NUREG/IA-0413 

International 

Agreement Report

Assessment of TRACE 5.0 

Against ROSA Test 3-1, 

Cold Leg SBLOCA

 

Prepared by:

S. Gallardo, V. Abella, G. Verdú and A. Querol

Universitat Politécnica de Valéncia

Camino Vera s/n

46022 Valencia, SPAIN

 

A. Calvo, NRC Project Manager

 

 

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

Manuscript Completed: March 2012 

Date Published: March 2012

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-0413 

International 

Agreement Report

Assessment of TRACE 5.0 

Against ROSA Test 3-1, 

Cold Leg SBLOCA

 

Prepared by:

S. Gallardo, V. Abella, G. Verdú and A. Querol

Universitat Politécnica de Valéncia

Camino Vera s/n

46022 Valencia, SPAIN

 

A. Calvo, NRC Project Manager

 

 

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

Manuscript Completed: March 2012 

Date Published: March 2012

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 purpose of this work is to overview the results provided by the simulation of a cold leg Small 

Break Loss-Of-Coolant Accident (SBLOCA) under the assumption of total failure of high pressure 

injection system in the Large Scale Test Facility (LSTF) via the thermal-hydraulic code TRACE5. 

The work is developed in the frame of OECD/NEA ROSA Project Test 3-1 (SB-CL-38 in JAEA). 

Test 3-1 simulated a PWR high-power natural circulation due to failure of scram during cold leg 

SBLOCA with a break size equivalent to 0.1% of the cold leg under assumption of total failure of 

High Injection System. 

A detailed model of the plant and the chronology of events, following these assumptions, has been 

developed with TRACE5. A comparison of the simulation of the test with the experimental results is 

provided throughout several graphs. An acceptable general behaviour is observed in the entire 

transient. In conclusion, this work represents a good contribution for assessment of the 

predictability of computer codes such as TRACE5.

v

FOREWORD 

Extensive knowledge and techniques have been produced and made available in the field of thermal-hydraulic 

responses during reactor transients and accidents, and major system computer codes have achieved a high 

degree of maturity through extensive qualification, assessment and validation processes. Best-estimate 

analysis methods are increasingly used in licensing, replacing the traditional conservative approaches. Such 

methods include an assessment of the uncertainty of their results that must be taken into account when the 

safety acceptance criteria for the licensing analysis are verified. 

Traditional agreements between the Nuclear Regulatory Commission of the United States of America 

(USNRC) and the Consejo de Seguridad Nuclear of Spain (CSN) in the area of nuclear safety research have 

given access to CSN to the NRC-developed best estimate thermalhydraulic codes RELAP5, TRAC-P, TRACB, and currently TRACE. These complex tools, suitable state-of-the-art application of current two-phase flow 

fluid mechanics techniques to light water nuclear power plants, allow a realistic representation and simulation 

of thermalhydraulic phenomena at normal and incidental operation of NPP. Owe to the huge required 

resources, qualification of these codes have been performed through international cooperation programs. 

USNRC CAMP program (Code Applications and Maintenance Program) represents the international 

framework for verification and validation of NRC TH codes, allowing to:

• Share experience on code errors and inadequacies, cooperating in resolution of deficiencies and 

maintaining a single, internationally recognized code version.

• Share user experience on code scaling, applicability, and uncertainty studies.

• Share a well documented code assessment data base.

• Share experience on full scale power plant safety-related analyses performed with codes (analyses of 

operating reactors, advanced light water reactors, transients, risk-dominant sequences, and accident 

management and operator procedures-related studies).

• Maintain and improve user expertise and guidelines for code applications.

Since 1984, when the first LOFT agreement was settled down, CSN has been promoting coordinated joint 

efforts with Spanish organizations, such as UNESA (the association of Spanish electric energy industry) as 

well as universities and engineering companies, in the aim of assimilating, applying, improving and helping the 

international community in the validation of these TH simulation codes1

, within different periods of the 

associated national programs (e.g., CAMP-España). As a result of these actions, there is currently in Spain a 

good collection of productive plant models as well as a good selection of national experts in the application of 

TH simulation tools, with adequate TH knowledge and suitable experience on their use.

Many experimental facilities have contributed to the today’s availability of a large thermal-hydraulic database 

(both separated and integral effect tests). However there is continued need for additional experimental work 

and code development and verification, in areas where no emphasis have been made along the past. On the 

basis of the SESAR/FAP2 reports “Nuclear Safety Research in OECD Countries: Major Facilities and 

Programmes at Risk” (SESAR/FAP, 2001) and its 2007 updated version “Support Facilities for Existing and 

Advanced Reactors (SFEAR) NEA/CSNI/R(2007)6”, CSNI is promoting since 2001 several collaborative 

international actions in the area of experimental TH research. These reports presented some findings and 

recommendations to the CSNI, to sustain an adequate level of research, identifying a number of experimental 

facilities and programmes of potential interest for present or future international collaboration within the safety 

community during the coming decade. 

CSN, as Spanish representative in CSNI, is involved in some of these research activities, helping in this 

international support of facilities and in the establishment of a large network of international collaborations. In 

 1 It’s worth to note the emphasis made in the application to actual NPP incidents.

2 SESAR/FAP is the Senior Group of Experts on Nuclear Safety Research Facilities and Programmes of NEA Committee on the 

Safety of Nuclear Installations (CSNI).
vi

the TH framework, most of these actions are either covering not enough investigated safety issues and 

phenomena (e.g., boron dilution, low power and shutdown conditions), or enlarging code validation and 

qualification data bases incorporating new information (e.g., multi-dimensional aspects, non-condensable gas 

effects). In particular, CSN is currently participating in the PKL and ROSA programmes.

The PKL is an important integral test facility operated by of AREVA-NP in Erlangen (Germany), and designed 

to investigate thermal-hydraulic response of a four-loop Siemens designed PWR. Experiments performed 

during the PKL/OECD program have been focused on the issues:

• Boron dilution events after small-break loss of coolant accidents.

• Loss of residual heat removal during mid-loop operation (both with closed and open reactor coolant 

system.

ROSA/LSTF of Japan Atomic Energy Research Institute (JAERI) is an integral test facility designed to simulate 

a 1100 MWe four-loop Westinghouse-type PWR, by two loops at full-height and 1/48 volumetric scaling to 

better simulate thermal-hydraulic responses in large-scale components. The ROSA/OECD project has 

investigated issues in thermal-hydraulics analyses relevant to water reactor safety, focusing on the verification 

of models and simulation methods for complex phenomena that can occur during reactor transients and 

accidents such as:

• Temperature stratification and coolant mixing during ECCS coolant injection

• Water hammer-like phenomena

• ATWS

• Natural circulation with super-heated steam

• Primary cooling through SG depressurization

• Pressure vessel upper-head and bottom break LOCA

This overall CSN involvement in different international TH programmes has outlined the scope of the new 

period of CAMP-España activities focused on:

• Analysis, simulation and investigation of specific safety aspects of PKL/OECD and ROSA/OECD 

experiments.

• Analysis of applicability and/or extension of the results and knowledge acquired in these projects to 

the safety, operation or availability of the Spanish nuclear power plants.

Both objectives are carried out by simulating experiments and plant application with the last available versions 

of NRC TH codes (RELAP5 and TRACE). A CAMP in-kind contribution is aimed as end result of both types of 

studies.

Development of these activities, technically and financially supported by CSN, is being carried out by 5 

different national research groups (Technical Universities of Madrid, Valencia and Cataluña). On the whole, 

CSN is seeking to assure and to maintain the capability of the national groups with experience in the thermal 

hydraulics analysis of accidents of the Spanish nuclear power plants.

_____________________________________

Francisco Fernández Moreno, Commissioner

Consejo de Seguridad Nuclear (CSN)
vii

CONTENTS

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

Page

FOREWORD................................................................................................................................. v

FIGURES....................................................................................................................................viii

TABLES....................................................................................................................................... ix

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

ACKNOWLEDGEMENTS...........................................................................................................xiii

ABBREVIATIONS ...................................................................................................................... xv

1 INTRODUCTION ................................................................................................................1-1

2 ROSA FACILITY DESCRIPTION.......................................................................................2-1

3 TRANSIENT DESCRIPTION..............................................................................................3-1

4 APPLIED METHOD: TRACE5 MODEL OF ROSA FACILITY...........................................4-1

5 RESULTS AND DISCUSSION...........................................................................................5-1

5.1 Steady-state...............................................................................................................5-1

5.2 Transient ....................................................................................................................5-2

5.3 System pressures ......................................................................................................5-3

5.4 Break..........................................................................................................................5-5

5.5 Primary loop mass flows ............................................................................................5-7

5.6 Vessel collapsed liquid levels ....................................................................................5-8

5.7 Maximum fuel rod surface temperature ...................................................................5-10

5.8 Hot and cold legs liquid levels..................................................................................5-11

5.9 Froude number calculation ......................................................................................5-14

5.10 Steam Generator relief valve flow rate.....................................................................5-16

5.11 Steam Generator Main Steam Isolation Valves mass flow rate...............................5-17

5.12 U-tubes collapsed liquid level ..................................................................................5-19

5.13 Steam generators secondary-side liquid level .........................................................5-20

5.14 Pressurizer liquid level .............................................................................................5-22

5.15 Hot and cold legs fluid temperatures. ......................................................................5-23

6 CONCLUSIONS .................................................................................................................6-1

7 REFERENCES ...................................................................................................................7-1
viii

FIGURES

Page

Figure 1 Model nodalization used for simulation.....................................................................4-1

Figure 2 3D Vessel nodalization and connections visualized with SNAP. ..............................4-2

Figure 3 Steam generator nodalization...................................................................................4-3

Figure 4 Primary and secondary pressures ............................................................................5-4

Figure 5 Primary and secondary pressures. ...........................................................................5-4

Figure 6 Break mass flow rate. ...............................................................................................5-5

Figure 7 Discharged inventory through the break...................................................................5-6

Figure 8 Primary loop A mass flow rate. .................................................................................5-7

Figure 9 Primary loop B mass flow rate. .................................................................................5-8

Figure 10 Core collapsed liquid level........................................................................................5-9

Figure 11 Upper plenum collapsed liquid level. ........................................................................5-9

Figure 12 Downcomer collapsed liquid level...........................................................................5-10

Figure 13 Maximum fuel rod surface temperature. .................................................................5-11

Figure 14 Collapsed liquid level in the hot leg A.....................................................................5-12

Figure 15 Collapsed liquid level in the hot leg B.....................................................................5-12

Figure 16 Collapsed liquid level in the cold leg A. ..................................................................5-13

Figure 17 Collapsed liquid level in the cold leg B. ..................................................................5-13

Figure 18 Froude number in the hot leg (0 to 0.06 NT). .........................................................5-15

Figure 19 SG A relief valve mass flow. ...................................................................................5-16

Figure 20 SG B relief valve mass flow. ...................................................................................5-17

Figure 21 SG A MSIV mass flow. ...........................................................................................5-18

Figure 22 SG B MSIV mass flow. ...........................................................................................5-18

Figure 23 SG U-tube up-flow side collapsed liquid levels in loop B........................................5-19

Figure 24 SG U-tube up-flow side collapsed liquid levels in loop A........................................5-20

Figure 25 Steam generator A. Secondary-side collapsed liquid level.....................................5-21

Figure 26 Steam generator B. Secondary-side collapsed liquid level.....................................5-21

Figure 27 Pressurizer collapsed liquid level (0 to 0.2 NT). .....................................................5-22

Figure 28 Hot leg fluid temperature in loop A. ........................................................................5-23

Figure 29 Hot leg fluid temperature in loop B. ........................................................................5-24

Figure 30 Cold leg fluid temperature in loop A........................................................................5-25
ix

TABLES

Page

Table 1 Control logic and sequence of major events in the experiment. ..................................3-1

Table 2 Predetermined core power decay curve (normalized to the Steady State power).......3-2

Table 3 Pumps relative rotational spee (normalized values). ..................................................3-3

Table 4 Core protection system logic........................................................................................3-3

Table 5 Number of heaters per heat structure. .........................................................................4-3

Table 6 Steady-state condition. Comparison between experiment and TRACE. .....................5-1

Table 7 Chronological sequence of events. Comparison between experiment and TRACE. ...5-2

xi

EXECUTIVE SUMMARY

Failure of scram in Small Break Loss-of-Coolant Accident (SBLOCA) scenarios produced in 

pressurized water reactors (PWR) has been proved to cause high core power for a long time and 

thermal-hydraulic response that hinders core cooling, gradually decreasing the primary coolant 

mass inventory. In such scenarios, there is a chance of natural circulation to occur, meeting 

different conditions as high core power with supercritical flow in the hot legs and counter-current 

flow limiting (CCFL) at the inlet of the steam generator (SG) U-tubes that holds a large amount of 

coolant in the tubes. In such context, detailed study of the transition between subcritical and 

supercritical flows is crucial for reactor safety analysis, being the Froude number (a dimensionless 

value that describes different flow regimes of open channel flow) an essential parameter to 

consider. 

In this study, developed within the OECD/NEA ROSA Project Test 3-1 (SB-CL-38 in JAEA), a cold 

leg SBLOCA under the assumption of total failure of high pressure injection system has been 

studied in the Large Scale Test Facility (LSTF) and simulated via the thermal-hydraulic code 

TRACE5. The test reproduces PWR high-power natural circulation due to failure of scram during 

cold leg SBLOCA with a break size equivalent to 1% of the cold leg, under assumption of total 

failure of high injection system. 

TRACE5 model of the plant has been applied, and the chronology of events set to imitate test 

conditions. Different thermal-hydraulic models have been tested in order to fully comprehend the 

high-power natural circulation phenomenology, such as the reflood, the CCFL or the critical flux 

model. Special attention is devoted to Froude number estimation in hot legs due to the prominence 

of the transition between subcritical and supercritical flows in high power natural circulation 

phenomenology. Assessment of TRACE5 leads to the conclusion that the code can successfully 

reproduce complicated conditions of High Power Natural Circulation, during a cold leg SBLOCA. 

Deviations from the experimental results are seen during the accumulator activation. Despite this 

fact, TRACE5 offers a good predictability in this type of accident scenarios

xiii

ACKNOWLEDGEMENTS

This paper contains findings that were produced within the OECD-NEA ROSA Project. The authors 

are grateful to the Management Board of the ROSA Project for their consent to this publication, and 

thank the Spanish Nuclear Regulatory Body (CSN) for the technical and financial support under the 

agreement STN/1388/05/748.

xv

ABBREVIATIONS

AFW auxiliary feedwater

AM accident management

CAMP Code Assessment and Management Program

CCFL counter-current flow limiting

CPU central processing unit

CSN Consejo de Seguridad Nuclear (Spanish Nuclear Regulatory Commission)

ECCS emergency core cooling system

FW feedwater

HPI high pressure injection

JAEA Japan Atomic Energy Agency

kg kilogram(s)

LSTF Large Scale Test Facility

l/s liter per second

m meter(s)

mm millimeter(s)

MPa megapascal

kg/cm2 kilogram per square centimeter

oC degrees Celsius

oK degrees Kelvin

MFW main feedwater

MSIV main steam isolation valve

MW megawatt(s)

MWe megawatt(s) electric

MWt megawatt(s) thermal

NPP nuclear power plant

NRC U.S. Nuclear Regulatory Commission

NT normalized to the total time of the transient

NV normalized to the steady state value

PCT peak cladding temperature

PV pressure vessel

PWR pressurized water reactor

PZR pressurizer

RV relief valve

s second(s)

SBLOCA small break loss-of-coolant accident

SG steam generator

SNAP symbolic nuclear analysis package

TRACE TRAC/RELAP Advanced Computational Engine

1-1

1 INTRODUCTION

Different Small Break Loss-of-Coolant Accident (SBLOCA) scenarios in a PWR have proved failure 

of scram to cause high core power for a long time and thermal-hydraulic response that hinders core 

cooling, gradually decreasing the primary coolant mass inventory. In such scenarios, natural 

circulation can occur gathering different conditions as high core power with supercritical flow in the 

hot legs and Counter-Current Flow limiting (CCFL) at the inlet of the Steam Generator (SG) U-tubes 

that hold a large amount of coolant in the tubes. 

Heat transfer via steam generators is mainly achieved by two-phase natural circulation after the 

beginning of the SBLOCA when the reactor coolant system is filled with sufficient water inventory. 

When the primary coolant inventory decreases, natural circulation ends and steam begins to be 

condensed in steam generator tubes and flows back to the reactor vessel through the hot legs. This 

is known as reflux condensation phenomenon. Reflux condensation and natural circulation become 

important in heat removal when the break size is 0.1% or lower [1]. Actually, the reflux 

condensation is one of the effective heat removal mechanisms during a SBLOCA. TRACE contains 

a unified heat transfer package that includes models for reflood heat transfer, which can be 

activated when reflood is expected to occur. 

The supply of cooling water into the reactor core is limited by the CCFL at the reactor vessel 

downcomer, pressurizer surge line and hot leg pipe. When CCFL occurs under the reflux 

condensation, it means that the pressure drop across the hot leg increases. This causes pressure 

increase in the upper plenum of the reactor vessel and the decrease of the liquid level in the reactor 

core, which results in an increase of the fuel temperature. Due to this, development of CCFL 

correlation is important for the prediction of system behavior and thermal-hydraulic phenomena.

A detailed study of the transition between subcritical and supercritical flows is also important for 

reactor safety analysis. In this frame, the Froude number is an important parameter to be 

considered. Froude Number is defined as the ratio between the mean flow velocity and the speed of 

an elementary gravity wave travelling over the water surface. Such ratio discriminates subcritical or 

supercritical flow depending on whether its velocity is higher or lower than the speed of a 

disturbance wave. There is extensive literature about Froude Number and heat transfer regimes [2, 

3, 4].

In this work, the purpose of authors is to describe the most relevant results achieved by using the 

thermal-hydraulic code TRACE5 in the frame of OECD/NEA ROSA Project Test 3-1 (SB-CL-38 in 

JAEA) [5, 6], to validate the code predictability. The experiment simulates a high-power natural 

circulation during 1% cold leg diameter SBLOCA, also considering the failure of the High Pressure 

Injection System (HPIS). The Test 3-1 of the OECD/NEA ROSA project was handled during 5th of 

December 2006 in the Large Scale Test Facility (LSTF) of the Japanese Atomic Energy Agency 

(JAEA). The LSTF [7] simulates a PWR reactor, Westinghouse type, of four loops and 3423 MW of 

thermal power, scaled to 1/48 in volume and two loops.

In this frame, TRACE5 code [8, 9] has been used in order to simulate a SBLOCA scenario, allowing 

testing its behavior for high power natural circulation. In addition, different thermal-hydraulic models 

have been tested in order to fully comprehend the high-power natural circulation phenomenology, 

such as the reflood, the CCFL or the critical flux model, giving a feedback for developers and 

programmers.

2-1

2 ROSA FACILITY DESCRIPTION

In this section, a brief description of the LSTF facility (in the Tokai Research Establishment of the 

JAERI) [7] is presented. The primary coolant system consists of the Pressure Vessel (PV), the 

primary loop A with the pressurizer (PZR) and the symmetrical primary loop B. Each loop contains a 

primary Coolant Pump (PC) and a Steam Generator (SG). The secondary-coolant system consists 

of the jet condenser (JC), the Feedwater Pump (PF), the Auxiliary Feedwater Pumps (PA) and 

related piping system in addition to two SG secondary systems. The ECCSs consist of the high 

pressure charging pump (PJ), the high pressure injection pump (PL), the Residual Heat Removal 

(RHR) system and the Primary Gravity Injection Tank (PGIT). The coolant discharged from the 

primary system is stored in the break flow Storage Tank (ST). The PV is composed of an upper 

head above the upper core support plate, the upper plenum between the upper core support plate 

and the upper core plate, the core, the lower plenum and the downcomer annulus region 

surrounding the core and upper plenum. LSTF vessel has 8 upper head spray nozzles (of 3.4 mm 

inner-diameter). 8 Control Rod Guide Tubes (CRGTs) form the flow path between the upper head 

and upper plenum. The maximum core power of the LSTF is limited to 10 MW which corresponds 

to 14% of the volumetrically scaled PWR core power and is sufficiently capable to simulate PWR 

decay heat power after the reactor scram. 

Regarding SG, each of them contains 141 U-tubes which can be classified in different groups 

depending on their length (an average length of 19.7 m can be considered, with a maximum height 

of 10.62 m and a minimum height of 9.156 m). U-tubes have an inner diameter of 19.6 mm and an 

outer diameter of 25.4 mm (with 2.9 mm wall thickness). As a consequence, the total inner and 

outer surface areas are 171 and 222 m2

, respectively. On the other hand, vessel, plenum and riser 

of steam generators have an inner height of 19.840, 1.183 and 17.827 m, respectively. The 

downcomer is 14.101 m in height.

3-1

3 TRANSIENT DESCRIPTION

The break unit is connected to the cold leg inner surface, and the orifice flow area corresponds to 

1% of the volumetrically-scaled cross-sectional area of the reference PWR cold leg. The transient 

starts at time zero with opening the break valve and increasing the rotational speed of the coolant 

pumps. The complete control logic of the transient is listed in Table 1.

Few seconds afterwards, the scram signal is generated. This signal produces the initiation of the 

core power decay curve (Table 2), calculated by considering the stored heat in fuel rods and 

delayed neutron fission power. The PWR core power decreased due to negative feedback induced 

mainly by an increase in the core void fraction during the transient.

Table 1 Control logic and sequence of major events in the experiment.

Event Condition

Break. Time zero.

Generation of scram signal. Time = 0.004 NT.

Pressurizer (PZR) heater off. Generation of scram signal or PZR liquid 

level below 0.306 NV.

 Initiation of core power decay curve simulation. Generation of scram signal.

Initiation of Primary Coolant Pump coastdown. Generation of scram signal.

Turbine trip (closure of stop valve). Generation of scram signal.

Closure of Main Steam Isolation Valve. Generation of scram signal.

Termination of Main Feedwater. Generation of scram signal.

Initiation of Auxiliary Feedwater. Generation of scram signal.

Initiation of accumulator system. Primary pressure decreases to a 

determined value.
3-2

Table 2 Predetermined core power decay curve (normalized to the Steady State power).

Normalized Time Normalized

Power Normalized Time NormalizedPower

0 1 0.058 0.3042

0.036 1 0.062 0.2763

0.040 0.8150 0.065 0.2423

0.044 0.5366 0.068 0.2263

0.046 0.4504 0.076 0.2079

0.050 0.3906 0.080 0.2000

0.054 0.3538 0.090 0.1913

Normalized Time Normalized

Power Normalized Time NormalizedPower

0.100 0.1832 0.600 0.0936

0.120 0.1577 0.800 0.0886

0.160 0.1487 1.000 0.0814

0.190 0.1342

0.256 0.1238

0.300 0.1096

0.400 0.1003

Scram signal produces the following actions: initiation of core power decay curve, initiation of 

primary coolant pumps coast down (Table 3), turbine trip, closure of Main Steam Isolation Valves 

(MSIV), termination of Main Feedwater (MFW) and initiation of Auxiliary Feedwater (AFW). When 

secondary pressure reaches a determined value, Relief Valve (RV) of SGs starts to open and close 

in order to maintain pressure almost constant. Finally, when primary pressure drops to a determined

value, the accumulator water system is activated. 
3-3

Table 3 Pumps relative rotational speed (normalized values).

Time

Normalized

Relative

rotational speed

Time

Normalized

Relative

rotational speed

Time

Normalized

Relative

rotational speed

0 1.000 0.006 0.280 0.016 0.125

0.0004 0.850 0.008 0.220 0.018 0.110

0.001 0.730 0.01 0.185 0.020 0.100

0.002 0.540 0.012 0.160 0.050 0.000

0.004 0.370 0.014 0.140 -- ---

In order to protect the facility, the core power is automatically decreased by the core protection 

system when the maximum fuel rod surface temperature excess 873 K, as it can be seen in Table 

4. 

Table 4 Core protection system logic.

Control of

core power to

Maximum fuel rod

surface temperature (K)

75% 873

50% 983

25% 903

10% 913

0% (core power trip) 923

4-1

4 APPLIED METHOD: TRACE5 MODEL OF ROSA FACILITY

TRACE5 code is designed to perform best-estimate analyses of LOCAs or operational transients. 

However, it has some limitations in use; for example, those transients in which one expects to 

observe thermal stratification of the liquid phase in the 1D components cannot be directly modeled. 

Furthermore, it is not appropriate for modeling situations in which transfer of momentum plays an 

important role at a localized level [9]. In this work, the LSTF has been modeled with 88 hydraulic 

components (7 BREAKs, 13 FILLs, 29 PIPEs, 2 PUMPs, 1 PRIZER, 21 TEEs, 14 VALVEs and 1 

VESSEL). In order to characterize the heat transfer processes, 48 Heat Structure components 

(Steam Generator U-tubes, core power, pressurizer heaters and heat losses) have been 

considered. Figure 1 shows the nodalization of the model using the Symbolic Nuclear Analysis 

Package software (SNAP) [10].

Figure 1 Model nodalization used for simulation.
4-2

In order to model the pressure vessel, a 3D–VESSEL component has been considered (Figure 2). 

A nodalization consisting of 19 axial levels, 4 radial rings and 10 azimuthal sectors has been 

selected. This nodalization characterizes with an acceptable detail the actual features of the LSTF 

vessel. Increasing the number of axial levels, azimuthal sectors or radial rings, does not improve 

significantly the agreement with experimental results, but increases CPU time. For each axial level, 

volume and effective flow area fractions have been set according to technical specifications

provided by the organization [7]. Active core is located between levels 3 and 11. Level 12 simulates 

the upper core plate. Levels 13 to 15 characterize the vessel upper plenum. The upper core support 

plate is located in level 16. Finally, upper head is defined between levels 17 to 19. 3D-VESSEL is 

connected to different 1D components: 8 Control Rod Guide Tubes (CRGT), hot leg A and B (level 

15), cold leg A and B (level 15) and a bypass channel (level 14). Control rod guide tubes have been 

simulated by PIPEs components, connecting levels 13 and 19 and allowing the flow between upper 

head and upper plenum.

Figure 2 3D Vessel nodalization and connections visualized with SNAP.

30 HTSTRs simulate the fuel assemblies in the active core. The component power manages the 

power supplied by each HTSTR to the 3D-VESSEL. Fuel elements (1008 in total) were distributed 

into the 3 rings: 154 elements in ring 1, 356 in ring 2 and 498 in ring 3 and also characterized by 

HTSTR components. In both axial and radial direction, peaking factors were considered. The power 

ratio in the axial direction presents a peaking factor of 1.495. On the other hand, depending on the 

radial ring, different peaking factors were considered (0.66 in ring 1, 1.51 in ring 2 and 1.0 in ring 3). 

The number of fuel rod components associated with each heat structure has been determined from 

the technical documentation given, taking into account the distribution of fuel rod elements in the 

vessel, as it can be seen in Table 5.

Bypass (upper 

plenum o 

downcomer)

1

3

2

4

10

9

6

7

5

8

Control rod guide tubes

Spray nozzles 

(downcomer –

upper head)
4-3

Table 5 Number of heaters per heat structure.

HTSTR Number of 

heaters HTSTR Number of heaters HTSTR Number of heaters

310 17 320 44 330 60

311 17 321 40 331 54

312 10 322 23 332 32

313 12 323 32 333 45

314 20 324 40 334 56

315 17 325 42 335 61

316 16 326 38 336 57

317 12 327 26 337 31

318 14 328 30 338 45

319 17 329 39 339 57

Figure 3 Steam generator nodalization.
4-4

A detailed model of SG (geometry and thermal features) has been developed, due to the fact that 

TRACE5 does not include any pre-determined steam generator component. A representation of the 

SG nodalization can be seen in Figure 3. Both boiler and downcomer components of secondaryside have been modeled by TEEs components. U-tubes have been classified into three groups 

according to each average length and heat transfer features. Steam-separator model can be 

invoked in TRACE5 setting a friction coefficient (FRIC) greater than 1022 at a determined cell edge, 

allowing only gas phase to flow through the cell interface. Heat transfer between primary and 

secondary sides has been performed by using HTSTR components. Cylindrical-shape geometry 

has been used to best fit heat transmission. Critical heat flux flag has been set in order to use an 

AECL-IPPE table, calculating critical quality from Biasi correlation [8, 9]. Inner and outer surface 

boundary conditions for each axial level have been set to couple HTSTR component to hydro 

components (primary and secondary fluids). Different models varying the number of U-tube groups 

were tested (1, 3 and 6 groups). It was found that results do not apparently change, using these 

models. However, in order to best fit the collapsed liquid level in U-tubes without drastically 

increasing CPU time, a 3-group configuration was finally chosen. Heat losses to environment have 

been added to secondary-side walls.

Regarding the break simulation, it is important to take into account the necessity of activating the 

Choke flow model in the break when critical flow conditions are expected to appear. Choke model 

predicts for a given cell the conditions for which choked flow is expected to occur, providing three 

different models in one: subcooled-liquid, two-phase and single-phase vapor model. The break has 

been simulated by means of a VALVE component connected to a BREAK component in order to 

establish the boundary conditions. This BREAK has been modelled following the recommendations 

of the TRACE5 user’s manual [9]. In this case, since the break is simulated to discharge in a big 

volume space (the Storage Tank), a dxin=1.0*10-6 (length) and a volin=1.0*106 (volume) has been 

selected with the purpose of providing a large area.
5-1

5 RESULTS AND DISCUSSION

5.1 Steady-state

Steady-state conditions achieved in the simulation were in reasonable agreement with the 

experimental values. In Table 6, the relative errors (%) between experimental and simulated results 

for different items are listed. It is important to remark that in any case, the maximum difference 

between experiment and simulation is 5%. In order to achieve the steady state conditions, the 

duration of simulation was stated to 3000 seconds. 

Table 6 Steady-state condition. Comparison between experiment and TRACE.

Item Relative Error (%)

(Loop with PZR)

Core Power 0.0

Hot leg Fluid Temperature 0.1

Cold leg Fluid Temperature 0.1

Mass Flow Rate 5.0

Pressurizer Pressure 0.2

Pressurizer Liquid Level 4.3

Accumulator System Pressure 0.2

Accumulator System Temperature 0.5

SG Secondary-side Pressure 0.7

SG Secondary-side Liquid Level 3.8

Steam Flow Rate 0.4

Main Feedwater Flow Rate 0.0

Main Feedwater Temperature 0.3

Auxiliary Feedwater Temperature 0.3
5-2

5.2 Transient

Table 7 lists the chronology sequence of events during the transient and the comparison in 

Normalized Time (1 NT = Total Transient Time) between the experiment and TRACE results. 

 Table 7 Chronological sequence of events. Comparison between experiment and 

 TRACE.

Event

Experiment 

Normalized 

time

TRACE5 

Normalized 

time

Valve open 0.000 0.000

Scram signal 0.004 0.008

Initiation of coastdown of primary coolant pumps 0.004 0.008

Primary coolant pumps stopped 0.049 0.052

Termination of continuous opening of SG RVs, 

termination of two-phase natural circulation, break 

flow from single-phase liquid to two-phase flow

About 0.054 0.061

Core power decrease by LSTF core protection 

system About 0.324 0.325

Max. fuel rod surface temperature About 0.364 0.400

Primary pressure lower than SG secondary-side 

pressure

About 0.400 0.405

Initiation of accumulator system (primary pressure 

lower than a determined value) About 0.420 0.421

Break valve closure 1.000 1.000

Variables presented in this section follow the requirements for an exhaustive analysis of the 

transient. These variables include: Pressures in both primary and secondary circuits, mass flow rate 

and inventory at the break, primary mass flow, vessel collapsed-liquid levels, collapsed-liquid levels 

in hot and cold legs, Froude number in hot leg, U-tubes liquid level and liquid level in SG 

secondary-side. Results have been normalized in each graph. 
5-3

5.3 System pressures

A comparison between primary and secondary pressures is presented in Figure 4 and 5. At time 

zero primary pressure starts to decrease due to the loss of coolant discharged through the break. 

The scram signal is generated few seconds after the break, along with the closure of the main 

Steam Isolation Valves (MSIV) of the Steam Generators (SG), termination of Main Feed Water 

(MFW), initiation of Auxiliary Feed Water (AFW) and the coast down of the primary coolant pumps. 

After that, only natural circulation occurs, contributing to cool down the core during the first part of 

the transient, until 0.06 NT, when the SG Relief Valves (RVs) are fully closed. 

Some seconds after the closure of the MSIVs, the secondary pressure of SGs increases reaching 

its maximum value at 0.05 NT. Simultaneously, primary pressure is recovered temporarily due to 

MSIVs closure and then starts to decrease along with the beginning of the power decay curve. 

When secondary pressure reaches a determined value, RV of SGs starts to open and close in order 

to maintain pressure almost constant. At 0.1 NT, RVs are definitively closed. 

The primary pressure becomes lower than the SG secondary-side pressure at about 0.4 NT. The 

accumulator system is initiated at about 0.42 NT when the primary pressure decreases to a predetermined value. At this moment, depressurization becomes effective due to steam condensation 

caused by the coolant injection in the cold legs. Most of the core is quenched at about 0.45 NT after 

its seal, induced by the condensation of vapour of the coolant injected by the accumulator for the 

first time, causing a sudden increase of the primary pressure due to the high production of vapour in 

the core. The primary pressure falls below the secondary pressure at about 0.47 NT after the 

automatic drop of core power.

In general, both primary and secondary-side pressures are successfully reproduced by TRACE5 in 

the whole transient, disagreeing only during the accumulator water injection. TRACE5 does not 

reproduce the small pressure peaks registered during the accumulator water injection.
5-4

0 0.05 0.1 0.15 0.2 0.25 0.3 0.4

0.5

0.6

0.7

0.8

0.9

1

Normalized Time

Normalized Pressure

Measured Primary Pressure

Measured Secondary Pressure SGA

Measured Secondary Pressure SGB

TRACE5 Primary Pressure

TRACE5 Secondary Pressure SGA

TRACE5 Secondary Pressure SGB

Figure 4 Primary and secondary pressures 

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

Normalized Time

Normalized Pressure

Measured Primary Pressure

Measured Secondary Pressure SGA

Measured Secondary Pressure SGB

TRACE5 Primary Pressure

TRACE5 Secondary Pressure SGA

TRACE5 Secondary Pressure SGB

Figure 5 Primary and secondary pressures.
5-5

Comparison of pressures between experiment and simulation suggests that condensation of vapour 

in cold leg produced by the water injection from accumulators is not properly simulated by TRACE5. 

From this moment on, it can be observed a slight difference in the primary pressure between 

experiment and simulation till the end of the transient. 

5.4 Break

Figure 6 and Figure 7 show the mass flow and the discharged coolant inventory, respectively, 

through the break. In order to adjust both break mass flow and the discharged primary coolant 

inventory with TRACE5, a sensitivity analysis varying the discharge coefficient was performed. A 

discharge coefficient of 1.0 (for single-phase liquid and two-phase), using Choked Flow model [8], 

fits successfully both mass flow break and discharged coolant inventory, at least until the 

accumulator water injection (at 0.6 NT), when most part of the coolant was discharged through the 

break. Steam single-phase coefficient cannot be changed since TRACE5 does not offer this 

possibility. TRACE5 estimates a change of flow from liquid single-phase to two-phase at around 0.1 

NT, which completely agrees with the experimental measurements. 

The CCFL model has also been used in the TEE component of the break, following the 

recommendations, with the default values of the coefficients for the Kutateladze model. A proper 

adjustment of the discharged mass flow inventory gives a high reliability to the TRACE5 model. 

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 -0.2

0

0.2

0.4

0.6

0.8

1

1.2

Normalized Time

Normalized Mass Flow Rate

Measured Break Flow Rate

TRACE Flow Rate

Figure 6 Break mass flow rate.
5-6

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 -0.2

0

0.2

0.4

0.6

0.8

1

1.2

Normalized Time

Normalized Discharged Inventory

Measured Discharged Primary Coolant Mass

TRACE5

Figure 7 Discharged inventory through the break.
5-7

5.5 Primary loop mass flows

Primary loop A and B mass flow rates are shown in Figures 8 and 9, respectively. In the first part of 

the transient the primary mass flow increases due to the higher angular speed of the pumps. At 

0.05 NT it starts to decrease simultaneously with the pumps coast down. After that only natural 

circulation occurs and it contributes to cool down the core during the first part of the transient, until 

0.06 NT, when the SG RVs are fully closed. 

Primary loop two-phase natural circulation continued about 0.06 NT when the RV SGs terminated 

continuous opening. Liquid accumulated in the SG upflow-side U-tubes and inlet plenum during 

reflux condensation mode. Flow in hot legs became supercritical during two-phase natural 

circulation due to high vapour and liquid velocity, causing the hot liquid level to be quite low.

TRACE5 results agree with the experimental values. The main discrepancies are found during the 

accumulator water injection period. As it can be seen in Figure 9, between 0.4, and 0.8 NT, 

experimental primary mass flow through loop B, increases three times. These increases of mass 

flow are due to the accumulator water entrance in cold legs. This effect has not been properly 

predicted by TRACE5. 

0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 -0.1

0

0.1

0.2

0.3

0.4

0.5

0.6

Normalized Time

Normalized Mass Flow Rate

Primary Loop A Mass Flow Rate

TRACE5 A

Figure 8 Primary Loop A mass flow rate.
5-8

0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

0.45

0.5

Normalized Time

Normalized Mass Flow Rate

Primary Loop B Mass Flow Rate

TRACE5 B

Figure 9 Primary Loop B mass flow rate.

5.6 Vessel collapsed liquid levels

The following Figures (10, 11 and 12) show a comparison between the collapsed liquid levels in the 

upper plenum, core and downcomer, respectively, for both experimental and TRACE results. All the 

tendencies are successfully reproduced by TRACE5. Slight discrepancies must be remarked. Core 

clearance is slightly later simulated by TRACE, even though the core quench is predicted with 

precision.
5-9

0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

Normalized Time

Normalized Core Collapsed Level

Measured Core Collapsed Level

TRACE5

Figure 10 Core collapsed liquid level.

On the other hand, TRACE perfectly simulates the upper plenum clearance and quench.

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 -0.2

0

0.2

0.4

0.6

0.8

1

1.2

Normalized Time

Normalized Upper Plenum Collapsed Level

Measured Upper Plenum Collapsed Level

TRACE5

Figure 11 Upper plenum collapsed liquid level.
5-10

Concerning the downcomer, TRACE5 also makes a good prediction, only having a slight 

discrepancy in the first part of the transient, before the clearance, being the level predicted slightly 

lower than in the experiment.

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

0.4

0.5

0.6

0.7

0.8

0.9

1

Normalized Time

Normalized Downcomer Collapsed Level

Measured Downcomer Collapsed Level

TRACE5

Figure 12 Downcomer collapsed liquid level.

5.7 Maximum fuel rod surface temperature

The maximum rod surface temperature starts to increase when the core clearance takes place. 

TRACE predicts such increment at 0.24 NT approximately, concurrent with the experiment. At 0.4 

NT approximately it reaches the maximum, activating the core protection system by power control, 

as can be seen in the following Figure 13.
5-11

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 450

500

550

600

650

700

750

800

850

900

950

Normalized Time

Normalized Temperature

Fuel Rod Surf T

Fuel Rod Surf T

Fuel Rod Surf T

TRACE5

Figure 13 Maximum fuel rod surface temperature.

5.8 Hot and cold legs liquid levels

The following Figures (14, 15, 16 and 17) show the liquid level in hot and cold legs, respectively. 

Experimentally, liquid level was obtained with a three gamma ray beam densitometer. In the 

experiment, the collapsed liquid level of both hot legs is almost the same during the Accident 

Management (AM) action, since the natural circulation phenomenon is produced in both loops. The 

flow in the hot legs becomes two-phase flow and is horizontally stratified with high liquid velocity 

after the break. Afterwards it becomes supercritical during two-phase natural circulation due to the 

high liquid and vapour velocity, causing a drastic drop of the level. At about 0.06 NT, when the twophase natural circulation ends, the flow in the hot legs becomes subcritical, recovering the liquid 

level. At 0.18 NT a significant drop of the liquid level is produced, after the inlet plenum of the steam 

generator is completely empty, and it recovers after the actuation of the accumulator injection 

system. No transition to supercritical flow is produced in the cold legs since the liquid velocity is not 

accelerated by the upper plenum. At 0.18 NT the liquid level of the cold legs drops and it 

temporarily recovers with the injection of the accumulator. 

In general, TRACE can reproduce all the tendencies and changes in both hot and cold legs. Most 

discrepancies are found in the last part of the transient because TRACE does not exactly reproduce 

the recovering of the liquid level in any of the legs, due to the rapid coolant discharge through the 

break from the accumulation injection on.
5-12

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.2

0.4

0.6

0.8

1

1.2

Normalized Time

Normalized Collapsed Level 

Hot Leg A Collapsed Level

TRACE5 A

Figure 14 Collapsed liquid level in the hot leg A.

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

Normalized Time

Normalized Collapsed Level

Hot Leg B Collapsed Level

TRACE5 B

Figure 15 Collapsed liquid level in the hot leg B.
5-13

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

Normalized Time

Normalized Collapsed Level

Cold Leg A Collapsed Level

TRACE5 A

Figure 16 Collapsed liquid level in the cold leg A.

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.2

0.4

0.6

0.8

1

1.2

Normalized Time

Normalized Collapsed Level

Cold Leg B Collapsed Level

TRACE5 B

Figure 17 Collapsed liquid level in the cold leg B.
5-14

5.9 Froude number calculation

The Froude number is defined as the ratio of inertia to gravity forces in the flow. This ratio may also 

be interpreted physically as the ratio of the mean flow velocity to the speed of an elementary gravity 

(surface or disturbance) wave traveling over the water surface. When the Froude number is equal to

one, the speed of the surface wave and that of the flow is the same. The flow is in the critical state. 

When the Froude number is less than one, the flow velocity is smaller than the speed of a 

disturbance wave traveling on the surface. Flow is considered to be subcritical. Gravitational forces 

are dominant. The surface wave will propagate upstream and, therefore, flow profiles are calculated 

in the upstream direction. When the Froude number is greater than one, the flow is supercritical and 

inertial forces are dominant. The surface wave will not propagate upstream, and flow profiles are 

calculated in the downstream direction.

Figure 18 shows the evolution of the Froude number in the time interval between 0 and 0.06 NT in 

both hot legs. Froude number is calculated using the following formula: 

gh

U Fr =

,

where U is the liquid velocity (m/s), h the liquid level (m) and g the acceleration due to gravity 

(m/s2

). Froude number graphic shows that flow is supercritical during the interval 0.0 to 0.007 NT, 

and between 0.017 to 0.06 NT, both in TRACE5 and in the experiment. 

Figure 18 shows that flow in hot legs is supercritical soon after the break, during the interval 0.0 to 

0.007 NT, and becomes subcritical during a short period (0.007 to 0.015 NT). Flow in hot legs is 

again supercritical during two-phase natural circulation (0.015 NT to 0.06 NT), due to the high liquid 

and vapor velocity by high core power, causing a drastic drop of the level. Finally, flow in hot legs 

returns to subcritical at 0.06 NT, recovering the hot leg liquid level, when two-phase natural 

circulation was terminated.

Liquid flow in the hot legs is accelerated by vapor flow and liquid head in the upper plenum when 

the upper plenum liquid is above the hot leg level. Such water acceleration mechanisms are not 

observed in the cold legs. No transition to supercritical flow occurs in the cold legs. In general, 

TRACE5 can reproduce all the tendencies and changes in both hot and cold legs.
5-15

0 0.01 0.02 0.03 0.04 0.05 0.06

0.6

0.8

1

1.2

1.4

1.6

1.8

2

2.2

Normalized Time

Froude Number

Froude Number

Froude number exp

TRACE

Figure 18 Froude number in the hot leg (0 to 0.06 NT).
5-16

5.10 Steam Generator relief valve flow rate

In the TRACE5 model, these valves are represented by means of components VALVE-73 (loop A) 

and VALVE-75 (loop B). A good agreement has been achieved between TRACE5 and the 

experiment, as can be observed in Figures 19 and 20. These figures perfectly show the periods 

corresponding to the RV actuation. It is important to remark that this section only shows the vapour 

mass flow through the RV.

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.2

0.4

0.6

0.8

1

1.2

Normalized Time

Normalized Mass Flow Rate

Mass Flow Rate Relief Valve A

TRACE

Figure 19 SG A relief valve mass flow.
5-17

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 -0.2

0

0.2

0.4

0.6

0.8

1

1.2

Normalized Time

Normalized Mass Flow Rate 

Mass Flow Rate Relief Valve B

TRACE

Figure 20 SG B relief valve mass flow.

5.11 Steam Generator Main Steam Isolation Valves mass flow rate

Main Steam Isolation Valves (MSIVs) are automatically closed when the scram signal is generated. 

This occurs at 0.004 NT in the experiment and at 0.008 NT in the simulation with TRACE5 (Figure 

21 and Figure 22). During the rest of the transient, the MSIVs remain closed and therefore the mass 

flow is zero.
5-18

0 0.02 0.04 0.06 0.08 0.1 0.12 0.14 0.16 0.18 0.2 -0.2

0

0.2

0.4

0.6

0.8

1

1.2

Normalized Time

Normalized Mass Flow Rate

SGA Steam Line Mass Flow Rate

TRACE5 

Figure 21 SG A MSIV mass flow rate.

0 0.02 0.04 0.06 0.08 0.1 0.12 0.14 0.16 0.18 0.2 -0.2

0

0.2

0.4

0.6

0.8

1

1.2

Normalized Time

Normalized Mass Flow Rate

SGB Steam Line Mass Flow Rate

TRACE5 

Figure 22 SG B MSIV mass flow rate.
5-19

5.12 U-tubes collapsed liquid level

U-tubes have been grouped according to similar lengths. In the TRACE5 simulation and due to the 

calculation time cost together with the obtained results, a 3-group classification has been adopted. 

Collapsed liquid levels obtained with TRACE5 are satisfactory, properly reproducing the clearance 

of the tubes. The collapsed liquid level of the U tubes in the steam generator indicates the 

beginning of a rapid increase of the void fraction before the core power starts to decrease. The 

accumulation of liquid in the upflow side and in the inlet plenum takes place during the reflux 

condensation, probably due to the CCFL in the inlet of the U tubes and in the bottom of the inlet 

plenum produced by the high vapour velocity caused by the core power. The accumulation of liquid 

in the upflow side and in the inlet plenum causes a small drop of the core liquid level between 0.06 

and 0.1 NT. Results are shown in Figures 23 and 24. 

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 -0.2

0

0.2

0.4

0.6

0.8

1

1.2

Normalized Time

Normalized values

SG U-tube upflow-side collapsed liquid levels in loop w/o PZR

SGB Tube1 Inlet

SGB Tube6 Inlet

TRACE5 group1

TRACE5 group2

TRACE5 group3

Figure 23 SG U-tube up-flow side collapsed liquid levels in loop B.
5-20

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 -0.2

0

0.2

0.4

0.6

0.8

1

1.2

Normalized Time

Normalized values

SG U-tube upflow-side collapsed liquid levels in loop with PZR

SGA Tube1 Inlet

SGA Tube6 Inlet

TRACE5 group1

TRACE5 group2

TRACE5 group3

Figure 24 SG U-tube up-flow side collapsed liquid levels in loop A.

5.13 Steam generators secondary-side liquid level

The following Figures (25 and 26) show the collapsed liquid level of the secondary side of the 

steam generator. Along with the closure of the MSIVs, the main vapour line flow to the turbines is 

interrupted and the Main Feed Water (MFW) line of both steam generators is closed.

From the generation of the SI signal on, the auxiliary feed water system begins actuating with the 

purpose of keeping the secondary liquid level of the steam generators.

The stopping of the relief valves is followed by an increase of liquid level from the secondary side of 

the steam generator due to the vapour condensation caused by the auxiliary feed water. The 

auxiliary water stops at 0.56 NT when the liquid level in the secondary side of the steam generator 

reaches over the top of the U tubes. The prediction with TRACE5 is satisfactory.
5-21

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

Normalized Time

Normalized Collapsed Level

SGA Wide Range

TRACE5 A

Figure 25 Steam generator A. Secondary-side collapsed liquid level.

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

Normalized Time

Normalized Collapsed Level

SGB Wide Range

TRACE5 B

Figure 26 Steam generator B. Secondary-side collapsed liquid level.
5-22

5.14 Pressurizer liquid level

Figure 27 shows the pressurizer water level. Liquid level starts to decrease immediately after the 

break. Experimentally, it becomes completely empty at 0.06 NT. On the other hand, both heaters of 

the PZR automatically connect to the maximum power after the break, according to the control logic 

of the depressurization of the transient. Both heaters disconnect at 0.067 NT. The liquid level of the 

PZR remains empty for the rest of the transient.

0 0.02 0.04 0.06 0.08 0.1 0.12 0.14 0.16 0.18 0.2 -0.2

0

0.2

0.4

0.6

0.8

1

1.2

Normalized Time

Normalized values

Measured Pressurizer level

TRACE5 PZR water level

Figure 27 Pressurizer collapsed liquid level (0 to 0.2 NT).
5-23

5.15 Hot and cold legs fluid temperatures.

In Figures 28 and 29 it can be seen fluid temperatures measured and simulated in the hot legs A 

and B, respectively. Fluid temperature in cold leg in loop A is represented in Figure 30.

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

0.65

0.7

0.75

0.8

0.85

0.9

0.95

1

Normalized Time 

Normalized value

Hot leg fluid temperature in loop with PZR

Hot Leg A

TRACE5

Figure 28 Hot leg fluid temperature in loop A.

Hot leg fluid temperature increases due to the superheated steam coming from the core once the 

core uncovery has begun. 
5-24

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

0.65

0.7

0.75

0.8

0.85

0.9

0.95

1

1.05

1.1

Normalized Time 

Normalized value

Hot leg fluid temperature in loop without PZR

Hot Leg B

TRACE5

Figure 29 Hot leg fluid temperature in loop B.

Hot leg fluid temperature increases when hot steam flow passes through the crossover leg at about 

0.44 NT after the loop seal clearing in the loop without pressurizer. The initiation of the accumulator 

coolant injection (at 0.4 NT approximately) causes the fluid temperature to decrease at the cold 

legs.

The main discrepancies between measured and simulated curves are observed in the cold leg loop 

A after the activation of the core protection system by power control (at 0.5 NT approximately).
5-25

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

0.65

0.7

0.75

0.8

0.85

0.9

0.95

1

1.05

Normalized Time 

Normalized value

Cold leg fluid temperature in loop with PZR

Cold Leg A

TRACE5

Figure 30 Cold leg fluid temperature in loop A.

6-1

6 CONCLUSIONS

The objective of this report was to obtain detailed thermal-hydraulic data in the simulation of the 

OECD/NEA ROSA Project Test 3-1, high-power natural circulation in a PWR, for the validation of 

the code TRACE5.

Results show that TRACE5 can successfully reproduce complicated conditions of High Power 

Natural Circulation, when a break flow in a cold leg is produced. TRACE5 adequately predicts the 

mass flow and the coolant inventory discharged through the break, and the changes between 

subcritic and supercritic flow. In fact, flow in hot legs became supercritical during two-phase natural 

circulation due to high vapor and liquid velocity. 

Results with TRACE5 confirm that both hot legs present a similar thermal hydraulic behavior. The 

only deviations from the experimental results occur during the period when the accumulator 

actuates. Primary pressure does not increase when coolant is injected from accumulator into the 

hot leg of the loop without PZR and therefore the coolant is more rapidly discharged through the 

break. Also hot leg in loop without PZR does not have a sudden increase of mass flow due to this 

fact, and all hot and cold legs begin to empty before the experimental.

7-1

7 REFERENCES

1. H. Y. Jeong, “Prediction of counter-current flow limitation at hot leg pipe during small-break 

LOCA”, Annals of Nuclear Energy, 29, 571-583 (2002).

2. K. W. Seo, M. H. Kim, M. H. Anderson, M. L. Corradini, “Heat Transfer in a Supercritical 

Fluid: Classification of Heat Transfer Regimes”, Nuclear Technology, 154, 335-349 (2006).

3. M. L. Corradini, “Transport Phenomena in Supercritical Fluids in GEN-IV Reactor Designs”, 

Nuclear Technology, 167, 145-156 (2009).

4. B. Freeze, S. Smolentsev, N. Morley, M. Abdou, “Characterization of the Effect of Froude 

Number on Surface Waves and Heat Transfer in Inclined Turbulent Open Channel Water 

Flows”, International Journal of Heat and Mass Transfer, 46, 3765-3775 (2003).

5. Thermalhydraulic Safety Research Group, Nuclear Safety Research Center, Final Data 

Report of ROSA/LSTF Test 3-1 (High Power Natural Circulation Experiment SB-CL-38 in 

JAEA), Japan Atomic Energy Agency, Japan (2008).

6. T. Takeda, H. Asaka, H. Nakamura, “RELAP5 analysis of OECD/NEA ROSA Project 

experiment simulating a PWR small break LOCA with high-power natural circulation”, 

International Conference on the Physics of Reactors “Nuclear Power: A Sustainable 

Resource”, Casino-Kursaal Conference Center, Interlaken, Switzerland, September 14-19 

(2008).

7. The ROSA-V Group, ROSA-V Large Scale Test Facility (LSTF) system description for the 

3rd and 4th simulated fuel assemblies, JAERI-Tech, Japan, 2003.

8. Nuclear Regulatory Commission, Division of Risk Assessment and Special Projects, Office 

of Nuclear Regulatory Research, TRACE V5.0. User’s manual. Volume 1: Input 

Specification, Nuclear Regulatory Commission, U. S (2007).

9. Division of Risk Assessment and Special Projects. Office of Nuclear Regulatory Research. 

U. S. Nuclear Regulatory Commission, TRACE V5.0. Theory manual. Field Equations, 

Solution Methods and Physical Models, Nuclear Regulatory Commission, US (2007).

10. Nuclear Regulatory Commission and Applied Programming Technology, Symbolic Nuclear 

Analysis Package (SNAP) (2007).




UNITED STATES 

NUCLEAR REGULATORY COMMISSION

WASHINGTON, DC 20555-0001

--------------------

OFFICIAL BUSINESS
NUREG/IA

-0413 Assessment of TRACE 5.0 Against ROSA Test 3-1, 

Cold Leg SBLOCA