NUREG/IA-0425 

International 

 Agreement Report

TRACE5 Assessment of 100% Direct Vessel Injection 

Line Break in ATLAS Facility 

Prepared by:

Byung-Gil Huh, Dong-Gu Kang, Chae-Yong Yang, Seung Hoon Ahn

Korea Institute of Nuclear Safety

Daejeon, Korea

 

 

 Ki-Yong Choi

Korea Atomic Energy Research Institute

Daejeon, Korea

 

A. Calvo, NRC Project Manager

 

 Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

Manuscript Completed: October 2012

Date Published: March 2013

Prepared as part of 

The Agreement on Research Participation and Technical Exchange

Under the Thermal-Hydraulic Code Applications and Maintenance Program (CAMP)

Published by

U.S. Nuclear Regulatory Commission
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NUREG/IA-0425 

International 

 Agreement Report

TRACE5 Assessment of 100% Direct Vessel Injection 

Line Break in ATLAS Facility 

Prepared by:

Byung-Gil Huh, Dong-Gu Kang, Chae-Yong Yang, Seung Hoon Ahn

Korea Institute of Nuclear Safety

Daejeon, Korea

 

 

 Ki-Yong Choi

Korea Atomic Energy Research Institute

Daejeon, Korea

 

A. Calvo, NRC Project Manager

 

 Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

Manuscript Completed: October 2012

Date Published: March 2013

Prepared as part of 

The Agreement on Research Participation and Technical Exchange

Under the Thermal-Hydraulic Code Applications and Maintenance Program (CAMP)

Published by

U.S. Nuclear Regulatory Commission

iii

ABSTRACT

The calculations using TRACE V5.0 patch2 code were conducted for 100% DVI line break test 

of ATLAS which is the first domestic standard problem (DSP-01). In the steady state conditions, 

the errors between the calculated and measured values are acceptable for most 

primary/secondary system parameters. From the transient calculations, the predicted sequence 

of events occurred some seconds later than that of experiment due to slow depressurization of 

the primary side after break. Before the LSC, the code could not predict well the behavior of 

downcomer and core collapsed water levels due to the under-prediction of the discharged mass 

and high flow rate returned to core from hot leg. After the LSC, the downcomer and core 

collapsed water levels drop rapidly and the code predicts relatively well the trend of downcomer 

and core water level. All RCPs suction legs or loop seals are cleared completely and suddenly. 

The predicted cladding temperature generally agrees well with the experiment except the 

peaking behavior. In the sensitivity studies, if the discharging mass increases arbitrarily, the 

downcomer and core water level before SIP injection shows better prediction result than that of 

base case. However, except for this, the overall behavior for main parameters shows the bad 

prediction results. For the effect of flow restriction from hot leg using the CCFL model, the 

predicted core water level was not significantly decreased compared to the base case. In 

conclusions, TRACE code has good capabilities to simulate the 100% DVI line break test of 

ATLAS. However, TRACE code including the choked flow and CCFL models needs to be 

improved and more detailed modeling is needed to predict more accurate results.

v

CONTENTS

Page

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

CONTENTS................................................................................................................................v

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

LIST OF TABLES ......................................................................................................................ix

EXECUTIVE SUMMARY ...........................................................................................................xi

FOREWORD ........................................................................................................................... xiii

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

2. ATLAS FACILITY AND TEST DESCRIPTION.....................................................................3

2.1 Overview of the ATLAS..................................................................................................3

2.2 Test Procedure ..............................................................................................................9

3. MODELING INFORMATION..............................................................................................13

3.1 Main Loop Modeling ....................................................................................................13

3.2 Reactor Vessel Modeling.............................................................................................16

3.3 Emergency Core Cooling System (ECCS) & Break Modeling......................................20

4. STEADY STATE ANALYSIS..............................................................................................25

5. TRANSIENT ANALYSIS....................................................................................................27

6. SENSITIVITY STUDY........................................................................................................41

6.1 Increase of Break Flow................................................................................................41

6.2 CCFL Model in the Upper Plenum...............................................................................43

7. CONCLUSIONS ................................................................................................................47

8. REFERENCES ..................................................................................................................49
vi
vii

LIST OF FIGURES

Page

Figure 2.1 Schematics of the ATLAS Facility...........................................................................4

Figure 2.2 Isometric Configuration of the ATLAS Facility.........................................................5

Figure 2.3 Arrangement of the Primary Loop of the ATLAS.....................................................6

Figure 2.4 Configuration of the Break Simulation System for the DVI Break Tests..................7

Figure 2.5 Schematics of the Containment Simulation System ...............................................8

Figure 2.6 Photograph of ATLAS ............................................................................................9

Figure 3.1 ATLAS Node for TRACE......................................................................................15

Figure 3.2 Reactor Vessel Modeling of ATLAS......................................................................19

Figure 3.3 Heat Structure Modeling for ATLAS Core.............................................................19

Figure 3.4 Emergency Core Cooling System (ECCS) in APR1400 .......................................23

Figure 3.5 TRACE Model for the Reactor Vessel and ECCS ................................................23

Figure 3.6 Plane Figure for the Connection between the Reactor and the ECCS .................24

Figure 5.1 Blowdown Phase (at 12.5 seconds after break) ...................................................29

Figure 5.2 Natural Circulation (at 102 seconds after break) ..................................................29

Figure 5.3 RCS Conditions after Loop Seal Clearing (at 140 seconds after break)...............30

Figure 5.4 Core Power..........................................................................................................31

Figure 5.5 Pressurizer Pressure ...........................................................................................32

Figure 5.6 Break Flow Rate ..................................................................................................33

Figure 5.7 Accumulated Break Mass ....................................................................................33

Figure 5.8 Core & Downcomer Collapsed Water Level .........................................................35

Figure 5.9 Core & Downcomer Collapsed Water Level in the Initial Period ...........................35

Figure 5.10 Pump Suction Leg Collapsed Water Level...........................................................36

Figure 5.11 Safety Injection Pump (HPSI) Flowrate................................................................37

Figure 5.12 Safety Injection Tank (SIT) Flowrate ....................................................................38

Figure 5.13 Hot Leg Massflow Rate........................................................................................39

Figure 5.14 Cold Leg Massflow Rate......................................................................................39

Figure 5.15 Heater Rod Surface Temperature ........................................................................40

Figure 6.1 Accumulated Break Mass ....................................................................................42

Figure 6.2 Downcomer Collapsed Water Level .....................................................................42

Figure 6.3 Core Collapsed Water Level ................................................................................43

Figure 6.4 Effect of CCFL Model on the Core Collapsed Water Level ...................................44

Figure 6.5 Effect of CCFL Model in RELAP5 on the Core Collapsed Water Level.................45

ix

LIST OF TABLES

Page

Table 2.1 Comparison of the Sequence of a 100% DVI Line Break.......................................10

Table 3.1 Applied Decay Power Level with the Variation of Time ..........................................18

Table 3.2 Mass Flow Rate of SIP with respect to the Downcomer Pressure..........................21

Table 3.3 Containment Conditions with Variation of Time......................................................22

Table 4.1 Steady State Calculation Results ..........................................................................26

Table 5.1 Sequence of Events ............................................................................................. 27

xi

EXECUTIVE SUMMARY

An integral effect test on the SBLOCA (Small-Break Loss of Coolant Accident) aiming at 100% 

DVI (Direct Vessel Injection) line break was conducted with the ATLAS in 2009 by KAERI (Korea 

Atomic Energy Research Institute). In this study, the calculations using TRACE V5.0 patch2 

code were conducted for 100% DVI line break test of ATLAS which is the first domestic standard 

problem (DSP-01) to assess TRACE code capability to simulate the transient thermal-hydraulic 

behavior for SBLOCA. 

For the modeling of ATLAS facility, the reactor vessel was modeled in three dimension using 

VESSEL component and flows within a coolant loop of the primary and secondary sides were 

modeled in one dimension using main hydraulic components such as PIPE and TEE in TRACE 

codes. Also, the emergency core cooling system, the heater rods and main control logics were 

properly modeled in order to consider the various conditions of this experiment. The steady 

state was determined by conducting a null transient calculation and the errors between the 

calculated and measured values are acceptable for most primary/secondary system 

parameters.

In an experiment, after initial steady-state conditions were reached, the DVI line break test was 

initiated by opening a break simulation valve at 199 seconds, and transient calculation was also 

conducted by setting up 199 seconds as the initial time. The predicted sequence of events 

occurred some seconds later than that of the experiment due to slow depressurization of the

primary side after break. Before the SIP starts, the predicted pressurizer pressure decreases 

slowly compared to the experimental results because of the under-estimation of break flow 

according to the characteristic of choked flow model. Before the LSC, the code could not predict 

well the behavior of the downcomer and core collapsed water levels due to the under-prediction 

of the discharged mass and high flow rate returned to core from hot leg. After the LSC, the 

downcomer and core collapsed water levels drop rapidly and the code predicts relatively well 

the trend of the downcomer and core water level. In this study, all RCPs suction legs or loop 

seals are cleared completely and suddenly and the code predicted the LSC ~ 25 seconds later 

than the measured value due to the slow depressurization rate. The predicted cladding 

temperature generally agrees well with the experiment except the peaking behavior. This peak 

of cladding temperature is directly related to core collapsed water level dip before the LSC. 

The sensitivity studies were performed to identify 1) the effect of discharging mass for the 

depressurization and 2) the effect of flow restriction from hot leg for the core collapsed water 

level. If the discharging mass increases arbitrarily, the downcomer and core water level before 

SIP injection shows better prediction results than that of the base case. However, except for 

this, the overall behavior of the main parameters shows the bad prediction results. For the effect 

of flow restriction from hot leg using the CCFL model, the predicted core water level was not 

significantly decreased compared to the base case. The results are different in comparison with 

that of RELAP5 code and we need to perform further studies to better understand this 

difference.

In conclusion, the TRACE code has good capabilities to simulate the 100% DVI line break test 

of ATLAS. However, there are some discrepancies in quantitatively predicting the primary 

pressure, break flow, the downcomer and core collapsed liquid level, and so on. Therefore, the 

TRACE code including the choked flow and CCFL models needs to be improved and more 

detailed modeling is needed to predict more accurate results.

xiii

FOREWORD

This report represents one of the in-kind contributions submitted to fulfill the bilateral agreement 

for cooperation in thermal-hydraulic activities between Korean Institute of Nuclear Safety (KINS) 

and the U.S. Nuclear Regulatory Commission (NRC) in the form of a Korean contribution to the 

NRC’s Code Assessment and Maintenance Program (CAMP), the main purpose of which is to 

validate the TRAC/RELAP Advanced Computational Engine (TRACE) Code.

Since 2006, the integral effect test facility, ATLAS (Advanced Thermal-Hydraulic Test Loop for 

Accident Simulation) has been operated by Korea Atomic Energy Research Institute (KAERI), 

which was constructed to simulate the reactor coolant system (RCS) behavior during transients 

in its reference designs, OPR-1000 and the APR-1400. After a series of the direct vessel 

injection (DVI) line break tests in the ATLAS was complete for four break sizes, 5%, 25%, 50%, 

and 100%, KAERI proposed the first domestic standard problem (DSP-1) in 2009, aiming at

assessing the capability of the existing codes for 100% DVI line break among the code users in 

Korea. At that time the TRACE code has not been used by any users. Just after termination of 

the DSP-1 program, KINS has reassessed the problem, using TRACE V5.0 patch2, with much 

information on the test specification provided by KAERI.

This report decribes the TRACE code assessment for 100% DVI line break in the ATLAS facility. 

According to a coordinated frame, NUSTEP for cooperation among Korean Utilities, KINS, 

research institutes, this report has been reviewed as the contribution by KINS and KAERI.

1

1. INTRODUCTION

A small break loss of coolant accident (SBLOCA) is characterized by relatively slow reactor 

coolant system (RCS) depressurization rates and by relatively slow mass loss from the RCS, 

compared to the design basis accident of large break loss of coolant accident (LBLOCA) [1]. 

Unlike the LBLOCA, the sequence of events following a small break in a LWR can evolve in a 

variety of ways. Operator actions, reactor design, ECCS set points, break size and location will 

have a bearing on how the SBLOCA scenario unfolds. Another principal difference is the 

domination of gravity effects in small breaks versus inertial effects in the large breaks [2]. In 

order to predict the thermal-hydraulic characteristics during SBLOCA adequately, code must 

have sufficient modeling capabilities to take these factors into account. 

The DVI-adopted plants treat a DVI line break as another spectrum among the SBLOCAs in 

their safety analysis because the DVI nozzle directly attached to the reactor vessel is vulnerable 

to a postulated break from a safety viewpoint. The thermal hydraulic phenomena in the RPV 

down-comer are expected to be different from the cold leg injection (CLI) mode during the 

postulated design basis accidents. In the event of a DVI line break, the vapor generated in the 

core is introduced to the reactor pressure vessel (RPV) downcomer through the hot legs, the 

steam generators and the cold legs. Then the vapor should pass through the upper part of the 

RPV downcomer to be discharged through the broken DVI nozzle. Therefore, the behavior of 

the two-phase flow in the upper annulus downcomer is expected to be complicated and relevant 

models need to be implemented into the safety analysis codes in order to predict these thermal 

hydraulic phenomena correctly. So far there is not enough integral effect test data for the DVI 

line breaks which can demonstrate the progression of the DVI line break accident realistically 

and can be used for an assessment and improvement of safety analysis codes.

KAERI has been operating an integral effect test facility, ATLAS (Advanced Thermal-Hydraulic 

Test Loop for Accident Simulation) for accident simulations for the OPR-1000 and the APR-1400 

[3]. After the ATLAS was modified to have a configuration for simulating the DVI line break 

accidents of the APR1400 at the beginning of 2008, sensitivity tests for different DVI line break 

sizes were performed by KAERI [3]. The Integral effect database for four break sizes were 

established; 5%, 25%, 50%, and 100%. The ATLAS has been used to provide the unique test 

data for the 2 (hot legs) x 4 (cold legs) reactor coolant system with direct vessel injection (DVI)

of emergency coolant. The 100% DVI line break test was selected for the first domestic 

standard problem (DSP-01) in 2009 to enhance the understanding on the behavior of nuclear 

reactor systems with the DVI and to assess existing thermal-hydraulic analysis codes such as 

MARS, RELAP and so on. The TRACE code has not been used to simulate a DVI line break in 

the DSP-01 program. 

The TRACE code is the thermal-hydraulic system code and has been developed by USNRC for 

a realistic analysis of thermal-hydraulics transients in pressurized water reactors [4]. TRACE 

has been designed to perform best-estimate analyses of loss-of-coolant accidents (LOCAs), 

operational transients, and other accident scenarios in pressurized light-water reactors (PWRs) 

and boiling light-water reactors (BWRs). It can also model phenomena occurring in experimental 

facilities designed to simulate transients in reactor systems. Models used include 

multidimensional two-phase flow, nonequilibrium thermo-dynamics, generalized heat transfer, 

reflood, level tracking, and reactor kinetics.
2

In this study, the calculations using TRACE V5.0 patch2 code released in the beginning of 2011 

and comparison with experimental data were conducted for 100% DVI line break test of the 

ATLAS which is the first domestic standard problem (DSP-01) to assess TRACE code capability 

to simulate the transient thermal-hydraulic behavior for the DVI line break.
3

2. ATLAS FACILITY AND TEST DESCRIPTION

2.1 Overview of the ATLAS

The ATLAS is a large-scale thermal-hydraulic integral effect test (IET) facility for advanced 

pressurized water reactors (PWRs), APR-1400 and OPR-1000. It can simulate a wide variety of 

accident and transient conditions including large and small break LOCAs. 

The ATLAS has the following characteristics: (a) 1/2-height, 1/288-volume, full-pressure 

simulation of the APR1400; (b) geometrical similarity with the APR1400, including 2 (hot legs) x 

4 (cold legs) reactor coolant loops, DVI of emergency core cooling water, integrated annular 

downcomer, etc.; (c) incorporation of specific design characteristics of the 1000-MW (electric) 

class OPR1000 such as a cold-leg injection and the low-pressure injection pumps, (d) a 

maximum 10% of the scaled nominal core power, and (e) simulation capability of broad 

scenarios, including the reflood phase of the large-break LOCA, small-break LOCA scenarios 

including the DVI line breaks, steam generator tube rupture, MSLB, mid-loop operation, etc [5]. 

Scientific design of the ATLAS was accomplished from the viewpoints of both a global and local 

scaling based on Ishii et al.[6]’s three-level scaling methodology. 

The ATLAS consists of a primary system, a secondary system, a safety injection system, a 

break simulating system, a containment simulating system, and auxiliary systems as shown in 

Figure 2.1 and 2.2. The primary system includes a reactor pressure vessel, 2 hot legs, 4 cold 

legs, a pressurizer, 4 reactor coolant pumps (RCPs), and 2 steam generators (SGs). The 

arrangement of the primary loop of ATLAS including DVI lines is shown in Figure 2.3. When a 

starting signal is generated to simulate a DVI line break of the ATLAS, a quick-opening valve in 

Figure 2.4 is fully opened to discharge the RCS inventories from the RPV into the containment 

simulator through a broken DVI nozzle. The secondary system of the ATLAS is simplified to be 

of a circulating loop-type. The steam generated at steam generators is condensed in a direct 

condenser tank and the condensed feedwater is again injected to the steam generators. Most of 

the safety injection features of the APR1400 and the OPR1000 are incorporated into the safety 

injection system of the ATLAS. The safety injection system of the ATLAS consists of four safety 

injection tanks (SITs), two high pressure safety injection pumps (SIPs) which can simulate 

safety injection and long-term cooling, a charging pump for charging auxiliary spray, and a 

shutdown cooling pump and a shutdown heat exchanger for low pressure safety injection, 

shutdown cooling operation and recirculation operation. The break simulation system consists of 

several break simulating lines such as large break LOCA (LBLOCA), direct vessel injection (DVI) 

line break LOCA, SBLOCA, steam generator tube rupture (SGTR), main steam line break 

(MSLB) and feedwater line break (FLB), etc. Each break simulating line consists of a quick 

opening valve, a break nozzle and instruments as shown in Figure 2.4. It is precisely 

manufactured to have a scaled break flow through it in the case of LOCA tests. The containment 

simulating system of the ATLAS has a function of collecting the break flow and maintaining a 

specified back-pressure in order to simulate containment as shown in Figure 2.5. Besides, the 

ATLAS has some auxiliary systems such as a makeup system, a component cooling system, a 

nitrogen/air/steam supply system, a vacuum system, and a heat tracing system. Secondary and 

auxiliary systems are designed as simply as possible since the main focus of the IET using the 

ATLAS will be on the simulation of primary-system transient and accidents, except for a MSLB 

and a FLB. Figure 2.6 shows a photograph of the front view of the ATLAS facility. The 

information on the ATLAS and some test results can be found in the literatures [7,8,9,10,11].
4

Figure 2.1 Schematics of the ATLAS Facility
5

Figure 2.2 Isometric Configuration of the ATLAS Facility
6

Figure 2.3 Arrangement of the Primary Loop of the ATLAS
7

Figure 2.4 Configuration of the Break Simulation System for the DVI Line Break Tests
8

Figure 2.5 Schematics of the Containment Simulation System
9

Figure 2.6 Photograph of ATLAS

2.2 Test Procedure

The 100% DVI line break of the ATLAS was conducted according to the following experimental 

procedure [11]. Basically, the experimental conditions for the present tests were determined by a 

pre-calculation with the best-estimate thermal hydraulic code, MARS3.1. First of all, a transient 

calculation was performed for the DVI line break of the prototypic plant, APR1400 to obtain the 
10

reference initial and boundary conditions. The safety injection system of the APR1400 has 4 

mechanically separated hydraulic trains. They are also electrically separated by 2 divisions, 

implying that each emergency diesel powers 2 hydraulic trains. The pre-calculation was 

conducted with the assumption of loss of off-site power simultaneously with the break and the 

worst single failure as a loss of a diesel generator, resulting in the minimum safety injection flow 

to the core. Furthermore, the safety injection flow to the broken DVI-4 nozzle was not credited. 

Therefore, the safety injection flow by the safety injection pump (SIP) was injected only through 

the DVI-2 nozzle opposite to the broken DVI-4 nozzle.

Table 2.1 Comparison of the Sequence of a 100% DVI Line Break

Events APR1400

(time, sec)

ATLAS

(time, sec)

Break open 0 0

Low pressurizer pressure 

trip (LPP) 20.9 If pressurizer pressure < 10.72 MPa

Pressurizer heater trip LPP+0.0 sec LPP+0.0 sec

Delay time is reduced 

by a square root of 2

Reactor scram & RCP trip LPP+0.5 sec LPP+0.35 sec

Turbine isolation LPP+0.1 sec LPP+0.07 sec

Main feedwater isolation LPP+10.0 sec LPP+7.07 sec

Safety injection pump 

start LPP+40.0 sec LPP+28.28 sec

Low upper downcomer 

pressure trip (LUDP) LUDP LUDP If downcomer pressure < 4.03 MPa

Safety injection tank (SIT) 

start LUDP+0.0 sec LUDP+0.0 sec

Low flow turndown of the 

SIT

If water level of the SIT 

is less than the 

specified set point

For the safety injection flow by the 4safety injection tanks (SIT), 3 SITs except for the SIT 

connected to the broken DVI-4 nozzle were available to provide the safety injection flow into the 

core. As for the core power, a conservative 1973 ANS decay heat curve with a 1.2 multiplication 

factor was used in the transient calculation. In the DVI line break, the containment backpressure does not affect the progression of this transient, because a choking condition is 

maintained throughout this transient. Therefore, the containment back-pressure was not an 

important control parameter in the present test.

Based on the calculated sequence of events of the DVI line break for the APR1400, the initial

and boundary conditions for the present integral effect test were determined. The delay time

required for an initiation of the safety injection systems such as the safety injection pump and

the safety injection tank were reduced by a square root of 2 according the scaling law of the

ATLAS. Also, the delay times for an isolation of the secondary feedwater or steam supply

systems were also scaled down by the scaling ratio of the ATLAS [5]. The detailed sequence of

events applied to the present test is summarized in Table 2.1.
11

The 100% DVI line break of the ATLAS was performed at the same pressure as the prototypic 

plant, APR1400. The temperature distribution along the primary loop was also preserved. The 

primary inventory was heated with core heaters to its specified steady state condition and was

pressurized by a pressurizer until the primary system reached a steady state condition. During

the primary heat-up process, the secondary system was also heated up to a specified target hot

condition by controlling the heat removal rate from the primary system. At a steady state

condition, the core power generated by electrical heaters was balanced by the energy removed

by the secondary system. The obtained steady state condition was maintained constant to

stabilize the system behavior of the ATLAS for more than 10 minutes.

During the heat-up process, several crucial components influencing the boundary conditions of 

the test were controlled by operators. The 4 core bypass flow control valves were controlled to 

have the predetermined stem positions to have a scaled core bypass flow rate. The initial water 

levels and pressures of the safety injection tanks were controlled to have these specified values. 

The refueling water storage tank (RWT) was filled with water to its initial level of 50% and the 

water inventory was electrically heated to its pre-determined temperature of 50o

C, and then the 

water was circulated through the injection line so as to preheat the line up to the same 

temperature as the water.

Subsequent to the heat-up process, several initialization procedures were taken to obtain the 

required initial and boundary conditions for the DVI line break. The primary coolant flow rate 

was reduced to 8% of the scaled value to have the same temperature distribution along the 

primary loop. Operation experience showed that the required 8% of the primary coolant flow 

rate was achieved in a natural circulation condition. The containment pressure control valve was 

opened fully to simulate an atmospheric pressure during the present test. All connecting pipe 

lines and three-way valves of the containment simulating system were aligned to measure the 

break flow rate separately.

After a steady state condition in the whole ATLAS system was maintained for more than 10

minutes, the transient test was commenced. First of all, data logging was initiated to log all

measurement points in a steady state condition. After the initial data logging was completed for 

about 200 seconds, the DVI line break test was initiated by opening a quick-opening break valve, 

OV-BS-03 at the break spool piece. A DVI line break was simulated by installing a break spool 

piece at one of the DVI nozzles. The configuration of the break spool piece is shown in Figure 

2.4. It consists of a quick opening valve, a break nozzle, a case holding the break nozzle, and a 

few instruments. A pressure transducer and 2 thermocouples were installed both upstream and 

downstream of the break nozzle. Detailed geometry of the break nozzle for the present DVI line 

break tests is shown in Figure 3.4. The break nozzle was installed vertically downward at the 

discharge line of the DVI nozzle. The quick opening valve was opened within 0.5 seconds by 

operators when the test was initiated. The break flow was discharged to the containment 

simulating system.
12

When the pressurizer pressure reached a specified pressure of 10.72MPa, the low pressurizer 

pressure (LPP) signal was automatically generated by embedded control logics. The heaters of 

the pressurizer and all tracing heaters in the primary system were tripped at the same time of 

the LPP signal. The RCP was automatically tripped with a time delay of 0.35 seconds after the 

LPP signal. The main steam and the main feed water line were isolated with a time delay of 0.07

seconds and 7.07 seconds after the LPP signal, respectively. The isolation of the secondary 

system requires a simultaneous actuation of several valves in the pipe line. It was done by the 

programmed control logics without operator intervention. Operation of the SI pump was 

triggered by the LPP signal with a time delay of 28.3 seconds. The initiation of the SI pump 

requires an alignment of the valves located in the supply line. This alignment was also

completed automatically by the control logic without any time delay.

When the downcomer pressure of the reactor vessel became lower than the specified pressure 

of 4.03MPa, the SIT started to deliver the high SI flow to the reactor vessel through the three 

DVI nozzles by fully opening the flow control valve. When the water level of the SIT reached a 

specified set point, the stem of the flow control valve was lowered to a specified position to 

supply a required low injection flow rate. When the water level of the SIT was lowered to a 

specified empty set point, the flow control valve was fully closed for the nitrogen gas, not 

injected into the reactor vessel. The transient was terminated with the end of the data logging 

when it was judged by operators that all the major phenomena have already taken place.
13

3. MODELING INFORMATION

TRACE code has been developed as the unified code for the reactor thermal hydraulic analyses 

in USNRC [4]. The Symbolic Nuclear Analysis Package (SNAP) is a suite of integrated 

applications designed to simplify the process of performing engineering analysis and it provides 

a flexible framework for creating and editing input for engineering analysis codes as well as 

extensive functionality for submitting, monitoring, and interacting with the codes [12]. The SNAP 

currently can support most analysis codes developed in USNRC, for example the CONTAIN, 

COBRA, FRAPCON-3, MELCOR, PARCS, RELAP5 and TRACE codes. In this study, the Model 

Editor which is one of SNAP client applications used to simulate the DVI line break test of 

ATLAS with TRACE codes. 

As mentioned previously, one of the objectives of the ATLAS development was to understand 

the complicated multi-dimensional phenomena like the downcomer boiling and the ECC water 

bypass which were important safety issues of the APR1400. The safety analysis codes such as 

RELAP5 have some limitation in order to predict these thermal hydraulic phenomena correctly 

since it is a kind of one-dimensional based codes. However, the TRACE code has the capability 

to simulate a 3D Cartesian- and/or cylindrical-geometry flow by using 3D components named 

VESSEL [4]. This is important in the prediction of the ECC bypass and the steam-water 

interaction in the downcomer during a LOCA. In this study, the reactor vessel including the core 

and the downcomer was modeled in 3 dimensions using the VESSEL component and flows 

within a coolant loop of the primary and secondary sides were modeled in one dimension using 

main hydraulic components such as PIPE and TEE in TRACE codes. The hot leg, cold leg, 

steam generator, steam line, and reactor coolant pump (RCP) were included in the primary and 

secondary systems. Also, the emergency core cooling system (ECCS) like the safety injection 

tank (SIT) and the safety injection pump (SIP) and the containment which is the boundary 

condition of LOCA were modeled. In order to simulate the heater rod in ATLAS, 2 heat 

structures, one was the average heaters and the other was the hot heater, were implemented in 

the each volume of core. Main control logics were modeled by the control block in TRACE. 

3.1 Main Loop Modeling

The APR1400 is the prototype of ATLAS basically and it has 2 (hot legs) x 4 (cold legs) reactor 

coolant loops. Therefore, the composition of main components and the connection between the 

components in ATLAS is almost equal to those in APR1400. The pressurizer is connected to 2-

side loop of primary system as shown in Figure 3.1. Each loop includes a hot leg, a steam 

generator, 2 reactor coolant pumps (RCPs), and 2 cold legs. Also, all passive heat structures 

including U-tubes of steam generator in each loop are modeled in this study. 

○ The hot leg is connected from the VESSEL component modeled as the core to the steam 

generator. The hot leg consists of 4 volumes which are modeled with 3 PIPE components. 

The surge line of pressurizer is connected to the hot leg as the cross flow junction and the 

hot leg is horizontally connected to the VESSEL component. 
14

○ The inlet/outlet plenum of SG is modeled as the PIPE component and is connected from 

the hot/cold leg to the U-tube. The U-tubes (V-340/V-440) in SG consist of 12 volumes in 

the PIPE component as shown in Figure 3.1. The riser (V-650/V-750) in the secondary 

side of SG is modeled as the PIPE component with 5 volumes and 4 volumes of riser 

exchange the heat with those in the primary side. The separator (V-660/V-770) in SG is 

modeled as the SEPARATORS component. 

○ The pump suction leg is modeled as the PIPE component by 5 volumes and is connected 

from the SG to the RCP. The angle of flow path in those properly considers the 

experimental data. Also, in order to predict the loop seal clearing (LSC) better, the height 

of intermediate legs is modeled in consideration of the location of instrument in 

experiments. 

○ The reactor coolant pump (RCP) is modeled as the PUMP component. The PUMP 

component in TRACE describes the interaction of the fluid with a centrifugal pump [4]. 

Since the PUMP component has the generic pump model including heat, torque, speed 

model, pump-curve data, etc., the pump behaviors like the pressure differential across the 

pump impeller and the pump impeller’s angular velocity for steady-state and transient 

conditions can be calculated by properly specifying the initial conditions. 

○ The cold leg is connected from the reactor vessel and the RCP and is modeled as 4 PIPE 

components. Also, it is horizontally connected to the VESSEL component.

○ The pressurizer is modeled by the PIPE component with 10 volumes. The empty time of 

pressurizer can influence the fuel temperature behavior due to the core rewetting which 

the core is quenched by the water in the pressurizer as it flows into the upper plenum 

through the hot leg. In the steady-state condition, the FILL component connected to the 

upper region of pressurizer is used to set the boundary condition of pressure.
15

Figure 3.1 ATLAS Nodalization for TRACE
16

3.2 Reactor Vessel Modeling

The modeling of reactor vessel is performed by the VESSEL component and it includes the 

downcomer, upper and lower plenum, upper head, core, etc. The reactor vessel of ATLAS 

consists of the lower plenum, active core, upper plenum and upper head along the axial 

direction as shown in Figure 3.2. It is also made up of the core region and the downcomer along 

the radial direction. All heat structures for the reactor vessel including the heater rods to 

simulate the fuel assembly are modeled in this study. The detailed information on the TRACE 

modeling for ATLAS can be found in the literature [13].

○ The downcomer is the outermost region of the VESSEL component and consists of 6 

volumes along the azimuthal direction and 26 volumes along the axial direction. In the 

calculations by using the one-dimensional codes like RELAP5, the downcomer is 

generally modeled as 2 ~ 6 separated PIPE components and the single junctions are used

to connect azimuthally between the neighboring PIPE components to simulate the 

complicated multi-dimensional behaviors in the downcomer. However, this may have the 

limitation to correctly predict these phenomena. In this study, the reactor vessel is 

modeled as the cylindrical multi-dimensional component and it can produce the better 

results for the multi-dimensional behaviors. The DVI line is modeled by considering the 

injection position of design data in ATLAS. 

○ The lower plenum is composed of 2 radial, 6 azimuthal and 5 axial volumes and the 

boundary of axial volumes agrees with the geometric boundary of ATLAS. The lower 

second volume contains the flow skirt in which the flow exchanges occur between the 

downcomer and the core. 

○ The upper plenum is located between the fuel alignment plate (FAP) and the upper guide 

structure support plate (UGSSP) and it is connected from the core exit to the hot leg and 

the upper head region in the vessel. The upper plenum consists of 2 radial, 6 azimuthal 

and 3 axial volumes and the middle of axial volumes is connected to the hot leg of each 

loop. The upper head is made up of 2 radial, 6 azimuthal and 5 axial volumes as shown in 

Figure 3.2 and the bypass flow between the downcomer and the upper head is also 

modeled. 

○ The active core consists of 2 radial, 6 azimuthal and 12 axial volumes and 12 flow 

channels is modeled at the specific axial level. The TRACE code has a tool to model the 

heat structure like the general thermal hydraulic codes such as RELAP5. The HTSTR 

component in TRACE evaluates the dynamics of conduction, convection and gap-gas heat 

transfer. Also, all fluid components, for example PIPE, TEE, etc. which include input for a 

pipe wall, internally have HTSTR components that provide for the simulation of the heat 

transfer from the pipe wall to the fluid [4]. Therefore, all fluid components including 

VESSEL and the heater rod in ATLAS is modeled by using the HTSTR component. There 

are 396 heater rods in the core of ATLAS. The two HTSTR components are modeled in 

each volume of 12 flow channels. One is to simulate the average heat structure to 

consider the number of heater rods and guide tubes in each volume as shown in Figure 

3.3. The other is to model the hot heat structure which has 1.4 times heat power as the 

average value of one heater rod. Table 3.1 shows the applied decay power level with the 

variation of time.
17

○ Several bypass flow paths in the reactor vessel are also modeled in this study. It is 

important to predict the bypass flow correctly in the steady-state calculation since the 

modeling of bypass flow may influence the transient results. From the reference 

calculation by the MARS code and the design of ATLAS, 4 bypass flow paths are 

considered as follows. 

- The bypass flow between the downcomer and the upper head

- The bypass flow from hot leg nozzle gap to the downcomer

- The bypass flow through the control element assembly (CEA) guide tubes

- The bypass flow through the core guide tubes
18

Table 3.1 Applied Decay Power Level with the Variation of Time

Time after break (sec) Total Avg. Rod Power (kW) Total Hot Rod Power (kW) Factor

0.0 1499.022 67.416 1.000

24.1 1499.022 67.416 1.000

25.5 1442.808 64.887 0.9625

26.9 1405.333 63.202 0.9375

28.3 1349.119 60.674 0.9000

35.4 1199.217 53.932 0.8000

42.5 1105.528 49.719 0.7375

49.5 1011.839 45.505 0.675

56.6 955.626 42.977 0.6375

63.7 918.150 41.292 0.6125

70.7 899.413 40.449 0.6000

84.9 843.199 37.921 0.5625

99.0 805.724 36.236 0.5375

113.2 768.248 34.550 0.5125

127.3 749.511 33.708 0.5000

141.4 730.773 32.865 0.4875

212.2 637.084 28.651 0.4250

282.9 599.608 26.966 0.4000

353.6 562.133 25.281 0.3750

424.3 543.395 24.438 0.3625

495.0 524.657 23.595 0.3500

565.7 513.415 23.089 0.3425

636.4 505.919 22.752 0.3375

707.1 490.929 22.078 0.3275

1000.0 468.444 21.067 0.3125

3510.0 312.546 14.056 0.2085

7046.0 245.090 11.022 0.1635
19

Figure 3.2 Reactor Vessel Modeling of ATLAS

Figure 3.3 Heat Structure Modeling for ATLAS Core
20

3.3 Emergency Core Cooling System (ECCS) & Break Modeling

The reference plant of the ATLAS is the APR1400 as mentioned before. The APR1400 adopts 

new safety injection features such as four mechanically independent safety injection systems, a 

DVI system, fluidic devices in each SIT, and an elimination of a low pressure safety injection 

system. Each of the four SITs of APR1400 has a fluidic device which passively controls the 

discharge flow rate into the reactor coolant system. In the APR1400, a high flow condition is 

changed to a low flow condition due to a fluidic device during operation of the SIT. Also, the 

APR1400 adopts the in-containment refueling water storage tank (IRWST) and it results in the 

improvement of ECCS reliability by skipping the action to switchover to the recirculation mode. 

The ATLAS incorporates all safety injection design features of OPR1000 and APR1400. The 

emergency core cooling system in the APR1400 is shown in Figure 3.4.

The 100% DVI line break experiment of the ATLAS was performed with the assumption of loss 

of off-site power simultaneously with the break and the worst single failure as a loss of a diesel 

generator, resulting in the minimum safety injection flow to the core. Therefore, the safety 

injection flow through the safety injection tank (SIT) and the safety injection pump (SIP) to the 

broken DVI-4 nozzle was not credited. Thus, the safety injection flow by the safety injection 

pump (SIP) was injected only through the DVI-2 nozzle opposite to the broken DVI-4 nozzle.

The safety water in the safety injection tank was injected through the 3 DVI nozzles except for 

the broken DVI-4 nozzle. 

In ATLAS modeling of TRACE, the safety injection tank is modeled as the PIPE component and 

the pipe type is selected as the accumulator. The major parameters (e.g. water level, volumetric 

flow, discharge volume) are automatically calculated in the accumulator type. Two valves are 

modeled at the downstream of SIT and then the start of SIT injection, switchover to low flow 

region due to a fluidic device and termination of SIT injection can be simulated by using proper 

control logics in the PIPE and VALVE components as shown in Figure 3.5. The switchover to 

low flow region can be modeled by changing the flow area of valve at a specific water level of 

SIT. The safety injection pump is modeled as the FILL component. The injected flow rate is 

adjusted according to the pressure of the downcomer volume injected and the injecting time is 

determined by the trip signal for safety injection due to the low pressurization of RCS. Table 3.2 

shows the mass flow rate of safety injection pump with respect to the downcomer pressure. The 

initial conditions (e.g., water temperature, water level, pressure, etc.) are determined as the 

experimental data. The double ended guillotine break of DVI-4 line is considered in transient 

calculations. 

The break system of the ATLAS is greatly simplified by using a single junction and BREAK 

component which is opened instantly at 199 seconds and the default critical flow model of 

TRACE, Ransom-Trapp model [14] is applied to the single junction [15]. The containment 

conditions are defined in a break table in the BREAK component. Table 3.3 shows the 

containment conditions with variation of time that is used in transient calculations. Figure 3.5 

shows the modeling of the downcomer and the ECCS and Figure 3.6 shows the plane figure for 

the connection between the reactor vessel and the ECCS.
21

Table 3.2 Mass Flow Rate of SIP with respect to the Downcomer Pressure

Downcomer Pressure (MPa) SIP Flow Rate (kg/sec)

0.10 0.323158

0.45 0.318567

0.79 0.313965

1.14 0.309055

1.48 0.304144

2.17 0.294039

2.86 0.283579

4.24 0.260868

5.62 0.235408

7.00 0.20624

8.38 0.170933

9.75 0.123685

10.10 0.107735

10.44 0.088879

11.13 0.0
22

Table 3.3 Containment Conditions with Variation of Time

Time after break (sec) Pressure (Pa) Quality

0.0 100930.0 1.0

3.0 155260.0 1.0

4.0 166990.0 1.0

6.0 161720.0 1.0

11.0 183760.0 1.0

13.0 195840.0 1.0

14.0 202100.0 1.0

17.0 215630.0 1.0

19.0 230260.0 1.0

23.0 239740.0 1.0

29.0 207910.0 1.0

35.0 194080.0 1.0

38.0 191060.0 1.0

42.0 188690.0 1.0

46.0 204010.0 1.0

56.0 181200.0 1.0

67.0 170200.0 1.0

79.0 168250.0 1.0

91.0 171990.0 1.0

167.0 140090.0 1.0

238.0 120720.0 1.0

261.0 116060.0 1.0

497.0 104370.0 1.0

1001.0 102870.0 1.0

2000.0 102490.0 1.0

1000000.0 100000.0 1.0
23

Figure 3.4 Emergency Core Cooling System (ECCS) in APR1400

Figure 3.5 TRACE Model for the Reactor Vessel and ECCS

CONTAINMENT

SIT

SIT

SIT

SIT

S/G RV S/G

SIP

SIP SIP

SIP

IRWST

RCP RCP

RCP RCP
24

Figure 3.6 Plane Figure for the Connection between the Reactor Vessel and the ECCS
25

4. STEADY STATE ANALYSIS

The steady state was determined by conducting a null transient calculation and the steady state 

calculation results which are the initial conditions of 100% DVI line break test are shown in Table 

4.1. 

The core power in the calculation was set to 1.566 MW which is the value subtracting the heat 

loss of 88 kW from the experimental core power, because adiabatic boundary conditions are 

applied in the system [16]. The hot leg and cold leg temperatures are little lower than in the 

experimental data. However, the calculated temperature difference between the hot leg and cold 

leg is ~ 33 K, and this value is almost same as that of experiment. Therefore, it is considered 

that energy conservation of experiment is well simulated. In the case of momentum 

conservation, the RCS flow rate in calculation is a little different from the experiment. If we do a 

simple energy balance calculation using the measured hot leg and cold leg temperatures, the 

RCS flow rate assuming 1.5666 MW core power is 2.05 kg/sec which is almost the same as the 

code predicted values. As mentioned previously, in experiment, the 8% of the primary coolant 

flow rate was achieved in a natural circulation condition without operating the reactor coolant 

pump (RCP). On contrary, forced circulation condition using the RCP should be maintained in 

calculation. It is considered that the difference of RCS flow rate between the experiment and 

calculation results from the type of flow in the system. 

As shown in Table 4.1, the differences between the calculated and measured values are 

acceptable for most of the primary/secondary system parameters. Some initial conditions of 

ECCS are considered conservatively in comparison to the experiment. The initial pressure of 

containment is exactly the same as the experimental data.
26

Table 4.1 Steady State Calculation Results

Parameter Experiment Calculation Remarks

Primary system

- Core power (MW) 1.644

(heat loss 88 kW)

1.654 

(w/ heat loss)

- PRZ pressure (MPa) 15.49 15.5

- Core inlet temp. (K) 564.0 564.4

- Core outlet temp. (K) 598.1 596.9

- Hot leg temp. (K) 599.2 596.2 Ave. of 2 HL

- Cold leg temp. (K) 566.2 562.7 Ave. of 4 CL

- RCS flow rate (kg/s) 2.2 2.05 Ave. of 4 CL

Secondary system

- Steam-dome pressure (MPa) 7.86 7.81 Ave. of 2 SG

- Steam temp. (K) 569.0 566.5 Ave. of 2 SG

- Feedwater (FW) temp. (K) 509.0 505.4

- FW flow rate to economizer (kg/s) 0.34 0.399 Ave. of 2 SG

- FW flow rate to downcomer (kg/s) ~0.0 0.044 Ave. of 2 SG

ECCS

- SIT pressure (MPa) 4.23 4.03 Ave. of 4 SIT

- SIT temp. (K) 323.7 322.1 Ave. of 4 SIT

- SIT level (m) 5.28 5.26 Ave. of 4 SIT

- RWT temp. (K) 316.7 322.0 Ave. of 2 SIP

Containment

- Pressure (MPa) 0.10 0.10
27

5. TRANSIENT ANALYSIS

In the experiment, after initial steady-state conditions were reached, the DVI line break test was 

initiated by opening a break simulation valve at 199 seconds, and transient calculation was also 

conducted by the opening trip in the FILL component at 199 seconds as the initial time. The 

predicted sequence of events is compared with that of the ATLAS experiment as shown in Table 

5.1. 

Table 5.1 Sequence of Events

Event ATLAS Experiment 

(sec)

TRACE Calculation 

(sec) Remarks

Steady state condition < 199.0 < 199.0

Break open 199.0 199.0

Low pressurizer pressure 

(LPP) 219.0 229.6 10.72 MPa

Turbine trip 219.1 229.1 LPP + 0.07 sec

Reactor trip by LPP 229.9 LPP + 0.35 sec

RCP trip 229.9 LPP + 0.35 sec

Decay power start 223.0 223.0

Main feedwater isolation 226.0 236.7 LPP + 7.07 sec

SIP injection signal 246.0 257.9 LPP + 28.28 sec

Max. PCT 290.0 -

Loop Seal Clearing (LSC) ~295/~297/~319/~320 ~321/~322/~322/~322

SIT injection start 431.0 426.0

When the pressurizer pressure decreases below 10.72 MPa, the reactor/RCP trip and main 

steam isolation occur logically by the low pressurizer pressure (LPP) signal. However, in the 

experiment, core power decay was initiated with time delay of 4.0 seconds after the break, and 

this condition was applied in calculation as shown in Table 3.1. The main feedwater is isolated 

with time delay of 7.07 seconds after the LPP signal, logically. SIP was injected with time delay 

of 28.28 seconds after LPP, and SIT injection was initiated when the downcomer pressure was 

reduced below 4.03 MPa. After the calculation was started at 199 seconds, the pressurizer 

pressure decreased and reached to a low pressurizer set pressure at ~ 229.6 sec. The LPP 

time of TRACE is ~ 10 seconds later than that of the ATLAS experiment. The decay of heater 

power in TRACE is modeled to start at same time (~ 223.0 seconds) with the experiment. The 

safety injection pump (SIP) injection is predicted ~ 10.0 seconds later than that of experiment 

due to the delay LPP time. The safety injection tank (SIT) starts to inject ~ 426 seconds 

because of rapid reduction of pressure after the loop seal clearing (LSC).
28

In the general, for small or intermediate break, the core is uncovered and the cladding 

temperature increases. At this calculation, we can identify the following specific scenarios of 

SBLOCA for the plant with a U-tube steam generator.

○ Blowdown phase : After break occurs abruptly, the primary pressure is reduced rapidly. As 

soon as it reaches the set point of low pressurizer pressure, the control rod drops and the 

reactor scrams. Also, the reactor coolant pump is tripped simultaneously. After the safety 

injection signal happens, the emergency cooling water is injected into the RCS. During the 

blowdown phase, most of the RCS is filled with the liquid phase and the break flow is 

discharged as the subcooled or saturated liquid phase. Consequently, the reduction of 

primary pressure slows down at the extent of slightly higher value than the secondary 

pressure. At this time, the primary and secondary pressures are determined according to 

the core decay heat, the heat removal of SG, the set point of safety valves of SG, etc. 

Figure 5.1 shows the typical blowdown phase in this calculation.

○ Natural circulation : Lastly in the blowdown phase, the RCS reaches the quasi-equilibrium 

condition. This can continue for several hundred seconds according to the break size. 

During this period, the loop seal is full in the liquid phase. The coolant is discharged 

continuously and the vapor begins to be generated from the upper U-tube of steam 

generator. The vapor generation is propagated into the upper plenum through the upper 

head and the phase separation is accelerated. The decay heat in the core is removed 

from the discharged flow and the steam generators. The steam in U-tubes does not find 

the effective path since the loop seal is filled with the liquid and the steam generated in 

core is trapped in the RCS. Therefore, the liquid with low quality is discharged from the 

break. Figure 5.2 shows the typical natural circulation phase in this calculation.

○ Loop Seal Clearing (LSC) : As the water level of cold leg side in SG approaches the upper 

suction line (loop seal) of RCS, the trapped steam in the RCS can be released from the 

break position. This is called loop seal clearing (LSC). Also, the loop seal clearing can 

occur abruptly as soon as the pressure in upper region of vessel becomes larger than the 

hydrostatic pressure of liquid phase in the loop seals. The loop seal clearing can happen 

in the partial or entire loops. The discharged flow is changed from the low quality liquid to 

the steam. At just before loop seal clearing, the core can be uncovered due to the 

decrease of the collapsed core level according to the vessel pressurization. After the loop 

seal clearing, the pressure unbalance is resolved and the collapsed water level is covered 

gradually. Figure 5.3 shows the RCS conditions after the loop seal clearing.
29

Figure 5.1 Blowdown Phase (at 12.5 seconds after break)

Figure 5.2 Natural Circulation (at 102 seconds after break)
30

Figure 5.3 RCS Conditions after Loop Seal Clearing (at 140 seconds after break)
31

The core power is shown in Figure 5.4. The deviations of the core power was regarded as 

acceptable because we subtracted 5.6% heat loss, which was not measured but estimated from 

a separate effect test, from the measured core power. [16]. Therefore, the same core power as 

the experiment was used in this study. 

Figure 5.4 Core power

The pressurizer pressure is shown in Figure 5.5. Generally, as soon as the initiation of break 

occurs at the DVI line, primary pressure rapidly decreases due to the sudden coolant loss and 

the coolant in the RCS remains in the liquid phase during this blowdown period. As time goes by, 

the coolant becomes steam by the flashing and the boiling occurs in the core, and steam can be 

identified in the upper head, upper plenum, and hot legs. After the initial rapid depressurization 

ends by the flashing and boiling, the primary pressure reaches a plateau just above the 

saturation pressure of the secondary side between the SIP injection and the loop seal clearing 

After the loop seal clearing, the primary pressure begins to decrease below the secondary side, 

and continues to decrease as the break flow continues. 

The predicted pressurizer pressure agrees relatively well with the experimental data as shown 

in Figure 5.5; however, before the SIP starts, the predicted pressurizer pressure decreases 

slowly compared to the experimental results. This is the reason that the break flow is underestimated as shown in Figure 5.6. Also, the SIP injection and loop seal clearing occurs slightly 

later than the measured time due to the under-estimation of break flow. After the SIT injection 

starts, the predicted pressure reduces more rapidly than the experiment due to the overestimation of steam condensation in the vessel. 
32

Figure 5.5 Pressurizer Pressure

The predicted break flow rate is compared with the experiment in Figure 5.6. The extension 

version of the Ransom-Trapp model is used as a default two-phase choked-flow model in 

TRACE [15]. Therefore, the sensitivity studies for break flow were performed to find the optimal 

subcooled and two-phase multipliers in the choked-flow model. However, even with the 

sensitivity study, the predicted pressure reduces more slowly than the experiment and then the 

predicted break flow shows large difference from measured data before LSC as shown in Figure 

5.6. Since the pressure difference between RCS and containment is substantial, the critical flow 

condition is maintained during the calculation. After the instance of break, the predicted break 

flow rate grows rapidly like in the experiment, while the value is lower than in the experiment. 

Actually, measuring of break flow is difficult in the experiment. As shown in Figure 5.6, there are 

fluctuations before LSC, and it is considered that large uncertainties for measurement of break 

flow in early period of the transient make these fluctuations [8]. Therefore, it is not proper to say 

if the break flow before LSC is under-estimated or over-estimated, though a rough guess is 

possible from the comparison of the integrated break flow. 

Figure 5.7 shows the accumulated break mass, and before LSC, predicted break mass is 

slightly smaller than that in the experiment due to the under-estimation of the break flow. 

However, because the predicted LSC starts at a higher pressure, the break flow and mass are 

over-predicted between LSC and SIT injection. After LSC, the code under-estimates the mass 

discharged, since the predicted break flow rate is slightly smaller than in the experiment by 

around 420 seconds. As shown in Figure 5.5, the primary pressure suddenly decreases as cold 

SIT water is injected around 420 seconds, therefore, break flow rate decreases during this 

period. 
33

Figure 5.6 Break Flow Rate

Figure 5.7 Accumulated Break Mass
34

During a PWR SBLOCA, there are three distinct core heat-up potentials, and the first heat-up is 

caused by loop seal formation and the manometric core liquid depression.1 This heat-up is 

naturally mitigated by the occurring loop seal clearing. Therefore, its timing is very important to 

core water level and cladding temperature. In this case, liquid pools trapped in the RCP side 

intermediate leg are the loop seal, and act as plugs for steam flow from SG. 

Figure 5.8 and 5.9 show the core and the downcomer collapsed water level in the experiment 

and TRACE calculation. In general, the core collapsed water level is tightly coupled with the 

downcomer collapsed water level. In this study, the behavior of the core and downcomer 

collapsed level can be analyzed better by dividing it into that before the loop seal clearing at ~ 

320 seconds and that after the loop seal clearing. Before the loop seal clearing, the code could 

not predict well the behavior of the core collapsed level. After the break, both the measured and 

predicted core collapsed water level decrease for the initial ~ 10 seconds due to the coolant loss 

through the break but the depth of reduction is under-predicted. In the experiment, the collapsed 

water level increases more than the initial level for ~ 15 seconds after the initial ~ 10 seconds, 

and it is considered to be a sensor detection error. After beginning of the SIP injection, the 

injected cold water contracted the downcomer water level in the experiment. This water 

contraction and the build-up of pressure in the upper region of the vessel continuously reduced 

the core water level until ~ 290 seconds. However, in the TRACE calculation, the steam 

generated in the core moves to a cold leg through a hot leg and a steam generator. The moving 

steam is blocked by the non-saturated water in the loop seal of each loop and then the pressure 

of the upper plenum slightly increases until the loop seal clearing as shown in Figure 5.10. 

Therefore, during the period between the SIP injection and the LSC, the downcomer water level 

shows the stagnation like the pressure behavior because of the interaction between the injected 

SI water in the downcomer and the build-up pressure in the core. The stagnation of the 

downcomer water level cannot reduce the core water level to the extent of the measured value 

and the pressure build-up is also not strong enough to shrink the core water level. Therefore, 

the predicted water level isn’t changed largely like the measured value and maintained as 2.0 m 

~ 2.25 m. 

At the time of loop seal clearing (~ 290 seconds), the measured core water level dropped rapidly 

for ~ 15 seconds and then increased for ~ 7 seconds. The measured behavior of core water 

level around the loop seal clearing time can be regarded as reasonable because a typical loop 

seal clearing promotes the steam venting to the break and results in a rapid decrease of core 

water level. In the TRACE calculations, after the LSC, both the downcomer and core water level 

are reduced rapidly until the SIT injection starts. After the injection of SIT, the core water level is 

almost identical to the measured value but the downcomer water level is under-estimated 

because of the low SIT injection water and the large inventory in the intermediate leg. 

 

1 The second heat-up occurs following the core quench caused by loop seal clearing and is 

caused by a simple core boil-off. The third heat-up can occur following depletion of the 

accumulator tanks and before LPIS injection begins.
35

Figure 5.8 Core & Downcomer Collapsed Water Level

Figure 5.9 Core & Downcomer Collapsed Water Level in the Initial Period
36

As mentioned before, the accurate prediction of loop seal clearing is very significant in the 

SBLOCA analysis since it can influence the core water level and cladding temperature. The loop 

seal clearing occurs when the pressure in the upper region of vessel has enough force to 

overcome the hydrostatic pressure of liquid pools in the loop seals.

The collapsed water level of the RCP suction leg is shown in Figure 5.10. As shown in this 

figure, all 4 pump suction legs or loop seals are cleared completely and suddenly at ~ 300 

seconds in the experiment. The predicted water levels in the 4 pump suction legs also drop 

rapidly and all loop seals are cleared completely ~ 25 seconds later than the measured value 

due to the difference of depressurization behavior. It is known that the sequence and location of 

loop seal clearing are influenced by the location and size of the break. Figure 2.3 shows the 

arrangement of the primary loop of the ATLAS including DVI lines. As shown in this figure, the 

break occurs at the DVI-4 nozzle, and the DVI-3 nozzle is the closest the DVI nozzle from the 

break and the DVI-1 nozzle is the next closest. However, in the sensitivity study calculations for 

the location of the break, it is seen that the location and the number of loop seal clearing has no 

trends and LSC appears like chaotic phenomena, not depending on the flow resistance [11]. 

Therefore, comprehensive studies to investigate the mechanism of loop seal clearing should be 

performed. However, the timing of LSC is more important than the location and number of that 

in a safety point of view. 

Figure 5.10 Pump Suction Leg Collapsed Water Level

The predicted safety injection flow rate is compared with the experiment in Figure 5.11 and 5.12. 

For the safety injection pump (SIP) flow, code calculation result is in good agreement with the 

experiment results before the SIT injection. After the SIT injection, predicted values are slightly 

larger than the experimental data since the measured flow is reduced due to the interaction of 

SIT and SIP flows injected through the same DVI nozzle as shown in circle mark of Figure 5.11. 
37

For the SIT flow, injection starting time agrees well with that of the test, and the prediction 

shows much higher oscillations than in the experiment. We are not sure, but highly oscillating 

interfacial heat transfer in the downcomer after the SIT injection may result in some oscillations 

in the downcomer pressure and subsequent oscillations of SIT flows. Also, high oscillations of 

SIT flows can occur due to some numerical problem. Differently from the experiment, at some 

regions after ~ 1,000 seconds, SIT water is not injected. From previous studies, it is reported 

that if SIT is modeled by the PIPE component (not ACCUM component), SIT injection flow rate 

does not have any oscillations [17]. Therefore, models of the ACCUM component, especially 

numerical scheme should be improved to predict a more accurate SIT injection flow rate. As 

mentioned in Figure 5.8, the predicted SIT flow is under-estimated and this results in the underprediction of the downcomer water level after the SIT injection.

Figure 5.11 Safety Injection Pump (HPSI) Flowrate
38

Figure 5.12 Safety Injection Tank (SIT) Flowrate

Figure 5.13 shows the comparison of hot leg flows in the experiment and calculation. The code 

significantly over-predicts the hot leg flow before ~ 250 seconds. It is considered that this overprediction of hot leg flow results in the under-prediction of the core collapsed liquid level before 

~ 230 seconds. After ~ 250 seconds, a large amount of water is predicted to flow from the steam 

generators to the core for ~ 75 seconds as shown in the circle mark of Figure 5.13. However, in 

the experiment, the water continuously flows from the core to SG. This negative flow from the 

steam generators to the core results in over-prediction of the core water level in a period 

between SIP injection and LSC. As in a previous study [18], it is reported that if the CCFL 

(Count-Current Flow Limitation) option is applied in a FAP (Fuel Alignment Plate), the backward 

flow from SG to core disappears when RELAP5 and MARS are used. Therefore, in the 

sensitivity study of Chapter 6, we identify whether or not the CCFL model used at FAP can 

restrict the flow from the steam generators to the core. After the LSC, the increase of flow to hot 

leg can cause the reduction of the core water level. 

The predicted and measured cold leg mass flow rates are shown in Figure 5.14. At the 

beginning of the transient, the code shows a tendency to over-predict cold leg flows, but the 

difference between the experiment and actual calculation is not as large as the case of hot leg 

flows. At the time of loop seal clearing (~ 320 seconds), the mass flow of each loop increases 

rapidly in the calculation. 
39

Figure 5.13 Hot Leg Massflow Rate

Figure 5.14 Cold Leg Massflow Rate
40

Figure 5.15 shows the measured heater rod surface temperature and predicted one. The 

measured peak cladding temperature is 632 K at ~ 290 seconds while the peaking for predicted 

cladding temperature is not shown in this study. The predicted cladding temperature generally 

agrees well with the experiment results except in the peaking behavior. This peaking behavior of 

cladding temperature is directly coupled with the core collapsed water level. The measured core 

collapsed water level decreases up to ~ 1.1 m but the predicted water level is over-estimated 

between the SIP injection and the LSC as shown in Figure 5.9. The heater rod surface 

temperature has a similar trend of saturation temperature and primary pressure. Better 

predictions of discharging flow and core water level should be made first to predict peak 

cladding temperature more accurately 

Figure 5.15 Heater Rod Surface Temperature
41

6. SENSITIVITY STUDY

6.1 Increase of Break Flow

As we explained in the prior chapter, after initiating the DVI line break, the predicted pressurizer 

pressure decreases slowly compared to the experimental results until the SIP starts. This is the 

reason that the break flow is under-estimated as shown in Figure 5.3. Low discharged mass at 

break also results in the over-prediction of the downcomer and core water levels during this 

period. Therefore, the sensitivity study for the break mass is performed to identify the effect of 

depressurization. The break area is arbitrarily modified to simulate a similar accumulated break 

mass with experimental data as shown in Figure 6.1. 

The downcomer and core water level predicted in this sensitivity run are compared to those of 

the base case calculation in Figure 6.2 and 6.3, respectively. Shown in the figure, if we increase 

the break flow, the downcomer water level decreases more rapidly and the stagnation period of 

the downcomer water level becomes short since the loop seal clearing occurs earlier around ~ 

300 seconds compared to the base case. Before the loop seal clearing, the downcomer water 

level shows the better prediction results with the measured values. After break, the predicted 

core collapsed water level in this sensitivity run decreases initially ~ 10 seconds and the depth 

of reduction is closer to the measurement due to the assumption of more coolant loss. The 

predicted core water level generally agrees well with the experiment results until ~ 260 seconds 

except for the abnormal increment of measured value at ~ 225 seconds that is considered as 

the sensor detection error. However, in the sensitivity run, the core water level still cannot be 

reduced to the extent of the measured value around 290 seconds due to the stagnation of the 

downcomer water level and weak pressure build-up in the core. Also, except for the periods 

before LSC, the overall behavior for main parameters shows the worse prediction results 

compared to the base case. 

Generally, there are the Henry-Fauske model and the Ransom-Trapp model as the choked-flow 

model of RELAP5 code [19] but the Ransom-Trapp model is only the two-phase choked-flow 

model in the TRACE code. The Henry-Fauske model has two input parameters that have to be 

given by the users. They are a discharge coefficient (CD) and a non-equilibrium factor (Neq). In 

order to find out the optimal CD and Neq, the sensitivity study is performed usually for the 

pressure behaviors. Both models in RELAP5 have been used widely in the real-plant and 

experiment calculations. Therefore, if the Henry-Fauske model would be implemented in the 

TRACE code, the diversity for the choked-flow model can be increased for a better prediction of 

break flow.
42

Figure 6.1 Accumulated Break Mass

Figure 6.2 Downcomer Collapsed Water Level
43

Figure 6.3 Core Collapsed Water Level

6.2 CCFL Model in the Upper Plenum

In the base case calculation, relatively large amount of water is predicted to flow from the steam 

generators to the core from ~ 250 seconds for ~ 75 seconds as shown in Figure 5.10 and it is 

believed that such a large liquid flow to the core is the main reason why the predicted core 

collapsed water level is not predicted to decrease much during the same period. The measured 

core water level reduced continuously during this period to reach its minimum at ~ 290 seconds 

(Figure 6.3). During this period, the steam generated in the core should flow out to the hot legs 

through the upper plenum. Therefore, there must be a steam/water counter-current flow in the 

upper plenum when the water flows from the hot legs to the core. At that time, the water flow 

from the upper plenum to the core can be restricted by the CCFL (Counter-Current Flow 

Limitation). 

In the base case, the CCFL model in the upper plenum is not considered at all. Therefore, we 

change the axial edge data of the upper inactive core volume (V2A20 in Figure 3.1) connecting 

the fuel alignement plate (FAP) to turn on the CCFL model. The Wallis model and Kutateladze

model are used by adjusting the ‘bankoff interpolation’ input.

If the CCFL model, as expected, restricted the liquid flow from the upper plenum to the core, the 

predicted core water level has to agree better with the measured value. In this study, despite 

using the CCFL model, the predicted core water level reduces a little more than that of the 

basecase and the significant decrease for the core water level isn’t shown from ~ 250 seconds 

to ~ 300 seconds. This shows the different results compared to that of the RELAP5 code. Figure 
44

6.5 shows the sensitivity results for the CCFL model of RELAP5/MOD3.3 which was performed 

in DSP2 [20]. Though the 6-inch cold leg break is analyzed in DSP2, we can find out that the 

overall thermal hydraulic behaviors are similar to those of the DVI line break. The Wallis model, 

which is the default of the CCFL model of RELAP5/MOD3.3, is applied to the upper core in 

Figure 6.5. In this case, since the CCFL model restricts the water flow to the core, the predicted 

core water level follows the measured level much better and it decrease continuously to reach 

its minimum value as shown in Figure 6.5. So, from more detailed studies, we need to identify if 

the CCFL model is operated properly or not in the upper core of the vessel component.

Figure 6.4 Effect of CCFL Model on the Core Collapsed Water Level
45

Figure 6.5 Effect of CCFL Model in RELAP5 on the Core Collapsed Water Level [Ref. 20]

47

7. CONCLUSIONS

Calculations using the TRACE V5.0 patch2 code were conducted for 100% DVI line break test 

of the ATLAS which is the first domestic standard problem (DSP-01). For the modeling of ATLAS 

facility, the reactor vessel was modeled in three-dimension, using the VESSEL component and 

flows within a coolant loop of the primary and secondary sides were modeled in one dimension 

using the main hydraulic components such as PIPE and TEE in the TRACE codes. Also, the 

emergency core cooling system, the heater rods and main control logics were properly modeled 

in order to consider the various conditions of this experiment. The steady state was determined 

by conducting a null transient calculation and the errors between the calculated and measured 

values are acceptable for most primary/secondary system parameters.

In the experiment, after the initial steady-state conditions were reached, the DVI line break test 

was initiated by opening a break simulation valve at 199 seconds, and transient calculation was 

also conducted by setting up 199 seconds as the initial time. The predicted sequence of events 

occurred some seconds later than that of experiment due to slow depressurization of the 

primary side after the break. Before the SIP starts, the predicted pressurizer pressure 

decreases slowly compared to the experimental results because of the under-estimation of 

break flow according to the characteristic of the choked flow model. Before the LSC, the code 

could not predict well the behavior of the downcomer and core collapsed water levels due to the 

under-prediction of the discharged mass and high flow rate returned to core from hot leg. After 

the LSC, the downcomer and core collapsed water levels drop rapidly and the code predicts 

relatively well the trend of the downcomer and core water level. In this study, all RCPs suction 

legs or loop seals are cleared completely and suddenly, and the code predicted the LSC ~ 25 

seconds later than the measured value due to the slow depressurization rate. The predicted 

cladding temperature generally agrees well with the experiment results except in the peaking 

behavior. This peak of cladding temperature is directly related to core collapsed water level dip 

before the LSC. 

The sensitivity studies were performed to identify 1) the effect of discharging mass for the 

depressurization and 2) the effect of flow restriction from hot leg for the core collapsed water 

level. If the discharging mass increases arbitrarily, the downcomer and core water level before 

SIP injection shows better prediction results than that of the base case. However, except for this, 

the overall behavior for the main parameters shows the bad prediction results. For the effect of 

flow restriction from hot leg using the CCFL model, the predicted core water level was not 

decreased significantly compared to the base case. The results were different in comparison 

with that of the RELAP5 code and we need to do further studies to better understand this 

difference. 

In conclusion, the TRACE code has good capabilities to simulate the 100% DVI line break test 

of the ATLAS. However, there are some discrepancies in quantitatively predicting the primary 

pressure, break flow, downcomer and core collapsed liquid level, and so on. Therefore, the 

TRACE code including the choked flow and CCFL models needs to be improved and more 

detailed modeling is needed to predict more accurate results.

49

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