NUREG-1744

Assessment of the 

TRAC-M Codes Using 

Flecht-Seaset Reflood 

and Steam Cooling Data 

U.S. Nuclear Regulatory Commission 

Office Nuclear Regulatory Research 

Washington, DC 20555-0001
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NUREG-1744

Assessment of the 

TRAC-M Codes Using 

Flecht-Seaset Reflood 

and Steam Cooling Data

Manuscript Completed: May 2001 

Date Published: May 2001 

Prepared by 

F. Odar 

Division of Systems Analysis and Regulatory Effectiveness 

Office of Nuclear Regulatory Research 

U.S. Nuclear Regulatory Commission 

Washington, DC 20555-0001
ABSTRACT

This report presents the results of an assessment of the capabilities of the TRAC-M(F77), 

Version 5.5.2.A, and TRAC-M(F90), Version 3.580, codes to predict reflood and steam cooling 

phenomena that may occur during a postulated large-break loss-of-coolant accident (LOCA). 

The assessment is predicated on test data from Flecht-Seaset Runs 31504 and 32753. The 

assessment shows that predictions of the reflood phenomena derived using both codes are 

inaccurate; however, it is judged that they can conservatively predict peak clad temperatures in 

heated rods since the code model expels more water from the test section than measured. The 

predictions of steam cooling in single-phase flow conditions are acceptable.

-iii-
TABLE OF CONTENTS 

Page 

Abstract iii 

Executive Summary. . xi 

1. Introduction. 1 

2. Reflood Phenomena and Code Requirements 3 

2.1 Description of the Phenomena 3 

2.2 Predictive Models and Experimental Facilities 9 

2.3 Discussion of Phenomena Observed in Flecht-Seaset Tests 10 

2.3.1 Bottom Quench, Transition Zone, Liquid Entrainment, 

Precursor Cooling, and Heat Transfer 11 

2.3.2 Liquid Carryout from the Core . . 19 

2.3.3 Fuel Assembly 19 

2.4 Summary of Code Requirements 19 

3. System Description . . . 23 

3.1 Facility Configuration 23 

3.1.1 Forced Reflood Tests and Testing Procedure 23 

3.1.2 Gravity Reflood Tests and Testing Procedure 25 

3.1.3 Steam Cooling Tests and Testing Procedure . 26 

3.2 Heater Rod Bundle 26 

3.3 Discussion of System Design Features 29 

3.3.1 Low-Mass Housing 29 

3.3.2 Bundle Differential Pressure Cells . 31
3.3.3 Pressure Control 31 

3.3.4 Facility Operation 31 

3.3.5 Power Measurement. 31 

3.3.6 Coolant Injection System 31 

3.3.7 Data Acquisition Systems 32 

3.4 Instrumentation . 32 

3.4.1 Heater Rod Bundle Instrumentation 32 

3.4.2 Test Section Instrumentation 33 

3.4.3 Carryover Tank Instrumentation 33 

3.4.4 Steam Separator Instrumentation 33 

3.4.5 Exhaust Pipe Instrumentation 33 

3.4.6 Accumulator Tank Instrumentation 33 

3.4.7 Injection Line Instrumentation 33 

4. TRAC-M Assessment with Flecht-Seaset Reflood Test Data, Run 31504 35 

4.1 Success Metrics 35 

4.2 TRAC Input Model for Flecht-Seaset Forced Reflood Run 31504 37 

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

Experimental Data from Flecht-Seaset Run 31504 43 

4.4 Discussion of TRAC Models 65 

4.5 Sensitivity Calculation without the Choking Model 71 

4.6 Conclusions 71 

5. TRAC-M Assessment with Flecht-Seaset Steam Cooling Test Data, 

Run 32753 77 

5.1 Discussion of Phenomena and Code Requirements 77 

5.2 Success Metrics 77

-vi-
5.3 TRAC Input Model for Flecht-Seaset Steam Cooling Run 32753 79 

5.4 Comparison of TRAC-M(F77) and TRAC-M(F90) Calculations 

with Experimental Data from Flecht-Seaset Run 32753 79 

5.5 Conclusions 80 

References . . 81 

FIGURES 

2.1 Sketch of Transition Boiling During Reflood. 4 

2.2 Flow Patterns During Reflood 5 

2.3 Flooding Rate vs Heat Transfer Coefficient 12 

2.4 Flooding Rate vs Clad Temperatures 13 

2.5 Flooding Rate vs Quench Time 14 

2.6 Flooding Rate vs Temperature Rise. 15 

2.7 Clad Temperature and Heat Transfer Coefficient in Different Zones 16 

2.8 Transition Front Relative to Quench Data in Slow Reflood 17 

2.9 Transition Front Relative to Quench Data in Fast Reflood 18 

2.10 Void Fraction Variation with Time During Reflood 20 

2.11 Carryout Fraction vs Time 21 

3.1 Flecht-Seaset System Configuration 24 

3.2 Flecht-Seaset 161 Rod Test Bundle. 27 

3.3 Heater Rod Thermocouple Locations 28 

3.4 Flecht-Seaset Axial Power Shape 28 

3.5 Decay Power Curve (ANS + 20%) 29 

3.6 Steam Probe Design 30

-vii-
4.1 Flecht-Seaset 31504 Exp. Clad Temperatures, z=1.22m 38 

4.2 Flecht-Seaset 31504 Exp. Clad Temperatures, z=1.98m 39 

4.3 Flecht-Seaset 31504 Exp. Clad Temperatures, z=2.59m 40 

4.4 Flecht-Seaset 31504 Exp. Clad Temperatures, z=3.35m 41 

4.5 TRAC-M Input Model Schematic for Flecht-Seaset Run 31504 42 

4.6 Initial Axial Clad Temperature Variation 45 

4.7 Initial Axial Steam Temperature Variation . . 46 

4.8 Comparison of Calculated and Experimental Clad Temps. @z=1.22m . 47 

4.9 Comparison of Calculated and Experimental Clad Temps. @z=1.98m . 48 

4.10 Comparison of Calculated and Experimental Clad Temps. @z=2.59m . 49 

4.11 Comparison of Calculated and Experimental Clad Temps. @z=3.35m . 50 

4.12 Calculated and Measured Quench Front Propogation 51 

4.13 Experimental Differential Pressure Measurements Along the Test Section 52 

4.14 Calculation of Axial Void Fraction Distribution by TRAC-M(F77) Code 53 

4.15 Comparison of Differential Pressures (Taps 1-2) 54 

4.16 Comparison of Differential Pressures (Taps 2-3) 55 

4.17 Comparison of Differential Pressures (Taps 3-4) 56 

4.18 Comparison of Differential Pressures (Taps 4-5) 57 

4.19 Comparison of Differential Pressures (Taps 5-6) 58 

4.20 Comparison of Differential Pressures (Taps 6-7) 59 

4.21 Steam Temperatures (Z=1.83 - 1.94m) 60 

4.22 Steam Temperatures (Z=2.82 - 3.03m) 61 

4.23 Steam Temperatures (Z=3.35 - 3.51 m) 62 

4.24 Comparison of Test Section Pressures 63

-viii-
4.25 Comparison of Liquid Carryouts 64 

4.26 Comparison of Steam Carryouts 65 

4.27 TRAC-M Reflood Flow Regime Map. . . 67 

4.28 Calculated Relative Velocities at Cell#10 68 

4.29 Calculated Relative Velocities at Cell#17 69 

4.30 Comparison of Clad Temperatures with and without Choking, Z=1.22m 72 

4.31 Comparison of Clad Temperatures with and without Choking, Z=1.98m 73 

4.32 Comparison of Clad Temperatures with and without Choking, Z=2.59m 74 

4.33 Comparison of Clad Temperatures with and without Choking, Z=3.35m 75 

5.1 Experimental Data, Clad Temperatures at Z=1.52m 78 

5.2 Comparison of Predicted and Measured Clad Temperatures 80

-lx-
EXECUTIVE SUMMARY

This report presents an assessment of the capabilities of the TRAC-M(F77), Version 5.5.2.A, 

and TRAC-M(F90), Version 3.580, codes to predict reflood and steam cooling phenomena for 

pressurized water reactors during a large-break loss-of-coolant accident. The assessment of 

the reflood predictive capability is performed using test data from Flecht-Seaset Run 31504, 

while the assessment of steam cooling is performed using test data from Flecht-Seaset Run 

32753. These tests simulate unblocked bundle forced reflood and steam cooling conditions in 

PWRs. 

The top-level requirement in the development of a thermal-hydraulic code is that the code 

should correctly model the physics of the phenomena. This assessment shows that both TRAC 

codes fail to correctly model flow regime maps during reflood and, consequently, they do not 

correctly predict interfacial areas between phases, vapor fractions, phasic velocities, and 

temperatures. Almost none of the important parameters (such as clad temperatures, steam 

temperatures, differential pressures, and liquid and steam carryouts) are consistently predicted 

throughout the range within the acceptable accuracy. The predictions in ranges where clad 

temperatures are predicted within accuracy limits are accidental. The model in the codes 

predict higher liquid carryout than measured in the test. This would indicate that the codes will 

predict higher clad temperatures than measured in the test data, thereby yielding conservative 

predictions. 

The Office of Nuclear Regulatory Research has a research program in place at Pennsylvania 

State University (PSU) to improve the agency's capability to model reflood phenomena using its 

system analysis codes. The findings in this report provide guidance on how future modeling of 

reflood should proceed. This report assesses the code against one particular Flecht-Seaset 

(F-S) run. However, the decision to initiate the PSU test program was based on several reflood 

cases at various facilities. These cases exhibit the overall trend that TRAC-M(F90) overpredicts 

peak clad temperature (PCT), even though the predictions are in reasonable agreement with 

runs made at the CCTF and SCTF facilities. The TRAC-P code series has been used by NRC 

to perform LOCA analysis because of this conservative trend. However, as the agency begins 

to risk-inform its regulations, new designs are submitted and licensees request further uprates. 

NRC would be better equipped to audit licensee submittals with a less conservative, more 

accurate reflood model. In order to model reflood accurately for all designs, conditions and test 

facilities, the proper physics must be modeled mechanistically. The author recommends that 

new flow regime maps for reflood be constructed and the selection criteria for the maps should 

also include effects of both phasic velocities. 

The assessment with the steam cooling test data shows that both TRAC-M codes correctly 

predict the single-phase convective flow heat transfer with acceptable accuracy. The 

agreement between predictions and the test data is "Excellent" to "Reasonable." It is judged 

that "Excellent" agreement throughout the whole region of valid testing may be obtained by 

modeling the input deck in more detail.

-Xi-
SECTION 1 

INTRODUCTION 

This report presents the results of an assessment of the capabilities of the TRAC-M(F90), 

Version 3.580, and TRAC-M(F77), Version 5.5.2A, codes to calculate reflood and steam cooling 

phenomena for pressurized-water reactors (PWRs). The reflood assessment was performed 

using test data from Flecht-Seaset Run 31504, while the steam cooling assessment was 

performed using test data from Flecht-Seaset Run 32753. These tests simulate unblocked 

bundle forced reflood and steam cooling conditions in PWRs. 

Section 2 of this report discusses the reflood phenomenon and requirements for the code's 

predictive capabilities. It begins by discussing the reflood phenomena, in general, as they 

would occur in a PWR during a postulated large-break loss-of-coolant accident (LOCA). It then 

discusses what part of the reflood phenomena would be simulated in Flecht-Seaset tests. 

Finally, it discusses the requirements that are necessary for thermal-hydraulic codes to predict 

the reflood phenomena. 

Section 3 of this report describes the configuration of the Flecht-Seaset facility and the test 

procedures that are used to conduct reflood and steam cooling tests. It then describes the 

heater bundle and various system components, as well as the instrumentation that is used in 

the system. 

Section 4 describes the results of the assessment that was performed using Flecht-Seaset 

Run 31504. This section assesses the capabilities of the TRAC-M(F77) and TRAC-M(F90) 

codes to predict reflood phenomena. It begins by describing the success metrics and expected 

uncertainties. Next, it compares calculations with the actual test data, reviews important 

aspects of the various code models that are used in modeling reflood phenomena, and 

presents conclusions on the basis of the requirements presented in Section 2. 

Section 5 describes the results of the assessment that was performed using Flecht-Seaset 

Run 32753. This section assesses the capabilities of the TRAC-M(F77) and TRAC-M(F90) 

codes to predict steam cooling phenomena. It begins by describing the success metrics and 

expected uncertainties. Next, it compares calculations with the actual test data, and presents 

conclusions.

-1-
SECTION 2 

REFLOOD PHENOMENA AND CODE REQUIREMENTS 

2.1 Description of the Phenomena 

The purpose of reflooding the core is to protect the integrity of the fuel cladding. During a 

large-break LOCA, cladding temperature changes as follows: 

Cladding temperature increases during blowdown from normal operating conditions of 

approximately 3250C to approximately 550-8000C (roughly 1000-1 5000F). 

During refill and reflood, cladding temperature varies because of the decay heat 

generation in fuel rods and heat removal by two-phase flow through the core. 

Cladding temperature increases as long as heat removal is less than heat generation. 

Cladding temperature decreases when heat removal is greater than heat generation. 

As reflood progresses, the cladding very rapidly quenches and cools to the saturation 

temperature as the cladding surface becomes wetted. 

The emergency core cooling (ECC) systems are designed to inject water into the primary 

system, and to recover or reflood a potentially uncovered core. Steam produced during the 

core quenching process creates a pressure that is balanced against the hydraulic height 

difference between the core and the downcomer. The pressure differential is limited by the 

height of the cold leg. 

As the reflood water enters the core, it is heated by the stored energy and decay power of the 

rods. Depending on the time during the transient and injection flow rate, the reflood water at 

the quench front can be subcooled or saturated, or the water can reach saturation below the 

quench front so that a two-phase mixture reaches the quench front. During the reflood, the 

hydraulic conditions on the rod surfaces change completely from unwetted to wetted conditions 

within a few seconds. At the quench front, the stored energy of the rod is released over a 

relatively short distance in which the local axial temperature gradient can be as large as 

500 K/cm. The transition boiling heat transfer regime exists during these few seconds. Correct 

modeling of the transition boiling is important in order to accurately predict core reflood 

behavior. 

A high heat flux over a short rod length is a very significant aspect of the reflood process. 

Figure 2.1 illustrates the variation in the rod temperature during a reflood test, along with the 

calculated surface heat flux obtained from the Semiscale Mod-3 test data (Ref. 1). During the 

1-second time period around 59 s, a large heat flux occurs as the rod returns to a wetted 

nucleate boiling condition. This would imply the occurrence of critical heat flux (CHF) at 59 s, 

i.e., at t 2 . Quenching starts slightly earlier, at t1 . The rod surface temperature at this time is 

at the quench temperature, T where the direct contact between the liquid and the wall starts.

-3-
1600

600-- 1200 

6800 

'9500

T r e4 e r n p ra w 0 0 

400 o 

375 AI 0 

so 57 58 59 60 61 62 rime (see) 

Figure 2.1 Sketch of Transition Boiling During Reflood 

The transition boiling region starts at t, and ends at t2 . In the transition boiling region after t1 , 

as the frequency and duration of direct contact between the liquid and the wall increase, the 

heat flux increases; and consequently, the rod surface temperature decreases. 

Based on the flooding rate and subcooling and quality of the incoming fluid various flow 

regimes may appear during a reflood. Figure 2.2 illustrates sketches of flow patterns observed 

in reflood tests with saturated and subcooled inlet conditions. Saturated conditions are 

associated with low inlet flows, about 1.0-in/s (0.0254m/s). In the case of a reflood with 

saturated low inlet flow, the vapor generated at the quench front sputters and causes a 

transition from an annular to a frothy flow regime followed by Dispersed-Flow Film Boiling, 

DFFB, regime. Details of the frothy regime will be discussed in Section 2.3.1. The steam 

generation from quenching produces very large vapor velocities which entrain and shear liquid 

filaments into droplets. During this process, the vapor fraction changes from low to moderate. 

These droplets are then swept into the upper region of the core producing dispersed flow film 

boiling. The cooling produced by these droplets is called "precursory cooling." During a large 

break LOCA this is the prevailing flow pattern during reflood since the incoming fluid flow to the 

inlet of the core, after the refill process is over, is usually saturated or close to saturation and 

the flow rates are generally low.

-4-
I I

- -Uquld 

Uquld 

OF 

Vapor , 

.0 CHF. 

Saturated Inlet Row Subcooled Inlet Flow 

CHF Critical Heat Flux 

OF Quench Front 

Figure 2.2 Flow Patterns During Reflood

-5-
At high inlet flows, about 6.0-in/s (0.15-m/s), with high subcooling, a different flow pattern 

emerges as the right sketch of Figure 2.2 illustrates. An Inverted-Annular Film Boiling, IAFB, 

takes place immediately downstream of the quench front. Just above the quench front, a stable 

film boiling regime develops. The flow pattern is characterized by a single-phase liquid core. At 

this location, the quality of the two-phase fluid is considered low. This implies that the liquid is 

still subcooled and most of the heat from the rods is absorbed by raising the liquid temperature. 

Further downstream, the liquid core is still surrounded by the vapor annulus and it is 

hydrodynamically unstable, Ref. 2. Heat transfer rates across the vapor film will vary greatly 

with surface wave motion at the vapor liquid interface. Further along the channel an unstable 

flow pattern transition regime develops resulting in a dispersed flow. Generally, heat transfer 

rates are high and the quench front quickly propagates up the bundle. During a large break 

LOCA, this type of flow pattern does not occur too often since the incoming flow to the inlet of 

the core is generally not subcooled and the flow rates are not very high. 

During a reflood, steam is mainly generated at the quench front, causing entrainment of water 

slugs, fragments, and droplets upward through the core. Knowledge of the extent of 

entrainment is necessary to predict the hydrodynamic and thermal response in the core. If the 

steam velocity exceeds a certain value, entrained water can be swept completely out of the 

core. At lower velocities, partial entrainment and levitation of the water may occur. As the 

entrained water travels upward through the core, water fragments and droplets impact the rod 

walls and cause local cooling. Entrained liquid also absorbs heat by radiation from the rods and 

by convection from superheated steam. Thermal non-equilibrium may exist with saturated 

water droplets and superheated steam. Therefore, there is a complex heat transfer process 

from cladding to steam and from steam to droplet where evaporation of the droplet can occur. 

In addition, there is a gradient of steam temperature from the quench front where mostly 

saturated conditions exist (with moderate qualities) to the top of the core where there is a high

quality mixture with high superheat. 

The heat transfer mechanism in this dispersed flow region is the most important mechanism 

during the reflood. After the blowdown before the reflood, the reactor core experiences an 

almost adiabatic heat-up. As the water enters the core, water droplets produced by liquid 

entrainment and fragmentation contact the hot rods and begin core cooling. Rod temperatures 

attain their maximum values, and then turn around. 

Spacer grids are structural members in the reactor core that support the fuel rods at a 

predetermined rod-to-rod pitch. All fuel assemblies have grid spacers at the same elevation 

across the core. Since the grid represents a perturbation to the flow channel, its thermal

hydraulic effects are important. The grid reduces the fuel assembly flow area by contracting the 

flow and then expanding it downstream of each grid. During normal operation, the flow is 

single-phase liquid. The flow contracts and accelerates within the grid and then expands 

downstream, thereby disrupting and reestablishing the fluid and thermal boundary layers on the 

fuel rod. This mechanism increases the local heat transfer within and downstream of the grid. 

Several single-phase experiments have shown that the single-phase heat transfer downstream 

of a grid spacer can be modeled as an entrance effect phenomenon in which the abrupt 

contraction and expansion result in the establishment of a new boundary layer downstream of 

the grid. During a large-break LOCA, the flow is no longer single-phase, and the 

thermal/hydraulic effects of the grid spacer must be evaluated in terms of two-phase flow. 

During reflood, grid spacers interact with the fuel rods and the two-phase flow in the flow 

channels. From various two-phase flow studies, it has been observed that grid spacers caused

-6-
enhancement of the heat transfer, mainly by desuperheating the vapor. The enhancement of 

heat transfer results from three mechanisms: 

(1) grid rewet 

(2) convective enhancement 

(3) droplet breakup 

Since the grids are unpowered, they can quench before the fuel rods. If the grids rewet, they 

create additional liquid surface area that can help to desuperheat the vapor in the non

equilibrium two-phase droplet flow. The region surrounding a wetted grid will have a higher 

interfacial heat transfer coefficient (compared to droplets) because the velocity for the vapor 

flow relative to the liquid film is larger. In addition to desuperheating the vapor, the liquid film 

will evaporate, resulting in higher steam flow and convective heat transfer. The increased 

interfacial heat transfer between the film on the grid and the vapor flow and the generation of 

additional saturated vapor from the liquid film on the grid will result in lower vapor temperatures 

downstream of the grids. In addition, the grids can also break up the entrained droplets into 

smaller ones, thereby increasing the surface area for evaporation. The evaporation of the 

smaller droplets will provide an additional steam source which increases the convective heat 

transfer coefficient. 

Thermal-hydraulic processes in a reactor during normal operation are basically one

dimensional. During the blowdown portion of a large-break LOCA, the process is one

dimensional. During reflood, multidimensional flow patterns occur in the vessel core and upper 

plenum for the following reasons: 

" Nonuniformities in core power distribution, intermittent positioning of the primary loops 

circumferentially around the vessel, and differences in resistance to flow through the 

intact and broken hot legs naturally tend to promote multidimensional effects. 

" As the flow passes through the upper plenum, it must flow around several structural 

elements. Furthermore, the flow behavior of a collected pool in the upper plenum may 

be highly three-dimensional because of the tendency of the water pool to collect in 

dead spaces and in regions near walls and structures and in certain portions of the 

upper plenum as a result of the flow pattern toward the hot legs 

Liquid deentrainment occurs when water entrained in the steam is removed from the steam flow 

at other places in the core, the upper plenum, or beyond. Deentrainment occurs as a result of 

gravitational or inertial forces. Deentrainment is enhanced when the flow slows down because 

of a flow area increase, or when the flow changes direction to pass around obstructions or 

structures or to turn out through a nozzle. This deentrainment removes entrained water from 

the two-phase flow mixture, and deentrained water that accumulates in the core and upper 

plenum provides a source of water for core cooling that supplements the cooling obtained by 

bottom reflood. This enhances core cooling near the fuel assembly grid spacers, and is also 

important for the upper regions of the core. 

The extent of a liquid carryover to the upper plenum is expressed by the liquid carryover rate 

fraction; that is, the ratio of a total liquid mass flow rate out of the core to the liquid mass 

entering the core. A liquid carryover reduces the rate of accumulation of water within the core, 

and provides a source of water for deentrainment in the upper plenum.

-7-
Pool formation refers to the collection or behavior of water in the upper plenum during reflood. 

The source of this water is entrained water carried up from the core, which is then deentrained 

in the upper plenum. The upper plenum contains a significant number of internal structures that 

act as a steam separator to deentrain liquid in the two-phase flow. In addition, the decrease in 

flow velocities from the core to the upper plenum allows gravitational separation of liquid from 

steam. The flow behavior of a collected pool may be highly three-dimensional because of the 

tendency of the water pool to collect in low-flow areas and in regions near the walls and 

structures. Upper plenum pool formation is important to LOCA/ECCS performance for the 

following reasons: 

(1) The pool provides a source of additional cooling water for the core. Additionally 

deentrainment and pool formation reduce the carryover to the steam generator. 

(2) The pool can reduce the flooding rate by creating a static head pressure drop in the 

upper plenum. 

(3) Water flows downward from the upper plenum into the core under the influence of 

gravity. This downward flow is opposed by steam flow upward from the core. The 

downward flow tends to be nonuniform or channeled as a result of upward steam flow. 

The radial power profile in the core and the variations in stored heat developed by 

uneven quenching cause development of nonuniform upward steam flow. 

Top-down quenching occurs when the top portions of the fuel rods quench before the portions 

just below. Top-down quenching provides additional cooling to the upper core and, in some 

cases, to the intermediate and/or lower core that primarily obtains the cooling from bottom 

reflood. In standard designs, top-down quenching may occur from the fallback of water that is 

deentrained at the top of the core or in the upper plenum. Different reactor designs (such as 

upper head injection and AP600) provide substantial top-down quench during reflood using 

different processes. This results in quenching of the uppermost part of the fuel rods sooner 

than would occur from the propagation of the bottom quench front. In standard designs, top

down cooling does not generally extend to the hot spot in the core, which is typically at about 

two-thirds of the core height. 

Radial power distribution causes radial gradients in steam generation and liquid entrainment. 

Thermal-hydraulic effects resulting from radial power distributions are important because they 

lead to crossflows between the low-powered assemblies and the high-powered assemblies. 

The variations in the quench front elevation across the core yield a radial profile that resembles 

the inverse of the radial power profile (i.e., lower quench front elevation in higher-powered 

bundles and conversely). Radial variations in the vapor generation cause variations in the 

entrainment and liquid fraction above the quench front (i.e., greater entrainment in higher 

powered bundles). There is relatively free lateral communication in the core. Horizontal 

pressure differences and crossflows are established by the radial power distribution. The net 

effect is an increase in the water flow rate and cooling of the higher-powered bundles 

(assemblies) compared to the cooling given the average flow rate entering the bottom of the 

core. Such crossflows tend to prevent some hot spots from developing. 

At the beginning of a reflood, a manometer-type oscillation can occur. The name derives from 

the general observation that the core and downcomer levels oscillate like the levels in a U-tube 

manometer that has been perturbed. The U-bend manometer is formed by the downcomer 

water column and the core water column, and is connected by the lower plenum. This

-8-
phenomenon was observed in several facilities including the loss-of-fluid test (LOFT), 

Semiscale, slab core test facility (SCTF), cylindrical core test facility (CCTF), and full-length 

emergency core heat transfer (FLECHT). Manometer oscillations are initiated by the sudden 

rapid introduction of ECCS water into the core inlet associated with accumulator injection. As 

water penetrates into the core, rapid generation of steam occurs and pressure begins to build 

up in the upper portion of the core. This tends to push water from the core into the downcomer. 

As the core water level goes down, the steam generation rate and the back-pressure are 

reduced. This reduced pressure, in conjunction with the increased downcomer level, causes a 

reversal and the core level rises again. The subsequent upward movement of the core level 

tends to overshoot the equilibrium location, leading again to increased steam generation. This 

causes an oscillatory movement of both the collapsed core water level and the downcomer 

level. The importance of manometer oscillations is a potential increase or decrease in the core 

cooling rate, depending upon which cycle of the oscillation the core is in. The increased cooling 

rate is caused by more rod surface area being covered with water during the upward surge of 

the oscillations. The decreased cooling occurs when the collapsed water level in the core is 

low. 

As ECC water enters the hot core, some of it will turn into steam. The resulting increase in 

pressure tends to suppress the reflooding velocity in the core. This retarding effect of steam on 

the reflood rate is referred to as "steam binding." Another source of back-pressure retarding the ECC injection rate is liquid carried to the steam generators and evaporated as a result of 

heat transfer from the secondary side to the primary side. Finally, the steam binding effect may 

occur as a result of the leakage of hot secondary liquid to the primary side through a concurrent 

steam generator tube rupture. 

In Babcock & Wilcox reactor designs, eight check valves are located in the core barrel between 

the upper plenum and the downcomer annulus above the inlet nozzles. These valves are held 

shut during normal reactor operation by the discharge pressure from the reactor coolant pumps. 

In the event of a LOCA, the vent valves open during the refill, and steam flows through the vent 

valves to the ECCS-induced condensation point in the downcomer annulus. During the reflood, 

steam generated in the core flows through the vent valves to the system break and, thus, 

bypasses the loop flow resistance. This relieves the steam pressure above the core, and allows 

an increased reflood rate. 

Appendix K to 10 CFR Part 50 requires that any effect of a fuel rod flow blockage must be 

explicitly accounted for in safety analysis calculations when the core reflood rate drops below 

1.0 in/s. The effect of a flow blockage on heat transfer is a combination of two competing 

thermal-hydraulic phenomena. Specifically, flow area reduction causes flow acceleration, 

droplet breakup, improved mixing, and steam desuperheating, and consequently increased 

heat removal rates. In addition, flow bypass reduces the flow rate through the blocked region. These two effects are dependent on blockage geometry and distribution. 

2.2 Predictive Models and Experimental Facilities 

As discussed in the previous section, many different phenomena are involved during a reflood 

process. Some of these phenomena occur because of different reactor designs. However, 

most of these phenomena occur in all designs because they are fundamental to the reflood 

process. Their dominance may be different in different reactors or different types of LOCAs. 

These phenomena are listed below:

-9-
1. bottom quench 

2. liquid entrainment and precursor cooling 

3. vapor superheat 

4. heat transfer 

5. liquid deentrainment 

6. top-down quenching 

7. radial power effects and three-dimensional flow 

8. manometer oscillations 

9. initial cladding temperature effects 

10. liquid carryover from the core 

11. upper plenum geometry effects and pool formation 

12. liquid carryover to the hot leg 

13. steam binding 

14. grid spacer effects 

15. flow blockage effects 

16. vent valve effects 

The models in thermal-hydraulic codes should be capable of predicting the above phenomena 

with acceptable accuracy. In most PWRs in the United States, the ECC is injected through cold 

legs, and passes down the downcomer into the lower plenum and up the core. Hence, "bottom 

reflood" of the core after a large-break LOCA is the predominant mode of quenching the core, 

and many experiments have been conducted to study the thermal-hydraulic phenomena of 

bottom reflood. Depending on the reflooding rate, different flow regimes may develop ahead of 

the quench front as the core is quenched from the bottom up. In most U.S.-designed reactors, 

the first 15 phenomena may occur during a postulated LOCA. In Babcock & Wilcox plant 

designs, the last phenomenon related to vent valve operation also becomes important. 

Much of the U.S. database for reflood heat transfer comes from the PWR FLECHT and Full

Length Emergency Core Heat Transfer System Effects and Separate Effects (Flecht-Seaset) 

programs (Refs. 3 and 4). The Semiscale program has also produced a data set on core 

thermal-hydraulics of reflood (Refs. 5 and 6). In Japan, the earlier reflood experiments and 

extensive CCTF and SCTF tests by the Japan Atomic Energy Research Institute (JAERI) 

provide a database of large-scale test facilities. The Japanese CCTF and SCTF test facilities 

are the largest reflood test facilities in the world. Similar smaller reflood test facilities exist in 

France (Ref. 7), Germany (Refs. 8 and 9), and the United Kingdom (Ref. 10). There are also 

small-scale test facilities at several universities, such as the University of California at Los 

Angeles (UCLA) (Ref. 11), University of New York at Stony Brook (Ref. 12), and the University 

of California at Berkeley (Ref. 13). These small-scale university test facilities help to develop 

models and understanding that can help in interpreting the behavior and performance of the 

larger-scale tests. 

2.3 Discussion of the Phenomena Observed in Flecht-Seaset Tests 

This report presents assessment studies of the TRAC-M code using the data from the Flecht

Seaset facility performed using the 161-rod unblocked test section. This section discusses 

which of the phenomena listed in the previous section can be assessed using Flecht-Seaset 

testing. The Flecht-Seaset facility simulates only one-dimensional reflood phenomena. In the 

Flecht-Seaset program, visual observations could be made through view ports located at 

several elevations of the bundle. Since the hydraulics and heat transfer are closely coupled, 

the interactions among system parameters (such as flooding rate, pressure, subcooling, initial

-10-
cladding temperature, and mass effluent fraction) have been studied. Among these 

parameters, the flooding rate was found to be the most influential. Different flooding rates gave 

rise to different flow regimes which, in turn, determined the increase in the cladding 

temperature and the quenching of the cores. A detailed analysis of the data was presented in 

Ref. 4, and the important observations and a description of the phenomena are summarized 

herein. 

2.3.1 Bottom Quench, Transition Zone, Liquid Entrainment, Precursor Cooling, and Heat 

Transfer 

Figures 2.3 and 2.4 illustrate the effect of variations in the flooding rate on the heat transfer 

coefficient and peak temperatures at the mid-plane elevation for some of the runs in the 161

rod bundle unblocked test section used in the Flecht-Seaset program. The figures show an 

orderly increase in heat transfer at all times as the flooding rate increases. Temperatures show 

a more rapid turnaround, and peak temperatures were lower as the flooding rate was 

increased. 

Figures 2.5 and 2.6 show the quench front progress. As expected, the higher the flooding rate, 

the faster the quench front moves up. Figure 2.6 shows the trend of increasing temperatures 

with decreasing flooding rates. These trends are expected to occur. 

For the low flooding rates, a transition zone, or a froth region exists above a quench front. In 

the transition zone the flow is not quite dispersed as shown in Fig. 2.2. Hence, the heat 

transfer mechanism in the transition zone differs from that in the dispersed flow and from that 

below the quench front. In order to apply the proper heat transfer mechanism when calculating 

heat transfer from the hot wall to the two-phase flow, knowledge of the length of this transition 

zone is required. Murao (Ref. 14) observed that temperature history curves during reflood 

show a sudden change to a steeper slope sometime before quench (see Period II in 

Figure 2.7). This slope change is related to the flow regime change ahead of the quench front. 

Similar behavior was observed in the FLECHT tests, and a method was developed to calculate 

the length of the transition zone above the quench front. It is possible to distinguish three 

different flow regimes as indicated in Figure 2.7. The first occurs when there is a dispersed 

droplet flow at the elevation (Period I), after a short period of single-phase steam flow. The 

second occurs when the transition zone above the quench front is at the elevation (Period II), 

and the third occurs when the quench front is above the elevation (Period Ill). These different 

flow regimes are distinguished in the heat transfer coefficient curves of Figure 2.7. The 

relatively moderate increases in heat transfer during Period I are followed by a rather sharp 

increase in heat transfer during Period II, and the quench produces a further sudden increase. 

It is observed that the transition zone could be about 0.3 m for a reflooding rate of 0.0254 m/s 

(1.0 in/s), and could increase to about 1 meter for a reflooding rate of 0.076 m/s. In these tests, 

the transition zone length is a direct function of the reflood rate. Figures 2.8 and 2.9 illustrate 

the progress of fronts relative to quench data for slow and fast reflood runs, respectively.

-11-
- - u r , 1 IVui , IOU ITm rU/ ko.i inJ •u ; r -

Run 31805, 20.6 mmlsec (0.81 injsec) 01N 

0 300 

S-50 R 

40 - 2i 

.AJ011 

oC I * 0 

-250 4M00 

Time (sec) 59441 

Figure 2.3 Flooding Rate vs Heat Transfer Coefficient 0-12
1750 I I 

3182

Run 31805, 20.6 mm/sec (0.81 in./sec) 

1500 - Run 31504, 24.6 mm/sec (0.97 in./sec) 

•---- Run 31203, 38.4 mm/sec (1.51 inJsec) 

Run 31302, 76.5 mm/sec (3.01 in./sec) 2500 

Run 31701, 155 mm/sec (6.1 inJsec) 

1250 

•1000 •_•i5-%- , - 2000 

1500 

500 -1000 

250- 1 

0 - I I 32 

0 100 200 300 400 500 

Time (sec) 69447 

Figure 2.4 Flooding Rate vs Clad Temperatures

-13-

I I i 3182 i
_0 A 0.39

Flooding rate (in.s) 

1.97

V

10 50 

Flooding rate (mtnis)

100 500

Peak Inlet Initial 

Pressure Power Subcooling Temperature 

Symbol Test IMPa(psia)] [kW/m (kW/ft)] IK(°F)] [K(°F) 

A FLECHT SEASET 0.28 (40) 1.3 (0.40) 351(140) 1155(1620) 

4. FLECHT SEASET 0.28 (40) 2.3 (0.70) 342(125) 1144(1600) 

o SKEWED 0.28 (40) 2.3 (0.70) 351(140) 1144(1600) 

o COSINE 0.28 (40) 3.1 (0.95) 351(140) 1144(1600) 

0 COSINE 0.14 (20) 4.1 (1.24) 295 (40) 1144(1600) 

* COSINE 0.10 (15) 2.3 (0.69) 286 (23) 1144(1600) 

* COSINE 0.28 (40) 2.4 (0.73) 351(140) 1144(1600) 

* COSINE 0.41 (60) 4.1 (1.24) 352(143) 1144(1600) 

6 9371 

Figure 2.5 Flooding Rate vs. Quench Time

-14-

i nnn r- ,

3.94 19.7

500 -

E 

o S1o

50 H

100 I

I
Flooding rate (in./sec) 

3.94

100 

Flooding rate (mrn/s)

Pressure 

Test [MPa (ps12)l 

FLECHT SEASET 0.28(40 

FLECHT SEASET 0.28(40) 

SKEWED 0.28(40) 

COSINE 0.28(40) 

COSINE 0.28(40) 

COSINE 0.28(40) 

COSINE 0.28(40) 

SKEWED 0.14 (20)

Peak Inlet 

Power Suboooling [kWlm(kW/ft)] [1(('1)] 

1.3(0.40) 351(140) 

2.3 (0.70) 342(125) 

2.3 (0.70) 351(140) 

3.1 (0.95) 295(40) 

3.1 (0.95) 295(40) 

2.4(0.73) 351(140) 

2.5(0.75) 351(140) 

2.3 (0.70) 323(90)

Initial 

Temperature Elevation 

[F(°F)] lm(in.)] 

1155(1620) 1.83 (72) 

1144(1600) 1.83(72) 

1144(1500) 3.05 (120) 

1144(1600) 1.83(72) 

1144(1600) 1.98(78) 

1144(1600) 1.83 (72) 

1144(1600) 1.98(78) 

1144(1600) 3.05 (120) 

6 9423

Figure 2.6 Flooding Rate vs. Temperature Rise

-15-

A

y

Soo 1-

0 

I-. 

E 

a.

1001-

39.4 

-958 

LL 

- 238 

E 

so0F- -148

"0 ~V 10

- 94

1000

Symbol 

0 

0 U 

0 

<>

m ! . |

I

20-

! !
1•,• I I J I I I2282 2000 

1000 

750 

1500 L 

C L 

0 

E. 

1000 CE 

10 

oo 

S o" 

250- 

-0 

05 I 5o13 

o 1 32 

.300 I 52.8 a 

250- Z 

Z' 200 -- 

- E - -0 

1- 

3 0 

150 

0 

0 

20 0 

100 

0 

C 

II 

0 0 

0 100 200 300 400 500 600 700 800 

'rime (Sec) 

6 9529 

Figure 2.7 Clad Temperature and Heat Transfer 

Coefficient in Different Zones

-16-
147.6 

120 

96 

48. 

.2 

24

0 10 

0 100 200 300 400 500 600 689 

Quench time (Sec) os O 

Figure 2.8 Transition Front Relative to Quench Data in 

Slow Reflood

-17-

3.75 

3.50 

3.00 

250 

C 

0 .50 

1.00 

0.60
4 AT7

3.75 1 1 1 1l ,I. 

3.50 o o 

0 c:oo 0 

3.O - l 120 

0 a0 

2.50 - m96 

.2 2.00 Quench data C 

0 72 

1.50 

0 48 

1.00 0 Transition front 

0.50 0 

24 

0 

0 1I I I I 0 

0 100 200 300 400 500 600 700 800 

Quench time (sec) 9464 

Figure 2.9 Transition Front Relative to Quench Data in Fast 

Reflood

-18-
Quench front propagation during reflood may also be examined from bundle average void 

fraction data that is reduced from pressure drop measurements. Figure 2.10 shows the void 

fraction changes for the 0.0254 m/s (1.0 in/s) reflooding case from the Flecht-Seaset tests. It is 

seen that the void fraction between 0.0 m and 0.6 m elevations rapidly decreases from a value 

of 1.0 to a value of zero, indicating that the quench front moves up very quickly and a liquid

only region is established at the bottom of the core. At higher elevations, the rate of change in 

the void fraction is lower than those at low elevations because of the accumulation of steam 

generated at low elevations. 

2.3.2 Liquid Carryout from the Core 

The liquid carryout is an important parameter during reflood since the liquid carried out has not 

been used in quenching the core. Figure 2.11 presents the variation of a time-integrated liquid 

carryout fraction versus time. The higher the flooding rate, the more rapid the carryout fraction 

approaches its asymptotic value. This is expected, simply because of the accelerated pace of 

events occurring at higher flooding rates. The same trend has also been observed in the 

previous FLECHT tests. The principal effect of increasing flooding rates is to proportionately 

increase the mass flow above the quench front. 

2.3.3 Fuel Assembly 

In the Flecht-Seaset program, the bundle geometry was changed from that in the FLECHT 

program with 15X1 5 PWR core geometry to reflect the 17X1 7 PWR core geometry. The 

program had several subprograms, including some tests with blocked bundles. The 161-rod 

unblocked bundle tests serve as a reference. for comparison with the blocked bundle, as well as 

for comparison with FLECHT which contained 15X1 5 core geometry. Comparison of test data 

indicated that the change from 15X15 to 17X17 bundle geometry did not change the reflood 

characteristics. Results indicate that uncertainties attributable to other parameters are greater 

than those introduced by changes in bundle geometry. 

2.4 Summary of Code Requirements 

The top-level requirement is, no doubt, that the equations in the code represent appropriate 

physics in order to predict all of these phenomena. An assessment of whether this requirement 

has been satisfactorily fulfilled is performed by reviewing the code models. If the review 

indicates that this requirement is satisfactorily fulfilled, the next level of requirements is that 

trends and important parameters should be predicted with acceptable accuracy. Primary trends 

are listed below: 

(1) The flooding rate has a first order effect on thermal-hydraulic phenomena. Clad 

temperatures and time of quenching decrease with increasing flooding rates. Flow 

regimes are dependent upon flooding rates and temperatures. They play an important 

role in determining cladding temperatures and time of quenching. At the quench front, 

the stored energy of the rod is released over a short length. Rod temperatures 

decrease suddenly as the rod quenches.

-19-
c 0.7 .1 

_; 1.52-1.83 m (60-72 in.) level 

0 %0.6 

>00.5 

0.4 1.22-1.52 m (48-60 in.) level 

0.3 

0.2 0.91-1.22 mn (31-48 in.) level 

0.1 0.61-0.91 m (24-36 in.) level 

0 - I I ! 

0 100 200 300 400 500 600 700 

Time (s) 6 940 

Figure 2.10 Void Fraction Variation with Time During Reflood

-20-
1.00

a/ 

0 

_ 0.50 

-- Run 31203, 38.4 mm/sec (1.51 in.Isec) 

-• / - -Run 31302, 76.5 mm/sec (3.01 in./sec) 

0.25 ! -- -I Run 31504, 24.6 mmlsec (0.97 in./sec) 

Run 31701, 155 mm/sec (6.1 in.isec) 

Run 31805, 20.6 mm/sec (0.81: in./sec) 

0 . I 

0 100 200 300 400 500 

Time (sec) 6 9318 

Figure 2.11 Carryout Fraction vs. Time

-21-
(2) The flow regimes ahead of the quench front are characterized by a transition or froth 

region. The length of this transition region is a direct function of the reflooding rate. It 

is observed that the higher the reflooding rate, the longer is the transition zone. The 

heat transfer mechanism in the transition zone is different from that in the dispersed 

flow regime zone and that in the quenching zone. 

(3) Bottom reflood progresses very quickly during the onset of reflood; however, intense 

steam generation soon retards the overall progression of the quench front to a 

relatively uniform progression. Nevertheless, good core quenching rates are achieved 

even for flooding rates of 1 inch per second. Some of the liquid is carried out, and is 

not available to cool the core. The amount of liquid carried out is affected by upper 

plenum geometry and other system effects. 

Important parameters are listed below: 

1. clad temperatures 

2. vapor fraction and differential pressures at different elevations 

3. steam temperatures 

4. liquid and steam carryout 

Acceptable accuracies can be judged from the spread of experimental data. An example of 

success metrics for reflood phenomena for Flecht-Seaset Run 31504 is developed in Section 4.

-22-
SECTION 3 

SYSTEM DESCRIPTION 

3.1 Facility Configuration 

The Flecht-Seaset test facility was developed by modifying the existing FLECHT test facility. A 

detailed description of the facility is presented in Ref. 3. Figure 3.1 shows the configuration of 

the modified facility used to conduct unblocked bundle forced reflood tests. The facility was 

further modified for gravity reflood tests by adding a downcomer. For steam cooling tests, the 

facility was again modified to take steam from outside sources. 

The facility modifications applicable to all tests are as follows: 

(1) a new low-mass housing test section and upper and lower plenums 

(2) an upper plenum baffle to improve liquid carryover separation 

(3) a 161-rod bundle and related instrumentation 

(4) a new pressurized-water supply accumulator and injection line with three rotameters 

and a turbine meter to measure injection rates from 10 mm/sec (0.4 in/sec) to 150 

mm/sec (6.0 in/sec) in forced flooding tests, and up to 6.49 kg/s (14.3 Ib/s) in gravity 

reflood tests 

(5) a close-coupled carryover tank connected to the test section upper plenum with 

a storage capacity of 65.8 kg (145 Ib) 

(6) a steam separator with a storage capacity of approximately 193 kg (425 Ib), a 

capacity of 0.315 kg/sec (2500 lb/hr), and a liquid collection tank with a volume 

of 9.5 kg (21 Ib) to collect liquid entrained in the exhaust line 

(7) exhaust piping with a system pressure control valve and an orifice plate 

flowmeter to measure exhaust steam flow rates 

(8) a commercially available electric steam boiler with a capacity of 0.016 kg/sec 

(125 lb/hr) of steam production to establish the initial loop pressure and 

temperature 

3.1.1 Forced Reflood Tests and Testing Procedure 

The following general procedure was used to establish initial test conditions and perform a 

typical Flecht-Seaset unblocked bundle reflood test. The accumulator was filled with water and 

heated to the desired coolant temperature of 530C (127 0F) nominal. The boiler was turned on 

and brought up to a nominal gage pressure of 0.62 MPa (75 psig). The carryover vessel, 

steam separator, separator drain tank, test section upper plenum, and test section outlet piping

-23-
STEAM 

BOI LER 

LINE

(

(D 

i-rJ 

(D 

I--) 

(D1 

(D 

t-I

Fr 

ct 

I

C0 

(1) 

-t

H

0•

LEGEND 

-EXHAUST 

0 

TER __ 

PPLY 

-i0

DIFFERENTIAL PRESSURE 

STATIC PRESSURE 

FLUID THERMOCOUPLE 

WALL THERMOCOUPLE 

TURBINE METER 

ORIFICE PLATE FLOW METER

- CZj3- ROTAMETER 

SI UNITS 

I GPM = 6,31 x 10-5

' 3 /sec 

I FOOT 0,305 METER

N

LIQ. LEVEL 

CARRYOVER TANK TEST SECTION' 

'ALL INSTRUMENTATION IS NOT SHOWN ROTAMETERS
(located before the separator) were heated while they were empty to slightly above the 

saturation temperature corresponding to the test run pressure. The exhaust line between the 

separator and exhaust orifice was heated to 2600C (5000F) nominal, and the test section lower 

plenum was heated to the temperature of the coolant in the accumulator. The above 

component heating was accomplished using clamp-on strip heaters. 

The test section, carryover vessel, and exhaust line components were pressurized to the 

desired system pressure of 0.14 to 0.41 MPa (20 to 60 psia) by valving the steam boiler into the 

system, and setting the exhaust line control valve to the desired pressure. The coolant in the 

accumulator was pressurized to 2.76 MPa (400 psia). Water was then injected into the test 

section lower plenum until it reached the beginning of the heated length of the bundle heater 

rods. Coolant was circulated and drained to ensure that the water in the lower plenum and 

injection line were at the specified temperature before the run. Power was then applied to the 

test bundle and the rods were allowed to heat up. When the temperature in any two designated 

bundle thermocouples reached a preset value between 260'C and 8710C (500°F to 16000 F), the 

computer automatically initiated flooding and controlled power decay. Solenoid valves (not 

shown in Figure 3.1) and a hydraulic control valve controlled the coolant injection into the test 

section. During a run, the boiler was valved out of the system, and the pressure was 

maintained at the desired level by the control valve located in the exhaust line. Liquid effluent 

leaving the test section was separated in the upper plenum and collected in a carryover tank. A 

baffle assembly in the upper plenum was used to improve liquid carryout separation and 

minimize liquid entrainment into the exhaust vapor. The separator located in the exhaust line 

was used to separate any remaining entrained liquid carryout from the vapor. Dry steam flow 

leaving the separator was measured by an orifice meter before exhausting to the atmosphere. 

In order to ensure that a single-phase flow measurement was made, the piping upstream of the 

orifice was heated to a temperature well above the saturation. After all designated heater rods 

were quenched, as indicated by the rod thermocouples, power to the heater rods and coolant 

injection to the test section were terminated. The entire system was depressurized by opening 

the exhaust control valve, and the Computer Data Acquisition System (CDAS) was deactivated. 

Water stored in all components was drained and weighed. 

After runs 30223 through 30817 had been conducted, it was noticed that some water entered 

the bundle steam probe lines. In order to prevent this occurrence in further tests, a modification 

was made to the above test procedure. This modification consisted of power pulsing the bundle 

to approximately 260'C (500'F) before applying the full power. Power pulsing vented the 

bundle steam probe lines before starting a new run. This procedure was used to dry out the 

bundle thimbles and steam probes. 

3.1.2 Gravity Reflood Tests and Testing Procedure 

During the test series, further modifications were made to the facility to conduct gravity reflood 

tests. Some of the modifications were made to simulate reactor system loop behavior. These 

modifications consisted of connecting a downcomer to the lower plenum, moving the injection 

line from the lower plenum to the bottom of the downcomer, installing a resistance orifice plate 

between the test section outlet pipe and the inlet flange of the entrainment separator to 

simulate the hot leg resistance, venting the top of the downcomer to the entrainment separator, 

and installing additional differential pressure cells. 

The procedure used to perform the forced reflood tests was also used to perform the gravity 

reflood tests, with one exception. Specifically, after flooding was initiated, the rate was

-25-
adjusted, if necessary, to ensure that the level in the downcomer did not go above the 4.88-m 

(192-in) elevation. 

3.1.3 Steam Cooling Tests and Testing Procedure 

To obtain steam flow heat transfer data in the rod bundle geometry, a series of low-temperature 

steam cooling tests was conducted. For these tests, the facility was modified by connecting the 

lower plenum of the test section to the Flecht-Seaset steam generator facility boiler with a 

steam line. 

The steam cooling tests were initiated by injecting steam at a high flow rate [approximately 0.3 

kg/sec (0.6 Ib/sec)] into the bundle. When the system reached the saturation condition, the 

steam flow rate was adjusted to the specified value. When the steam flow was stabilized, a 

constant bundle power was applied, and the CDAS was activated. The test was terminated 

when a steady-state condition was achieved. Both transient and steady-state data were 

recorded. 

3.2 Heater Rod Bundle 

A cross-section of the test bundle is shown in Figure 3.2 with the heater rod instrumentation 

groups. The bundle comprised 161 heater rods (93 uninstrumented and 68 instrumented), 

4 instrumented thimbles, 12 steam probes, 8 solid triangular fillers, and 8 grids. The triangular 

fillers were welded to the grids to maintain the proper grid location. The fillers also reduced the 

amount of excess flow area from 9.3 to 4.7 percent. Locations of electrically heated rods, 

thimbles, and steam probe tubes are identified by alphanumeric coordinates indicated in 

Figure 3.2. 

Each heater rod is identified by a circle. Some heater rods are instrumented by thermocouples 

welded inside the cladding. If a rod is instrumented, it is identified by a group number (shown in 

the circle), which identifies the locations of various thermocouples inside the rod. A cross

section of a heater rod with thermocouple locations is shown in Figure 3.3. The heater rods 

were designed to provide an axial cosine-shaped power profile. As a run started, the computer 

provided decay power to the rods. Figures 3.4 and 3.5 show the axial power shape and 

variation of decay power with time, respectively. 

Steam probe tubes and thimbles are identified by double circles in Figure 3.2. Steam probe 

tubes contain steam probes for measurement of superheated steam temperatures. Steam 

probes were located in the bundle and in the test section outlet pipe. The bundle steam probes 

were located in the thimble tubes (Figure 3.6). The probe was designed to separate moisture 

from the high-temperature steam, and aspirate the steam across the thermocouple and into an 

ice bucket. Because of a large number of probes, the probes were controlled to limit the 

amount of steam aspirated from the test section. This was accomplished by manifolding the 

aspiration lines for common elevations of steam probes together, and closing the lines when the 

particular elevation had quenched.

-26-
The probe in the outlet pipe was installed in the elbow of the test section outlet pipe. This 

probe, which measured the temperature of the steam leaving the test section, was designed to 

measure steam temperature in the same manner as the bundle probes. The end of the 

thermocouple was formed so that it did not touch the walls of the 6.4-mm (0.25-in.) diameter 

tubing. Thermocouples were welded to the inside of thimble walls to measure their 

temperature.

A B C D E F G H

WINDOW 

1800

1 

2 

3 

4 

5 

6 

7 

8 

9 

10 

11 

12 

13 

14 

15

00

I J K L M N 0

PRESSURE 

TAP

HEATER ROD

WINDOW

Figure 3.2 Flecht-Seaset 161-Rod Test Bundle

-27-
AS VIWE FROM BOTO EN GOF O 

T/C~ 8GR.2 /C6GI 

T/C~~ ~~ 6 RR IC3GR 

" T/C GR.1 • T/ T/ 4GR.5 

ASVEWE FRO BOTOMEN OFMS-0OD 

Figure 3.3 Heater Rod Thermocouple 

Locations 

ELEVATION (ft) 

0 I 2 3 5 C 6 7 8 9 I0 II 2 

IAO 8I I I O6 

1.66 T IFLECH HEATER ROD 

Si.I% 

0': 1 2.3 3.3 4 7 8 9 0 11 

1.2 

1.6 

C5) 

0.88 

C 0.8 

CL 

FLECOT REFERENCE COSINE 

0.2 

0.0 

0.00 0.25 0.50 0.75 1.00 1.25 1.50 1.75 2.00 2.25 2.50 2.75 3.00 3.25 3.50 

ELEVATION (m) 

Figure 3.4 Flecht-Seaset Axial Power Shape

-28-
1.0

0.9 

0.8 

o 0.7 

0.6 

0.5 

0.0 

0 100 200 300 400 500 600 700 800 900 1000 

TIME AFTER FLOOD (SEC) 

Figure 3.5 Decay Power Curve (ANS + 20%) 

3.3 Discussion of System Design Features 

The following paragraphs describe the major design features that were used for the Flecht

Seaset unblocked bundle test series. 

3.3.1 Low-Mass Housing 

A low-mass housing design was used in the previous FLECHT test series to minimize the 

housing effects. The effect of the housing on heater rod temperatures was studied by 

performing two tests with the same initial conditions except for the housing wall initial 

temperatures. The effect of the cold and hot housing on the rod surface temperature and the 

corresponding heat transfer coefficient, excluding the peripheral rods, was found to be 

negligible. In addition, the hot and cold housing did not affect the quench front along the 

bundle. It was concluded that performing reflood tests with an unheated low-mass housing was 

acceptable, since a cold housing did not significantly affect the reflood heat transfer and 

hydraulic behavior of the rod bundle.

-29-
SHIELD

THIMBLE 

WATER DRAIN LINE

N Il II 

SECOND SHIELD 

Figure D-6. Steam Probe Design 

Figure 3.6 Steam Probe Design

-30-

h,

1I uu.3 . 9

N

STEAM 

FLOW i

STEAM 

FLOW I
The housing was constructed of 304 stainless steel, and is rated for 0.52 MPa (60 psig) at 

816°C (1500°F). The design allowed for 1,000 pressure and temperature cycles. Two quartz 

sight glasses were located 180 degrees apart at the 0.91-, 1.83-, and 2.74-m (36-, 72-, and 

108-in) elevations for viewing and photographic study in some selected runs. The sight glass 

configuration allowed backlighting as well as frontlighting for the purpose of photographic 

studies. The sight glasses had clamp-on heaters to raise the quartz temperature at the 

initiation of reflood to approximately 260'C (500°F). This eliminated the formation of a liquid 

film on the quartz during a run. 

3.3.2 Bundle Differential Pressure Cells 

The test section differential pressure cells provided data used to calculate a mass balance and 

bundle void fraction. Low-range [+ 0.0069 MPa (+ 1 psi)] pressure transducers were used to 

improve the accuracy of the data. The cells were located every 0.30 m (12 in) along the test 

section. The differential pressure cell manifold was carefully bled to eliminate any trapped air 

and, thus, improve the repeatability of the readings. 

3.3.3 Pressure Control 

Maintaining a constant upper plenum pressure had been a difficult problem in previous 

FLECHT test programs. In an effort to reduce the pressure oscillations, several modifications 

were made to the facility for the FLECHT test series. The first modification was to increase the 

volume of the steam separator to help reduce the magnitude of the oscillations. The second 

modification was to replace the air-operated globe exhaust control valve with an air operated V

ball control valve. The V-ball valve had a larger flow loss coefficient, and a more linear 

operating characteristic. 

3.3.4 Facility Operation 

The test facility was designed for automatic operation whenever critical functions required a 

high degree of sophistication, safety, or repeatability. The Computer Data Acquisition System 

(CDAS) was the heart of the operation. It monitored, protected, and controlled the facility 

operation, and collected data. It also controlled pressure, power, and flow during the test. In 

addition, the CDAS software monitored critical safety parameters during the test, and used 

corresponding outputs to run the test. 

3.3.6 Power Measurement 

The technique of bundle power measurement used in the FLECHT test series was used again 

for the present unblocked bundle test series. The bundle power measurement systems were 

improved by the addition of a secondary independent power measurement system, and the 

adoption of a system calibration. Details of the secondary power measuring system are 

presented in Ref. 3. 

3.3.6 Coolant Injection System 

The facility configuration (Figure 3.1) shows the coolant injection system, which consisted of a 

hydraulic valve for programmed flow and a turbine meter in series with three rotameters. The 

flow out of the rotameters went either to the lower plenum of the test section for forced reflood, 

or to the bottom elbow of the downcomer for gravity reflood. Solenoid valves were used to

-31-
initiate flooding and to channel the flow through the desired rotameter. A bidirectional turbo

probe in the downcomer crossover pipe measured the flow between the downcomer and the 

test section in the gravity reflood tests. In the steam cooling configuration, the steam was 

supplied to the lower plenum from the boiler. The steam flow rate was measured by a vortex 

meter used on the steam generator facility. 

3.3.7 Data Acquisition Systems 

Data acquisition was accomplished by the following systems. 

Computer Data Acquisition System (CDAS) 

The CDAS, the primary data collecting system used on the Flecht-Seaset unblocked bundle 

facility, consisted of a PDP-1 1 computer and associated equipment. The system could record 

256 channels of analog input data representing bundle and system and fluid temperatures, 

displaying temperature directly in degrees Fahrenheit. A Fluke recorder also recorded millivolt 

data from the test section differential pressure cells, allowing the operator to monitor the 

operation and repeatability of the differential pressure cells. 

Multiple-Pen Strip Chart Recorders 

Six Texas Instruments strip chart recorders were used to record 1) bundle power; 2) selected 

bundle thermocouple readings; 3) reflood line rotameter and turbine meter flows; 4) turbo probe 

flows; 5) accumulator, separator drain tank, housing, and carryover tank levels; 6) exhaust 

orifice differential pressure; and 7) selected bundle steam probe thermocouple readings. These 

recorders gave the loop operators and test directors immediate information on the test progress 

and warning in the event of any system anomalies. 

3.4 Instrumentation 

The instrumentation on the unblocked bundle facility was designed to measure temperature, 

pressure, flow, liquid level, and power. The temperature data were recorded by type K 

(chromel-alumel) thermocouples using 66°C (150'F) reference junctions. The liquid level and 

pressure data, both static and differential, were measured by balanced bridge strain gage 

transducers. Power input to the bundle was measured by Hall-effect watt transducers, which 

produce a direct current electrical output proportional to the power input. The fluid flow 

measurements were made by a turbine meter and rotameters for coolant injection, an orifice 

with a differential pressure transducer for exhaust steam flow, and a bidirectional turbo-probe 

for downcomer crossover flow during gravity reflood. 

3.4.1 Heater Rod Bundle Instrumentation 

The bundle instrumentation is described in Section 3.2. Not all of the available instrumentation 

in the bundle was recorded because of the 256-channel limit of the CDAS. However, 205 

thermocouples, including those pertaining to heater rods, thimbles, and steam probes, were 

recorded by the CDAS.

-32-
3.4.2 Test Section Instrumentation

The upper and lower plenums had fluid thermocouples. The upper plenum had a differential 

pressure transducer to measure liquid collection rates, a static pressure transducer to measure 

the test section pressure, and a static pressure transducer (not recorded) to provide a control 

signal to the V-ball exhaust valve. The low-mass housing had differential pressure transducers 

located every 0.30 m (12 in.) along the heated length of the bundle and one differential 

pressure transducer to obtain a liquid level and average bundle void fraction over the entire 

heated length. 

3.4.3 Carryover Tank Instrumentation 

The carryover tank was instrumented with two wall thermocouples and one fluid thermocouple. 

A differential pressure transducer was used to measure the liquid collected in the tank as a 

function of time. 

3.4.4 Steam Separator Instrumentation 

The steam separator was instrumented with three wall thermocouples and a differential 

pressure transducer for measuring the entrained liquid carryout rate. The steam separator 

drain tank was instrumented with two wall thermocouples, one fluid thermocouple, and a 

differential pressure transducer for measuring the entrained liquid carryout rate. 

3.4.5 Exhaust Pipe Instrumentation 

The exhaust pipe was equipped with a differential pressure transducer connected across an 

orifice plate for measuring single-phase steam flow. A fluid thermocouple and a static pressure 

transducer were used to determine exhaust steam thermodynamic conditions. Two wall 

thermocouples were used to record exhaust line temperatures during heat-up and the 

experimental run. A steam probe located in the exhaust line measured the temperature of the 

steam leaving the upper plenum. 

3.4.6 Accumulator Tank Instrumentation 

Accumulator tank instrumentation included a fluid thermocouple and a differential pressure 

transducer for measurement of liquid level. 

3.4.7 Injection Line Instrumentation 

Injection line instrumentation consisted of a 3.78 x 10-5 to 3.78 x 10'3 m3 /sec (0.6 to 60 gal/min) 

turbine meter to measure coolant injection with three rotameters, 0 to 3.78 x 104, 0 to 1.14 x 

10-3, and 0 to 6.31 x 10" m3 /sec (0 to 6, 0 to 18, and 0 to 100 gal/min), as a redundant 

measurement. Two fluid thermocouples were used to measure coolant injection temperatures.

-33-
SECTION 4 

TRAC-M ASSESSMENT 

WITH 

FLECHT-SEASET REFLOOD TEST DATA, RUN 31504 

4.1 Success Metrics 

Validation of thermal-hydraulic codes requires comparison of code predictions to experimental 

data. The success metrics are used to judge how well the code is validated. There are 

qualitative and quantitative metrics. The recently completed RELAP5 adequacy assessment 

effort presents important concepts that define these criteria. Part of these concepts is repeated 

here for convenience. Additional concepts on providing "User Guidelines" are added here. 

(1) "Excellent agreement" applies when the code exhibits no deficiencies in modeling a 

given behavior. Major and minor phenomena and trends are correctly predicted. The 

calculated results are judged to agree closely with the data. With few exceptions, the 

calculations will lie within the specified or inferred uncertainty bands of the data. The 

code may be used with confidence in similar applications. The term "major 

phenomena" refers to phenomena that influence key parameters, such as rod cladding 

temperature, pressure, differential pressure, mass flow rate, and mass distribution. 

Predicting the major trends means that the prediction shows the significant features of 

the data. Significant features include the magnitude of a given parameter through the 

transient, slopes, and inflection points that mark significant changes in the parameter. 

(2) "Reasonable agreement" applies when the code exhibits minor deficiencies. Overall, 

the code provides an acceptable prediction. All major trends and phenomena are 

correctly predicted. Differences between calculated values and data are greater than 

those that are deemed necessary for "excellent agreement." The calculation 

frequently lies outside but near the specified or inferred uncertainty bands of the data. 

However, the correct conclusions about trends and phenomena would be reached if 

the code were used in similar applications. The code models and/or facility model 

noding should be reviewed to see if improvements can be made. 

(3) "Minimal agreement" applies when the code exhibits significant deficiencies. Overall, 

the code provides a prediction that is not acceptable. Some major trends or 

phenomena are not correctly predicted, and some calculated values lie considerably 

outside the specified or inferred uncertainty bands of the data. Incorrect conclusions 

about trends and phenomena may be reached if the code were used in similar 

applications. If the agreement is "Minimal," the model noding must be reviewed, and 

sensitivity studies using different noding or options should be performed. The 

selection of nodalizations or options should not be arbitrary, and it should be aimed to 

model the phenomena more accurately. If the agreement can be improved with 

different noding or options, it can be reclassified as "Reasonable" or "Excellent." In 

this case, new "User Guidelines" should be developed. If the agreement cannot be 

improved, it should be reclassified as "Insufficient agreement," as described below.

-35-
(4) "Insufficient agreement" applies when the code exhibits major deficiencies, and 

provides an unacceptable prediction of the test data because major trends are not 

correctly predicted. Most calculated values lie outside the specified or inferred 

uncertainty bands of the data, and incorrect conclusions about trends and phenomena 

are almost certain if the code is used in similar applications. An appropriate warning 

must be issued to users. Selected code models should be reviewed and modified 

and, if necessary, new models should be developed and assessed before the code 

can be used with confidence in similar applications. 

Quantitative metrics define requirements for accuracy of code predictions of the test data. The 

accuracy depends on two elements. The first is the "goodness" of code models. Accuracy of 

the predictions depends on how realistically code models simulate phenomena. The second is 

the experimental data spread. Discussions below show how the goodness of the code models 

and test data characteristics affect the accuracy of predictions. 

The assessment of the TRAC-P code, an earlier version of the TRAC-M code, concluded that 

the modeling of physics of the grid spacers was inadequate. Therefore, the option to include 

grid spacer models should not be used (Ref. 15). Analysis of the reflood phenomena in 

Section 2 indicates that grid spacers substantially contribute to cooling of the rod cladding. Any 

calculation without the grid spacer effect should indicate hotter cladding temperatures for the 

rods. TRAC-M calculations will be biased because of the absence of grid spacer effects. The 

best way to quantify this bias is to compare experimental data of cladding temperatures with 

and without grid spacers. The effect of grid spacers on heat transfer has been studied for 

reflood conditions in the Flooding Experiments with Blocked Arrays (FEBA) test series, which 

shows that the effect of grid spacers on cladding temperatures is approximately 100 K. Hence, 

we should expect approximately 100 K of bias in code predictions. 

The next important element in determining a quantitative metric is the experimental data 

spread, which differs from the instrument uncertainty that addresses the accuracy in measuring 

with an instrument. The experimental data spread (the uncertainty band of the test data) 

occurs because of the random nature of the phenomena. In the Flecht-Seaset reflood 

experiments, the reflood is considered to be one dimensional since the geometry of the test 

section and the rod bundle is one-dimensional. Hence, the input deck is built on the basis of 

one-dimensional assumptions. The code will calculate one cladding temperature at a specific 

location of the rod at each time step of the calculation. The rod modeled in the input deck 

represents all similar rods. However, in reality, the two-phase flow in the test section has some 

randomness in that all water slugs are not of the same length or shape, all droplets are not of 

the same size or shape, and their concentrations vary from location to location within the test 

section. Hence, one should neither expect that maximum cladding temperatures occur in all 

rods at the same time in the same horizontal plane although the power shape for each rod is 

the same, nor expect that all rods quench at the same time in the same horizontal plane. There 

will be a normal spread of experimental data attributable to the random nature of the two-phase 

flow. The code will predict only one average value for the maximum rod surface temperature. 

For an "Excellent" agreement, the prediction can be at the edge of the spread of the maximum 

rod surface temperature data. This spread will form part of the quantitative metric.

-36-
Figures 4.1 through 4.4 show clad temperatures measured at different elevations and at 

different rods. The lowest spread occurs at lower levels, and the spread increases with 

increasing elevation. At higher elevations, droplets move randomly in many directions, and 

their sizes and shapes vary. Clad temperatures at a horizontal plane can vary as much as 

100 K. Thermocouples are located at different rods in a horizontal plane. These give multiple 

measurements at a given elevation. These measurements not only show that a spread of 

100 K can occur in clad temperatures; they also indicate that quenching times at high 

elevations may vary up to 120 s. Thermocouples affected by the housing wall and those giving 

clearly erroneous readings are not included in these plots. 

However, similar multiple measurements are not available for other parameters such as 

pressure drops. For pressure drop measurements, there is only one pair of pressure taps 

available to measure each pressure drop between two elevations. Hence, for pressure drops it 

is not possible to determine the data spread or uncertainty bands caused by randomness of 

two-phase flow. Similarly, steam probe measurements are made at few unpowered rods at 

selected elevations. It is assumed that these measured quantities give the average quantities 

for a horizontal plane. Hence, assessment of these parameters can only be made qualitatively. 

In summary, a quantitative metric for clad temperature predictions is a bias of 100 K and a 

spread of 100 K. This would make a total of 200 K difference in predictions. The quenching 

time spread is 120 s. The assessment of the remaining parameters will be made qualitatively. 

4.2 TRAC Input Model for Flecht-Seaset Forced Reflood Run 31504 

Figure 4.5 shows the TRAC input model schematic used to model forced flow reflood in Flecht

Seaset Run 31504. The input deck contains a single one-dimensional vessel component 

(Component 5) modeling the test section and plenums, one fill component (Component 1) 

where the flow rate is defined, a break component (Component 4) where the system pressure is 

defined, and two pipe components (Components 3 and 2) connecting the break and fill 

components, respectively. The fuel (heater) rods are modeled using a rod component 

(Component 6), and the test section housing wall is modeled using another rod component 

(Component 7). 

A one-dimensional vessel component is used since the reflood phenomenon that occurs in the 

Flecht-Seaset test section is one-dimensional. That component contains 18 cells, the lower 

plenum is modeled with cells 1 and 2, the rod bundle occupies cells 3 through 17, and the 

upper plenum is modeled with cell 18. The 161 heater rods are modeled as a single rod 

component. The input deck does not model grid spacers since the effects of grid spacers on 

thermal-hydraulic phenomena are not correctly modeled in the TRAC code (Ref. 15). Coolant 

injection was applied to the lower plenum through the pipe and fill components, and the flow 

exits from the break component. 

The axial power shape modeling using 18 axial cells accurately duplicated the axial power 

shape as constructed in heater rods including all steps indicating different power levels. This 

was done in order to accurately compare predictions of temperatures with measured values. 

This is much finer nodalization than that used for standard reactor fuel rod modeling. Normally, 

fuel rods are modeled using three or four axial elevations. When a coarse nodalization is used, 

there will be an additional uncertainty in calculations that should be accounted for.

-37-
1200 

1000 

v, 800 

CL 

E 

o 600 

400 

200

Flecht-Seaset Forced Reflood 31504 

Exp. Data, Clad Temps. z=48in (1.22m)

Time (s) 

Figure 4.1 Flecht-Seaset 31504 Exp. Clad Temperatures, z=1.22m

-38-

0 200
Flecht-Seaset Forced Reflood 31504 

Exp. Data, Clad Temps. z=78in (1.98m)

1400 

1200 

1000 

800 

600 

400 

200 

0

0 200 400 

Time (s)

Figure 4.2 Flecht-Seaset 31504 Exp. Clad Temperatures, z=1.98m

-39-

0 

CL. 

E 

!-
Flecht-Seaset Forced Reflood 31504 

Exp. Data. Clad Temps. z=102in (2.59m)-all data

1 400 

1200 

1000

E 

E 

I--

800 

600

400 

200

0 0 200 400 600

Time (s)

Figure 4.3 Flecht-Seaset 31504 Exp. Clad Temperatures, z=2.59m

-40-
Flecht-Seaset Forced Reflood 31504

1000 

800

600

400

Exp. Data Clad Temps. z=132in (3.35m)

200 L 

0 200 400 600 

Time (s)

800

Figure 4.4 Flecht-Seaset 31504 Exp. Clad Temperatures, z=3.35m

-41-

a

E 

I-
Core Section 

of Vessel Wall Fuel Rods 

ROD ROD

0 0D

(• El BREAK 

4 

3 PIPE 

3

End 

of Heating

VESSEL 

Cartesian Geometry 

(x=1, y=l, z=18) 

Height (m) 

4

3.6578 

3.1090 

2.9261 

2.7432 

2.5603 

2.3774 

2.1946 

2.0117 

1.6459 

1.4630 

1.2802 

1.0973 

0.9114 

0.7315 

0.5486

0

4.2926 

4.0640

18 

17 

16 

15 

14 

13 

12 

11 

10 

9 

8 

7 

6 

5 

4 

3 

2 

2 

F1 

1 F-1

Grid Spacer Locations 

S4.2926 

< 3.5568 

- - 3.0226 

< 2.5146 

o 1.9812 

*-1.4478 

K 0.9398 

< 0.4064

FILL

Figure 4.5 TRAC-M Input Model Schematic 

Seaset Run 31504

for Flecht-

-42-

Start 0.0 of Heating

3.5154 

3.3325 

3.1496 

2.9667 

2.7838 

2.6010 

2.4181 

2.0523 

1.8694 

1.6866 

1.5037 

1.3208 

1.1379 

0.9550 

0.4064 

0.1778 

0.0 

PIPE
4.3 Comparison of TRAC-M(F90) and TRAC-M(F77) Calculations with Experimental Data 

from Flecht-Seaset Run 31504 

Calculations were performed with TRAC-M(F77), Version 5.5.2A, and TRAC-M(F90), Version 

3580. Both calculations used the newrfd=1 option, which specifies the bottom-up quenching 

reflood model developed in 1992. A top-down quenching model was added in 1999. In the 

TRAC-M(F77) code, if newrfd=3 is selected, this option will activate the combined model of 

bottom-up and top-down quenching. The TRAC-M(F90) code does not have this option. As 

stated earlier, the grid spacers are not modeled, and the expected uncertainty in predicting clad 

temperatures is on the order of 200 K. 

Flecht-Seaset Run 31504 is one of a series of forced reflood tests conducted in the facility. 

Initial test conditions and key results from the test are presented in Table 1. 

TABLE I 

CONDITIONS AND KEY RESULTS FOR RUN 31504 

Parameter Test Condition Value 

Upper Plenum Pressure 0.28 MPa 

Rod Initial Clad Temperature at the 1.83-m Level 1136 K 

Rod Peak Power 2.3 kW/m 

ECCS Injection Rate -0.37 kg/s 

Coolant Temperature 324 K 

Radial Power Distribution Uniform 

Initial Temperature of Hottest Rod 1093 K 

Elevation of Maximum Temperature of Hottest Rod 1.98 in 

Maximum Temperature of Hottest Rod 1423 K 

Temperature Rise 330 K 

Turnaround Time of Hottest Rod 130 s 

Quench Time of Hottest Rod 325 s 

Bundle Quench Time 594 s 

The input deck reflects the initial conditions reasonably well. A run following the experimental 

procedure; i.e., with application of a constant power without any inlet flow, has been performed. 

The rod surface temperature and steam temperature profiles were determined as the rod 

surface temperature at the location of the maximum heat generation reached the specified 

value at which the inlet flow and power decay would be initiated during experiments. These

-43-
calculated temperature profiles were used as an input to the second deck to simulate initial test 

conditions. Figures 4.6 and 4.7 compare the calculated initial heater rod cladding and steam 

temperatures along the test section with experimental data, respectively. Calculated 

temperatures agree with experimental data very well. 

The uncertainty of experimental measurements was on the order of 100 K in a two-phase flow. 

Experimental data shown in Figure 4.6 include all thermocouples including those giving 

erroneous readings. The data spread seems to be larger than 100 K. (It is on the order of 

150 K.) One reading that seems to be an outlier can be thrown out. The data spread is 

probably attributable to natural circulation of the steam before the initiation of testing. Natural 

circulation would occur because of heat losses from the test section wall, and nonuniformity of 

the axial power shape. This is also substantiated in Figure 4.7, which depicts the initial axial 

variation of steam temperatures. The spread of experimental data is on the order of 100 to 

150 K, and there is an outlier that can be thrown out. Calculations were performed using a one

dimensional model where the natural circulation is not modeled. Considering this limitation, the 

agreement between the calculated temperatures and measured data is Excellent for initial 

conditions. 

The test was initiated by applying the decay power and forcing the specified inlet flow 

simulating the ECCS injection into the test section when the rod temperature reached the 

specified level. The test section was initially full of steam. As the inlet flow progressed, water 

slugs and droplets provided precursory cooling to the rods. As the flow progressed, the rods 

were quenched. Comparisons of calculated clad temperatures and experimental data are 

illustrated at different elevations of the test section in Figures 4.8 through 4.11. Clad 

temperatures are calculated with both TRAC-M(F77) and TRAC-M(F90) codes. Only bounding 

experimental temperature curves are shown. At t=0.0, the clad temperatures are at their initial 

values. ,4

Temperatures initially increase because precursory cooling is not suffic' nt to immediately cool 

down the rods. As the time progresses, the quench front propagates5 precursory cooling 

increases, and decay power and rod power decrease. Ian lov~tion is pocificd, clad 

temperatures first increase, attain their maximum values, turn around as precursory cooling 

increases, and finally drop to saturation value as the rod quenches. Figure 4.12 illustrates the 

quench front propagation. Predictions of the quench time with the TRAC-M(F77) code during 

the first 200 s of the transient are slightly delayed, but they meet the acceptance criterion of 

120 s spread. After 200 s, but before 500 s, predictions indicate substantially earlier quench 

that does not meet the acceptance criterion. Between 500 and 620 s, quenching occurs within 

the acceptance criterion. Predictions of the quench time with the TRAC-M(F90) code are 

always delayed. Except for the short time period in the beginning of the transient (about the 

first 150 s of the transient), the delay is beyond the acceptance criterion. 

Figures 4.8 through 4.11 illustrate clad temperature predictions with both codes and bounding 

experimental data for various elevations of the test section. The bounding experimental data 

traces bound the spread of the experimental data, which is 100 K. Since the tests were run 

with grid spacers, and they are not modeled in the codes, the codes should predict hotter 

temperatures, up to 100 K, which is the expected bias. Figures 4.8 through 4.11 show that the 

predictions of maximum clad temperatures with the TRAC-M(F77) code meet the acceptance 

criterion for all elevations. Predictions with the TRAC-M(F90) code meet the acceptance 

criterion for low elevations (below -2.00 m), but at higher elevations, they do not meet the 

acceptance criterion.

-44-
1200 

1000 

S800 

E 

600 

400 

0

Flecht Seaset Forced. Reflood 31504 

Comparison of Calculated Initial Clad Temperatures with Experimental Data

1 2 3 

Axial Height - Heater Rod Length (m)

Figure 4.6 Initial Axial Clad Temperature Variation

-45-

4
1000 

• 800 

(D 

CL E 

E 

600 

400 0

Flecht Seaset Forced Reflood 31504 

Comparison of Calculated Initial Axial Steam Temps. with Exp. Data

1 2 3 4

Axial Height - Heater Rod Length (m)

Figure 4.7 Initial Axial Steam Temperature Variation

-46-
1200 

1000 

800 

C. 

E (D 

600 2= 6O00 

400 

200

Flecht-Seaset Forced Reflood 31504 

Comparison of Calculated and Experimental Clad Temps z=1 .22m

0 100 200 

Time (s)

300

Figure 4.8 Comparison of Calculated and Experimental Clad Temps. 

@z=l.22m

-47-
Flecht-Seaset Forced Reflood 31504 

Comparison of Calculated and Experimental Clad Temps z=1.98m

0 100 200 300 400 

Time (s)

Figure 4.9 Comparison of Calculated and 

@z=1.98m

Experimental Clad Temps.

-48-

1750

7 

C. 

E a, 

-'i 

C) 

U)

1250 

750

250 500
1750 

1250 

750 

250

Flecht-Seaset Forced Reflood 31504 

Comparison of Calculated and Experimental Clad Temps z=2.59m

0 200 400 600 

Time (s)

Figure 4.10 Comparison of Calculated and Experimental Clad Temps. 

@z=2.59m

-49-

C,) 

E

2 

.,2
Flecht-Seaset Forced Reflood 31504 

Comparison of Calculated and Experimental Clad Temps z=3.35m

E 

a) 

2

1400 

1200 

1000 

800 

600 

400 

200 

0

0 200 400 600 800 

Time (s)

1000

Figure 4.11 Comparison of Calculated and Experimental Clad Temps. 

@z=3.35m

-50-
Flecht Seaset Forced Reflood 31504 

Comparison of Calculated and Experimental Quench Elevations

x N 

x,

4 

3 

Eý 

.2 

w 

0 

0 

0 200

X

S X X 1.0 

)OC 

-.x SXX 

"0-0 TRAC-M (F77) 

o. - OTRAC-M (F90) 

X Experimental Data

400 

Time (s)

600 800

Figure 4.12 Calculated and Measured Quench Front Propagation 

-51-

k 

-e 

>0

X

X 

X

J

I
Figure 4.13 illustrates experimental differential pressure data taken along the test section. 

These data were taken at 1-foot increments starting from the 0-foot reference and ending at the 

10-foot height. Observe that the data indicate orderly and monotonic progression of a sharp 

quench front suggested by rapid increases of pressure drops at lower elevations, from 0 to 

4 feet. The sharpness of the quench front disappears at higher elevations; however, 

progression of the quench front is still orderly and monotonic; i.e., the height of the quench front 

increases with increasing time. 

Figure 4.14 illustrates calculations of vapor fractions along axial distance with the TRAC-M(F77) 

code. Although calculations show predictions of quench fronts at different times, they indicate 

that progression is not orderly and monotonic; e.g., quenching at lower elevations occurs at 

later times. Calculations and animation of the results with the XTV code show that the quench 

front and vapor fractions along the test section continuously oscillate. This shows that the code 

does not correctly predict the phenomena. 

Flecht-Seaset Forced Reflood 31504 

Experimental DPs along Test Section 

0---1 BU-3-4 [a-- .. U-0-1 

& -ABU-1-2 

-- -BU-2-3 

p- -s BU-4-5 

A-A AB U-5-6 '--- BU-1r-7 

- BU-7-8 

BU-8-9 -- TOUL-9•-10 

, 2 

Sn41 

n 100A 200 300 400 500 600

Time (s) 

Figure 4.13 Experimental Differential Pressure 

Measurements Along the Test Section

-52-
1 

0.8 

.c 0.6 

0 

14. 

0 

0.4 

0.2 

0

Flecht Seaset Forced Reflood 31504 

Calculated Axial Void Fraction Distribution at Different Times

0 1 2 3 4 5

Axial Distance (m) 

Figure 4.14 Calculation of Axial Void Fraction Distribution by 

TRAC-M(F77) Code

-53-
Figures 4.15 through 4.20 present comparisons of code predictions of differential pressure 

drops with experimental data. Both codes predict large-amplitude, high-frequency oscillations, 

while the measurements show relatively smooth conditions. The predictions of large-amplitude, 

high-frequency oscillations are not considered realistic. Qualitatively, there is "Insufficient 

agreement" between the calculations and test data. 

Flecht-Seaset Forced Reflood 31504

Comparison of Differential Pressures, DP 1-2

4 

3.

to 

0

0 500

Time (s)

Figure 4.15 Comparison of Differential Pressures (Taps 1 - 2)

-54-
Flecht-Seaset Forced Reflood 31504 

Comparison of Differential Pressures, DP 2-3

4 

IraI 

iI 

P 

C . 

.BU-2-3 

_ I!DP(2-3) 2- - TRAO-M (F90 

G-e DP(2-3) TRAC-M( 

0 500 

Time (s) 

Figure 4.16 Comparison of Differential Pressures (Taps 2 - 3)

-55-
Flecht-Seaset Forced Reflood 31504 

Comparison of Differential Pressures, DP 3-4

3 

CL

0 500

Time (s)

Figure 4.17 Comparison of Differential Pressures (Taps 3 - 4)

-56-
5

C

C. 

0

0

Flecht-Seaset Forced Reflood 31504 

Comparison of Differential Pressures, DP 4-5

0 500 

Time (s)

1000

Figure 4.18 Comparison of Differential Pressures (Taps 4 - 5)

-57-

-- 9 BU-4-5 Exp. Damt 

a- - 2 DP(4-5) TRAC-M (F90) 

9-0 OP(4-5) TRAC-M (F77) 

I I 

I i I 

I !i
5 

CO 

a

0 

.m 

t0 

03 

0

0

Flecht-Seaset Forced Reflood 31504 

Comparison of Differential Pressures, DP 5-6

0 500 1000

Time (s) 

Figure 4.19 Comparison of Differential Pressures (Taps 5 - 6)

-58-
Flecht-Seaset Forced Reflood 31504 

Comparison of Differential Pressures, DP 6-7

II!

JilI

J ii 

II 

I II 

I I

ii I 

Ii I

I, i1 J'f t' 'ii.' jITzL,tf I f.i i'J 

I ' I' 

Il 

II 

II

o.---* BU-6-7 Exp. Data 

a- -- DP(6-7) TRAC-M (F90) 

3---O DP(6-7) TRAC-M (F77) 

: I ' 

I : • II 

Ii " I n i ) l I 

(t )'; 't) • ~r • .t••

II

500 

Time (s)

I* 

I j 

tj i'I It i

1 1 1 

lj I'j 

' I 

I' I II 

Figure 4.20 Comparison of Differential Pressures (Taps 6 - 7) 

-59-

4F

a30a 

03 

(I, a) 

CO

-1

II I3 

.&

ItI

0 1000

I11

l';l,•ti.Jr (
Figures 4.21 through 4.23 present comparison of calculated steam temperatures with 

experimental data at some elevations. There are very few steam temperature measurements; 

hence, an estimate of their uncertainties cannot be made. However, comparisons indicate that 

calculated values differ by several hundred degrees from the experimental data. Hence, the 

agreement is Insufficient. 

Flecht-Seaset Forced Reflood 31504

1300 

1100

900 

700

500

300 ' -100

Comparison of Exp. Steam Measurements at z=1.83-1.94m with Calculations 

I*--- TRAC-M(F'7). Cefl#1 

l-U - .SPIOL-8 

TRAC-M{F9V), 0e11#1 SP71-6

Figure 4.21

100 300 500 

Time (s) 

Steam Temperatures (Z=1.83 - 1.94m)

-60-

21 

a) 

E 

0) 

a,
Flecht-Seaset Forced Reflood 31504 

Comparison of Exp. Steam Measurements at z=2.82-3.05m with Calculations

100 300 500 

Time (s)

Figure 4.22 Steam Temperatures (Z=2.82 - 3.03m)

-61-

1300 

1100

0 

1.. 

0 a

E 

E 

E 

Ca 

2 

(IO

900 

700

500

300 ' 

-100
Flecht-Seaset Forced Reflood 31504 

Comparison of Exp. Steam Measurements at z=3.35-3.51 m with Calculations

500 

300 

-100 100 300 500 

Time (s)

Figure 4.23 Steam Temperatures (Z=3.35 - 3.51m)

-62-

1300 

1100

2 

(D 

C

E 

E 

(Z a) 

Ul)

900 

700
These comparisons indicate that the codes are not correctly predicting phenomena. Although 

peak clad temperature predictions with the TRAC-M(F77) code are excellent, the rest of the 

clad temperature predictions Minimally agree with the data. There is Insufficient agreement 

between predictions of differential pressures and steam temperatures and test data. One can 

conclude that predictions of peak clad temperatures are fortuitous. 

A big surprise in code calculations was that differences between TRAC-M(F77) and TRAC

M(F90) predictions were very high. Examination of output files indicated that the TRAC-M(F77) 

code was predicting choked flow at the exit of the test section. There were some corrections to 

the choked flow model in the TRAC-M(F90) code; therefore, the code was not predicting 

choked conditions. Figure 4.24 demonstrates the effects of choked flow predictions. The 

choking of the flow produces high pressures in the test section as predicted by the TRAC

M(F77), code while the prediction of the test section pressure with the TRAC-M(F90) code 

remains almost at a constant value of 40 psia (the pressure level at which the experiments were 

conducted). 

Flecht-Seaset Forced Reflood 31504 

Comparison of Test Section Pressures 

3e+06 

G----OTRAC--M F77 

-TRAC-M F(9o• 2 Exp,. Pressu,0 

2&+06

£ 

03 

a, 

ci, 

0O 

a.

le+06 

01 

0 200 400 

Time (s)

Figure 4.24 Comparison of Test Section Pressures

-63-
Choking should not have been predicted since this phenomenon does not occur during reflood. 

It is clear that the code predictions are completely wrong, and the codes are predicting different 

phenomena. Prediction of choking with the TRAC-M(F77) code indicates that predictions of 

two-phase velocities are also extremely high. Corrections of the choking model in the TRAC

M(F90) code does not change this conclusion. Since choking is not predicted in the TRAC

M(F90) code, prediction of two-phase velocities should be higher. It is expected that 

predictions with both codes would indicate substantially more liquid carryout than in testing. 

This would lead to higher clad temperature predictions, since less liquid would contact the 

surfaces of the rods and less cooling would occur. Figures 4.25 and 4.26 compare the 

predicted liquid and steam carryouts (integrated liquid and steam outflows) with experimental 

data. 

Flecht Seaset Forced Reflood 31504

200 

0 

• 100 

E 

0-L 

0

Comparison of Calculated and Experimental Uquid Flows I' •'I I

200 400 600 

Time (s) 

Figure 4.25 Comparison of Liquid Carryouts

800

-64-
200[ 

150 

0 

E 

2 100 

E 

50 

0 

0

Flecht Seaset Forced Reflood 31504 

Comparison of Calculated and Experimental Steam Flows

200 400 600 

Time (s) 

Figure 4.26 Comparison of Steam Carryouts

800

As expected, Figure 4.25 demonstrates predictions of excessive liquid carryout by both codes. 

Excessive liquid carryout leads to predictions of less energy removal and higher clad 

temperatures. It also leads to production of less steam than measured during experiments. 

This is confirmed in Figure 4.26. 

4.4 Discussion of TRAC Models 

Excessive liquid carryout has been predicted in many reflood tests (Ref. 15). Excessive liquid 

carryout is an indication that the interfacial area used by the code is larger than in tests. This 

raises the question of how well flow regime maps have been modeled in the code.

-65-
Figure 4.27 shows the flow regime map used in the TRAC code during reflood when the inlet 

conditions are subcooled or low quality. The flow regime map is an "Inverted Annular Flow" 

regime, which is also depicted in Figure 2.2. Figures 4.28 and 4.29 show the relative velocities 

calculated with TRAC-M(F77) and TRAC-M(F90) codes at Cell #10 (about at the middle of the 

test section) and at Cell #17 (almost at the end of the test section). Note that the relative 

velocities are very high. High steam velocities may tear the liquid away from the liquid core, 

and also carry large liquid slugs out of the test section without permitting the liquid to contact 

the heated walls for a sufficient length of time. High relative velocities explain why the liquid 

carryout from the test section is higher than the measured level, and why calculated clad 

temperatures are generally higher than experimental values. 

The TRAC flow regime maps are predicated on the work performed by Ishii and his coworkers 

(Ref. 16). The reference indicates that these flow regime maps exist at relatively low relative 

velocities. For example, it states that if the relative velocity is increased to 2.0 m/s, smooth and 

rough wavy sections are nonexistent. The relative velocities shown in Figures 4.28 and 4.29 

are much higher than 2.0 m/s. They go up to 20.0 m/s. This would lead to the conclusion that 

at these predicted velocities, flow regime maps used by the TRAC-M(F77) of TRAC-M(F90) 

codes cannot be consistent with the flow regimes that occurred in tests. 

According to Ref. 16, the inverted annular flow regime has four different regions with distinct 

characteristics (smooth, rough-wavy, agitated flow, and dispersed flow regions). Figure 4.27 

depicts these regions, as well as the transition flow region. This last region refers to the 

transition boiling region between the CHF point and the Tq. The tests in Ref. 16 were 

performed under conditions that did not simulate the transition flow region. 

Ref. 16 defines boundaries for four different regions. A break-up length for each region is 

defined by a correlation. Each correlation is founded on the Capillary number, and has its own 

region of applicability, which is predicated on both the Weber number and the vapor fraction 

measured at the inlet of the test section. The Weber number is founded on the relative velocity 

between the liquid and gas. An examination of the documentation on the TRAC code (Ref. 17) 

indicates that although the same Ishii correlations for break-up lengths have been used, the 

same regions of applicability have not been used. The code uses different regions of 

applicability that are defined by the vapor fraction alone, and are not predicated on the Weber 

number. Table 2, below, shows the Capillary number criteria and the Weber number 

restrictions in Ref. 16 and TRAC code application in Ref. 17. 

In Ref. 16 Z is the average break up length measured from the inlet of the test section, D is 

the liquid jet hole diameter, Ca is the Capillary number defined using fluid jet properties 

(viscosity, velocity and surface tension) at the inlet, Ca = .t V / a , We is the Weber number 

11 1 

defined using fluid jet and gas properties at the inlet, We = p V )D / c1 ,and c is the 

v f 

void fraction at the inlet of the test section.

-66-
ZAG

ZRW

ZSM 

Ivrted 

Annular 

Flow 

(Smooth) R 

Transition Flow 

I ZCHF 

Bubbly 

FRow 

Figure 4.27 TRAC-M Reflood Flow 

Regime Map

-67-
Flecht Seaset Forced Reflood 31504 

Comparison of Relative Velocities, (Vg-VI), at CeI#1 0

200 400 600 

Time (s)

800

Figure 4.28 Calculated Relative Velocities at Cell #10

-68-

20 

15

(A, 

5: 

0, 

$ 

0 

.2 

a, 

a, 

a,

10 

5 

0

-5 

-10

0
30 

20 

U) 

5 

-0 10 

0 

0 0 

-10 

0

Flecht Seaset Forced Reflood 31504 

Comparison of Relative Velocities, (Vg-VI), at.Cell#17

200 400 600 

Time (s)

800

Figure 4.29 Calculated Relative Velocities at Cell #17

-69-
TABLE 2

FLOW REGIME TRANSITION CRITERIA

Regime Ishii's Correlation Region of Applicability TRAC Application 

Smooth Inverted 0.05((x (.3 

Annular Flow (IAF) Z / D = 60oka We / a2i < 10-2 

Rough Wavy IAF Z/D= 2951Ca We o 2 i _ 0.3(cx(0.4 

Agitated IAF Z/ D = 595-0. We7/5 < 3.575 

Dispersed Section Z / D = 595ai& We / ci _ 10 0.75(ox(O.98 

The applicability of correlations in Ref. 16 to simulate the reflood phenomenon during a LOCA 

may be very limited. Experiments were conducted at a steady state, and flow regimes varied 

axially. The void fraction in correlations was determined at the inlet of the section, and its value 

was determined from the test section geometry alone, and remained constant for a given test 

section. Hence, the break-out lengths depended only on liquid velocities (i.e., Capillary 

numbers). The region of applicability was determined by the correlations with the Weber 

number and the void fraction, both of which were evaluated at the inlet of the test section. The 

break-out lengths did not vary with the amount of power applied to the test section. This can be 

true only if the amount of heating did not change during tests. Hence, the power or power 

profile did not enter into correlations. However, in reactor applications, power and power profile 

can be a major factor. 

The TRAC modeling is different in that the Capillary number is calculated using fluid properties 

at the CHF point. The criteria of applicability of Ishii correlations (i.e., break-out lengths) are 

predicated on local void fractions. Details of calculations of the local void fractions are 

presented in Ref. 18. There is no restriction on the basis of the Weber numbers that are 

calculated using relative velocities. 

It is clear that implementation of the flow regime maps modeled by Ishii (Ref. 16) has not been 

correctly performed in the TRAC code (Refs. 17 and 18). On the other hand, Ishii correlations 

in Ref. 16 may have important shortcomings in that break-out lengths are independent of the 

power and power profile of the test section, and correlations are developed using parameters 

measured at the inlet. They are not local quantities. Determination of the type of reflood 

modeling in the TRAC code and the type of experiments needed for modeling reflood are 

outside of the scope of the present assessment. 

The Office of Nuclear Regulatory Research has a research program in place at Pennsylvania 

State University (PSU) to improve the agency's capability to model reflood phenomena using its 

system analysis codes. The findings in this report provide guidance on how future modeling of 

reflood should proceed. This report assesses the code against one particular Flecht-Seaset 

(F-S) run. However, the decision to initiate the PSU test program was based on several reflood

-70-
cases at various facilities. These cases exhibit the overall trend that TRAC-M(F90) overpredicts 

peak clad temperature (PCT), even though the predictions are in reasonable agreement with 

runs made at the CCTF and SCTF facilities. The TRAC-P code series has been used by NRC 

to perform LOCA analysis because of this conservative trend. However, as the agency begins 

to risk-inform its regulations, new designs are submitted and licensees request further uprates. 

NRC would be better equipped to audit licensee submittals with a less conservative, more 

accurate reflood model. In order to model reflood accurately for all designs, conditions and test 

facilities, the proper physics must be modeled mechanistically. The author recommends that 

new flow regime maps for reflood be constructed and the selection criteria for the maps should 

also include effects of both phasic velocities. 

4.5 Sensitivity Calculation without the Choking Model 

To evaluate the effect of choking on predictions, a sensitivity study was performed with the 

TRAC-M(F77) code. The choking model was artificially turned off. Figures. 4.30 through 4.33 

compare calculations of the TRAC-M(F77) code with and without the activation of the choking 

model and those of the TRAC-M(F90) code with the experimental data. As expected, the 

predictions of the TRAC-M(F77) code without the choking model are very close to those of the 

TRAC-M(F90) code. There are some small differences, but they are expected because of bug 

fixes or error corrections in the TRAC-M(F90) code. This comparison shows that the 

conversion to the TRAC-M(F90) code for this type of application is successful. 

4.6 Conclusions 

The top-level requirement in the development of a thermal-hydraulic code is that the code 

should correctly model the physics of the phenomena. This assessment shows that the TRAC 

code does not correctly model flow regime maps during the reflood and, consequently, it cannot 

correctly predict interfacial areas between phases, and vapor fractions, phasic velocities, and 

temperatures. Almost none of the important parameters (such as, clad temperatures, steam 

temperatures, differential pressures, and liquid and steam carryouts) are consistently predicted 

throughout the range with acceptable accuracy. The predictions in ranges where clad 

temperatures are predicted within accuracy limits, are accidental. The predictions of clad 

temperatures are higher than the experimental data because predictions of liquid carryout by 

the codes during a reflood have consistently been high. (See also Ref. 15.) Hence, these 

predictions are conservative. 

The comparison of TRAC-M(F77) and TRAC-M(F90) codes and the sensitivity study performed 

with the TRAC-M(F77) code show that the conversion of the TRAC-M(F77) code to the TRAC

M(F90) code has been successful for this type of application.

-71-
Flecht-Seaset Forced Reflood 31504 

Comparison of Calculated and Experimental Clad Temps z=1.22m

1200 

1000

800 

600

400 

200

0 100 200 

Time (s)

300

Figure 4.30 Comparison of Clad Temperatures with and without 

Choking, Z=1.22m

-72-

0. 

E 

I

"0
1750 

1250

E 

CO 

"0

750 

250

Flecht-Seaset Forced Reflood 31504 

Comparison of Calculated and Experimental Clad Temps z=1.98m

0 100 200 300 400 

Time (s)

500

Figure 4.31 Comparison of Clad Temperatures with and without 

Choking, Z=l.98m

-73-
1750 

1250

£ 

E 

0 

I

"0 

(5

750 

250

Flecht-Seaset Forced Reflood 31504 

Comparison of Calculated and Experimental Clad Temps z=2.59m

0 200 400 600 

Time (s)

Figure 4.32 Comparison of Clad Temperatures with and without 

Choking, Z=2.59m

-74-
Flecht-Seaset Forced Reflood 31504 

Comparison of Calculated and Experimental Clad Temps z=3.35m

1400 

1200 

1000

800 

600 

400 

200 

0

0 200 400 600 800 

Time (s)

1000

Figure 4.33 Comparison of Clad Temperatures with and without 

Choking, Z=3.35m

-75-

_.@ 

E
SECTION 5 

TRAC-M ASSESSMENT 

WITH 

FLECHT-SEASET STEAM COOLING TEST DATA, RUN 32753 

5.1 Discussion of Phenomena and Code Requirements 

The phenomenon that occurs during steam cooling is rather simple. During some part of the 

transient, the reactor core is assumed to be full of steam, and no liquid is present. The incoming 

flow from the inlet is also assumed to be steam. The question is whether the incoming steam 

flow can cool the fuel assemblies. Hence, the capability required from the codes is proper 

modeling of heat transfer in a convective single-phase flow. The top-level requirement is that 

codes should model appropriate physics to predict heat transfer in a single-phase flow. The next 

level of requirements is that trends and important parameters should be predicted with 

acceptable accuracy. 

Steam cooling tests were conducted in the Flecht-Seaset test facility in order to simulate steam 

cooling conditions in a PWR. These tests were initiated by pressurizing and preheating the rod 

bundle and associated piping with steam that was slightly superheated. Once the test section 

was heated above the saturation temperature, the rod power was turned on to a preset value 

and was kept constant. Tests were run until steady-state conditions were obtained. Hence, the 

expected primary trend is that clad temperatures will increase and asymptotically approach a 

steady-state value. The important parameter is the clad temperature. 

5.2 Success Metrics 

As stated in Section 4.1, validation of thermal-hydraulic codes requires comparison of code 

predictions with experimental test data. The same definitions will be used for "Qualitative" 

metrics, as described in Section 4.1 (i.e., "Excellent Agreement," "Reasonable Agreement," 

"Minimal Agreement," and "Insufficient Agreement"). 

The quantitative success metric will be determined on the basis of the experimental spread of 

the clad temperature data, which will be plotted and examined. Figure 5.1 shows the scatter of 

thermocouple readings at Z_1.52m elevation in Run 32753. The spread of the data is within 

70 K, and there is one outlier trace that is giving wrong measurements. This thermocouple trace 

will not be included in the uncertainty band. The spread is attributed to the natural circulation 

occurring within the test section because of heat loss from the walls and effects of unheated 

tubes where steam probes and instruments are located. Locations of unheated rods are shown 

in Figure 3.2. 

If clad temperatures are predicted within 70 K, predictions will be considered "Excellent." It 

should be noted that data above Z=1.52m elevation were not used for comparisons. Severe rod 

bundle distortions occurred above this elevation, and a majority of the heater rod temperatures 

did not attain their steady-state values according to Ref. 19. Because of distortions, many of the 

runs were not usable.

-77-
Flecht-Seaset Steam Cooling Test 32753 

Exp. Data, Clad Temps. at z=60in (1.52m)

480 

460 

440

420 

400

0 200 400 600 800

Time (s) 

Figure 5.1 Experimental Data, Clad Temperatures at Z=l.52m

-78-

0 

1CL 

E
5.3 TRAC Input Model for Flecht-Seaset Steam Cooling Run 32753

The TRAC input model schematic used to model steam cooling Flecht-Seaset Run 32753 is 

similar to the one used Forced Reflood Run 31504, except that the power applied to the test 

section remained constant with time and the heater rod was.finer nodalized than in Run 31504 in 

order to bring the node interfaces to thermocouple locations for accurate comparisons of code 

predictions with test data. As in the case of Run 31504, the input deck does not model heat 

losses from the test section to the environment. 

5.4 Comparison of TRAC-M(F77) and TRAC-M(F90) Calculations with Experimental Data 

from Flecht-Seaset Run 32753 

Calculations were performed with TRAC-M(F77) and TRAC-M(F90) codes and compared with 

experimental data. Table 3 lists initial test conditions for Run 32753. Comparison of predictions 

by the TRAC-M(F77) and TRAC-M(F90) codes and the test data are presented in Figure 5.2. As 

stated above, because of severe rod distortions, comparisons are made only up to the elevation 

of 1.52 m. Comparisons show that agreements between the predictions and the test data at 

different elevations are well within the uncertainty band of 70 K, up to the elevation of 1.52 m. At 

this elevation, the differences between code predictions and the thermocouple trace grow to 

approximately 70 K. The thermocouple trace represents the average of the traces in Figure 5.1. 

Hence, the agreement is slightly out of the bound of uncertainty, but will be considered 

"Reasonable" agreement. One reason for this slight difference is that the input deck does not 

model heat losses from the test section. A slight bias for hotter predictions is expected. 

Examination of Figure 5.2 leads to the conclusion that the trend of the test data (i.e., the 

asymptotic approach to the steady-state values at different elevations) is very well predicted. 

Code-to-code differences in predictions are approximately null, as expected. 

TABLE 3 

INITIAL CONDITIONS FOR RUN 32753

-79-

Parameter Test Condition Value 

Upper Plenum Pressure 0.28 MPa 

Initial Rod Wall Temperature 408 K 

Rod Peak Power 0.205 kW/m 

Inlet Flow Rate 0.36 kg/sec 

Coolant Temperature 405 K 

Bundle Radial Profile Uniform
Flecht-Seaset Steam Cooling Test 32753

Comparison of Clad Temperatures

500 

480 

460 

E 

2 440 

420 

400

0 500 1000 

Time (s)

1500

Figure 5.2 Comparison of Predicted and Measured Clad Temperatures 

5.5 Conclusions 

This assessment shows that both TRAC-M codes correctly predict the single-phase convective 

flow heat transfer within acceptable accuracy. The agreement between predictions and the test 

data is "Excellent" to "Reasonable." It is judged that "Excellent" agreement throughout the whole 

region of valid testing may be obtained by modeling the input deck in more detail. 

The comparison of TRAC-M(F77) and TRAC-M(F90) codes show that the conversion of the 

TRAC-M(F77) code to the TRAC-M(F90) code has been successful for this type of application.

-80-
REFERENCES 

1. G.G. Loomis and R.W. Shumway, "Transition Boiling Heat Transfer in the Semiscale Mod-3 

Core During Reflood," Nuclear Technology, "56, 426, March 1982. 

2. G. DeJarlais, "An Experimental Study of Inverted Annular Flow Hydrodynamics Utilizing an 

Adiabatic Simulation," NUREG/CR-3339, ANL-83-44, March 1983. 

3. M.J. Loftus et al., "PWR FLECHT-SEASET Unblocked Bundle Forced and Gravity Reflood 

Task Data Report," NUREG/CR-1532, Vols. 1 and 2, WCAP-9699, 1980. 

4. N. Lee et al., "PWR FLECHT-SEASET Unblocked Bundle, Forced and Gravity Reflood 

Task Data Evaluation and Analysis Report," NUREG/CR-2256, WCAP-9891, 1981. 

5. G.E. McCreery, G.G. Loomis, H.R. Bruestle, "Thermal-Hydraulic Analysis of the Semiscale 

Mod-1 Reflood Test Series (Gravity Feed Tests)," TREE-NUREG-1 010, Jan., 1977. 

6. A.C. Peterson, G.G. Loomis, L.L. Chen, "Thermal and Hydraulic Response of the 

Semiscale Modl Core During Forced Feed Reflood Tests," TREE-NUREG-1 001, Oct., 

1976. 

7. R. Deruz and J.M. Veteau, "Pericles Reflood Experiments," presented at 111' Water Reactor 

Safety Information Meeting, Gaithersburg, MD, 1983, NUREG/CP-0048, Vols. 1-6, 

January 1984. 

8. P. Ihle and L. Rust, "FEBA-Flooding Experiments with Blocked Arrays, Data Report 1, Test 

Series I through IV," KfK 3658, March 1984. 

9. P. Ihie and L. Rust, "FEBA-Flooding Experiments with Blocked Arrays, Data Report 2, Test 

Series V through VIII," KfK 3649, March 1984. 

10. K. G. Pearson et al., "Reflooding Experiments on a 49-Rod Containing a Long 90% 

Blockage, AEEW-R-1591, January 1983. 

11. V.J. Dhir and M. Fruckler, "Studies of a Single and Two-Phase Heat Transfer in a Blocked 

4-Rod Bundle," EPRI Report NP-3485, June 1984. 

12. S.L. Lee, S.K. Cho, H.J. Sheen, "Reentrainment of Droplets from Grid Spacers in Mist Flow 

Portion of a LOCA Reflood of a PWR," NRC/ANS Topical Meeting on Basic Thermal

Hydraulic Mechanisms in LWR Analysis, September 1982, NUREG/CP-0043, 1983. 

13. Seban et al., "UC-B Reflood Program: Experimental Data Report," EPRI/NP-743, 

April 1978. 

14. Y. Murao et al., "Experimental and Analytical Modeling of the Reflood Phase During PWR 

LOCA," paper presented at the 195 National Heat Transfer Conference, Orlando, Florida 

(ASME), July 27-30, 1980.

-81-
15. B.E. Boyack, J.F. Lime, D.A. Pimentel, J.W. Spore, and J.L. Steiner, "TRAC-M:FORTRAN 

77, Version 5.5, Developmental Assessment Manual, Volume 1: Nonproprietary 

Assessment Sections," LA-UR-99-6480, December 1999. 

16. N.T. Obot and M. Ishii, "Two Phase Flow Regime Transition Criteria in Post-Dryout Region 

Based on Flow Visualization Experiments," Int. J. Heat Mass Transfer, Vol. 31, No.12, 

pp 2559-2570, 1988. 

17. R.A. Nelson, D.A. Pimentel, S.J. Jolly-Woodruff, and J.W. Spore, "A Phenomenological 

Thermal-Hydraulic Model of Hot Rod Bundles Experiencing Simultaneous Bottom and Top 

Quenching and Optimization Methodology for Closure Development," LA-UR-98-3043, 

April 1998. 

18. J.W. Spore, J.S. Elson, S.J. Jolly-Woodruff, T.D. Knight, J.C. Lin, R.A. Nelson, 

K.O. Pasamehmetoglu, R.G. Steinke, and C. Unal of Los Alamos National Laboratory and 

J.H. Mahaffy and C. Murray of Pennsylvania State University, "TRAC-M/FORTRAN 90 

(Version 3.0), Theory Manual," LA-UR-00-910, March 2000. 

19. S. Wong and L.E. Hochreiter, "Analysis of Flecht-Seaset Unblocked Bundle Steam Cooling 

and Boiloff Tests," NUREG/CR-1533, January 1981.

-82-
NRC FORM 335 U.S. NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER 

(2-89) (Assigned by NRC, Add VoL, Sup., Rev., 

NRCM 110Z and Addendum Numbem, if amy.) NRCM 110, 

201, E3= BIBLIOGRAPHIC DATA SHEET 

(See insmci•obs on the reverse) • NUREG-1 744 

2. TITLE AND SUBTITLE 

Assessment of TRAC-M Codes Using Flecht-Seaset Reflood 

and Steam Cooling Data 3. DATE REPORT PUBLISHED MONTH ! YEAR 

May 2001 

4. FIN OR GRANT NUMBER 

5. AUTHOR(S) 6. TYPE OF REPORT 

F. Odar Technical 

7. PERIOD COVERED (Indusive Dates) 

8. PERFORMING ORGANIZATION - NAME AND ADDRESS (If NRC, pmvida Diision, Office or Region, U.S. Nuclear Regulatory Commissio, and ma•tm address; siOdi-, 

provde name and mnaing address-) 

Division of System Analysis and Regulatory Effectiveness 

Office of Nuclear Regulatory Research 

U.S. Nuclear Regulatory Commission 

Washington, DC 20555-0001

9. SPONSORING ORGANIZATION - NAME ANU ADDREuS (fI NRU, ye -,ame asaoove;ff a r, P-,,, .... .U. ...... ........... 

and maietg addressm)

Same as 8. Above.

10. SUPPLEMEN[A?•T NUOIT

11. ABSTRACT (200 words orless)

This report presents the results of an assessment work performed using TRAC-M(F 77), Version 5.5.2A and TRAC-M(F90), 

Version 3.580 codes. The report assesses capabilities of both cod.es to predict reflood and steam cooling phenomena which 

may occur during a postulated Large Break Loss of COolant Accident, LBLOCA. The assessment is ased on Flecht-Seaset 

Run 31504 and 32753 test data. The assessment shows that oth codes predictions of he redood phenomena are not 

accurate; however, it is judged that they conservatively predict peak clad temperatures in heated rods since the model expels 

more water from the test section than measured. The predictions of steam cooling in single phase flow conditions are 

acceptable.

12. KEY WORDS/DESCRIPTORS (List words or phrases that vwi assist researchers in /ocating the report) 13. AVAILABILITY STATEMENT 

Reflood 

unlimited 

Steam Cooling 14. SECURITY CLASSIFICATION 

Thermal Hydraulics (Ms pa") 

Computer Codes unclassified 

Code Assessment ahis Rep) 

Code Validation uThi asRified unclassified 

15. NUMBER OF PAGES 

16. PRICE 

NRC FORM 335 (2-~~NEEEE Nis rm i a eeruisyM.ý

NRC FORM 335 (2-89) This fonrm was ell•:Worlliclty pr-ouý= by •IUrR = rms., .r-
J rnted 

on recycled• 

paper 

Federal Recycling Program
NUREG-1744 ASSESSMENT OF THE TRAC-M CODES USING FLECHT-SEASET REFLOOD AND 

STEAM COOLING DATA

UNITED STATES 

NUCLEAR REGULATORY COMMISSION 

WASHINGTON, DC 20555-0001 

OFFICIAL BUSINESS 

PENALTY FOR PRIVATE USE, $300

MAY 2001