this study, a MgO density of 2650 kg/m3 and 

corresponding specific heat were used for the Run 719 analysis. 

To ensure that TRAC was correctly calculating the adiabatic heatup, a TRAC calculation that 

could be verified with a hand calculation was performed. The SCTF Core-III model was used 

with the following changes. The axial and radial relative power profiles were set to unity, all rod 

material properties were set equal to the properties for Nichrome, and the core power was set at a 

constant value of 10 MW. The heatup period was 20 s. The hand-calculated lumped parameter 

energy balance predicted a uniform temperature of 488.5 K (no radial distribution); TRAC 

calculated a heater rod centerline temperature of 487.7 K, and a cladding temperature of 485.5 K. 

The difference represents the slight heat losses from the heater rod to the steam environment 

during the adiabatic heatup. Because the heater rod material properties are important in 

determining the cladding temperature, further effort may be required to ensure that the correct 

material properties are being used in other assessment analyses.

Page 4-87
4.4.5.3.3 Hot-Leg and S/W Separator

The hot leg and S/W separator noding are shown in Fig. 4.4.71. The hot leg is modeled with a 

PIPE component containing four cells, and the outer surface of the hot leg is assumed to be 

insulated. 

The S/W separator is modeled with a 3-D VESSEL component with four axial levels, one 

radial sector, and two azimuthal sectors. The hot leg connection to the S/W separator is in the 

outer radial face of cell 1 at level 3. The respective intact and broken cold leg connections to the 

S/W separator are made at level 2 to the outer radial faces of cells 1 and 2. 

4.4.5.3.4 Intact Cold Leg 

The intact cold leg noding is shown in Fig. 4.4.72. The intact cold leg is modeled with a TEE 

component consisting of seven cells in the primary side, and one cell in the secondary side. The 

secondary side of the TEE component is connected to the primary side at cell 7. The primary side 

includes the loop seal and the pump simulator. The intact cold leg orifice and pump simulator 

orifice are located at cell faces 2 and 6. The secondary side models the ECC injection port. 

4.4.5.3.5 Broken Cold Leg (Pressure Vessel Side) and Containment Tank 1 

The broken cold leg noding between the pressure vessel and containment tank 1, shown in 

Fig. 4.4.73, is modeled with a VALVE component containing six cells. The break valve is at the 

interface between cells 3 and 4. The opening and closing actions of the valve are modeled with 

the trip and control logic in TRAC. For Run 719, the valve was set fully open and not actuated. 

Containment tank 1 is modeled as a BREAK component using the measured containment 

pressure as the boundary condition. 

4.4.5.3.6 Broken Cold Leg (S/W-Separator Side) and Containment Tank 2 

The broken cold leg noding between the S/W separator and containment tank 2, shown in 

Fig. 4.4.74, is modeled with a VALVE component containing eight cells. The break valve for this 

segment of the broken cold leg is at the interface between cells 3 and 4. The opening and closing actions of the valve are modeled with the trip and control logic in TRAC. For Run 719, the valve 

was set fully open and not actuated. Containment tank 2 is modeled as a BREAK component using the measured containment pressure as the boundary condition. 

4.4.5.4 Comparison of TRAC-M(F77) and TRAC-M(F90) Calculations 

Calculations were performed using both TRAC-M(F77) and TRAC-M(F90). Figs. 4.4.75 

through 4.4.83 compare the clad temperatures calculated at different elevations in the core. In 

these figures, experimental data are not shown since the data are obtained under 2-D/3-D program and the data are not publicly available. When an agreement is reached with other partners to 

release these data, this report will be revised to include comparisons between the data and the 

code predictions.*

Page 4-88

* See footnote on page 4-79.
As shown in Figs. 4.4.75 through 4.4.83, the agreement between the code predictions is 

"Excellent" to "Reasonable." Differences between the predictions are small, considering that 

some error corrections were implemented in the TRAC-M(F90) code. Consequently, we consider 

the conversion from TRAC-M(F77) to TRAC-M(F90) to be successful for this application. 

4.4.5.5 Conclusions 

The conversion TRAC-M(F77) to TRAC-M(F90) has been successful for this application. 

REFERENCES 

4.4.1 Tuzla, K., C. Unal, 0. Badr, S. Neti, and J. C. Chen, "Thermodynamic Nonequilibrium 

in Post-Critical-Heat-Flux Boiling in a Rod Bundle," U.S. Nuclear Regulatory 

Commission Document NUREG/CR-5095, Vols. 1-4, July 1987. 

4.4.2 Obot, N.T., and M. Ishii, "Two-Phase Flow Regime Transition Criteria in Post-Dryout 

Region Based on Flow Visualization Experiments," Int. J. Heat Mass Transfer, 

Vol. 31, No. 12, pp. 2559-2570, 1988. 

4.4.3 Hirano, K., et al., "Data Report on Large-Scale Reflood Test-14, CCTF Test Cl-5 

(Run 014)," Japan Atomic Energy Research Institute Document JAERI-memo 57-214, 

August 1982. 

4.4.4 Odar, F., NUREG-1744, "Assessment of the TRAC-M Codes Using FLECHT

SEASET Reflood and Steam Cooling Data," NRC: Washington, D.C. May 2001. 

4.4.5 Loftus, M.J., et al., "PWR FLECHT-SEASET Unblocked Bundle Forced and Gravity 

Reflood Task Data Report," Westinghouse Electric Corporation Report, WCAP-9699, 

NUREG/CR-1532, Vols. I and 2, 1980. 

4.4.6 Nelson, R.A., D.A. Pimentel, S.J. Jolly-Woodruff, and J.W. Spore, "A 

Phenomenological Thermal-Hydraulic Model of Hot Rod Bundles Experiencing 

Simultaneous Bottom and Top Quenching and Optimization Methodology for Closure 

Development,"Los Alamos National Laboratory Report, LA-UR-98-3043, April 1998. 

4.4.7 Damerell, P.S., and J.W. Simons, Eds., "2-D/3-D Program Work Summary Report," 

U.S. Nuclear Regulatory Commission Document NUREG/IA-0126, June 1993. 

4.4.8 Sugimoto, J. and Y Murao, "Evaluation Report on CCTF Core I Reflood Tests C1-5 

(Run 14), C1-7 (Run 16) and C1-14 (Run 23) - Effects of Initial Clad Temperature," 

Japan Atomic Energy Research Institute Document JAERI-M-83-026, February 1983. 

4.4.9 Aksan, N., F. D'Auria, H. Glueser, R. Porchard, C. Richards, and A. Sjoberg, 

"Separate Effects Test Matrix for Thermal-Hydraulic Code Validation, Volume I -

Page 4-89
Phenomena Characterization and Selection of Facilities and Tests," Organization for 

Economic Cooperation and Development Nuclear Energy Agency Document 

NEA/CSNI/R(93)/Part. 1/Rev., and "Volume II - Facility and Experiment 

Characteristics," Organization for Economic Cooperation and Development Nuclear 

Energy Agency Document NEA/CSNI/R(93)/Part 2/Rev., September 1993. 

4.4.10 "CSNI Integral Test Facility Validation Matrix for the Assessment of Thermal

Hydraulic Codes for LWR LOCA and Transients," Organization for Economic 

Cooperation and Development Nuclear Energy Agency Draft Document 132, Rev. 6, 

June 1995. 

4.4.11 Motley, F., "Research Information Report Results from TRAC Analysis of Cylindrical 

Core Test Facility Core I Test Series," Los Alamos National Laboratory Report, 

LA-20130-TN-85-5, January 1986. 

4.4.12 Akimoto, H., "Analysis of TRAC-PF1 Calculated Core Heat Transfer for a CCTF 

Test," Nuclear Engineering and Design 88, 215-227, 1985.

Page 4-90
Table 4.4.1 Predicted and Measured Quench Times

Table 4.4.2 Predicted and Measured Quench Velocities

Page 4-91

Location (M) Experimental Predicted Quench Time, Predicted Quench Time, 

Quench Time (s) TRAC-M(F77) (s) TRAC-M(F90) (s) 

0.15 65.5 31.5 27.5 

0.30 114.7 111.0 100.0 

0.45 178.0 200.0 184.0

Location (m) Experimental Predicted Quench Predicted Quench 

Quench Velocity Velocity, TRAC-M(F77), Velocity, TRAC-M(F90), 

(mInm/s) (mm/s) (mm/s) 

0.15 2.3 2.5 1.8 

0.30 3.0 1.4 1.3 

0.45 2.4 2.9 2.5
Table 4.4.3 Conditions and Key Results for Run 31504

Table 4.4.4 Flow Regime Transition Criteria

Page 4-92

Parameter Test Condition Value 

Upper Plenum Pressure 0.28 MPa 

Rod Initial Clad Temperature at the 1.83 m Level 1136 K 

Rod Peak Power 2.3 kW/m 

ECCS Injection Rate -0.37 kg/s 

Coolant Temperature 324 K 

Radial Power Distribution Uniform 

Initial Temperature of Hottest Rod 1093 K 

Elevation of Maximum Temperature of Hottest 1.98 m 

Rod 

Maximum Temperature of Hottest Rod 1423 K 

Temperature Rise 330 K 

Turnaround Time of Hottest Rod 130 s 

Quench Time of Hottest Rod 325 s 

Bundle Quench Time 594 s

Regime Ishii's Correlation Region of TRAC 

Applicability Application 

Smooth Inverted Annular Z/D = 60,rJ'a We/c 2 i < 10- 2 0.05 < a < 0.3 

Flow (IAF) 

Rough Wavy IAF Z/D = 295JC We/c1 2 < 10-1 0.3 < cc < 0.4 

Agitated IAF Z/D = 595 N/U& We/oC2i < 3.5 0.4 < (x < 0.75 

Dispersed Section Z/D = 595 F-- We/ 2 i< 10 0.75 <cc < 0.98
Table 4.4.5 Operating Parameters for CCTF Run 14

Page 4-93

System Pressure 2.02 kg/cm2 

Total Power 9.36 MWt 

Radial Power Distribution 1.07:1.0:0.82 

Downcomer Wall Temperature 1820C 

Primary Piping Wall Temperature 1200C 

Steam-Generator Secondary-Side Liquid Temperature 2620 C 

Peak Clad Temperature at ECC Injection Initiation 5020C 

ECC Liquid Temperature 114 0C 

Lower-Plenum Liquid Level 0.87 m 

ACC Injection Rate 278 m3/h 

Low-Pressure Coolant Injection Rate 30.2 m3

/h
Table 4.4.6 CCTF Component Scaled Dimensions

Component 

Pressure Vessel 

Vessel inside diameter (mm) 

Vessel thickness (ram) 

Core barrel outside diameter (mm) 

Core barrel inside diameter (mm) 

Thermal shield outside diameter (mm) 

Thermal shield inside diameter (mm) 

Downcomer length (mm) 

Downcomer gap (mum) 

Downcomer flow area (m2? 

Lower-plenum volume (mi) 

Upper-plenum volume (mi3) 

Fuel (heater-rod) Assembly 

Number of bundles 

Rod array 

Rod heated length (mm) 

Rod pitch (mm) 

Fuel rod outside diameter (mm) 

Thimble tube diameter (mm) 

Instrument tube diameter (mm) 

Number of heater rods 

Number of nonheated rods 

Core flow area (m2) 

Core fluid volume (m3) 

Primary Loop 

Hot-leg inside diameter (mm) 

Hot-leg flow area (m2

) 

Hot-leg length (mm) 

Pump-suction inside diameter (mm) 

Pump-suction flow area (M2

) 

Pump-suction length (mm) 

Cold-leg inside diameter (mm) 

Cold-leg flow area (M2) 

Cold-leg length (mm)

PWR 

4394 

216 

3874 

3760 

4170 

4030 

4849 

114.3 

4.23 

29.6 

43.6 

193 

15x15 

3660 

14.3 

10.72 

13.87 

13.87 

39372 

4053 

5.29 

17.95

(173 in.) 

(8.5 in.)

736.6 (29 in.) 

0.426 

3940 

787.4 (31 in.) 

0.487 

9750 

698.5 (27.5 in.) 

0.383 

5600

Page 4-94

CCTF 

1084 

90 

961 

929 

4849 

61.5 

0.197 

1.38 

2.04 

32 

8x8 

3660 

14.3 

10.7 

13.8 

13.8 

1824 

244 

0.25 

0.915

155.2 

0.019 

3940 

155.2 

0.019 

7950 

0.019 

5600

Ratio

1/1 

1/21.44 

1/21.44 

1/21.44

1/1 

1/1 1/1 

1/1 

1/1 

1/21.58 

1/18.09 1/21.2 

1/19.6

1/4.75 

1/22.54 

1/1 

1/5.07 

1/25.77 

1/1 

155.21/4.50 

1/20.26 

1/1
Table 4.4.6 (cont) CCTF Component Scaled Dimensions

Component 

Containment tank I (mi3) 

Containment tank 2 (mi3 ) 

Storage tank (M3) 

Accumulator tank (M3) 

Saturated water tank (mi3 ) 

Steam-generator Simulator 

Number of tubes 

Tube length (average) (m) 

Tube outside diameter (mm) 

Tube inside diameter (mm) 

Tube wall thickness (mm) 

Heat-transfer area (M2) 

Inlet-plenum volume (M3) 

Outlet-plenum volume (im3) 

Primary-side volume (m3) 

Secondary-side volume (M3)

PWR

3388 

20.5 

22.225 (0.875 in.) 

19.7 (0.05 in.) 

1.27 

4784 (51 500 ft2) 

1.03 

4.25 

4.25 

30.50 (1077 ft3) 

157.33 (5556 ft3)

Page 4-95

CCTF Ratio

30 

50 

25 

5 

3.5

158 

15.2 

25.4 

19.6 

2.9 

192 

0.048 

0.198 

0.198 

2.4 

4.9

1/21.44 

1/1.35 

1/1 

1/24.92 

1/21.44 

1/21.44 

1/21.44 

1/15.25 

1/26.22
Table 4.4.7 Component Elevations of the CCTF

Component PWR CCTF Discrepancy 

Bottom of heated region in core (mm) 0 0 0 

Top of heated region in core (mm) 660 3660 0 

Top of downcomer (mm) 4849 4849 0 

Bottom of downcomer (umm) 0 0 0 

Centerline of cold leg (mm) 5198 4927 -271 

Bottom of cold leg (inside) (mm) 4849 4849 0 

Centerline of loop-seal lower end (mm) 2056 2047 -9 

Bottom of loop-seal lower end (mm) 1662 1959 +297 

Center of hot leg (num) 5198 4927 -271 

Bottom of hot leg (inside) (mm) 4830 4849 +19 

Bottom of upper core plate (mm) 3957 3957 0 

Top of lower core plate (mm) -108 -50 +58 

Bottom of tube sheet of steam-generator 7308 7307 -1 

simulator (mm) 

Lower end of steam generator 5713 5712 -1 

simulator plenum (mm) 

Top of tubes of steam generator 17952.7 14820 

simulator (avg.) (mm)

Page 4-96
Table 4.4.8 Initial Test Conditions

Parameter Test Condition 

Total power 7.87 MW 

Linear power 1.18 kWm1 

Radial power distribution A B C 

1.37 1.20 0.76 

Containment pressure 0.2 MPa 

Steam generator secondary pressure 5.3 MPa 

Downcomer wall temperature 468 K 

Vessel internals temperature 423 K 

Primary piping temperature 406 K 

Lower plenum liquid temperature 394 K 

ECC liquid temperature 310 K 

ECC injection type Lower plenum and cold leg injection 

Pump K factor 15 

ECC injection rates, durations, and injection Accumulator flow: 0.104 m3s-1 from 81.0 to 

locations 94.0 s into lower plenum* 

Accumulator flow: 0.088 m3 s-1 from 94.0 to 

107.0 s into cold legs* 

LPCI flow: 0.0116 m3s-1 from 107.0 to 979.0 s 

into cold legs* 

Initial water level in lower plenum 0.86 m 

Initial water level in steam generator 7.4 m 

secondary side 

Power decay Constant, from 0 to 90.5 s* 

ANS 1.0 + actinide (30 s after scram), from 90.5 

to 690.5 s* 

Linear decay, from 690.5 to 898.0 s* 

Reflood initiation time 90.5 s* 

Peak clad temperature at reflood initiation 1074 K at TE31Y17

* Time in this table is defined as time after test initiation.

Page 4-97
Table 4.4.9 Chronology of Testing Events for CCTF Run 54

Event 

Test initiated 

(Heater rods power on) 

(Data recording initiated) 

Accumulator injection initiated 

Power decay initiated 

(Bottom of core recovery) 

(Reflood initiated) 

Accumulator injection switched 

from lower plenum to cold leg 

Accumulator injection ended and 

LPCI injection initiated 

Maximum PCT time 

All heater rods quenched 

Linear power decay initiated 

Power off 

LPCI injection ended 

Test ended (Data recording ended)

Time After Test 

Initiation (s) 

0

Time Relative to 

Reflood Initiation (s) 

-90.5

81.0 

90.5 

94.0 

107.0 

123.0 

552.5 

690.5 

898.0 

979.0 

1032.0

-9.5 

0.0 

3.5 

16.5

32.5 

462.0 

600.0 

807.5 

888.5 

941.5

Page 4-98
Table 4.4.10 Summary of Test Conditions for SCTF Run 719 (S3-15)

Parameter Test Condition 

System pressure 0.2 MPa 

Initial total core power 7.12 MW 

Radial power distribution 1.36/1.2/1.1./1.0 

(bundle 1/2/3/4/5/6/7/8) 0.91/0.86/0.81/0.76 

Decay power 1.02 x ANS + 1.02 x actinides 

Initial peak cladding temperature 1134 K 

(at initiation ECC injection) 

ECC injection mode Lower plenum injection 

Upper head injection None 

UCSP injection None 

Lower plenum injection flow rate 37 kg s-1 for 13 s 

37 to 13 kg s- 1 over 15 s 

13 to 4.1 kg s-1 over 70 s 

4.1 to 3.75 kg s-1 over 10 s 

3.75 kg s- 1 for 500 s 

Water temperature 353 K for 36 s 

393 K for remainder of test 

Injection period 608 s 

Cold leg injection None 

Lower plenum liquid level 1.2 m, saturated at 393 K

Page 4-99
45 kwo,AC "SIGHT GLASS 

BOILER 

FLOW 

METER 

METER FLOW 

Figure 4.4.1 Two-Phase Heat Transfer Test Loop

Page 4-100
44.5 mm

E " __ Shroud 

E ( E \ 9 RodS E6 

In 

Cell I 

Cell 11 

'Figure 4.4.2 Cross-Sectional View of the Lehigh Test Bundle

Page 4-101
Ports for Vapor 

Temperature and 

Pressure Probes 

M -A_ Copper Block 

LO 

Hot Patch 

C 

Hot Rods 

Strainer 

CO 

Lower Seal 

Test Section Inlet 

(All dimensions are in millimeters) 

Figure 4.4.3 Lehigh Rod Bundle Test-Section Schematic

Page 4-102

C 

0 

U) 

I-
Vapor Probe 

Location _

Spacer 

Grid No 

Location 

Vapor Probe 

Location 

Pressure 

Port

EXIT PIPE Elevation 

_ 1.219

INJECTION PIPEI 

FILL

fComponent 5 

Component 4 

Component 1 

Component 2

Component 3

Figure 4.4.4 Lehigh Rod Bundle Facility Input Model Schematic

Page 4-103
Lehigh Reflood Test 02/24/85-20

Comparison of Calculated and Measured Clad Temps. Z=0.1 5m

100 200 

Time (s) 

Figure 4.4.5 Comparison of Calculated and Measured Clad Temperatures (Z = 0.15 m) for Lehigh 

Reflood Te-1"12/24/85-20

1000 

800

CD 

4k

-D 

CL 

E 

CO 

0

600

400 

200

0 300

K
(.

0 100 200 

Time (s) 

Figure 4.4.6 Comparison of Calculated and Measured Clad Temperatures (Z = 0.30 m) for Lehigh 

Reflood Test 02/24/85-20

Lehigh Reflood Test 02/24/85-20 

Comparison of Calculated and Measured Clad Temps. Z=0.30m

1200 

1000

(

Cz 

0.

eQ 

0

800 

600

400 

200

300
Lehigh Reflood Test 02/24/85-20

Comparison of Calculated and Measured Clad Temps. Z=0.45m

e-0 TRAC-M(F77) 

0 ..... Exp. Data TW Z=0.45m 

I+ - -E TRAC-M(F90), rftn-006001001006@i

:1

200 

0 100 200 300

Time (s)

Figure 4.4.7 Comparison of Calculated and Measured Clad Temperatures (Z = 0.45 m) for Lehigh 

Reflood Test 02/24/85-20

1200 

1000

Iq 

0 

0%

CL 

E 

I

"10

800 

600

400
Lehigh Reflood Test 02/24/85-20

Comparison of Calculated and Measured Clad & Vapor Temps. Z=0.60m

100 200 

Time (s)

Figure 4.4.8 Comparison of Calculated and Measured Clad Vapor Temperatures (Z = 0.60 m) for 

Lehigh Reflood Test 02/24/85-20

1200 

1000

Cb 

C -4

800 

600

CI 

E 0

400 

200

0 300
Lehigh Reflood Test 02/24/85-20

Comparison of Calculated and Measured Clad & Vapor Temps. Z=0.965m

100 200 300

Time (s)

Figure 4.4.9 Comparison of Calculated and Measured Clad & Vapor Temperatures (Z = 0.965 m) for Lehigh Reflood Test 02/24/85-20

k.

1100 

1000

900

0 -I." 

E5 ai) 

CL

800

700 

600

500 

400 

300

0
STEAM 

Oil Tý I P VALVE LEGEND 

Oo EXHAUSTDRFERNTALE 

D/P,12 DIFFERENTIAL PRESSURE 

ORIFICESTATIC PRESSURE 

0(--)) 

FLUID THERMOCOUPLE 

AT 'Jj WALL THERMOCOUPLE 

FOiWATER UR NME R 

FRO 0/P SUPPLYTUBNMER 

,'I z , 12.19. LEVEL 

LEV "- ORIFICE PLATE FLOW METER /P STEAM SEPARATOR 

rA 0-100 GPM LEVEL F-- ROTAMETER I F -= -D. CONTROL 

8 I GPM 6.31 x 10- 51

3/sec 

LIQ. LEVEL 0-6 6PM I FOOT 0.305 MEIER -- :=. TURBINE METER 

CARRYOVER TANK TEST SECT ION' 

oI:::1 'ALL ALL IROTAMETERS INSTRUMENTATION IS NOT SHOWN
A B C D E F G H I J K L M N 0

WINDOW 1800 

1 000Q 

2 003000002Q000 

7 Q OOQO000 

00 R 

PRESSUR 

4 OIC0 18(DM •• S•( oo 

e°AP 

5 0 00 1 (DO511- 00 

"6 0 06 2 G )00 40, 

"0?0000 A0 OQOQO7O103 C.00 

00 (DO ®QQ 

WINDOW 

Figure 4.4.11 FLECHT-SEASET 161 Rod Test Bundle

Page4-110
ELEVATION (ft) 

0 I 2 3 , 5 6 7 8 9 10 I1 12

i.8 

1.6 

1.4, 

1.2 

1.0 

0.8 

0.6 

0.1

0.2 

0.0

0.00 0.25 0.50 0.75 1.00 1.25 1.50 1.75 2.00 2.25 2.50 2.75 3.00 3.25 3.50 

ELEVATION (m) 

Figure 4.4.12 FLECHT-SEASET Axial Power Shape

Page 4-111

r.4 

a-. 

0 

I

C,1 

",4
1.0 

0.9 

0.8 

o 0.7 

0.6 

0.5 

0.0 

0 100 200 300 400 500 600 700 

TIME AFTER FLOOD (SEC) 

Figure 4.4.13 Decay Power Curve (ANS + 20%)

800 900 1000

Page 4-112
Core Section 

of Vessel Wall Fuel Rods 

ROD ROD

End 

of Heafng

VESSEL 

Cartesian Geomnetry 

(x=l, y=1, z=1 8) 

.(0)

heIft (m) 

t4

F-1 BREAK 4 

(iIF-1 PIPE 

3 

Fý 4.2926 

3.6578 V 4.0640

3.1090 

2.9261 

2.7432 

2.5603 

2.3774 

2.1946 

2.0117 

1.6459 

1.4630 

12802 

1.0973 

0.9114 

0.7315 

0.5486

17 

16 

15 

14 

13 

12 

11 

10 

9 

8 

7 

6 

4

3.5154 

3.3325 

3.1496 

2.9667 

2.7838 

2.6010 

2.4181 

2.0523 

1.8694 

1.6866 

1.5037 

I13208 

1.1379 

0.955m

-0.40U 

0.1778 

S0.0 

2 

I - PIPE 

ILE

Grid Spacer Locations 

-- 4.2926 

- 3.56 

-r 3.0226 

*-2.5146 

S-1.9812 

S-1.4478

14- 0.9396

-W VU0qu

Figure 4.4.14 TRAC-M Input Model Schematic for FLECHT-SEASET Run 31504

Page 4-113

of Heatin

3

AA
Flecht Seaset Forced Reflood 31504

Comparison of Calculated Initial Clad Temperatures with Experimental Data

1 2 3 4 

Axial Height - Heater Rod Length (m)

Figure 4.4.15 Initial Axial Temperature Variation

(

1200

1000

UIQ

.I

CL 

E L) I.-

800

600 

400

0
Flecht Seaset Forced Reflood 31504 

Comparison of Calculated Initial Axial Steam Temps. with Exp. Data

1 2 3 4 

Axial Height - Heater Rod Length (m)

Figure 4.4.16 Initial Axial Steam Temperature Variation

1000

800 

600

(D 

Q... 

4-'! 

E C

E 

CO 

U)

It

400

0

\
Flecht-Seaset Forced Reflood 31504 

Comparison of Calculated and Experimental Clad Temps z=1.22m

100 200 

Time (s)

Figure 4.4.17 Comparison of Calculated and Experimental Clad Temperatures @ z = 1.22 m

1200

1000

00 

4

800 

600

Q

E 

CO~

400 

200

0 300
(' ( 

Flecht-Seaset Forced Reflood 31504 

Comparison of Calculated and Experimental Clad Temps z=1.98m 

1750 

G-O TRAC-M(F77ý 

----- TRAC-M(F90 

X---e 3 -078 

R .-... 0 6J-078 

1250 

oeQ E • , 

:3 

4-C. 

V., 

750 

250 0 100 200 300 400 500 

Time (s)

Figure 4.4.18 Comparison of Calculated and Experimental Clad Temperatures @ z = 1.98 m
Flecht-Seaset Forced Reflood 31504 

Comparison of Calculated and Experimental Clad Temps z=2.59m

200 400

Time (s) 

Figure 4.4.19 Comparison of Calculated and Experimental Clad Temperatures @ z = 2.59 m

1750 

1250

CL 

E 

Ca) 

L. 

4-'

to

750

250

0 600
(

Flecht-Seaset Forced Reflood 31504 

Comparison of Calculated and Experimental Clad Temps z=3.35m

200 400 600 800 

Time (s)

Figure 4.4.20 Comparison of Calculated and Experimental Clad Temperatures @ z = 3.35 m

1400 

1200

1000 

800 

600

£ 

D. 

E 

cj) 

a) 

L.. 

4-. C.) 

4-' 

C,)

400 

200 

0

0 1000
Flecht Seaset Forced Reflood 31504

Comparison of Calculated and Experimental Quench Elevations

X

)WCX

X X

X

X 

x

I

XX

XX

1-0

1.3"

)3K

X

0 -.

/

0-o TRAC-M (F77) 

a-- 10 TRAC-M (F90) 

X Experimental Data

400 600 800

Time (s) 

Figure 4.4.21 Calculated and Measured Quench Front Propagation

4 

3

10 

03 ,Q

E 

€.. 

0 

ca

X - 1-e

.O"

.0

1 

0

X

0 200

X -/-.

(;1

/
/

Flecht-Seaset Forced Reflood 31504 

Experimental DPs along Test Section 

5 i i 

0-a BU-3-4 

El----El BU-0-1 

A- -A BU-1-2 

A-- -0 BU-2-3 

' • -m BU-4-5 

-A BU-5-6 

4 V-----V BU-6-7 

X--X BU-7-8 

*-4BU-8-9 

v'.... vBU-9-10 

23_ :.'" - -a-- - •- - __•- ..--.. z-•--,., , =0 =-. • --. = .•• -- .•-.--0 -- _ 

cQ / 

:1= 

I1C1 I ' 'l 

5 II , /• €/,,(,• , 

I , . , A ' i) %/ '•' f,...,,'.,'- ".'t..f. • , i.'. , 

0 0100 200 300 400 500 600

Time (s)

Figure 4.4.22 Experimental Differential Pressure Measurements Along the Test Section

(
Flecht Seaset Forced Reflood 31504 

Calculated Axial Void Fraction Distribution at Different Times

0 1 2 3 4 5

Axial Distance (m)

Figure 4.4.23 Calculation of Axial Void Fraction Distribution by TRAC-M(F77)

k.

i

1 

0.8

fJ'

" 0.6 

LL 

0 Q. 

> 0.4

0.2 

0
Flecht-Seaset Forced Reflood 31504 

Comparison of Differential Pressures, DP 1-2 

4 

LIt I 

S,. ... . ..e . ..o -; . . o i ... . ............ ------ .. .... ..to .......... .. .. .. .... . .. ... 

CIO• P' 6, RA , 

2,11" . - BU-1-2 Exp. Data 0L o , I... n DP(1-2) TRAC-M (F90) ""Al1 -1, 

G--O DP(1-2) TRAC-M (F77) 

ci' 

0 

0 500 

Time (s)

Figure 4.4.24 Comparison of Differential Pressures (Taps 1 - 2)

I
Flecht-Seaset Forced Reflood 31504 

Comparison of Differential Pressures, DP 2-3 

4 

C J if 

II 

II 

0 L I .I II II II 

22 ,I,,,r' 

I IIv * .... BU-2-3 

'~'Ii~'~ P 0 0DP(2-3) TRAC-M (F90~ 

eoDP(2-3) TRAC-M (F77)1 

hI 

-,, II 

0 ; 

0 500 

Time (s)

Figure 4.4.25 Comparison of Differential Pressures (Taps 2 - 3)
Flecht-Seaset Forced Reflood 31504 

Comparison of Differential Pressures, DP 3-4

0

Time (s)

Figure 4.4.26 Comparison of Differential Pressures (Taps 3 - 4)

4 

3

Cb

C,,) 

"C2 

U). 

ci 

0-

1 

0

500
Flecht-Seaset Forced Reflood 31504 

Comparison of Differential Pressures, DP 4-5

Time (s) 

Figure 4.4.27 Comparison of Differential Pressures (Taps 4 - 5)

5

crO 

t,

CO 

U) 

Co 

CD 

eI 

L.

0

0

1000 500
Flecht-Seaset Forced Reflood 31504 

Comparison of Differential Pressures, DP 5-6

0 500 

Time (s)

Figure 4.4.28 Comparison of Differential Pressures (Taps 5 - 6)

(

5

(

"-4~

Cn 

U) 

0, 

13... 

U) 

0D

0

1000
Flecht-Seaset Forced Reflood 31504 

Comparison of Differential Pressures, DP 6-7

500 

Time (s) 

Figure 4.4.29 Comparison of Differential Pressures (Taps 6 - 7)

4

,-0 

00

as 1I... 

a. 

0) 

13

-1

0 1000
(

Flecht-Seaset Forced Reflood 31504 

Comparison of Exp. Steam Measurements at z=1.83-1.94m with Calculations

1300 

1100

100 300 500 

Time (s)

Figure 4.4.30 Steam Temperatures (Z = 1.83 - 1.94 m)

E 

E 

ci) 4CO 

U,

900 

700

500

300 L 

"-100

\
Flecht-Seaset Forced Reflood 31504

Comparison of Exp. Steam Measurements at z=2.82-3.05m with Calculations

100 300

Time (s) 

Figure 4.4.31 Steam Temperatures (Z = 2.82 - 3.03 m)

1300 

1100

2' 

E 

E 

Cz (I,

go

900 

700

500

300 

-100 500

! k
(

Flecht-Seaset Forced Reflood 31504 

Comparison of Exp. Steam Measurements at z=3.35-3.51 m with Calculations

1300 

1100

100 300 

Time (s)

Figure 4.4.32 Steam Temperatures (Z = 3.35 - 3.51 m)

(

C) 

a) 60. 

E CO 

CD)

900 

700

500

300 

-100 500

/

10 

to
Flecht-Seaset Forced Reflood 31504 

Comparison of Test Section Pressures

200 400

Time (s) 

Figure 4.4.33 Comparison of Test Section Pressures

(

3e+06

2e+06

w

a3. 

a2.

1 e+06 

0

0
Flecht Seaset Forced Reflood 31504 

Comparison of Calculated and Experimental Liquid Flows

200 400 600 

Time (s)

Figure 4.4.34 Comparison of Liquid Carryouts

1'

200

"4-0 

>3 

0 

*0 

CU 

E

100

0 L 

0 800
Flecht Seaset Forced Reflood 31504 

Comparison of Calculated and Experimental Steam Flows

200 400 600 

Time (s) 

Figure 4.4.35 Comparison of Steam Carryouts

200 

150

rJQ (b

100

0 

E 

ca~ 

(D) 

E 

00

50 

0

0

t.

800

(
(

Flecht-Seaset Forced Reflood 31504 

Comparison of Calculated and Experimental Clad Temps z=1.22m

100 200 

Time (s)

Figure 4.4.36 TRAC-M(F77) Clad Temp. Predictions with and without Choking @ z = 1.22 m

1200 

1000

(

800 

600

CL 

E 

a) 

i

5CO 

vi

400 

200

0 300
Flecht-Seaset Forced Reflood 31504 

Comparison of Calculated and Experimental Clad Temps z=1.98m

100 200 300 400 

Time (s)

Figure 4.4.37 TRAC-M(F77) Clad Temp. Predictions with and without Choking @ z = 1.98 m

1750

1250

ON

E 

ca) 

0

750 

250

0 500
(

Flecht-Seaset Forced Reflood 31504 

Comparison of Calculated and Experimental Clad Temps z=2.59m

0 200 400 

Time (s)

600

Figure 4.4.38 TRAC-M(F77) Clad Temp. Predictions with and without Choking @ z = 2.59 m

1750 

1250

E 

0

750 

250

S
Flecht-Seaset Forced Reflood 31504

Comparison of Calculated and Experimental Clad Temps z=3.35m ' I 'rI

"A\

..... O 0 11E-132 ., . . 0 0- TRAC-M(F77), With Chok aU- -0 111-132 

,-A - 11 K-1 32 

0-0 TRAC-M(F90) 

)0 ......... TRAC-M(F77), No Choke 

0 0 200 400 600 800 1000

Time (s)

Figure 4.4.39 TRAC-M(F77) Clad Temp. Predictions with and without Choking @ z = 3.35 m

1400 

1200 

1000

00 

CD 

00

E 

"U) 

H

800 

600

4C 

20
ZAG

ZRW 

Inverted 

Annular 

Flow 

* (Rough-wavy) 

ZSM 

Inverted 

Annular 

Flow 

(Smooth) 

ZTR iTransition 

ZCHF 

F Bubbly Flow 

Figure 4.4.40 TRAC-M Reflood Flow Regime Map

Page 4-139
Flecht Seaset Forced Reflood 31504 

Comparison of Relative Velocities, (Vg-VI), at Cell#1 0

200 400 600

Time (s) 

Figure 4.4.41 Calculated Relative Velocities at Cell 10

20 

15

0 

Cci a1)

hv 

,-o

10 

5 

0

-5 

-10 

0 800

k.
It' -7 09d

Relative Velocity, Vg-VI, (m/s) 

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

N-)---- -

C- ; 

-- -. - _-------- 0 On 

-------- CD 

_CD gcD 

-- ~( -- C -nCD 

0~ 

0 0 

IC - -

0,0 0 0 

>0 C 

-0 - - - -0
•tank a 

Sftank ConFeinment I F 

Figure 4.4.43 Isometric View of CCTF

K
I!

K

Exhaust Pipe

I') "- ...... ECC water Injection locations: Cold Leg Cold leg: ECCS (D2)(3) and ) 

Pump Simulator Lower plenum: ECCS (5) 

Upper plenum: ECCS (1)

Pump 

Simulator 

20* 

90 

Pump Simulator

Figure 4.4.44 Top View of Primary-Loop Piping Layout

K
Medium-power zone 

(B region)

Intact 

hot leg 3

Broken' cold leg \ Ina\ Intact l lcold leg 

Boundary of bundle

Figure 4.4.45 Cross-Section of the CCTF Pressure Vessel 

Through the Core-I Vessel Core and Primary-Loop Nozzles

Page 4-144
Upper 

Plenum

Component No.: 3 

Vessel

Core

L 

ý_Lower Plenum

Elevation 

6.747 -

5.537

4.717 

4.217 

3.980 

3.860 

3.7325 

3.480 

3.250 

3.045 

2.845 

2.640 

2.435 

2.235 

2.030 

1.825 

1.625 

1.420 

1.215 

1.015 

0.810 

0.580 

0.3275 

0.200 

0.000

Component Component 

No.: 4 No.: 5 

Exit Pipe Break 

.Component No.: 6 

Rod

-Component No.: 7 

Slab

No.: 100

fInjection Pipej I Fill I 

Component No.: 2 Component No.: 1

Figure 4.4.46 CCTF Input Model Schematic

Page 4-145

q m
3.66 

3.5

3.0 

2.5 

1. 4 1 91.4 2 

.0 -( 1.417 

Cz 

>~ 1..1 CD 

1.0 ~ 1.17 

0.815 

T 1 0 0.5 1.0 L5 

S , Axial Power Profile (Q/QAv.) 

E 4 Elo 

Figure 4.4.47 TRAC Model of the CCTF Run 14 Power Profile and Instrument Locations

Page 4-146
(

CCTF Run#14, Reflood Test

Comparison of TRAC-M(F77) and TRAC-M(F90) Calculations, Z=0.38m

- TRAC-M(F77) n 

TRAC-M(F90), rftn-006001001006@0.38000q

I, 

I / 

it

I I

100 200 300 400

Time (s) 

Figure 4.4.48 CCTF Run 14, Reflood Test 

Comparison of TRAC-M(F77) and TRAC-M(F90) Calculations (Z=0.38m)

600 

550

500 

450 

400

0) 

I.

"*0

350 

300

0 500

4• 

...j
CCTF Run#14, Reflood Test 

Comparison of TRAC-M(F77) and TRAC-M(F90) Calculations, Z=1.1 05m 800 

TRAC-M(F77) 

--- TRAC-M(F90), rtn-00600100:10006@ 11.1:050J0 

700 

2.\600 /\ 

/\ 

0 E 

4600 

I3500 

400 """ 

0 100 200 300 400 500 

Time (s) 

Figure 4.4.49 CCTF Run 14, Reflood Test Comparison of TRAC-M(F77) and TRAC-M(F90) Calculations (Z=1.105m)

t
(

CCTF Run#14, Reflood Test

Comparison of TRAC-M(F77) and TRAC-M(F90) Calculations, Z=1.83m

100 200 300 400 500

Time (s)

Figure 4.4.50 CCTF Run 14, Reflood Test 

Comparison of TRAC-M(F77) and TRAC-M(F90) Calculations (Z=1.83m)

1000 

800

"eJ

0 

L.• 

E 

0

H

600

400 

200

0

f
CCTF Run#14, Reflood Test 

Comparison of TRAC-M(F77) and TRAC-M(F90) Calculations, Z=2.44m

100 200 300 400 500

Time (s) 

Figure 4.4.51 CCTF Test Run 14, Reflood Test 

Comparison of TRAC-M(F77) and TRAC-M(F90) Calculations (Z=2.44m)

1200 

1000

LA 

0

800 

600

CL 

0.

400 

200

0

t