NUREG/CR-2730

SAND82-1150

AN

Printed August 1982

Hydrogen Burn Survival:

Preliminary Thermal Model

and Test Results

W. H. McCulloch, A. C. Ratzel, S. N. Kempka, D. T. Furgal,

J. J. Aragon

This report documents part of the Hydrogen Burn Survival (HBS)

Program being conducted by Sandia Natibnal Laboratories

Prepared by

Sandia National Laboratories

Albuquerque, New Mexico 87185 and Livermore, California 9455(

for the United States Department'6f Energy • !i

under Contract DE-AC04-76DPOO7,89

"m t ill,, 

it(

3 tý " 31 ,

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SFJ °°. 8

NUCLEAR REGULATORY COMMISSION
NOTICE

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NUREG/CR-2730

SAND82-1150

AN

Printed August 1982

HYDROGEN BURN SURVIVAL: PRELIMINARY THERMAL MODEL

AND TEST RESULTS

W. H. McCulloch, A. C. Ratzel, S. N. Kempka,

D. T. Furgal and J. J. Aragon

Sandia National Laboratories

Albuquerque, New Mexico 87185

Operated by

Sandia Corporation

for the

U.S. Department of Energy

Prepared for

Severe Accident Assessment Branch

Division of Accident Evaluation

Office of Nuclear Regulatory Research

and

Chemical Engineering Branch

Division of Engineering

Office of Nuclear Reactor Regulation

Washington, D.C. 20555

Under Interagency Agreement DOE 40-550-75

NRC Fin Nos. A-1270 and A-1306
ABSTRACT

This report documents preliminary Hydrogen Burn Survival

(HBS) Program experimental and analytical work conducted through

February 1982. The effects of hydrogen deflagrations on safetyrelated equipment in nuclear power plant containment buildings

are considered. Preliminary results from hydrogen deflagration

experiments in the Sandia Variable Geometry Experimental System

(VGES) are presented and analytical predictions for these tests

are compared and discussed. Analytical estimates of component

thermal responses to hydrogen deflagrations in the upper and

lower compartments of an ice condenser, pressurized water reactor are also presented.

ii
EXECUTIVE SUMMARY

The Hydrogen Burn Survival (HBS) Program is focused on the

study of hydrogen deflagration (non-explosive burn) effects on

the performance of safety-related (Class 1E) equipment in nuclear

power plant containment buildings. A major objective of this

program is to develop analytical methods to aid the Nuclear

Regulatory Commission (NRC) in the evaluation of component survivability analyses. The analytical models address a loss-of-coolant

accident (LOCA) scenario in which the reactor core is uncovered

and a significant amount of hydrogen is released into containment.

The released hydrogen builds up to concentrations high enough to

allow for accidental or deliberate deflagrations in the containment.

The HBS Program at Sandia proceeds from the hydrogen research

work characterizing containment building deflagration environments.

The first phase of this program consisted of studies to assess the

complexity of the equipment survivability problem. This report

summarizes the activities and presents the results of those early

scoping efforts. As advances in the experimental program and

analytical model development are made, the information contained

in this report will be supplemented.

In support of model development, a series of deflagration

tests with varying hydrogen concentrations were conducted in the

Sandia Variable Geometry Experimental System (VGES). The VGES

tests were designed to provide data for characterizing the physical mechanisms involved in hydrogen deflagrations. Black aluminum

cubes were also exposed to hydrogen combustion environments to

acquire data on the thermal response of three-dimensional objects.

These thermal response data were compared to analytical model

predictions. The CORASPN and CINDA-3G computer codes were'used

to simulate the experiments. CORASPN, developed at SNLA, models

the convective and radiative heat transfer from the products of

combustion. CINDA-3G is a general purpose heat transfer code

used for modeling conduction heat transfer in multi-dimensional

geometries.

The results of experimentation and analysis to date have

indicated that under some conditions the hydrogen burn environment is a credible threat to the survival of safety-related

equipment. (As an interim criterion for survival in this program,

it is assumed that there is reason for concern if expected component temperatures exceed those encountered in LOCA qualification

tests. While some components could certainly withstand temperatures beyond this limit, testing in addition to present LOCA

qualification programs would be required to demonstrate the continued performance of equipment during and after such exposures.)

Analytical estimates of temperatures for equipment fully-exposed

to hydrogen deflagrations in containments indicate that components

may exceed LOCA qualification temperatures (approximately 433 K

(320 0F)).

iii
Equipment is not usually installed so that it would be "fullyexposed" to hydrogen burns, i.e., components would frequently

realize significant thermal protecton from nearby structures and

hardware. Experience to date suggests that consideration of such

protection in the safety analyses of critical equipment may alleviate any need to redesign components to tolerate higher temperatures or to demonstrate (with "qualification" tests) survival at

higher temperatures.

Future HBS tasks will be focused towards determining more

precisely the environmental conditions and component responses

produced by hydrogen deflagrations and identifying those accident

conditions which produce potentially threatening environments.

It is important to note that the results presented in this report

are based on single hydrogen burns. In reality, multiple burns

are expected for plants equipped with deliberate ignition systems.

Depending upon their magnitudes and timing, multiple hydrogen

burns could make the threat more severe than indicated by the

single burn analyses. Development of the capability to analyze

multiple burns will be a part of the future HBS effort. The

overall effort will require: 1) additional experimentation to

characterize hydrogen burn phenomena and 2) continued development

and verification of analytical models used to calculate response

of safety-related components to hydrogen deflagrations in containment. The analytical models will also be used to calculate

expected input heat flux profiles for components in full scale

reactors. The calculated heat fluxes will be used in the SNLA

Radiant Heat Facility to subject some safety-related components

to simulated thermal environments typical of postulated accidents.

iv
ACKNOWLEDGEMENTS

The Hydrogen Burn Survival Program involves personnel from

several SNLA organizations. The authors gratefully acknowledge

Susan Dingman, John Cummings, Allen Camp, Michael Wester, Martin

Sherman and Marshall Berman of Division 9441; Lou Cropp, Vince

Dandini and Beth Richards of Division 9445; and Bill Benedick

and Charles Daniel of Division 1131 for their contributions to

this effort.

V - vi
.1
TABLE OF CONTENTS

Section Page

1.0

2.0

3.0

INTRODUCTION .. ..................

EXPERIMENTAL PROGRAM DESCRIPTION

2.1 Test Specimens ......

2.2 Test Information and Insights

THERMAL ANALYSIS ....... ...........

3.1 Analytical Methods .......

3.2 Code Application ........

COMPARISON OF ANALYSIS AND EXPERIMENTS

CONCLUSIONS .............

REFERENCES ......... ..............

1

* . . . 2

2

.... 06

* . . . 10

* . . . 10

* . . . 11

* . . . 217

* . . . 25

.... 25

4.0

5.0

vii
LIST OF FIGURES

Figure Page

2-1 Variable Geometry Experimental System: (a) 3

Location of Components and (b) Thermocouple

Array

2-2 Flat-Plate Calorimeter 4

2-3 Fully Instrumented Test System 9

3-1 Sequoyah Ice Condenser Pressurized Water 12

Reactor Containment Building

3-2 Temperature Profiles of Gas and Hollow 14

Aluminum Cube Resulting From a 10% Hydrogen

Burn in the Sequoyah Upper Compartment

3-3 Temperature Profiles of Gas and Hollow 14

Aluminum Cube Resulting From a 10% Hydrogen

Burn in the Sequoyah Lower Compartment

3-4 Temperature Profiles of Gas and Solid 15

Aluminum Cube Resulting From a 10% Hydrogen

Burn in the Sequoyah Upper Compartment

3-5 Temperature Profiles of Gas and Solid 15

Aluminum Cube Resulting From a 10% Hydrogen

Burn in the Sequoyah Lower Compartment

3-6 Temperature Profiles of Gas and Pressure 16

Transducer Resulting From a 10% Hydrogen Burn

in the Sequoyah Upper Compartment

4-1 Gas Temperature Profile Resulting From a 19

15% Hydrogen Burn in VGES - Predicted vs.

Thermocouple Measurements

4-2 Calorimeter Temperature Profile Resulti-ng 19

From a 15% Hydrogen Burn in VGES - Predicted

vs. Flat-Plate Calorimeter Measurements

5-1 Comparison of Temperature Profiles Obtained 22

From VGES Experiments and Sequoyah Upper

Compartment Analysis

5-2 Comparison of Hollow Aluminum Cube Thermal 22

Response Obtained From VGES Experiments and

Sequoyah Upper Containment Analysis

5-3 Predicted Gas and Component Temperature 23

Profiles Resulting From a 10% Hydrogen Burn

in the Sequoyah Containment

viii
LIST OF TABLES

Table Page

2-1 Component Description and Pertinent Property 5

Data

2-2 Test Descriptions and Maximum Calorimeter 7

Temperatures

3-1 Sequoyah PWR Pre- and Post-Combustion LOCA 13

Conditions Assumed in Modeling Effort

4-1 Initial Conditions for VGES Experimental and 18

Analytical Comparative Studies

ix

1.0 INTRODUCTION

A possible consequence of a serious nuclear power plant

accident is the release of hydrogen into the reactor containment

building. In some postulated accident scenarios, the quantity

of hydrogen released into-the containment building is sufficiently

large to create concentrations which could result in a hydrogen

explosion or fire. During the 1979 Three Mile Island accident,

a hydrogen deflagration (i.e. non-explosive burn) did occur in

containment. Although the over-pressure created by the deflagration was not sufficient to breach containment, there was evidence

of damage as a result of the severe thermal environment [1l.

In a proposed modification to its regulations [2] regarding

improved hydrogen control capability in light water reactors, the

Nuclear Regulatory Commission (NRC) addressed means of minimizing

the threat to containment integrity caused by hydrogen release.

One proposed method of alleviating the dangers of containment

failure for ice condenser pressurized water reactors (PWRs) is

to deliberately ignite the hydrogen/air mixtures using glow plugs.

This method is intended to burn off hydrogen at low concentrations

(< 10% hydrogen by volume) rather than allowing the hydrogen concentration to build up to levels which might cause containment

failure if ignited. Licensee analyses indicate that glow plug

activation during an accident would subject the equipment inside

containment to repeated hydrogen burns. For this reason, it is

important to determine whether safety-related components can

function adequately when exposed to hydrogen combustion and postcombustion environments.

To address this problem, the Hydrogen Burn Survival (HBS)

Program was initiated at Sandia National Laboratories, Albuquerque,

New Mexico (SNLA), in the last half of FY '81. Most of the effort

to date has been directed towards developing experimental and analytical techniques for predicting the response of safety-related

equipment in ice condenser PWR containments when exposed to hydrogen deflagrations. The interactive and concurrent analytical and

experimental efforts are a unique feature of the SNLA program.

Experimental data obtained in small-scale test chambers can be

used to develop and validate analytical models before they are

applied to larger containment volumes.

The purpose of this report is to present:

1. Preliminary comparisons of analytical and experimental

results obtained for small-scale hydrogen/air deflagrations

2. Preliminary temperature profiles for fully exposed equipment in the post-combustion environment in a nuclear

reactor containment building

-1-
3. Component thermal response estimates which demonstrate

important differences between deflagrations in small

test vessels and large-scale containments.

This report focuses on the insights and conclusions reached

through February 1982. It should be noted that the results to be

discussed here illustrate the development of a hydrogen burn

survival analysis methodology. Further, this work is not intended

to define specific criteria for the survivability of safety-related

equipment in hydrogen deflagration environments.

2.0 EXPERIMENTAL PROGRAM DESCRIPTION

The experimental investigation of component response to hydrogen burns began with a series of tests conducted in the Variable

Geometry Experimental System (VGES) at SNLA during November 1981.

The VGES facility, shown schematisally in Figure 2-1, is a partially buried steel vessel (5.7 m volume) instrumented with:

1) pressure transducers, 2) an array of thermocouples to detect

flame front arrival and 3) several calorimeters for monitoring

the thermal response of components to hydrogen/air deflagrations

[3]. The tank is a vertical cylinder with domed ends, approximately 4.9 m (16 ft) high and 1.2 m (4 ft) in diameter. One of

two ignition sources, a glow plug or a spark plug, was used for

each of the six HBS experiments. The ignition source was located

on the tank axis 1.2 m (4 ft) from the bottom. Figure 2-1 indicates the location of the calorimeters and other instrumentation.

Steam is an important constituent in the containment atmosphere before, during and after hydrogen/air burns. However,

because the VGES tank is partially buried and is not insulated,

the tank remains very near the soil temperature so that steam

added prior to ignition would condense. For this, reason, carbon

dioxide was injected into the VGES tank as a non-condensing

substitute for steam in two of the six tests. The tank is also

equipped with fans to thoroughly mix the gases before each test.

The fans may be left on during testing to increase turbulence

during burns.

2.1 Test Specimens

At the time the tests were conducted, samples of typical

safety-related components found in nuclear reactors were not

available. Instead, several flat-plate and three-dimensional

hollow and solid calorimeters were exposed in the VGES tank.

These calorimeters were representative of equipment in containment and were selected for their simple geometries, known thermophysical properties, and availability. In addition, several

thin, flat-plate, brass calorimeters (Figure 2-2) from another

project at SNLA were readily available. These calorimeters were

originally used as passive measurement instruments and therefore

had several temperature-indicating tapes affixed to the back

(protected) surface. Thermocouples (RdF model 20112) were also

-2-
3.68m r3. Om

DYNAMIC

PRESSURE

2. 1m

TC OUTPUT 1.2m

0.6m

0

-0. 6m

-1.2m

BRASS CALORIMETER

LI HOLLOW ALUMMIUM CUBE

SOLID ALUMINUM CUBE

(a) Location of Components

II I I I

-0. 6m 0 0. 6m

[-J SLOW-fESPONSE GAS THERMOCOUPLES

* QUICK-RESPONSE GAS TIERMOCOUPLES

(b) Thermocouple Array

Figure 2-1. Variable Geometry Experimental System (VGES)

-3--
Figure 2-2. Flat-Plate Calorimeter

-4-
Table 2-1 Component Description and Pertinent Property Data

Length

Scale Used

In Convective

Modeling

(m)

Thermal

Conductivity

(W/m-k)

Thermal

Diffusivity

(m/sec 2 )

Surface

Type Treatment

Calorimeter #1

Brass plate

w/insulation

backing

Calorimeter #2

Brass plate

w/insulation

backing

Component #1

Hollow aluminum cube

Component #2

Solid aluminum cube

Steel

Containment

Vessel

Location

Along cylinder

surface 1.52m,

from base

Centered at

top of test

section

Along cylinder

surface 1.52m

from base

Along cylinder

surface 1.52m

from base

Lower 3/4 of

tank buried

Size

15.24 cm on a

side, 0.635mm

thick with

0.635cm insuration backing

15.24cm on a

side, 0.635mm

thick with

0.635cm insulation backing

10.16cm on a

side with

0.3175cm thick

walls

10.16cm on a

side

1.22m diameter,

4.88m long with

1.59cm thick

walls

0.1524

0.1524

0.1016

0.1016

2.44

128.0

128.0

206.0

206.0

36.0

3.4(10-5) Pyromark

paint

3.4(10-5)

8.42(10-5

8.42(10-5)

1.39(10- 5)

Pyromark

paint

Anodized

Aluminum

Anodized

Aluminum

Highly

oxidized

surface

Thermal

Emissivity

0.95

0.95

0.95

0.95

0.70
attached to the back side of these flat-plate calorimeters. The

exposed faces were spray painted with a high-temperature, flatblack paint to maximize heat transfer (absorption of the radiative

flux) from the combustion gases.

The three-dimensional calorimeters were fabricated to represent components which might be found in reactor containments.

Their simple geometries were selected to facilitate analysis of j

the test data. A hollow aluminum cube (-3.2 mm (- 0.13 in)

wall thickness) 10.2 cm (4.0 in) on a side was used to characterize

components with low heat capacitance. A solid aluminum cube was

fabricated to represent the more massive components expected to

be found in containment. To maximize heat transfer (i.e., the

absorption of radiative flux) these calorimeters were black

anodized. Additional geometry and thermophysical property data

for the calorimeters and components used in the experimental

modeling effort are summarized in Table 2-1.

2.2 Test Information and Insights

Six tests were conducted with varying hydrogen and carbon

dioxide concentrations, ignition sources, and gas mixing as indicated in Table 2-2. Selected maximum temperature and pressure

data recorded during the tests are also shown in the table.

Several observations can be made based on the results given in

Table 2-2. These observations are summarized below.

1. obviously, burning at higher hydrogen concentrations

produces higher gas temperatures and pressures and this

results in higher component temperatures.

2. The presence of carbon dioxide significantly lowered

the measured peak gas temperature but had little impact

on reducing the maximum temperatures reached by the components.' (A possible explanation is that though the

carbon dioxide absorbed some of the energy from combustion (i.e. energy sink) holding the gas temperature down,

it also increased the post-combustion radiation heat

transfer from the gas to the tank walls and to the components because it increased the emittance of the gas.

Noncondensing steam in large containment volumes would

be expected to have a similar effect on component temperatures.)

3. Operating the mixing fans during combustion had no discernable impact on the peak gas temperatures but significantly increased the measured gas pressure. (We would

have expected the gas temperatures to increase when the

fans were operating. There were not enough redundant

gas temperature measurements (from quick response thermocouples) to verify this expectation).

-6-
Table 2-2 Test Descriptions and Maximum Calorimeter Temperatures

B88 B89 B90 B91 B92 B93

% H2 10.5 10.1 10.1

10.1

15.2 15.0

% CO 2 0 0 10.0

ON OFF

0 0

Mixing Fans OFF ON OFF ON

Igniter Spark Spark Glow

Plug

Glow

Plug

Glow

Plug

Spark

MAXIMUM TEMPERATURES (K)*

Gas 1200

Flat Plate

Hollow Cube

Solid Cube

420

325

310

1200

420

330

305

1060

430

330

305

1075

400

335

1505

460

340

1475

455

345

310 305

MAXIMUM PRESSURE (ATM)

Transducer 2.6 3.6 2.6 2.8 3.5 4.9

* The temperatures measured were obtained from unreferenced,

uncalibrated chromel-alumel thermocouples of different wire

diameters and different thermocouple junction sizes. The

accuracy of these measurements is unknown.

-7-
4. The temperatures measured for the flat-plate calorimeters exposed to an environment resulting from a 10%

hydrogen deflagration in VGES are very close to LOCA

qualification temperatures. As will be discussed later,

the time required for the gas to cool-down in these

tests was very short compared to that expected in containment deflagrations. This means that the environment

of the small test chamber is much less severe than that

in large containment buildings. Therefore, we are concerned that deflagrations in large scale containments

will produce component temperatures in excess of LOCA

qualification specifications.

The VGES facility was designed to study hydrogen burn propagation and the included thermocouple array was utilized to indicate

flame front location. Since the data of interest were the arrival

times of the hydrogen flame (as indicated by the onset of an

increase in the thermocouple output), the thermocouples were made

of relatively heavy gauge wire to withstand repeated exposures.

In the HBS Program there is more concern for the actual gas temperatures, suggesting smaller, more responsive thermocouples.

The VGES facility included one such thermocouple mounted near

the wall at the top of the tank (see Figure 2-1). Its proximity

to the wall may have caused significant deviation from the actual

gas temperature. Future experimental activities will include

efforts to identify more suitable thermocouples and mounting

techniques for the measurement of gas temperatures.

In analyzing the experimental data, it became apparent that

a heat flux gauge capable of responding to rapid thermal transients is needed. Instrumenting calorimeters with additional

thermocouples would also assist in verifying the accuracy of

component thermal responses through redundant measurements.

These data are necessary for estimating the energy transfer to

the components from the combustion gas products to provide inputs

for future Radiant Heat Test Facility experiments.

Future experimental efforts will be conducted in the Fully

Instrumented Test System (FITS) at SNLA (shown in Figure 2-3).

The use of the FITS facility provides the following extensions

to the experimental program:

* More extensive instrumentation

* Observation of the condensation phenomenon

* Inclusion of steam injection in the initial gas mixture

at saturated or superheated conditions

* Elevated initial gas and wall temperatures more typical

of actual conditions resulting from reactor accident

scenarios

-8-
Steam Inlet

Detonation Tube

Chamber I

Static Pressure and TemperatureI I

I I

- Fog Water Supply

Removable Head

Access Walk

Thermocouple

and

Gas Sampling Probes

Membrane Pressure

Main Containment

350 psi W. P.

Termo-~~Z

v i!flz I q~[-Dynamic Pressure - - -f Shadowgraph

-" Camera

-- Igniter

Drain and Purge

Hydrogen I

1.5 m (4.9 ft) diameter 3.4 m (11.2 ft) length

5.6 m3 (195 ft3 ) volume

Figure 2-3. Fully Instrumented Test System (FITS)

-9-
* Possible implementation of gas velocity measurement

techniques.

In general, the VGES tests provided valuable information for

understanding the hydrogen combustion phenomena and associated

heat transfer processes. The tests have also provided insight

into the experimental procedures and apparatus required to obtain

more meaningful data in future tests. Further discussion pertaining to VGES results will be presented in Section 5 of this

report.

3.0 THERMAL ANALYSIS

3.1 Analytical Methods

The modeling effort used an SNLA developed computer code,

CORASPN, which predicts post-deflagration thermal responses in a

containment vessel. CORASPN assumes a uniform isothermal gas

volume for radiative and convective energy transfer and accounts

for possible steam condensation on containment and component

surfaces. All components and surfaces comprising the containment

are modeled as flat plates with adiabatic back surfaces. The

net energy transfer from the gas is assumed to be uniform over

each surface area. CORASPN models the transient energy processes

as quasi-steady-state energy exchanges during each time step and

uncouples the radiative, convective and conductive effects in

obtaining a total energy balance. Note that CORASPN is the first

generation computer simulation model from which the code HECTR

has evolved [3]. The latter code will be used in future analytical modeling efforts.

The most important early time energy transfer mechanism

following combustion is radiative exchange from the combustion

products to the containment surfaces. Radiative transfer calculations for the gaseous participating media are accomplished

using the exponential wide band models for steam and carbon

dioxide summarized by D. K. Edwards in a monograph entitled

"Molecular Gas Band Radiation" [4]. Necessary inputs to these

models include the steam and carbon dioxide partial pressures

and temperatures and the effective 'beam' (path) lengths between

surfaces. Though the enclosure surfaces are significantly cooler

than the gas in the early times of the analysis, radiative transfer

from the enclosure surfaces is included for gray diffuse walls

using the net radiation method described by Siegel and Howell [5].

Convective energy transfer may also be included in CORASPN,

depending upon user preference. Natural and forced convective

heat transfer coefficients are computed for each time step and

the larger contribution is used. A major shortcoming in the

convective analysis requires that the user must provide an estimated uniform free-stream gas velocity since the hydrodynamics

are omitted from modeling. In addition, the convective heat

transfer correlations used may be inappropriate when the free

-10-
stream gas temperature is significantly different from the

enclosure temperature. Steam condensation is accounted for

using an infinite diffusion rate, air-steam Nusselt film condensation model. The effect of noncondensables and the omission of

possible drop-wise condensation are major shortcomings of this

condensation model.

Finally, in order to account for multi-dimensional effects

in the hollow and solid cubes and in an actual pressure transducer

commonly used in many reactors, a series of CINDA 3-G [6] thermal

models were developed. CORASPN was modified into a subroutine

package to provide gas temperatures and radiative and convective

heat inputs to the surfaces of the three-dimensional component

models. By utilizing the CINDA 3-G models, the unrealistic backside adiabatic boundary condition is eliminated in the analyses

and temperature distributions within the components can be predicted.

3.2 Code Application

The conditions modeled by CORASPN result from a hydrogen

deflagration in a S 2 D accident sequence in the Sequoyah ice

condenser PWR. The "S 2 D" designation refers to a specific

loss-of-coolant accident (LOCA) where coolant is lost through a

small break in the primary cooling system and the emergency core

cooling system is inoperative. The pre-combustion containment

environment and the conditions resulting from an isochoric (constant volume), adiabatic (no heat loss) combustion of the gas

mixture are summarized in Table 3-1. These pre-ignition conditions were obtained at SNLA from an S 2 D LOCA simulation using

the computer code MARCH [7] developed by Battelle Laboratories.

The isochoric, adiabatic combustion environment given in Table

3-1 provided the initial conditions in the CORASPN analyses.

Our analyses considered hydrogen deflagrations in both the

upper and lower compartments of the Sequoyah containment. A

schematic of this reactor containment is given in Figure 3-1.

These analyses assumed that the spray cooling system was inoperative. Expansion of lower compartment combustion gases through

the ice condenser units and the influx of cool (310 K (99 0 C))

upper compartment air into the lower compartmentwere simulated

in the lower compartment analysis. An upper compartment deflagration represents the more severe accident scenario since there

is no gas recirculation (constant volume process).

-11-
Figure 3-1. Sequoyah Ice Condenser Pressurized Water Reactor

Containment Building

-12-
Table 3-1 Sequoyah Pre- and Post-Combustion LOCA Conditions Assumed in Modeling Effort

Conditions Prior to Ignition Post Combustion Conditions

-310 K gas temperature 1353 K gas temperature

-1.5 atm gas pressure 6.6 atm gas pressure

-0.05 steam fraction 0.158 steam fraction

0.10 hydrogen fraction

Time after accident initiation 1.3 hours.,

Upper compartment surface area 3781 m2

Lower compartment surface area 9693 m2

Upper compartment total volume 20000 m3

Lower compartment total volume 10980 m3

To estimate component response to the Sequoyah upper and

lower compartment burns described above, the solid and the hollow

aluminum cubes described in Section 2.1 were analyzed using

CORASPN and CINDA-3G codes. The results of these analyses are

shown in Figures 3-2 through 3-5. In the CINDA-3G analyses, the

entire surface area of each component is exposed to a uniform

heat flux obtained from CORASPN calculations. Figure 3-2 shows

that the hollow cube in the upper compartment, with sprays inoperative, experiences a maximum temperature of 722 K (840 0 F). In the

lower compartment analysis the hollow cube experiences a lower

maximum temperature of 580 K (584 0 F) as shown in Figure 3-3

resulting from the gas recirculation. The gas temperature profiles to which these components are exposed are also shown in

the figures. When the solid cube is exposed to the same postcombustion environments, peak temperatures of 437 K (327°F) in

the upper containment and 372 K (211*F) in the lower compartment

result (see Figures 3-4 and 3-5). These analyses suggest that

a component in the Sequoyah containment fully-exposed to a 10%

hydrogen burn could experience a wide range of maximum temperatures (between 372 K and 722 K (211 and 840*F)) depending upon

its thermal mass and its location in containment (i.e., upper or

lower compartment). Note that if the spray systems were operative these maximum temperatures would be significantly reduced.

A post-combustion thermal analysis was also applied to a

typical pressure transducer found in containment. This piece of

equipment was comprised of a 0.6 cm (0.25 in) thick carbon steel

casing which encloses a printed electronic circuit board and a

bourdon gauge. Figure 3-6 shows the thermal response of this

pressure transducer in the upper compartment, subject to deflagration conditions defined in Table 3-2. The computer model predicted

that the carbon steel casing and printed circuit board would

reach temperatures of 477 K (400°F). Again, these results were

obtained assuming that all surfaces were exposed to the post-

-13-
K1365 12000

1155 1 1620

0451 1240

I

w

I735

520

310

660

480

100

100 200 300 400 500

TIME AFTER DEFLAGRATION (sec)

Figure 3-2. Temperature Profiles of Gas and Hollow Aluminum Cube Resulting

From a 10 % Hydrogen Bum in the Sequoyah Upper Compartment

K 'F

1355 2000

1155 1 1620

945 1 1240 wU

IU

w

A.

U

735

520

310

550

480

100

0 100 200 300 400 500

TIME AFTER DEFLAGRATION (sac)

Figure 3-3. Temperature Profiles of Gas and Hollow Aluminum Cube Resulting

From a 10% Hydrogen Bum in the Sequoyah Lower Compartment

-14-
K 'F

1365 12000

1155 11620

945 1 1240

cc

735 860

520 1 480

310 100

0 100 2UU 300 400 800

TIME AFTER DEFLAGRATION (sec)

Figure 3-4. Temperature Profiles Gas and Solid Aluminum Cube Resulting From a

10 % Hydrogen Burn in the Sequoyah Upper Compartment

K

1365

-F

2000

1155 1620

945 11240

4

I735 1860

520 1 480

4 310 1 100

0 100 200 300

TIME AFTER DEFLAGRATION (sec)

400 500

Figure 3-5. Temperature Profiles of Gas and Solid Aluminum Cube Resulting From

a 10% Hydrogen Burn in the Sequoyah Lower Compartment

-15-
K 'F

1365 2000

1145 1 1800

920 1 1200

'U

I700

480

255

800

400

0

0 100 200 300 400 500

TIME AFTER DEFLAGRATION (9ee)

Figure 3-6. Temperature Profiles of Gas and Pressure Transducer Resulting From

a 10% Hydrogen Burn in the Sequoyah Upper Compartment

-16-
combustion environment. Thermal shielding of any transducer

surface would be expected to reduce the maximum temperatures

reported here.

Overall, these analyses indicate that the continued operation

of safety-related equipment could be impaired if these components

are fully-exposed to thermal environments resulting from a 10%

hydrogen/air combustion. CORASPN simulations of component thermal

responses have demonstrated two factors to consider in component

survival analyses. First, component casings can provide substantial protection to internal component instrumentation if the

thermal mass of the casing is large. Low thermal mass (i.e.,

thin-wall) components, conversely, could be expected to exceed

LOCA qualification temperatures (approximately 433 K (320 0 F)).

Second, the component location (upper or lower compartment) is

significant. Analyses for the same component in the two compartments resulted in a higher maximum temperature for the upper

compartment case since the environment there is more severe.

Note that the local proximity of these components to containment

structures has not been included in this work and hence thermal

shielding effects have not been taken into account.

4.0 COMPARISON OF ANALYSIS AND EXPERIMENTS

As with any analytical method, validation of predicted results

through experiments is extremely desirable. In this instance,

however, test facilities approaching the volumes of reactor containment buildings are not readily available for experimentation.

Therefore, verification of analytical (numerical) models must be

conducted in scaled-down facilities to gain confidence in phenomenological modeling. The heat transfer phenomena involved

(principally radiation) are notsubject to conventional scaling

laws. Hence results obtained from small-scale tests cannot be

extrapolated directly to larger containment volumes (the ratio

of containment volumes to the volume of the test facilities like

VGES are of the order of 4000-15000 to 1). Rather, they serve

only to verify the analytical models.

Experimental data obtained in VGES tests were compared with

CORASPN results. Initially, experimental results from the entire

test series were evaluated to determine thermocouple response uniformity from test to test. Thermocouple performance appeared

consistent under similar fan circulation conditions during the

tests (fans on/fans off). It was observed that gas circulation

(fans on) during the post-deflagration period tended to increase

the rate of gas temperature cool-down as.would be expected. Following this cursory evaluation, the 15% hydrogen deflagration test

conducted on November 12, 1981 was selected for detailed comparison

with CORASPN predictions. Initial conditions for the experiment

and the associated input conditions to CORASPN are summarized in

Table 4-1.

-17-
Table 4-1 Initial Conditions for VGES Experimental and Comparative Studies

Initial Experimental Conditions for VGES Test of

November 12, 1981

Gas Pressure 0.971 ATM

Gas Temperature 297 K

Hydrogen Content by Volume 0.1504

Moisture Content 0.001 humidity ratio

(assumed)

Ignition Time 3:25 PM

Igniter System Used 70 V glow plug

Circulating Fans Turned off Prior to ignition

Post-Combustion Input Conditions to CORASPN

Isochoric Temperature

Isochoric Pressure

Steam Volume Fraction

Average Gas Velocity

Component and Containment

Temperature

Spray Mitigation Scheme

1803.5 K

5.43 ATM

0.163

0.3 m/sec (assumed)

297 K

Inoperative

Figure 4-1 shows a comparison of CORASPN-calculated gas

temperature profiles and measured temperature profiles for a 15%

hydrogen deflagration in the VGES chamber. The gas temperatures

calculated by CORASPN approach experimental results after 20-25

seconds. However, even though the adiabatic assumption precludes

heat loss during'combustion, shortly after ignition, the calculations underestimate actual measurements. At the beginning of the

burn, the dominant energy transfer mechanism is radiative cooling.

Since radiation is proportional to the fourth power of temperature,

underestimating the gas temperature during this period could produce significantly lower peak component temperatures. The difference between predicted and measured temperatures for the two

brass flat-plate calorimeters (described in Table 2-1) is illustrated in Figure 4-2. The measured maximum temperatures significantly exceed the CORASPN calculated maximum. Similar trends

were noted in analyzing the solid and hollow aluminum cubes,

though the differences between experimental and analytical results

were smaller.

In interpreting these discrepancies, it appears that significant energy is prematurely removed from the hot combustion gases

in the analytical model immediately after the deflagration is

completed. We believe this may be attributed to post-combustion

initial conditions and/or to the condensation model employed by

-18-
K °F

1145 1600F

1033 11400

920 11200

810 11000

<

I700 1 800

5

4.

3'

2.

90 600

75 400

65 200

55 0 I I I I I I

0 10 -20 30 40 so 60 70 80 90 100

TIME (sec)

Figure 4-1. Gas Temperature Response Resulting From a 15% Hydrogen

Burn in VGES - Predicted vs. Thermocouple Measurements.

K

480

4F

400

460 l- 365

440 F- 330

420 - 295

I..

4

400 - 260

380 t- 225

360 - 190

340 - 155

320 l- 120

300 H a5

285 L- 50 0 10 20 30 40 50 s0 70 s0 90 100 110

TIME (sec)

Figure 4-2. Gas Temperature Response Resulting from 15% Hydrogen BurnPredicted vs. Flat- Plate Calorimeter Measurements.

-19-
CORASPN. Initially, the uniform gas temperature in CORASPN is

equal to the adiabatic, isochoric combustion temperature. Although

it has been reported in the literature that this temperature may

be used as an average gas temperature following combustion, there

may be local maximums substantially exceeding the adiabatic, isochoric combustion temperature. For example, Garforth and Rallis

[8] found that temperatures for stochiometric methane combustion

in an 8-cm (3.2 in) radius spherical vessel varied by + 500 K

(440'F) from the adiabatic isochoric temperature. This nonuniformity is attributed to compressive heating of the gas during

the combustion process through the volume. Previous work (e.g.,

Takeno [9]) has demonstrated that for combustion in spherical

vessels, the gas temperature near the ignition source in the

absence of heat loss from the gas will be 100-300 K (180-540°F)

higher than at the flame front location. We suspect that this

variation may be even more significant for nearly planar flame

propagation such as is expected in the VGES test chamber. It

should be noted that while the gas temperature may vary in the

test chamber, the pressure created by the burn will be essentially

uniform as assumed in CORASPN and measured in the VGES experiments.

In general, the adiabatic, isochoric combustion temperature

used as the source temperature for heat transfer may not represent

the maximum credible threat to components in containment. If a

component is near the ignition site, it could be exposed to gases

substantially hotter (hundreds of degrees Kelvin) for short

periods of time. We note that most of the previous work on

hydrogen deflagrations in reactor containments has addressed the

pressure rise and decay phenomena where this temperature discrepancy is of no major consequence.

Added difficulties in interpreting the analytical and experimental results arise from uncertainties pertaining to the condensation process. The VGES facility has a large surface area to gas

volume ratio andthe walls are cool; thus condensation effects

are critical. The condensation model incorporated in CORASPN

predicts immediate steam condensation following combustion. In

the analytical model, the condensate is instantaneously diffused

to the containment walls, blanketing the surfaces with a highly

absorbing film (water absorptivity is 0.94). This condensation

phenomenon removes significant quantities of thermal energy from

the gas and causes the gas temperature and pressure to decrease

rapidly. The radiative heating of the calorimeters and other

components is correspondingly reduced. It should be noted that

the condensation problem is not trivial, as the process depends

not only on the gas and wall temperatures, but also on the gas

hydrodynamics, on the steam concentrations and on the quantity

of noncondensable gases. Additional review of the condensation

process is planned to determine if a more suitable model may be

developed.

As a programmatic note, it should be pointed out that the

HBS Program tasks were originally defined based on the premise

-20-
that the hydrogen burn environment had been well defined and

characterized. As indicated above, the understanding of relevant

mechanisms and the representation of the phenomena in computer

codes are actually in a state of continuing development. This

will have some impact on the format and content of the analytical

procedures (i.e., algorithm) eventually produced by this effort.

Specifically, the algorithm must include flexibility to encompass

variables previously thought to be relatively fixed.

5.0 CONCLUSIONS

During the development and evaluation of HBS Program results,

several conclusions became apparent. They are summarized below.

1. In comparing analytical results from VGES and Sequoyah

modeling, we noted that gas transient cool-down periods

differed by several orders of magnitude, as is shown in

Figure 5-1. This implies that a component in a nuclear

reactor containment would be exposed to elevated gas

temperatures for a much longer period of time (depending

on the size of the containment volume) and would thus

reach substantially higher temperatures than in a smaller

test facility. This phenomenon is illustrated in Figure

5-2 which shows a component thermal response given the

experimental and containment environments depicted in

Figure 5-1. The apparent scale effect must be considered

when evaluating data from small-scale test facilities.

2. Analyses using the CORASPN code estimated component thermal responses in a nuclear reactor containment subject

to a 10% hydrogen deflagration with inoperative spray

systems. The range of computed maximum and minimum gas

and component temperatures (taken from Figures 3-2 and

3-5) are sui'marized in Figure 5-3. The component thermal

response results were obtained assuming that the components were fully exposed to a single, hydrogen/air deflagration in the Sequoyah containment. The upper limit

(Curve #2) is the calculated response of a low-thermal

mass component in the upper compartment (a severe thermal

environment with no sprays and no recirculation). The

lower limit (Curve #4) is the response of a high-thermal

mass component in the lower compartment (less severe

environment with gas recirculation). The wide range of

temperature extremes is the result of covering contingencies in component mass, component location and availability of sprays. The conclusion is that under some

conditions a low-thermal-mass component could reach

temperatures well above the LOCA qualification specifications. Furthermore, depending upon their magnitudes and

timing, multiple burns could make the threat even worse.

-21-
K 0 13651 2000F

1155 11620

I-945

'a

I-

*735

520

310

1 1240

860

480

100

100 200 300 400 500

TIME AFTER DEFLAGRATION (sec)

Figure 5-1. Comparison of Temperature Profiles Obtained From VGES

Experiments and Sequoyah Upper Compartment Analysis

K *F

,3..l '000

1155

I"946

735

520

310

820

640

460

280

100

100

TIME AFTER DEFLAGRATION (see)

Figure 5-2. Comparison of Hollow Aluminum Cube Thermal Response

Obtained From VGES Experiments and Sequoyah Upper

Compartment Analysis

-22-
K

1365

OF

2000

1165 1 1620

9451 1240

7

785

'I I

LEGEND

1. UPPER COMPARTMENT TEMPERATURE

2. HOLLOW CUBE, UPPER COMPARTMENT

T3. LOWER COMPARTMENT GAS TEMPERATURE

4. SOLID CUBE, LOWER COMPARTMENT

~2

\

~4

a I , I , r" -- ---- -- -

860

5201 480

3101 100

0 100 200 300 400 500

TIME AFTER DEFLAGRATION (sec)

Figure 5-3. Predicted Sequoyah Gas and Hypothetical Component Profiles

Resulting From a 10% Hydrogen Burn in the Sequoyah Containment

-23-
3. It is possible that localized areas in containment will

experience temperatures in excess of those calculated in

these analyses if the gas temperature field is non-uniform.

The analytical models used to characterize hydrogen burn

environments usually assume that the maximum temperatures

and pressures achieved are those resulting from adiabatic,

isochoric deflagrations. While this approach yields

satisfactory average gas temperatures and pressures, comparison of such calculations with VGES test data gives

reason to suspect that temperatures in certain localized

areas could be in excess of the isochoric combustion temperature. Therefore, until this possibility has been

given further consideration, the analyses presented should

not be regarded as an absolutely conservative upper limit

of expected component temperatures. Also, note that in

our analyses every component surface was exposed to the

post-combustion thermal environment. Thermal shielding

from nearby structures and equipment would be expected

to limit the exposure of at least some component surfaces

to these environments. This would result in reduced

maximum component temperatures. The degree of protection

would be determined by detailed thermal analyses of each

application.

In summary, preliminary HBS Program efforts suggest that

under certain conditions component temperatures higher than

LOCA qualification specifications can be expected as a result of

single or multiple hydrogen deflagrations. The response of components to these challenging environments is dependent upon their

location and configuration in containment and must be evaluated

with detailed specific analyses. Continuing HBS Program analytical and experimental efforts will be directed towards the definition of conditions under which components might be threatened

and towards the continued development of reliable thermal models

for analyzing safety-related components in hydrogen deflagration

environments.

-24-
REFERENCES

1. Alvares, N. J., Beason, D. G. and Eidam, G. J., "Investigation of Hydrogen Burn Damage in the Three Mile Island Unit

2 Reactor Building", GEND-INF-023, Vol. 1, June 1982.

2. "Interim Requirements Related to Hydrogen Control", Federal

Register, Vol. 46, No. 246, December 23, 1981, pp. 62281-

62285.

3. Berman, M., "Light Water Reactor Safety Research Program

Semiannual Report, April-September 1981", Sandia National

Laboratories, NUREG/CR-2481, SAND82-0006, February, 1982.

4. Edwards, D. K., "Molecular Gas Band Radiation", Advances in

Heat Transfer, Vol. 12, T. F. Irvine and J. P. Harnett, ed.,

Academic Press, New York, 1976.

5. Siegel, R. and Howell, J. R., Thermal Radiation Heat Transfer,

Second Edition, McGraw Hill Book Co., New York, 1981.

6. TN-AP-67-287, "Chrysler Improved Numerical Differencing

Analyzer (CINDA) for Third-Generation Computers", October 20,

1967.

7. Prassinos, P., Sandia Report in preparation, MARCH analysis

of S 2 D accident in the Sequoyah PWR.

8. Garforth, A. M. and Rallis, C. J., "Gas Movement During Flame

Propagation in a Constant Volume Bomb", Acta Astronautica,

Vol. 3, p. 879, April 23, 1982.

9. Takeno, T. and Iijima, T., "Theoretical Study of Nonsteady

Flame Propagation in Closed Vessels", Progress in Astronautics

and Aeronautics, Vol. 76, pp. 578-595.

-25-
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