ML20094P266

Rev a to MV-6570-001, Probabilistic Evaluation of Crystal River Power Station Containment Performance for Beyond Design Basis Conditions

ML20094P266
Person / Time
Site:	Crystal River
Issue date:	03/01/1991
From:	Hadiditamjed, Nakaki D, Wesley D ABB IMPELL CORP. (FORMERLY IMPELL CORP.)
To:
Shared Package
ML20094P270	List: ML20094P266
References
MV-6570-001, MV-6570-001-RA, MV-6570-1, MV-6570-1-RA, NUDOCS 9511280390
Download: ML20094P266 (157)

Text

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MV-6570-001, Rev. A PROBABILISTIC EVALUATION OF THE CRYSTAL RIVER POWER STATION CONTAINMENT PERFORMANCE FOR BEYOND DESIGN BASIS CONDITIONS l

Prepared for:

RISK MANAGEMENT ASSOCIATES 2309 Dietz Farm Road N.W.

i Albuquerque,NewMexico 87107 C

Prepared by:

ABB IMPELL CORPORATION 27401 Los Altos Suite 480 MissionViejo,Califomia 92691 D. A. Wesley D. K. Nakald H. HadicsTamjed s.Lu i-l March,1991 M

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e 9511290390 951122 PDR ADOCK 05000302 M W Corporata l

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SN91 1

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mme REPORT APPROVAL COVER SHEET l

RISK MANAGEMENT ASSOCIATES CLIEP(T:

Crystal River Unit 3 Containment Overpressure Evaluation PROJECT:

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JOB NUMBER (S):

PROBABILISTIC EVALUATION OF THE CRYSTAL RIVER POWER STATION REPORT TITLE:

CONTAINMENT PERFORMANCE FOR BEYOND DESIGN BASIS CONDITIONS MV-6570-001 REPORT NUMBER:

REVISION RECORD REY.

PREPARED REVIEWED APPROVED DATE A

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TABLE OF CONTENTS 1

SECTION DESCRIPTION PAGE Title Page Report Approval Cover Sheet Table d Contents i

Record d Revisions lii Ust d Tables iv i

ABSTRACT v

1 INTRODUCTION 1-1 2

PROBABluSTIC DESCRIPTION OF CAPACITY 2-1 2.1 Development d Failure Probabilities 2-1 2.2 Variability in Material Properties 2-3 2.2.1 Prestressing Tendons 2-4 l

2.2.2 Reinforcing Bars 2-5 l

2.2.3 Uner Plate 2-5 l

l 2.2.4 Concrete 2-6 2.3 ModelingVariability 2-6 i

3 ANALYSIS OF REACTOR BUILDING CAPACITY 3-1 3.1 Membrane Failure d the Reactor Building Shell 3-2 3.1.1 Hoop Membrane Failure in the Cygnder Wall 3-3 3.1.2 Meridional Membrane Failure in the Cylinder Wall 3-5 3.1.3 Dome Membrane Failure 3-6

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SECTION DESCRIPTION PAQf 3.2 Flexural Failure at the Wall-Base Junction 3-7 3.3 Failure of the Basemat 3-10 l

4 REACTOR CAVITY ACCESS TUNNEL DOORS 4-1 5

CORRELATION OF FAILURE MODES 5-1 REFERENCES R-1 APPENDIX A - CHARACTERISTICS OF THE LOGNORMAL A-1 DISTRIBUTION L

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LIST OF TABLES 1

TABLE TjTLE PA.Q..E 3-1 Pressure Capacities for the Reactor Building Structural Fail-ure Modes -Temperature Case 1 3-12 3-2 Pressure Capacities for the Reactor Building Structural Fail-ure Modes -Temperature Case 2 3-13 3-3 Pressure Capacities for the Reactor Building Structural Fail-ure Modes -Temperature Case 3 3-14 4-1 DF-2 and DF-3 Door Pressure Capacities 42 i

51 Failure Mode 5-2 5-2 Correlation Between Failure Modes (Strength and Modeling 5-3 Uncertainty) t 1

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ABSTRACT t

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A probabilistic evaluation of the Crystal River Unit 3 Nuclear Power Station con-tainment performance has been conducted in support of a probabilistic safety assessment program. Potential failure modes of the reactor building due to temperature and pressure 7

loadings well beyond the design basis conditions were considered in this evaluation. In all l

cases, failure was defined to be incipient leakage or a breach of the pressure boundary. The failure modes considered were gross structural failures of the concrete containment structure.

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Smaller, local failures at the containment penetrations were not considered. The capacities of the various failure modes were described in' terms of the internal containment pressure required to induce failure. It was assumed that the pressure capacdies associated with al of l-the failure modes could be treated as quasi-static. No dynamic ampBlication effects of the i

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pressure loading were considered. Three temperature ' cases were considered in which the

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Interior containment temperatures were assumed to be 300", 50(r, and 800' F. The temperature conditions considered in this investigation were assumed to correspond to steady state j

rnatorial temperatures.

Tofitwithinthecontextof aprohahilia* safetyassessment,thepressurecapacities of the respective failure modes were treated as lognormally distributed random variables, with the distribution of each pressure capacity descrbed by a median (50th percentue) pressure and a logarithmic standard deviation. The loganthmic standard deviation represented the uncertainty in the median value due to variabilty in the material properties and analytical modeling. This formulation allows for the probability of failure to be estimated as a func6an of pressure for each of the failure modes. Thus, in this irn::yk i, the median pressure j

capacities and logarithmic standard deviations are estimated for the potential faNure modes.

Several potential feture modes of the concrete containment tructure were eval-s usted. The failure modes of the reactor building shell and the basemat included: membrane tension failures of the cylinder and dome pcrdens of the sher, flexural failure at the base of l

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MV 6570-001, Rev. A i

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the containment wall, and flexure and shear failures of the basemat. All of these failure modes corresponded to gross structural failures and were assumed to lead to large leak areas and rapid depressurization of the containment.

Overall, by ranking the potential failure modes by their median pressure capacities, the controlling failure mode identified in this study was associated with the flexural failure of the wall-basemat Junction. However, by ranking the potential failure modes by their High Confidence of Low Probability of Failure (HCLPF) capacities, the controlling failure mode for the first two temperature cases was the basemat flexural failure, while that for the third tem-perature case was the wall-basemat junction flexural failure.

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1. INTRODUCTION i

3-1 A probabilistic safety assessment is being conducted for the Crystal River Nuclear Power Station to estimate the risk arising from a release of radioactivity from the site. ABB

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impell Corporation is under contract with Risk Management Associates to evaluate the capacity

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l of the Crystal River Unit 3 containment structure for elevated temperature and pressure loadings. Consistent with the nature of the probabilistic safety assessment, the evaluation methodology is based on estimating the capacity of the containment structure in terms of L

probabHistic parameters.

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The Crystal River containment structure is a pre-stressed concrete shell having the L

form of a circular cyundrical sheH capped by a torispherical dome. The inside radius of the i

concrete cylinder and the dome is 65 feet. The cylinder wat is 3.5 feet thick, while the dome is 3.00 feet thick. The springline for the dome is located 157 feet above the top of the basemat.

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1he containment is uned with a 3/8 inch thick mild, ductile steel plate. In the containment sher, the primary structural resistance is provided by unbonded, post-tensioned tendons in both the hoop and meridional directions. The basemat is a 12.5 feet thick conventional reinforced i

concrete foundation mat.

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This report documents the evaluation of the probabilistic capacity of the contain-

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ment structure. Several potential failure modes were investigated for the containment in which

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failure was defined as incipient leakage or a breach of the pressure boundary. Failure resulting from direct pressure induced modes such as membrane buiding shou failure was considered only. Failure from indirect, deformation induced modes such as pipe penetration failure

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resulting from the relative deformation of the bulldog sheE and the wrj stationary pipe support, either inside or outside the containment, was not considered. Thus, the failure modes included only gross structural failures and not amener leak failurea. These failure modes are l

evaluated for temperature conditions welin~ excess of me design accident temperature.

Median (50th percentile) failure pressures and their associated variabilities are estimated.

l Using these values, the probability of failure can be actimated as a function of pressure for the i

sae.veig failure modes.

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MV-6570@1, Rev. A For the investigation reported here, it was assumed that the failure pressures associated with all modes of failure could be treated as quasi-static (i.e., pressure rise times of atleast several seconds are assumed). Effects such as dynamic amplification of the pressure pulse on the reactor building shell or internal pressure wave loading on cables and equipment -

are not considered here. In addition, all temperatures in the materials were assumed to cor-respond to steady state conditions.

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2. PROBABILISTIC DESCRIPTION OF CAPACITY j

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1 In order to fit into the probabilistic safety assessment context, a probabilistic j

i description of the pressure capacity of the containment structure is required. Also, since the pressure capacity is treated as a random variable, it is possible for more than one failure mode to significantly contribute to the overall risk. Therefore, several potential failure modes must

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be evaluated. This section discusses how the capacities are described in a probabilistic form.

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2.1 Development of Failure Probabilities l

The pressure capacities are evaluated using limit state analyses for the various failure modes considered. in this investigation, failure is interpreted as incipient leakage due

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to a large catastrophic rupture. The calculated capacities are dependent on several factors, I

including the material properties, modeling assumptions, and the postulated failure criteria. A major source of uncertainty in the failure criteria ls the expected strain resulting in failure. Biaxial strains, gage length effects, and strain concentrations can greatly reduce the expected strain i

at failure when compared to the elongation data developed from standard specimen ultimate l

l tests. Considerable variability is introduced, not only in the failure criteria but in analytical i

modeling and other assumptions as well. In view of the expected variability in the base parameters and the analytical methods, the pressure capacity for any failure mode is con-l sidered to be a random variable. It is assumed that the pressure capacities are characterized by a lognormal probability distribution. Tho 'cgrcireal distribution is a mathematically tractable 4

distribution and has been shown to be a valid desc.fiption for the variability in material strengths.

In addition, for a random variable that can be expressed as the product and quotient of several I

random variablea, the distribution of the dependent variable tends to be lognormal regardless of the distributions of the independent base varinoles.

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l With the pressure capacity for a given failure mode assumed to be'a lognormal random variable and denoting it as P, the probability of failure occurring at a pressure less than a specific value p is expressed as I

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(2-1)

Prob (P 5 p)

P f

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P = probability that failure occurs at a pressure P $ p f

i P = random pressure capacity l

p = logarithmic standard deviation of P i

P = median (50th percentile) pressure capacity 4(-) = cumulative distribution function for a standard normal random variable C

in equation (2-1), the pressure capacity for a given failure mode is probabilistically described by the following expression fMS (2-2)

P

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In the above equation, f, the median pressure capacity, represents the intamal pressure level for which there is a 50% probability of failure, or equivalently, the best estimate of the pressure capacity. M is a lognormally distributed random variable having a unit median value and a logarithmic standard deviation, Su, which represents the uncertainty due to analytical modeling. S is also a lognormally distributed random variable having a unit median value and the logwiiiwi standard deviation, S,, representing the uncertainty hM with the material properties. The overall uncertainty in the pressure capacity is obtained by taking the square root of the sum of the squares of puandS,.

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-(E*u+E$)"*

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As a result, the probab0listic pressure capacity is described by three parameters: the median

_ prest,ure, the modeling uncertainty. u. and the material strength uncertainty, 3 I~

The logarithmic standard deviations, p u and p3, quantify the variability due to a 4

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ladt of knowledge resulting from differences between the lanalyticat model and the real

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situcture. Modeling uncertainties are introduced by assumptl,:ns used to develop analytical niodels and their abiMy to adaquately represent the failure cend tion. The strength uncertainties l

correspond to variabilities related to material resistance. Examples of the sources of such i

uncertainties include: variability in concrete strength, steel yield strength, stress-strain rela-i

. tionships, and the influence of temperature on material strength.

2.2 Varit.bility in Material Propert&s Test data on the strengths of the structural materials of the Crystal River nuclear j

l power station are presented in References 4 and 5. These data included the sample data

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points for the liner (Ref. 4), and the concrete (Ref. 5). Using these data, the sample statistics l

(mean and standard deviation) were evaluated for the characteristic material strengths. The j

f median values and logarithmic standard deviations were then evaluated, with the assumption that all of the material strengths could be characterized by a lognormal detribution. Appendix A contains a brief discussion on the features of the lognormal distribution and the relationship i

between the median and the logarithmic standard deviation and the sample statistics. For the tendons and the reinforcing, the information provided in Reference 1 was supplemented by l

generic data available in the literature (References 3,6, and 8).

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In References 4, and 5, the material strength test data corresponded to room i

temperature conditions. Since higher temperstures are considered in this irr. : :"ytion, it was necessary to estimate the influence of elevated temperatures on the Crystal River materials.

l This was acccing,;;.;,sd by estimating temperature dependent strength reduction factors from other available data on the same materials. Given a median charactedelic material strength at room temperature,3,w.7, the median material strength at a higher temperature was esti-mated by i

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A r $,u.r (2-4)

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i where A ris the median temperature strength reduction factor. Noting that there is uncertainty l

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in both the material strength at room temperature and in the reduction factor, the overall logarithmic standard deviation of the material strength at the elevated temperatures was j

f estimated as

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(S n*r + S$v.7)"

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in which 4,7 is the logarithmic standard deviation associated with the temperature strength 4

reduction factor and p,u.7 s the logarithmic standard deviation of the material strength at i

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room temperature.

The following subsections discuss the estimated median values and thelogarithmic standard deviations for the various structural materials.

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2.2.1 Prestressing Tendons L

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The typical prostressing tendon consists of 163 7 mm diameter low relaxation wires. These wires conform to ASTM A 421 65 Type BA (Reference 1) with a minimum ultimate tensile stress of 240 ksi. The characteristic minimum ultimate strength of the typical tendon is 2333.5 kips. No sample test data were avaitahia.

The median failure criterion for the tendons was estimated as 3% strain, which is 75% of the code minimum elongation reported in Reference 10 for A 421 low relaxation tendons.

The median strength of the tendons at 3% strain was estimated to be 2205.2 k at room tem-l j

perature, with a logarithmic standard devisdon of 0.03.

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Since the elevated temperature concmions are considered in this irc.z"i" r, the i

loss in strength at higher temperatures was estimated using data contained in Reference 2.

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2.2.2 Reinforcing Bars Low tensile strength, deformed reinforcing bars were specified for the Crystal River reactor building. The mild steel reinforcing used was ASTM A 61548 Grade 40 bars with minimum yleid point of 40 ksi and minimum tensile strength of 70 ksi(Reference 1). Based on test data for Grade 40 bars (Reference 11), the median yield strength is estimated as follows:

J, - 1.2f y,,,, - 1.2(40) - 48 ksi Hence, the uncertainty associated with the median value are given by:

3 p, = 1.6S '"(

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l In other words, the code minimum values are considered to represent 95% lower bound values.

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in order to estimate the reduction in strength and the added variability with elevated temperatures, the data in Reference 2 was used.

i 2.2.3 Uner Plate The reactor building wallis lined with a 3/8 in thick steel plate in the cylinder and j

dome portion which becomes 1/4" on the base mat floor. The transition knuckle plate between 1.

the wall and the base mat is 3/4" thick. The lir.or material conformed to the ASTM A 283 Grade I

C steel. Test data on the yield and tensile strengths of the liner plate was provided by Chicago l

Bridge & Iron Company (Referece 4). Based on these test results, a median yield stress of 44.6 kai and an anewinteri logarithmic standarc' deviation of 0.08 were estimated for the liner in the room temperature.

The appropriate reductions in the yield stress for elevated temperatures were estimated from the values given in the ASME Boiler and Pressure Vessel Code (Reference 3).

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2.2.4 Concrets 3

The concrete specified for the reactor building shell and basemat has a minimum design 28 days compressive strength of 5000 psi (Reference 1). Typically, the actualin-place concrete will have a compressive strength substantially higher than the specified strength.

f There are two primary reasons for this. First, the contractor mixing the concrete will try to produce a concrete batch such that the average compressive strength is somewhat greater than the specified strength. Second, as concrete ages, it gains strength, i

Test data were available on the in-place compressive strength of the dome concrete (Reference 5). Using this data, mean and coefficient of variation of the concrete compressive strength were calculated to be 6301 psi and 0.08, reepecCd. Based on a lognormal dis-

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tribution, these in turn result in the median and logarithmic standard deviation values of 6280 psi and 0.08, respectively. Note that these values include the aging effects on the concrete strength, since the data were based on core samples frota the dome.

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The effect of temperature on the concrete compressive strength was based on the results reported in Reference 7. According to these results, concrete retains at least 80% of its original compressivo strength at temperatures up to 8007. Therefore, the median concrete 4

I compressive strength at higher than room temperatures was taken as the 80% of the median l

value at room temperature or 5020 psi. In addition, a logarithmic standard deviation value of l

0.10 was used to represent the additional uncertainties introduced due to the elevated tem-l peratures resulting in total uncertainty of 0.13 as follows:

i (0.08* + 0.10 )"*

2 0.13

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f it is assumed that the above median compressive strength and uncertainty are representativ6 of the concrete in the reactor building wall, dome, and basemat.

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2.3 Modeling Variability i

Uncertainties will exist in the estimated pressure capacities due to differences

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between the analyticalidealization of the structure and the real conditions. There are numerous possible sources of modeling uncertainties. Examples of the possible sources include:

assumptions used to develop internal force distributions, failure criteria, and use of empirical j

formulae. Moreover, since they are dependent on the particular failure mode under consid-j eration, they must be evaluated on a case-by-case basis. However, in many instances, the evaluation of these uncertainties would require very detailed analysis and/or extensive data 1

which may not be available. As a result, it was necessary to'use subjective evaluation and i

engineering judgment to estimate these uncertainties. The evaluation of the modeling i

uncertainties associated with the median pressure capacities is included in the discussion of l

the specific failure modes in the subsequent sections.

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3. ANALYSIS OF REACTOR BUILDING CAPACITY i

h The capacity of the Crystal River reactor building is estimated based on incipient i

leakage as the failure criterion. In this section, several potential failure modes are investigated.

The controlling failure modes are ranked according to their respective median pressure capacities, in which the loads considered include temperature, pressure, and dead load. The l

effects of temperature are treated by including the reduction in material strengths with elevated j

temperatures. The failure modes examined include:

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1. Membrane failures of the containment shell 4

2. Failure at the containment wall - basemat junction

3. Failure of the basemat All of the failure modes evaluated here were considered to be quasi-static. In other words, the pressure rise times were assumed to be of sufficient duration such that the dynamic response characteristics of the containment shen could be neglected arxi the temperatures 1

in the material were assumed to have reached steady state. The failure modes evaluated in this section are associated with gross structural failure leading to rapid depressurization of the containment.

j The median pressure capacities and the associated variabilities were evaluated l

for three temperature conditions; temperatures at the inside face of the wall of 300',500', and j

800* F. The corresponding temperatures at the outside face of the concrete waB were assumed j

to be 70",100', and 200* F, respectively. The interior temperatursa were asiected to cover a i

range from the design basis temperature up to the temperatures postulated for severe acci-

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. dents in typical PWR containments. Note that the waH temperatures are actual material tem-j; peratures rather than gas temperatures. In the fogowing reeW the temperature cases j!

will be denoted as cases 1,2, and 3 correspondag to the interior temperatures of 300',50(r, and 800" F, respectively. Since the maximum interior temperature under the design basis acc dent was 281* F (Reference 1), the temperatures considered in this invW% are well r

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beyond the design basis conditions.

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i MV-6570-001, Rev. A 3.1 Membrane Failures of the Reactor Building Shell in order to estimate the membrane capacities of the containment shell, several issues must first be addressed, which include: the median failure criteria for the structural elements, the initialin-service state of stress and strain in the tendons, the stress strain behavior of the reinforcing and liner, the temperature distribution through the concrete wall, and the thermal strains in the liner. Since the failure modes are associated with membrane tension, only the liner, the tendons, and the bonded reinforcing are assumed to provide resistance to the internal pressure. No credit was taken for the tensile strength of the concrete.' Although r

design codes such as Division 2 of the ASME Boiler and Pressure Vessel Code do not permit -

inclusion of the liner strength in the containment design, it was included in the evaluation of the actual ultimate pressure capacities considered here.

The membrane strength of the reactor building shell is provided primarily by the pre-stressing tendons. The median failure strain for the tendons was taken to be 3%. By comparison, the expected uniaxial elongation capacity of the liner and the bonded reinforcing

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are both substantially larger than that for the tendons. Therefora, the median strain limit for

.the tendons was selected as the critical failure criterion. Noting that the median failure strain for the tendons was taken as 3%, the strains in the liner and the bonded reinforcing are expected to be between their respective yield and ultimate strains. However, to evaluate the contnbution of the liner and the reinforcing to the overall membrane strength, the effect of strain hardening was neglected, in other words, at failure, both the liner and the reinforcing were assumed to be at their respective yield stresses. This greatly simplifies the calculations while introducing l

a relatively small amount of conservatism in the results.

i The primary effect of the elevated temperatures was a reduction of the material strengths. Given the assumed temperatures at the inside and outside faces of the concrete I

wall for the different temperature cases, the Ener, the tendons, and the bonded reinforcing will be at different temperatures. To estimate the temperatures of the different materials, a linear temperature gradient was assumed through the concrete was and each material was treated as a distinct layer. ~ In this way, the temperature of each material could be estimated by linear interpolation and an appropriate temperature dependent material strength reduction factor could be applied to the characteristic material stress.

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l Because the interior temperatures considered in this investigation are sigruficantly

. higher than the exterior temperatures, the liner temperatures are much higher than the average temperature of the concrete wall. As a result, the thermal expansion of the liner is restrained by the concrete and compressive strains are induced in the liner due to temperature effects l

alone. There are also additional liner compressive strains due to prestressing and concrete creep and shrinkage. This tends to counteract the strains induced in the liner by the internal l

pressure, since the pressure will produce membrane tension in the liner. However, these

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compressive strains are small in comparison with the final liner strains at tendon rupture.

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3.1.1 Hoop Membrane Failure in the Cylinder Wall 1

The critical section of the cylindrical portion of the reactor building wall was taken to be approximately at midheight of the cylinder. This section is the least influenced by the end effects caused by the basemat and the dome. At the midheight of the wall, the bonded reinforcement in both the hoop and meridional directions is provided by #8 bars at 12 inches, at both the inside and outside faces.

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h The hoop tendons have a typical spacing of 1'-1" and span approximately 120" l

circumferential!y between the anchorage points at the buttresses. The hoop tendons are

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staggered such that, at any vertical cross section, two tendons act over a tributary length of 3'-3".

At tendon failure, the reinforcing bar forces were taken as the yield strength of the bars, since they were in uniaxial (hoop) strain. At the tendon failure condition, the liner is in biaxial tension and the hoop stress in the liner was estimated from the von Mises stress ( an equivalent uniaxial stress ), which is given by o,-

[(o, - o.)' + (o, - o,)* + (o, - o,)']"'

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4 MV-6570-001, Rev. A Making the assumption of a stress distribution equal to that for an ideal cylinder

subjected to internal pressure, the relationship between the liner hoop stress and the meridional stress was taken to be a, - h c. and o, = Q Substituting these values into Equation (3-1) led to a relationship between the liner hoop stress and the von Mises stress as o,

1.1550,

=

The von Mises stress, o,, was estimated as the uniaxial liner yield stress. In turn, the liner hoop stress was estimated.

l The hoop membrane capacity of the containment wall was then estimated from

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the hoop tensile resistance provided by the tendons, the bonded reinforcing, and the liner. It

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was found that the hoop tendons provide approximately 80% of the total hoop resistance. The median pressure capacities for the three temperatJre cases are simwn in Tables 3-1,3-2, and i

3-3.

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The modeling variabilities included uncertainty in the tendon failure strain, and i

uncertainty in the temperature and internal force distribution. A logarithmic standard deviation l

of 0.06 was used to represent the modeling uncertainty in the failure criterion. This was based on the tendons constituting about 80% of the total hoop resistance and that tendons typically i

exhibit a relatively small amount of strain hardening behavior as well as relatively small scatter in their ultimate strength capacities. The uncertainty in the temperature and the internal force distribution was estimated to have a p value of 0.08. Therefore, the overall modeling uncertainty for the hoop membrane fai!ure mode was estimated as (0.06 * + 0.08 *)"*

Su

=

i 0.10

=

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The material strength variability has contributions from each of the ccmporwas providing hoop resistance. The reinforcing and the liner have larger uncertainties than the tendons. Since the tendons resisted about 80% of the hoop load, it would be overly con-

. servative to directly ccmb;rie the individual variabilities of the malarials by the square root of 3-4 i[

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the sum of the squares (SRSS) method. Therefore, the overall material strength uncertainty j

was estimated by weighting the individual p values ( including the uncertainty due to tem-perature ) by the relative contributions of the respective materials to the total hoop resistance.

l Using this approach, the net variability in the hoop pressure capacity due to material strength was approximately equal to the uncertainty in the tendon strength. Thus, for each of the temperature cases, the logarithmic standard deviation associated with material strength, p,,

was estimated to be 0.06 as shown in Tables 3-1,3-2, and 3-3.

Also shown in Tables 3-1 through 3-3 are pressure levels corresponding to a high confidence of low probability of failure (HCLPF). These pressure levels correspond to values of ultimate pressure which would be expec%o from a conservative deterministic analysis of the individual failure modes. These HCLPF values were determined based on the rather arbitrary assumption of a 95% confidence level. That is, based on the estimated median pressure capacity and the associated logarithmic standard deviations for a given failure mode, there is an estimated 95% confidence that the random pressure capacity is greater than the HCLPF value, or equivalently, an estimated 5% con 6dence that the random pressure capacity is less than or equal to the HCLPF value.

g 3.1.2 Meridional Membrane Failure in the Cylinder Wall l

As with the hoop failure evaluation, the critical section of the wall was taken to be j

approximately at midheight of the cylinder ( midheight between the springRne and the top of the basemat ). The meridional bonded reinforcing was provided by #8 bars at a 12" spacing l

at both the inside and outside faces of the war. The meridional tendons in the cylinder run vertically and are anchored at the tendon gaRory below the basemat and at the ring girder located at the springHne. Through any horizontal cross section in the cylinder, there are a total of 144 vertical tendons, which are set on a circle having a radius of 67'-3 3/8" with a circum-ferential spacing of 2.93 ft.

1 Failure was considered to havs occurred at a meridional tendon strain of 0.03. At this strain level some strain hardening could be expected in both the Ener and the reinforcing.

However, in evaluatmg the meridional pressure capacity, it was assumed that the Ener and the reinforcing were at their respectNo yield stresses ( Le., strain hardening was neglected ). This 4

3-5 e

i e

MV-6570-001, Rev. A I

greatly simplified the calculation while introducing a slight conservative bias. The conservatism l

l introduced is small since the liner and the reinforcing provide less than 20% of the total meridional resistance and the contribution of strain hardening constitutes a small fraction of j

that percentage.

i l

l Similar to the method used to estimate the liner hoop stress, the meridional stress i

i in the liner at the meridional tendon failure was estimated from the von Mises stress. The same i

i relationship between the liner membrane stresses was assumed (i.e., o, - j a, and a, = 0).

4 i

.With the conservative assumption that the liner.is at yield at failure, the von Mises stress, o,,

is equal to the median liner yield stress at the liner temperature. Using Equation (31), the liner j

meridional stress was estimated as I

i o,

0.577 Sy

=

j where d y is the median liner yield stress.

I l

b The median pressure capacities for meridional membrane failure for the three temperature cases are shown in Tables 3-1,3-2, and 3-3. The median capacities are sub-j stantially greater than the hoop capacities. As with the hoop capacity evaluation, logarithmic

{

standard deviation of 0.10 was used to account for the modeling uncertainty due to tendon failure criterion and the temperature and internal force distnbutions. The material strength

{

variability was estimated by weighting the individual material variabilities by the relative con-tributions to the total meridional resistance. The modeling and material strength uncertainties are shown with the median capacities and the associated HCLPF values in Tables 31,3-2, and 3-3.

j 3.1.3 Dome Membrane Failure The torispherical dome has a reduced waI thickness of three feet. Based on the results reported in Reference 5, the dome capacity is controRed by the meridional c.fiicity of the dome at the junction of the two spherical segments of the dome. Based on information in Reference 1, the dome reinforcing was assumed to have a reinforcement ratio of 0.15% in 36 l-y

+

-+

. m

MV-6570-001, Rev. A the meridional direction at the inside face of the wall. De meridional dome reinforcing at the outside face of the wall was provided by #9 bars at 15' plus #11 bars at 12". As with the membrane capacity evaluation for the cylinder, tendon strain level of 3% was regarded to constitute failure of the dome.

As with the evaluation of the meridional capacity of the cylinder, at the tendon failure strain limit, the liner and the reinforcing were assumed to be at their respective yield stress and strain hardening effects were neglected. Again, this simplified the calculation while introducing a slight conservative bias. To estimate the membrane stress in the liner, the von Mises stress was used. For an ideal spherical shell subjected to intemal pressure, the rela-tionship between the principal stresses is given by c = a, and a, = 0. Substituting this into Equation (3-1) gives the following relation between the liner membrane stress and the von Mises stress a' y 0,

=

(

in which the von Mises stress is equal to the median yield stress at the liner temperature, since strain hardening effects were neglected.

The median pressure capaedies, variabilities, and HCLPF values are shown in Tables 3-1 through 3-3. Note that the dome capacities fall between the hoop capacity and the meridional capacities of the cylinder. The modeling and material strength uncertainties were evaluated by the same methods as for the cylinder membrane capacities. The modeling i

uncertainty included the urMuidrity associated with the tendon failure criterion as well as the uncertainty asrxisted with the temperature and internal force distributions. The material strength uncertainty was estimated by weighting the values for the individual materials by their relative contributions to the overaR membrane capacity.

1-3.2-Flexural and Sheer FaNro of the Wau-Basemat Junction Both flexural as well as shear failure of the wall-basemat junction were evaluated.

l The flexural cam was irwestigated by d @,g a failure envelope for combined merid-l

]'

37 i

i 4'

I

i-t

- MV-6570-001, Rev. A i

l ional membrane tension and meridional sheu moment as a flexure-tension interaction diagram.

The radial shear capacity was estimated using ultimate strength principles (Reference 10) in j

j which the shear is resisted by the concrete and reinforcing steel.

I l

The flexure-tension interaction was described by plotting the meridional moment versus the meridional membrane force. The failure envelope for the wall-basemat junction was f

represented by a straight line joining the two points (M,, 0) and (0, N ) on the interaction -

diagram; where M.and N are the pure bending and pure membrane ultimate capacities at l

the junction. In calculating the pure membrane ultimate load capacity, N., the liner and the l

bonded reinforcing were taken to be at their yield stress. The tendon force was taken at 3%

strain. In calculating the moment capacity, the ACI ultimate flexural strength principles were used. Due to the anchorage of the liner and the capability for strain compatibility, the liner was j

included along with the meridional bonded reinforcing in providing the flexural tensile force.

However, due to their placement in the wau cross section, the meridional tendons were not l

considered to contribute to the flexural capacity. As the internal pressure increases from an initial state corresponding to the Dead Load (D.L) + Prestress loading, the moment and

-{'

tension at the junction were assumed to increase proportionally until reaching the failure envelope. The slope of this loading line, i.e., the change in meridional moment relative to the 4

I change in membrane tension due to the pressure loading, was estimated from the junction l

moment and meridional tension values reported in Reference 1 due to i) D.L + Prestress and ii) D.L + Prestress + 82.5 psi.

l After the meridional moment and membrane tension at the failure condition are l

calculated, the median pressure capacity is estimated as the increase in pressure required to i

raise the moment and membrane force from their initial values under D.L + Prestress loading to the values at failure.

It was found that the failure cor'dition was predominantly a flexure fature. As a

[

result, the uncertainties calculated for M.were used to estimate the uncertamties meerv4m**d 4

with the median pressure capacities. For the material strength uncertainty, S,, the uncertainty

. values corresponding to the yield stress of the bonded in-Juicir.g were used for both M.and A For the modeling uncertainty, a S value of 0.11 was evaluated for the uncertainty in the 3

(

i i

,-.,,,-,n

,a

MV-6570-001, Rev. A flexural capacity equation. Additional modeling uncertainty values were estimated to account for the temperature and force distribution effects, resulting in a total modeling uncertainty of 0.16 for all three temperature cases.

The radial shear capacity at the containment wall-basemat junction is provided by the concrete and the #7 wall stirrups. In addition, some radial shear force is also resisted by the horizontal component of the inner layers of the meridional reinforcing at the knuckle region at the base of the wall ( which are oriented at about 15* from the vertical ). The radial shear capacity was taken as the sum of the contributions of the concrete and reinforcing steel. Since the effects of the pressure loading can result in net meridional membrane tension at the base of the wall, it was necessary to account for the reduction in the shear capacity of the concrete due to the presence of the tension acting on the cross section. This effect was included by using the provisions in Reference 10 which address the shear capacity of concrete with axial tension. As a result, the radial shear capacity was dependent on the meridional membrane force and, hence, the internal pressure. The radial shear capacity could then be expressed as a function of the yield stress of the reinforcing steel, the square root of the compressive

{

strength of the concrete, and the internal pressure. Using the results for the radial shear at the base of the wall as reported in Reference 1, an expression was developed for the radial shear demand force as a function of pressure. The median failure capacity was obtained by equating the radial shear demand force and the shear capacity, then soMng for the pressure.

' The variabilities associated with the median pressure capacity at which shear failure j

at the junction is predicted are evaluated as follows. Since the total shear capacity was taken l

as the sum of the shear capacities of the steel and the concrete, the material strength uncertainty for the pressure capacity has contributions from the variability in the yield stress of the rein-forcing and from the variability associated with M ( since the concrete shear strength is a function of M). This lead top, values of 0.12,0.12, and 0.14 for the three temperature cases, respectively. The modeling uncertainty of 0.19 for the three temperature cases includes contributions from the variabWty meanrMed with the shear capacity equation and the internal 4

force distribution. The median pressure capacities cini: g-Mng to the junction sheer failure, l

associated variabilities, together with the HCLPF values are listed in Tables 3-1, 3-2, and 3-3

- for the three temperature cases. For all three temperature cases, the pressure capacity of the shear failure mode is higher than that of the flexural failure mode. Therefore, the flexural failure mode is the i.Emiv ue g failure mode for the wal basemat junction.

U v

3-9 N

I f

MV-6570-001, Rev. A 1

3.3 Failure of the Basemat i

t The pressure capacity of the concrete basemat slab was determined for both j

flexural and shear failures. For both failure modes, median pressure capacities and the associated uncertainty values were determined for the three temperature cases. The tem-I~

perature inside the containment for the three temperature cases were the same as those used l

for the cylinder wall analysis. However, the temperature at the bottom of the 12.5 feet basemat was taken as 50" F for all three temperature cases. Again the temperature was assumed to i

. vary linearly through the thickness of the basemat.

Since the mat is 12.5 feet thick, the bottom reinforcing was considered to be at f

about room temperature for all the three temperature cases. This, plus the fact that the same i.

concrete compressive strength was used for all temperatures above room temperature ( see

{

Section 2.2.4 ), results in the basemat median ultimate moment capacity for radial bending to be the same for all three temperature cases. Hence, the basemat flexural pressure capacPJes, associated variabilities, and HCLPF values for all three temperature cases are identical.

(

The basemat median ultimate moment capacity at the center of the mat where the f

moment is maximum was estimated to be 5890 k ft/ft. The bending moment at the conter of the basemat was evaluated at various intemal pressure values. An iterative procedure was used to estimate the bending moment demand at the center of the basemat at various internal pressures. The basemat was assumed to rest on an elastic foundation with a subgrade modulus of 411 kips per cubic inch (Reference 9). The loading and deflections due to dead load, prestress, pressure as well as soil pressure were superimposed. An iterative approach was required because the soil reaction profile ( assumed to be parabolic ) was dependent on 4

the deflected shape of the mat. The iteration was based on assuming deflection at the center of 9)e becomat which determined the magnitude of the soil reaction proflie and the radius i

where upilft is predicted. The maximum radius where contact is maintained between the mat and the eat was estimated by equating the total soE reaction force with the total containment deed load. It was found that at an internal pressure of 142 pel, the moment at the center of ;

the basemat is 5776 k-ft/ft. Hence,142 psi was used as the median pressure capacity of the basematin flexural failure mode.

l 4

5 4

i I

3-10

!' }

MV-6570 001, Rev. A

~

Since the effect of the variation of the concrete compressive strength on the flexural capacity is small, the variability due to the material strength, S 3, was taken as the 3 value of the reinforcing at room temperature ( 3 = 0.11). The modeling uncertainty has contributions from the estimated variability of the predicted moment capacity of the basemat and the internal 2

force distribution. An overall modeling uncertainty, Du, of 0.23 was estimated. Tables 3-1, 3-2, and 3-3 show the median pressure capacities for the flexural failure mode of the basemat, the corresponding p values, and the HCLPF values for the three temperature cases.

For shear failure, the critical section of the basemat is located at about 12 feet from the junction with the containment wall. The shear capacity is provided by the cone:ete and the #11 bar shear reinforcing which are inclined at 45* with respect to the radial reintacing.

A relationship between the intemal pressure and the basemat shear at the entical sectior, was developed. Using this relationship, the median pressure capacity ( l.e., the pressure at wuch the shear demand equals the ultimate shear capacity of the basemat ) was estimated. The materie', strength and modeling uncertainties associated with the shear failure of the basemat wer9 estimated in a similar manner to that of the shear failure of the wall-basemat Junction.

Fct all three temperature cases, the median pressures for shear failure were higher than those

{'

for flexural failure. Therefore, the flexural failure mode was judged to be the critical failure mode for the basemat. The median pressure capacities, the corresponding S values, and the HCLPF values associated with the shear failure of the basemat for the three temperature cases are shown in Tables 3-1, 3-2, and 3-3 for comparison.

I 3-11 I

i

i MV 6570-001, Rev. A Table 3-1 Pressure Capacities for the Reactor Building Structural Failure Modes -

Temperature Case 1 i

i interie Tenperature = 300* F.

Failure Mode P

pu Ds s

HCLPF (psig)

(psig)

/

Wall Basemat Junction Flexure 140 0.16 0.13 0.21 99 3

Basemat Flexure 142 0.23 0.11 0.25 94 Wall-Basemat Junction Shear 147 0.19 0.12 0.22 102

)

b k er Basemat Shear 152 0.19 0.07 0.20 109 f

Cylinder Hoop Membrane 172 0.10 0.06 0.12 141 4

Dome Meridional Membrane 199 0.10 0.06 0.12 163 p

Cylinder Meridional Membrane 216 0.10 0.06 0.12 177 4

i i

e i

4

^

3-12

l 1

MV-6570-001, Rev. A Table 3 2 Pressure Capacities for the Reactor Building Structural Failure Modes -

Temperature Case 2 Interior Temperature = 500* F.

Failure Mode Du Ds p

HCLPF (psig)

(psig)

Wall-Basemat Junction Flexure 134 0.16 0.14 0.21 95 Basemat Flexure 142 0.23 0.12 0.25 94 Wall-Basemat Junction Shear 146 0.19 0.12 0.22 102

(

Basemat Shear 151 0.19 0.08 0.21 107 j

Cylinder Hoop Membrane 167 0.10 0.06 0.12 137 Dome Meridional Membrane 194 0.10 0.06 0.12 159 i

Cylinder Meridional Membrane 211 0.10 0.06 0.12 173 6

3-13 i

MV-6570-001, Rev. A Table 3-3 Pressure Capacities for the Reactor Building Structural Failure Modes -

Temperature Case 3 Interior Temperature = 800* F Failure Mode N

pu Ds HCLPF (Psig)

(psig)

Wall-Basemat Junction Flexure 122 0.16 0.16 0.23 83 Wall-Basemat Junction Shear 139 0.19 0.14 0.24 94 Basemat Flexure 142 0.23 0.11 0.25 94

(

Basemat Shear 150 0.19 0.09 0.21 106 Cylinder Hoop Membrane 157 0.10 0.06 0.12 129 Dome Meridional Membrane 166 0.10 0.08 0.13 134 Cylinder Meridional Membrane 199 0.10 0.06 0.12 163 l

.t I

3-14 o

MV-6570-001, Rev. A

4. REACTOR CAVITY ACCESS TUNNEL DOORS Two doors which provide reactor cavity access were evaluated for expected pressure capacity. Doors DF-2 and DF-3 are located at top of concrete elevations 95'-0* and 94'-0", respectively. The doors are located in series (i.e., the failure of both doors is required prior to leakage).

Both doors are similar in size and construction. The doors are constructed of 1.25' A36 steel plate and are pressure loaded against heavy angle section frames. One quarter inch thick EPDM E603 gaskets provide the seats between the doors and frame. Door hinges are slotted to provide allowance for additional gasket compression resulting from cavity pressure. The outside dimensions of the doors are 2'-10' by 2'-4' for DF-2 and 2'-4.5" by 2'-2.5" for DF-3.

C Pressure capacities for both doors are significantly greater than those determined for the various structural failure modes discussed in the previous section. Median pressure capacities for both doors at several temperatures are shown in Table 4-1. At high temperatures, some deterioration of the EPDM gasket will occur. However, the gasket is expected to continue to function provided it is continuously loaded by the internal pressure. Concrete deformations in the regions around the doors are not expected to result in uncovering the gasket. Thus, the door pressure capacity is expected to be contro5ed by deformation of the doors them-selves.

4 4-1 a

MV-6570-001, Rev. A Table 4-1 DF-2 and DF-3 Door Pressure Capacities Temperature

. DF-2 DF-3 (F)

Median p

HCLPF Median p

HCLPF Pressure (psig)

Pressure (psig)

(psig)

(psig) j 70 296 0.14 235 361 0.14 287 300 262 0.14 206 320 0.14 254

(

500 239 0.15 187 292 0.15 228 800 197 0.15 154 240 0.15 187 l

5 I

1 i

l t

4 j

MV-6570-001, Rev. A I

i

5. CORRELATION OF FAILURE MODES i

i l

For the purpose of estimating the correlation between structural failure modes, the uncertainty was subdivided into two independent components:

1. Uncertainty in modeling j

2. Uncertainty in strength These uncertainties may be represented by two independent random factors with j

logarithmic standard deviations. S uand,, respectively. The combined coefficient of variation

)

is then given by:

i p-#bsi i

(

The advantage of splitting the uncertainty into these two components is that for a given pair of failure modes the uncertainty factor for one of the components may be correlated for both modes, while the other is independent. However, for Crystal River, it was found in all cases that if the uncertainty in modeling for a given failure mode is expected to be correlated j

with the uncertainty in modeling for a different mode, the uncertainty in the strengths of the j

same two modes were also expected to be correlated. For example, the strengths of the various shell membrane capacities are then gov 6med largely by the tendon capacities for l

which common strength values were used and hence a high degree of correlation in the uncertainty of the strength parameters is expected. Similarly, for these same failure modes, j

similar modeling assumptions were used including failure strain criteria, simRar load redistri-l bution due to tendon friction, etc. Ukewise, while the strength and modeling assumptions used to evaluate the DF-2 and DF-3 doors are correlated, these assumptions are completely different from those used for the membrane @tes. Similar considerations for every pair e

of failure modes lead to the correlation matrix shown in Table 5 2. In this table, the uncertainty factors are assumed to be alther perfectly correlated, in which case, a cross is placed in the l

[

appropriate box, or perfectly uncorrelated. Perfect correlation is assumed whenever the 4

degree of correlation is estimated to be more than one-half.

r t

i 51

.y

!I J

-.--n

l l

MV-6570-001, Rev. A Table 5-1 Failure Mode Failure Mode Number Failure Mode 1

Wall-Basemat Junction Flexure 2

Wall-Basemat Junction Shear 3

Basemat Flexure 4

Basemat Shear 5

Cylinder Hoop Membrane 6

Dome Meridional Membrane 7

Cylinder Meridional Membrane 8

Door DF-2 9

Door DF-3 2

a l

C 5-2

1 1

MV-6570 001. Rev. A i

e 4

I Table 5-2

}

Correlation between Failure Modes 1

(Strength and Modeling Uncertainty) 4 i

i 1

1 i

1 2

3 4

5 6

7 8

9 1

l 1

X X

{

2 X

3 X

X 4

X 5

X X

X 6

Symmetric X

X

{

7 X

l 8

X X

9 X

ii i

l i

l 1

4 e

4 4

i

_.~_ __

1 1

4 MV 6570-001, Rev. A i

i J

i i

REFERENCES I

\\

4 j

1.

Final Safety Analysis Report (FSAR) for the Crystal River Unit 3 Nuclear Power Plant, i

2.

Holmes, M., Anchor, R.D., Cook, G.M.E., and Crook, R.N., "The Effects of Elevated Temperatures on the Strength Properties of Reinforcing and Prestressing Steels,"Ihg Structural Engineer. Volume 608, No.1, March,1982.

3.

"ASME Boiler and Pressure Vessel Code, Rules for Construction of Nuclear Power Plant i

Components,"Sectionill, Appendices,1989.

j 4.

" Test Results on the Crystal River Unit 3 Uner Material," Chicago Bridge & Iron Company,

. July 28,1969.

f 5.

" Crystal River Unit 3, Reactor Building Dome Delamination Report," Gilbert A=*rv4ataa, Inc., Report No.1913 December 1976.

?

b 6.

Park, R., and Paulay, T.. Reinforced Concrete Structures. John Wiley,1975.

I 7.

Abrams, M. S.," Compression Strength of Concrete at Temperatures to 1600 *F,"Imm-perature and Concrete. American Concrete Institute, SP-25,1971.

3 8.

" Steel Reinforcement - Physical Properties and U.S. Availability,' ACI Materials Journal.

January-February 1989.

l 9.

" Crystal River Unit No. 3 Reactor Building Shell Cale's, Book 11:01.1 to 1:01.3,' December 5,1973.

j

10. ' Building Code Requirements for Reinforced Concrete (ACl 318-89)," American Concrete i

institute, November 1989.

11.

Mirza, S.A. and MacGregor, J.G., "Vartsbility of Mechanical Properties of Reinforcing Bars," Jaumal of Structural Division. ASCE, Vol.105, No. ST5, May,1979.

3, m_

y i-

12. " Median Maledel Properties and Variabstles," ABB impell Corporation, Calculaden CA-6570-001-001, Rev. 0.

p.

. j"

'(

L;.

R-1

,l-

MV4570 001, Rev. A i

13.

" Containment Structure Membrane Capacities," ABB Impell Corporation, Calculation CA4570001-002, Rev. O.

14.

"Basemat and Wall - Basemat Junction Capacities," ABB impell Corporation, Calculation CA4570-001-003, Rev. O.

15.

" Reactor Cavity Access Tunnel Doors DF-2 & DF-3,' ABB Impell Corporation, Calculation CA4570-001-004, Rev. O.

(

I a

1 i

R-2

1 i

s 0106-015-R001, Rev. 0 I

s APPENDIX A i

CHARACTERISTICS OF THE LOGNORMAL DISTRIBUTION J

l i

h Some of the characteristics of the lognormal distribubon which are useful to keep

)

in mind when generating estimates of P, p u, and p, are summarized in References A-1 and l

A-2. A random variable X is said to be lognormally distributed if its natural logarithm Y, given by:

Y=

In(X)

(A-1) 1 i

is normally distributed with the mean of Y equal to in(A), where Xis the median of X, and with l

the standard deviation of Y equal to p, which will be defined herein as the logarithmic standard deviaNon of X. The coefficient of variation of X, COV,is given by the relationship:

(

1 1

lexp(p*)- 1

( A-2) l COV l

For p values less than about 0.5, this equation becomes approximately i

p

( A-3)

COV

=

and COV and p are used interchangeably.

l l

For a lognormal distribution, the median value is used as the characteristic parameter of central tendency (50 percent of the values are above the median value and 50 j

percent are below the median value). The logarithmic standard deviation, S, or the coefficient of variation, COV, is used as a measure of the c5spersion of the distribution.

l The relationship between the median value, f, logarithmic standard deviation, p, and any value x of the random variable can be expressed as 1

3 A-1

i 0106-015-R001, Rev. 0

)

/'

x f.e 8

( A -4) a where q is the standardized Gaussian random variable (with'zero mean and unit standard deviation). Therefore, the frequency that X is less than any value x' equals the frequency that j

qis less than n'where In(x*/f)

(g_g) q.

E o

1 Because q is a standardized Gaussian random variable, one can simply enter q

standardized Gaussian tables to find the frequency that a is less than n' which equals the probability that X is less than x'. Using the cumulative distribution tables for the standardized Gaussian random variable, it van be shown that the f e' value of a lognormal distribution corresponds to the 84th percentile value (i.e.,84 percent of the data fall below the +S value).

The f e'8 corresponds to the value for which 16 percent of the data fall below.

C One implication of the usage of the lognormal distribution is that if A, B, and C are independent lognormally distributed random variables, and if 1

A' B' q

( A-6)

L D

=

c,

! ~

where q, r, s, and t are given constants, then D is also a lognormally distributed random.

I variable. Further, the median value of D, denoted by b, and the logarithmic variance, S i, which is the square of the logiudiinic standard deviation, S o, of D, are given by O " Ar.ps 4

( A -7) g, i

.and 8 2 a

S$ = r S.,rSj+t gs (A-8) a p.

A-2

,i

0106-015-R001, Rev. 0 I

where A, B, and C are the median values, and pa, p,, and pc are the logarithmic standard deviations of A, B, and C, respectively.

References A-1 Benjamin, J. R. and Comell, C. A., Probability. Statistics and Decision for Civil Enal-i neera, McGraw-Hill, Inc.,1970.

1 l

A-2 Kennedy, R. P. and Chelapati, C. V., ' Conditional Probability of A Local Flexural Wall Failure of A Reactor Building As A Result of Aircraft impact,' Holmes and Narver, Inc.,

prepared for General Electric Company, San Jose, California, June,1970.

i A-3 4

1 CR-3 IPE Level 2 Appendices 4

s i

8 i

I 4

o i

J u

t f

APPENDIX A

(

CR3 K3BA-S1 LBLOCA all debris in Cell 1 l

i 491

Table A.1-1:. Fission Products Distribution I

?

I

)

1 492

$1MPLE -F35Sl0N PRODUCT MASSE $ (KG) la CELL 1 AT TIME o 40000.000 (5)

MOST TYPE NAME 1 CSI 2 CSON 3 TE o sa 5 BA 6 LA

.2 AE8050L. U02 2.80260E-(5 2.80260E 45 3.86011E 041.56831E 07 2.92680C 06 7.84650E 07

'O TOTAL-WALL-4.17648E 051.74064E 04 4.42677E 03 2.31355E*02 2.07518E-021.30216E 02 0 LOW CELL

• ouMMY 1.15317E 01 6.91905E 01 2.64370E+00 1.31118E+01 1.20852E+01 ' 5.96872E+00 TOTAL 1.15359E 01 6.92079E 01 2.64851E +00 1.31350E+01 1.21060E+01 5.98174E+00 M0$f TYPE NAME 7 CE 8 RU.

9 ' P!

2 AERO$0L Uo2 7.45403E 07 2.35775E 119.89590E 03 0 TOTAL WALL 1.43409E 021.41835E 07 5.27013E 01 0 LOW CELL

• DumY 7.16724E+00 6.93879E 05 4.57085E+02 TOTAL 7.18158E+00 6.95298E 05 4.57622E+02 MOST MATEa!AL INF0eMATION IN CELL 1

AT TIME = 40000.000 (s)

HOST TYPE NAME POWER (WATTS) TEMPERATURE (K) 2 AEROSOL 002 0.00000E+00 3.66800E+02 0 TOTAL WALL 0.00000E+00 0.00000E+00 0 LOW CELL + DUMMY. 0.00000E+00 0.00000E+00 TOTAL 0.00000E+00 3

f i

t_.

i e

$!NPLE F15510N PRODUCT Masses (KG) 1*J CELL 2 AT TIME a 40000.000 (S) 15T TYPE NAME 1 C$1 2 C50H 3 TE 6 SR 5 8A 6 LA 2 AER050L U02 1.40130E 441.40130E 44 2.38035E 03 9.71481E 071.81821E 05 4.83019E 06 0 ' TOTAL WALL 5.50807E*04 3.17038E 03 1.06238E 02 3.21661E 02 2.54512E 02 3.52930E 02 0 TOTAL-Floot 1.30601E 03 7.72442E 03 5.41163E 02 3.71002E 013.05151E 01 1.65316E 01 TOTAL 1.85682E 031.08948E 02 6.71205E 02 4.03169E-013.30620E 012.00613E 01 MOST TYPE NAME 7 CE 8 RU 9 PI 2 AER050L UO2 4.58860E 061.45860E 10 6.10023E 02 0 TOTAL WALL 3.17877E 02 3.73500E 07 6.06071E 01 0 TOTAL FLOOR 2.49365E 01 2.26159E 06 6.55095E+00 TOTAL 2.81157E 012.63523E 06 T.21803E+00 Most MATERIAL INFORMATION IN CELL 2 AT TIME = '40000.000 ($)

HOST TYPE NAME POWER (WATTS) TEMPERATURE (K) 2 AEROSOL 002 0.00000E+00 3.66628E+02 0 TOTALL WALL

'O.00000E+00 0.00000E+00 0 TOTAL FLOOR 0.00000E+00 0.00000E+00 TOTAL 0.00000E+00 l

7

5$

- SIMPLE F15510N PRODUCT MASSES (KG) IN CELL 4 AT TIME o 40000.000 (S)

=1ST TYPE NAME 1 Csl 2 C50N 3 TE 4 SR 5 BA 6 LA

'2 AEROSOL UO2 0.00000E+00 0.00000E+001.89093E 04 7.72344E 081.44623E 06 3.83591E-07 0 TOTAL WALL 1.48423E 01 8.88012E-01 1.01642E 01 2.05361E 03 7.91680E 03 3.62106E 04 0 LOW CELL

• DUMMY 1.91492E+01 1.15015E+02 8.27173E+001.17052E+00 2.21046E+00 3.30278E 01 TOTAL ~

1.92976E+01 1.15903E+02 8.37356E+001.17257E+00 2.21838E+00 3.30641E 01 MO$f TYPE NAME 7 CE 8 RU 9 Pt 2 AEnos0L -uo2 3.64405E 07 1.15940E 11 4.84566E 03 0 TOTAL WALL 4.71966E-04 4.13538E 09 8.17745E-01 0 LOW-CELL + DUMMY 4.05143E 01 3.60052E 06 1.51388E+02 TOTAL 4.05616E 013.60467E 061.82211E+02 MOST MATERIAL INFORMATION IN CELL 4 AT TIME a 40000.000 ($)

NOST TYPE NAME POWER (WATTS) TEMPERATURE (K) 2 AERos0L 002 0.00000E+00 3.33580E+02 0 TOTAL WALL.

0.00000E+00 0.00000E+00 0 LOW CELL + DUMMY 0.00000E+00 0.00000E+00 TOTAL 0.00000E+00 i

I

.$1MPLE FIS$10N PRODUCT MASSES (KG) IN CELL

.5 AT TIME a 40000.000 ($)

M T TYPE NAME:

1 - CSI 2' C$0H 3 TE 4 SA 5 BA 6 LA 2 AERo$0L' U02 1.40130E 451.40130E 451.40302E 05 5.74504E 091.07750E 07 2.84347E 08 0 TOTAL-WALL 4.69204E 02 2.80631E-013.20581E 02 7.45573E 04 2.58473E 031.72542E 04 TOTAL 4.69204E 02 2.80631E 013.20721E 02 7.45579E 04 2.58484E 031.72571E-04 No$T TYPE.

NAME-7 CE 8 RU 9 PI 2 'AER0$0L' UO2 2.70123E 08 8.61982E 13 3.59463E-04 0 -TOTAL WALL 2.07058E 041.79444E*09 2.54044E-01 TOTAL.

2.07085E 04 1.79530E 09 2.54404E 01 NOST MATERIAL INFORMATION IN CELL 5 AT TIME = 40000.000 (S)

N0$f TTPE-NAME POWER (WATTS) TEMPERATURE (K) 2 Atmos 0L U02 0.00000E+00 3.32998E+02

-0 TOTAL WALL 0.00000E+00 0.00000E+00 TOTAL 0.00000E+00 f(

l l

1 i

)

1 2

$1nPLE Fl5$10N P4000CT MASSES (EG) IN CELL 6 AT TIME e 40000.000 (S) 1$7 TTPE uAME-1 Csl 2 C50N 3 TE 4 sa 5 BA 6 LA 2 Aerosol U02 0.00000E+00 0.00000E+00 8.65718E 05 3.53685E 08 6.62381E 071.75603E 07 0 TOTAL-WALL-1.42311E 02 8.64111E 02 9.39763E 03 3.84184E 04 9.75044E 041.77192E 05

-TOTAL 1.42311E 02 8.64111E 02 9.48421E 03 3.84219E 04 9.75706E 04 1.78948E 05 Mo$f TYPE NAME 7 CE.

'8 RU 9 Pt 2 AEROSOL U02 1.66819E 07 5.30898E 12 2.21844E 03 0 TOTAL WALL 4.11746E 05 1.41502E 10 8.86628E 02 -

TOTAL' 4.13414E 05 1.46811E 10 9.08812E 02 NOST MATERIAL INFORMAfl0N IN CELL 6 AT TIME a 40000.000 (s)

NOST TTPE NAME POWER (WATTS) TEMPERATURE (K) 2 AEROSOL U02 0.00000E+00 3.33650E+02 0 TOTAL WALL 0.00000E+00 0.00000E+0A TOTAL 0.00000E+00 1

i

t F

$1MPLE FIS$10N (90 DUCT MASSES (KG) IN CELL 7 AT TIME a 40000.000 (5)-

M T TYPE NAME 1 CSI 2 CSON 3 TE 4 $R 5 BA 6' LA 2 AERos0L. 002 2.66247E 44 2.66247E 44 4.82489E 041.97535E 07 3.70445E 06 9.77907E*07 0 -TOTAL.

ROOF 4.76490E 03 2.90771E 02 3.156240 031.63462E 04 3.61757E 04 9.47787E-06 O TOTAL.

WALL-

• 5430E 021.73334E 01 1.89603E-02 8.11361E 04 2.00551E-03 4.40818E 05 TOTA.

3.1079E 02 2.02411E 012.25991E-02 9.75021E 04 2.37097E 03 5.45376E 05 N0$f TTPE NAME e CE 8 RU 9 Pt l

2 AEROSOL UO2 9.28992E 07 2.96390E 11 1.23618E 02 0 TOTAL ROOF 2.05411E 05 8.27787E 113.03635E 02 i

'O TOTAL WALL-9.89023E 05 3.77551E 101.74659E 01 TOTAL 1.20372E-04 4.e9969E 10 2.17385E 01 5

l Nolf MATERIAL INFORMATION IN CELL 7 AT TIME = 40000 '.L1 (5)

NOST TTPE NAME POWER (WATTS)- TEMPERATURE (E)

2. AEROSOL UO2 0.'iC000E+00 3.32041E+02 0 TOTAL ROOF

'0.00000E+00 0.00000E+00 0 TOTAL WALL 0.00000E+00 0.00000E+00 l

TOTAL 0.00000E+00 i

a P

l' A

i

$lsePLE. FIS$10N PRODUCT MASSES (EG) IN CELL 8 AT TIME a 40000.000 ($)

MT TYPE NAME 1 CSI 2-CSON 3 TE 4 ER 5 sA 6 LA 2'AER050L U02 4.20390E 45 8.40779E 45 4.82407E 041.98094E 07 3.72216E 06 9.76670E 07 0 TOTAL WALL 2.95257E 031.75604E 021.90002E 031.75845E 031.98977E 03 2.99448E 04 O LOW CELL + DUMMY ' 2.57282E+01 1.52850E+02 1.00457E+01 3.61857E+00 4.72099E+00 9.01801E 01 TOTAL 2.5 7312E+01 1.52868E+02 1.00481E+01 3.62033E+00 4.72298E+00 9.02102E-01 MOST TYPE NAME 7 CE 8 RU 9 P!

2 AEROSOL U02 9.27820E 07 2.97129E 11 1.23566E 02 0 TOTAL WALL 4.79920E 04 3.40905E 09 6.91741E 02 0 LOW CELL + DUMMY 1.32320E+001.20293E 05 3.39147E+02 TOTAL 1.32368E+00 1.20327E 05 3.39229E+02 MOST MATERIAL INFORMATION IN CELL 8 AT TIME = 40000.000 (s)

NOST TYPE NAME POWER (WAfft) TEMPERATURE (K) 2 AEROSOL U02 0.00000E+00 3.44956E+02 0 107AL - WALL 0.00000E+00 0.00000E+00 0 LOW CELL + DUMMY 0.00000E+00 0.00000E+00 TOTAL 0.00000E+00 i

m

l l

1 SIMPLE FIS$10s PRODUCT MAS $ES (KG) IN CELL 9 AT TIME e L0000.000 ($)

%T TYPE NAME -

.1 C$1 2 CSON 3 TE 4 SR 5 BA 6 LA 2 AER0$0L 002 1.80138E 15 9.82980E 15 3.24917E-06 2.61741E 08 2.93760E 08 7.76884E 10 0' TOTAL FLOOR 1.78251E 051.06455E 041.03306E 05 2.87391E 06 3.76692E 06 6.49865E 07 TOTAL 1.78251E 051.06455E 041.03631E 05 2.90008E 06 3.79629E 06 6.50642E-07 MOST TYPE hAME 7 CE 8 Ru 9. PI 2 AERo$0L' U02 1.63078E-09 9.09113E 15 3.64908E 06 0 TOTAL FLOOR 9.75544E 07 8.67538E 12 2.27186E 04 TOTAL 9.77175E-07 8.68447E 12 2.30835E-04 r

hosi MATERIAL luf0AMATION IN CELL 9 AT TIME s 400f-J.000 ($)

Mosi TYPE-NAMF POWFR (WATTS) TEMPERATURE (K) 2 AEROSOL 002 0.00000E+00 3.00000E+02 0 TOTAL FLOOR

'O.00000E+00 0.00000E+00

[

TOTAL 0.00000E+00

~

r l

B

p.-

CR3 K3BA-S1 (LBLOCA all debris cell 1) 27.00 -

26.00 -

25.00 -

in b

24.00 -

ts 23.00 -

to N

OC A

nu-A

~

A h

21.00 -

20.00 -

19.00 -

d 18.00 i

i i

i i

i i

i 0.00 100.00 200.00 300.00 400.00 500.00 600.00 700.00 800.00 900.00 1000.00 TIME SINCE START OF ACCIDENT, MIN Figure A-1: CELL PRESSURE 11/25/91

e CR3 K3BA-S1 (LBLOCA all debris cell 1) o.90 -

o.80 -

Cell l '

o.7o -

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' O -----..._,_ p h

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i.

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i o.co 100.oo 200.00 300.00 me.oo 500.00 600.00 700.00 800.00 900.00 1000.00 l

TIME SINCE START OF ACCIDENT, MIN Figure A-5: CELL STEAM FRACI' ION 11/25/91

D O%'

\\

CR3 K3BA-S1 (LBLOCA' all debris cell 1) 0.90 -

OJO -

ZO

' 7' -

O.60 -

W Zg 0.50 -

g 0.40 -

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0.00 r

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0.00 100.00 200.00 300.00 400.00 500.00 600.00 700.00 800.00 900.00 1000.00 TIME. SINCE START OF ACCIDENT, MIN Figure A-6: CELL HYDROGEN FRACTION 11/25/91-

- ~.

.m CR3 K3BA-S1 (LBLOCA all debris cell 1) 200.00 -

ffe 180.00 -

1l,/

l i j\\;sf's, 160.00 -

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40.00 -

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20.00 -

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0.00

,y-i 0#

100.00 200.00 300.00 TIME SINCE START, SEC.

Figure A-7: CELL OXYGEN FRACTION 1I/25/91 f

r CR3 K3BA-S1 (LBLOCA all debris cell.1) 3,,, _

9.00 -

Z

(

s.00 -

E n4 7.00 -

I d

Q^

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x 0*

5.00 -

e O-Hy Nv 4.00 -

Hg 3.00 -

a 4

2.00 -

m b

^

1.00 -

l l

0.00 i

l 0.00 100.00 200.00 300.00 400.00 500.00 600.00 700.00 800.00 900.00 1000.00 l

TIME SINCE START OF ACCIDENT, MIN Figure A-8: GAS FLOW RATE TO OUTSIDE I1/25/91

\\

t CR3 K3BA-S1 (LBLOCA all ' debris cell 1) 20.00 -

18.00 -

p m

A

~

g 16.00 -

QCw f

14.00 -

m i

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A 10.00 -

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0 2a-o.00 i

I i

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0.00 100.00 200.00 300.00 400.00 500.00 600.00 700.00 sw.m m

im TIME SINCE START OF ACCIDENT, MIN Figure A-9: CSI SUSPENDED IN CELL ATMOSPHERE 11/25/91

CR3 K3BA-S1 (LBLOCA all debris cell 1) 90.00 -

80.00 -

M m

4 70.00 -

',, Cell 8 j

,/

A

/

y 60.00 -

,/

4 Z

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!I II ll 10.00 -

5 1

0.00 0.00 100.00 200.00 300.00 400.00 500.00 600.00 700.00 800.00 900.00 1000.00 TIME SINCE START OF ACCIDENT, MIN i

Figure A-10: CSI DEPOSITED IN CELL 11/25/91

n CR3 K3BA-S1 -(LBLOCA all debris cell 1) 5m-i 495.00 -

',.. Cell 7 y

m a

440.00 -

Ed h

385.00 -

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-4 '

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^

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l:

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0.00 100.00 200.00 300.00 400.00 500.00 600.00-700.00 800.00 900.00 1000.00 TIME SINCE START OF ACCIDENT, MIN Figure A-11: SR SUSPENDED CELL ATMOSPHERE I1/25/91 i

.r CR3 K3BA-S1 (LBLOCA all debris cell 1) 59.m -

~

ox 495.00 -

j

\\

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---M~~7 ^ 7""7""]" ^ 7""7'""7 0.00 j

0.00 100.00 200.00 300.00 400.00 500.00 600.00 700.00 800.00 900.00 1000.00 TIME SINCE START OF ACCIDENT, MIN Figure A-13: FP MASS RELEASED TO ENVIRONMENT 11/25/91

p a

CR3 K3BA-S1 (LBLOCA all-debris cell 1) 4,, _

e00.00 -

N 3500.00 -

-N

~

H 3000.00 -

~

g 2500m -

w H

2000.00 -

1500.00 -

M 1

c Q

1000M -

1 0

500.00 -

d m

0.00 -

t

-se.m 0.00 100.00 200.00 300.00 400M 500.00 600.00

'/00 M 800.00 MM 1000 M TIME SINCE START OF ACCIDENT, MIN Figure A-14: HEA Y O IDE LAYER TEMPERATURE 11/25/91'

O T

.CR3 K3BA-S1 (LBLOCA all debris cell 1) m00.00 -

3500.00 -

W 5

H 3000.00 -

~

g 2500.00 -

N H

2000.00 -

1500.00 -

N Q

1000.00 -

~

O k

500.00 -

a A

0.00 -

-500.00 0.00 100.00 200.00 300.00

1. 0.00 500.00 M.00 700.00 800.00 900.00 1000.00 TIME SINCE START OF ACCIDENT, MIN '

Figure A-16: LIGHT OXIDE LAYER TEMPERATURE 11/25/91l

i CR3 K3BA-S1 (LBLOCA all debris cell 1)-

900.00 -

800.00 -

700.00 -

a

[

600.00 -

m

.)p 500.00 -

Q O

I 400.00 -

e 300.00 -

Uv 200.00 -

100.00 -

0.00 0.00 100.00 200.00 300.00 400.00 500.00 600.00 700.00' 800.00 900.00 1000.00 TIME SINCE START OF ACCIDENT, MIN Figure A-17: CCI H2 PRODUCTION 11/25/91

I m.,

CR3 K3BA-S1 (LBLOCA all debris cell 1) 2em -

18.00 -

16.00 -

14.00 -

g A

ius fx nm-O m

Z 10.00 -

M2

........ q.....................

-A sm-

~.

b

~

.................. ~

' Depth 6.00 -

d 4m -

t 2.00 -

8#

i i

i i

i i

om loom 200.00 3com me.oo 500.00 600.00 voom soom 9oom icoom TIME SINCE START OF ACCIDENT, MIN Figure A-21: CAVITY DIMENSIONS 11/25/91

K3BA (LBLOCA) Contain 2sm -

27.00 -

um-1s m -

d m

24.00 -

r.J M

23m -

m c

W 5

M A

22.00 -

J I

a w

y 21.00 -

3 2

I 20.00 -

a b

"\\

19.00 -

1sm om loom zoom aoom 4com soom som 7eom soom soom locom TIME SINCE START OF ACCIDENT, MIN 10/21/91

APPENDIX B DCH Analysis Results t

524

Table B.1-1: Fission Products Distribution SIMPLE FISSION PRODUCT MA$$ES (KG) IN CELL 1

AT TIME = 90000.000 (s)

MOST TYPE NAME 1 CSI 2 CSON 3 TE 4 $R 5 BA 6 LA 2 AEROSOL U02 2.97202E 08 7.16163E-09 2.16615E 04 1.44098E 06 4.12289E 06 2.25221E 07 0 TOTAL ROOF 4.76773E 04 2.08931E 031.50405E 04 2.97015E-04 2.99675E-04 8.91387E 05 0 TOTAL WALL 3.29824E 031.34675E 021.59925E 03 5.21227E 03 5.10516E 031.65177E 03 0 TOTAL FLOOR 1.90007E 01 9.56478E 01 4.93591E 02 1.09754E 01 1.29817E 01 1.92109E-02 TOTAL 1.93782E 01 9.72035E 01 5.13254E 02 1.15264E 01 1.35226E 01 2.09521E 02 HOST TTPE NAME 7 CE 8 RU 9 PI 2 AEROSOL UO2 2.17329E 07 7.71241E 08 3.06698E-03 0 TOTAL R00F 8.04978E 05 2.19177E 08 9.76946E-03 0 TOTAL WALL 1.42751E-031.47985E 071.29050E 01 0 TOTAL FLOOR 1.77313E 02 9.34016E 07 7.10945E+00 TOTAL 1.92395E 02 1.18104E 06 7.25134E+00 HOST MATERIAL INFORMATION IN CELL 1

AT TIME = 90000.000 (S)

H0$f TYPE NAME POWER (WATTS) TEMPERATURE (K) 2 AERos0L 002 0.00000E+00 1.41518E+03 0 TOTAL ROOF 0.00000E+00 0.00000E+00 0 TOTAL WALL 0.00000E+00 0.00000E+00 0 TOTAL FLOOR 0.00000E+00 0.00000E+00 TOTAL 0.00000E+00

(

j 1

$25 i

SIMPLE FIS$10N PRODUCT MA55ES (KO) IN CELL 2 Af f!ME =

90000.000 (s)

N057 TYPE NAME 1 C$1 2 C50N 3 TE 4 se 5 BA 6 LA 2' AEROSOL U02 1.73984E 07 4.19224E 08 9.48751E 04 8.31163E 061.92648E 051.06150E 06 0 f0fAL WALL 9.50002E 03 3.4724TE 021.11934E 021.02771E-021.07291E 02 2.52987E 03 0 TOTAL FLOOR 3,71928E 021.47901E 014.06854E 02 9.42759E-02 9.80830E 021.68227E 02 TOTAL 4.66930E 021.82626E-015.28275E 021.04561E 01 1.08831E-012.13536E-02 HOST TYPE NAME 7 CE 8 RU 9 PI 2 AEROSOL 002 1.02735E 06 2.97560E 071.43562E 02 0 total WALL 1.47624E 031.71353E 06 6.74977E 01 0. TOTAL FLOOR 1.52214E-02 2.18437E-06 5.01102E+00 TOTAL 1.66986E 02 4.19546E 06 5.70035E+00 Mosi MATERIAL INFORMATION IN CELL 2 AT TIME = 90000.000 (s)

NOST TYPE NAME POWER (WATTS) TEMPERATURE (K) 2 AEROSOL 002 0.00000E+00 5.03473E+02 0 TOTAL WALL 0.00000E+00 0.00000E+00 0 TOTAL FL0ct 0.00000E+00 0.00000E+00 T07AL 0.00000E+00 t

l 1

j l

\\

526 o

$1MPLE FIS$10N PRODUCT MASSES (KG) IN CELL 3 AT TIME = 90000.000 (s) 40$f TYPf NAME 1 C$1 2 CSON 3 TE 4 $2 5 BA 6 LA 2 AEROSOL UO2 2.68919E 06 6.47937E 071.46663E 021.28460E 04 2.97794E 041.64087E 05 0 TOTAL WALL 4.99151E-02 1.86072E 01 4.73255E 02 5.65103E 02 5.8 %55E 02 1.39796E 02 0 TOTAL FLOOR 8.80347E 013.65522E+00 8.45159E 01 1.96678E+00 2.06055E+00 3.85393E 01 TOTAL 9.30265E 013.84129E+00 9.07150E 012.02342E+00 2.11982E+00 3.99389E 01 Most TYPE NAME 7 CE 8 RU 9 P!

2 AEROSOL 002 1.58807E 05 4.60020E 06 2.21915E 01 0 TOTAL WALL 9.07215E 03 6.47682E 06 3.24951E+00 0 TOTAL FLOOR 3.16266E 01 4.26222E 05 1.05393E+02 TOTAL 3.25354E-01 5.36992E-05 1.08864E+02 Hosi MATERIAL INFORMAfl0N IN CELL 3 AT TIME a 90000.000 (5)

NO$f TYPE NAME POWER (WATTS) TEMPERATURE (K) 2 AERO$0L UO2 0.00000E+00 5.17865E+02-0 TOTAL WALL 0.00000E+00 0.00000E+00 0 TOTAL FLOOR 0.00000E+00 0.00000E+00 TOTAL 0.00000E+00 t

t I

i l

1 C

527

$1MPLE F15$10N PRODUCT MASSES (KG) IN CELL 4 AT TIME = 90000.000 (5)

N05T TYPE NAME 1 C$1 2 CSON 3 TE 4 st 5 BA 6 LA 2 AEROSOL' U02-4.61973E 061.10991E 06 2.53799E-02 2.19941E-04 5.14228E 04 2.83472E 05

.0 total WALL-7.19393E 01 3.70498E+00 1.18061E 01 1.36417E *01 1.6552M *01 3.35486E 02 0 TOTAL FLoom 4.26454E+00 2.28632E+01 1.01101E+001.98645E+00 2.51247E+00 3.17363E 01 i

fotAL 4.98394E+00 2.65682E+01 1.15445E+00 2.12308E*00 2.67851E+00 3.50940E 01 Most TYPE NAME 7 CE 8 RU 9 P1

2. AEROSOL UO2 2.74313E-05 7.99584E 06 3.82991E 01 0 TOTAL WALL 2.48294E 02 6.98366E 06 9.26586E+00 0 TOTAL FLOOR 2.84885E 01 1.84294E 05 2.14917E+02 TOTAL 3.09742E 01 3.34089E 05 2.24565E+02 HOST MATERIAL INFORMATION IN CELL 4 AT TIME = 90000.000 (S)

NOST TYPE NAME POWER (WAffS) TEMPERATURE (K) 2 AEROSOL Uo2 0.00000E+00 6.34099E+02 0 TOTAL WALL 0.00000E+00 0.00000E+00 0 TOTAL FLOOR 0.00000E+00 0.00000E+00 TOTAL 0.00000E+00 4

• t 528

1

$1MPLE flS$10N PRODUCT MASSES (KG) IN CELL 5 AT TIME = 90000.000 ($)

g HOST TYPE NAME 1 CSI 2 CSON 3 TE 4 SR 5 BA 6 LA 2 AEROSOL 002 1.89494E 06 4.56217E 071.03523E-02 9.04372E 05 2.10094E-041.15777E 05 0 TOTAL. WALL 2.11283E 01 1.06565E+00 4.49992E 02 5.45779E 02 6.48391E 021.37388E 02 TOTAL-2.11285E 01 1.06565E+00 5.53516E 02 5.46683E 02 6.50492E 021.37503E 02 HOST TTPE NAME 7 ' CE 8 AU 9 PI 2 AEROSOL 002 1.12048E 05 3.25059E 06 1.56537E 01 0 TOTAL WALL 9.96590E-03 3.32162E 06 3.68991E+00 TOTAL 9.97710E 03 6.57220E 06 3.84644E+00 HOST MATERIAL INFORMATION IN CELL 5 AT TIME = 90000.000 ($)

MOST TYPE NAME POWER (WATTS) TEMPERATURE (K) 2 AERos0L 002 0.00000t+00 6.09881E+02 0 TOTAL WALL 0.00000E+00 0.00000E+00 i

TOTAL 0.00000E+00 P

d g.

d a

4 J

I e

A 4

k i

529 i

J

$1MPLE FIS$10M PRODUCT MASSES (EG) IN CELL 6

AT TIME = 90000.000 ($)

MOST TYPE-NAME 1 C$1 2 CSON 3 TE 4 SR 5 BA 6 LA

'2 AEROSOL 002 2.21467E 06 5.31911E 07 1.21621E 02 1.05406E-04 2.46443E 04 1.35854E 05 0 TOTAL WALL 1.27330E 015.65126E 014.03826E 02 7.22361E-02 7.97990E 02 1.85437E 02 TOTAL 1.27332E 015.65127E 015.25447E 02 7.23415E 02 8.00454E 021.85573E 02 MOST TYPE

'mAME 7 CE 8 RU 9 PI 2 AEtos0L u02 1.31464E 05 3.83126E 06 1.83533E 01 0 TOTAL WALL 1.34282E 02 3.40578E 06 3.82542E+00 TOTAL 1.34414E 02 7.23704E-06 4.00895E+00 Mo$T MATERIAL INFonMAfl0N IN CELL 6 AT TIME = 90000.000 ($)

MOST TYPE NAME POWER (WATTS) TEMPERATURE (K) 2 AEROSOL U02 0.00000E+00 6.26123E+02 0 TOTAL WALL 0.00000E+00 0.00000E+00 TOTAL 0.00000E+00 4

(.L 530

$1MPLE FIS$10N PRODUCT MA$$ES (KG) IN CELL 7

AT TIME = 90000.000 (s)

MOST TYPE NAME 1 CSI 2 C$0N 3 TE 4 SR 5 BA 6 LA 2 AEROSOL U02 5.67498E 051.41430E E5 3.21024E-012.80356E 03 6.51458E 03 3.59004E-04 0 TOTAt 2007 1.49101E 02 6.54605E 02 9.07962E 04 2.33952E 04 6.51128E 04 4.82406E-05 0 TOTAL WALL 1.26829E 015.77667E 01 1.81906E 021.75129E-02 2.41164E-02 3.56665E 03 -

TOTAL 1.41798E 016.43142E 013.40122E 012.05504E 02 3.12821E 02 3.97389E 03 HOST TYPE NAME 7 CE 8 RU 9 P!

2 AEROSOL 002 3.47440E 041.00813E 04 4.85378E+00 0 TOTAL ROOF 2.89081E 05 6.47085E 08 9.18698E 02 0 TOTAL-WALL 3.01389E 03 1.63024E-06 1.55448E+00 TOTAL 3.39024E-031.02508E 04 6.50014E+00 Most MATERIAL INFORMAfl0N IN CELL T

AT TIME = 90000.000 ($)

MOST TYPE NAME POWER (WATTS) TEMPERATURE (K) 2 Atto$0L Uo2 0.00000E+00 6.18730E+02 l

0 TOTAL ROOF 0.00000E+00 0.00000E+00 0 TOTAL WALL 0.00000E+00 0.00000E+00 TOTAL 0.00000E+00 a

7 4

w 4

531

'l l

$IMPLE FIS$10N PRODUCT MA$$ES (KG) IN CELL 8 AT TIME = 90000.000 ($)

MOST TYPE NAME 1 C51 2 CSON 3 TE 4 $R 5 BA 6 LA 2 AEROSOL. 002 2.60811E 05 6.41928E 061.40137E 01 1.25935E 03 2.84964E 031.58000E 04 0 TOTAL WALL 8.18229E 02 3.56996E 01 1.81606E 02 3.16128E 02 3.52135E 02 9.21412E 03 0 TOTAL FLOOR 1.12233E+01 5.61161E+01 4.54524E+00 1.16661E+01 1.31806E+01 1.93050E+00 foiAL 1.13051E+01 5.64731E+01 4.70354E+00 1.16989E+01 1.32187E+01 1.93987E+00 MO$f TYPE NAME 7 CE 8 RU 9 PI 2 AEROSOL U02 1.52954E 04 4.30976E 05 2.13835E+00 0 TOTAL WALL 5.85682E-03 1.82834E-06 1.66853E+00 0 TOTAL FLOOR 1.72677E+00 5.93265E 05 7.07830E+02 TOTAL 1.73278E+001.04252E 04 7.11637E+02 Most MATERIAL INFORMATION IN CELL 8 AT TIME = 90000.000 ($)

HOST TYPE NAME POWER (WATTS) TEMPERATURE (K) 2 AER0$0L 002 0.00000E*00 5.82851E+02 0 TOTAL WALL 0.00000E+00 0.00000E+00 0 total FLOOR 0.00000E+00 0.00000E+00 TOTAL 0.00000E+00 l

j b

532

k SIMPLE FIS$10N PRODUCT MASSES (KO) IN CELL 9 AT TIME = 90000.000 (5)

MOST ffPE NAME 1 CSI 2 C50N 3 ft 4 $8 5 BA 6 LA 5.70876E+00 2.14876E+016.16851E 01 2.13763E+00 2.30872E+00 8.69210E 01 2 AER050L 002 1.46785E+00 5.52660E+001.19499E 015.32757E 015.76213E 012.21142E 01 0 TOTAL FLOOR 7.17661E +00 2. 70142E+01 7.36350E 01 2.67039E +00 2. 88494E +00 1.09035E +00 TOTAL

1. 0$7 ffPE NAME 7 CE 8 RU 9 PI 2 AEROSOL 002 3.48630E 01 3.09864E 05 6,33975E+01 0 TOTAL FLOOR 8.78861E 021.09137E 061.43678E+01 total 4.36516E 01 3.20T77E 05 7.77652E+01 MOST MATERIAL INFORMATION IN CELL 9 AT TIME = 90000.000 ($)

HOST TYPE NAME POWER (WAffS) TEMPERATURE (K) 2 AERos0L UO2 0.00000E+00 3.00755E+02 0 TOTAL FLOOR 0.00000E+00 0.00000E*00 TOTAL 0.00000E+00 I

i i

f 1

533 I

4 CR3 BLACKOUT CONTAIN RESULTS DCH 51 m -

4600.00 -

4100.00 -

CELL 1 N

3600.00 -

g W

,.C E L L 2 Q

g 3100.00 -

i,,,,-

I CELL 4 b'

f 2600.00 -

C E L L 6

,C E L L 3

/

l w

2100.00 -

[

, CELL 5 j

-W lt i

/

H l

/

/

4 1600.00 -

/

J ll l

M

\\

/

CELL 7

/

, '.f.g.

1100.00 -

\\

!,&~

'.. ' '^ -..

600.00 -

I

\\

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' CELL 8

" * " - ~

i 100.00 480.00 482.00 484.00 486.00 488.00 490.00 492.00 494.00 4 %.00 498.00 500.00 TIME SINCE START OF ACCIDENT, MIN Figure B-1: CELL TEMPERATURE 10/16/91

p.

CR3 BLACKOUT CONTAIN RESUUfS DCH 1509.00 -

1359.00 -

-1209.00 -

1959.00 -

m 909.00 -

Y 759.00 -

Cr)N 04 A

609.00 -

A

1 h

459.00 -

309.00 -

159.00 -

9.00 j

480.00 482.00 484.00 486.00 488.00 490.00 492.00 494.00 4 %.00 498.00 500.00 TIME SINCE START OF ACCIDENT, MIN Figure B-4: CELL PRESSURE 10/16/9i

O CR3 BIACKOUT CONTAIN RESULTS DCH 1509.00 -

us9.00 -

1209.00 -

1959.00 -

d eA m.u -

A

~

m 759.00 -

mW ma A

609.00 -

A

~

JW g

459.00 -

l 1

309.00 -

~

159.00 -

1 i

9.00

~,

i 480.00 482.00 484.00 486.00 488.00 490.00 492.00 494.00 496.00 498.00 500.00 TIME SINCE START OF ACCIDENT, MIN 1

Figure B-5: CELL PRESSURE 10/16/91

O CR3 BIACKOUT CONTAIN RESULTS DCH 1509.00 -

1359.00 -

1209.00 -

1059.00 -

m A

909.00 -

LE i

759.00 -

m Er.1 og A

609.00 -

A

~

a 459.00 -

309.00 -

159.00 -

a 9.00 i

i i

i i

i i

i i

'7 480.00 482.00 484.00 486.00 488.00 490.00 492.00 494.00 4 %.00 498.00 500.00 TIME SINCE START OF ACCIDENT, MIN Figure B-6: CELL PRESSURE 10/16/91

^

p..

s CR3 BLACKOUT CONTAIN RESULTS DCH 1509.00 -

1359.00 -

BA 1209.00 -

1959.00 -

d m

A 909.00 -

N 759.00 -

m N

cc:

A 609.00 -

A

^

A h

459.00 -

309.00 -

159.00 -

9.00

~

1 i

i i

i i

i i

i i

480.00 482.00 484.00 486.00 488.00 490.00 492.00 494.00 496.00 498.00 500.00-TIME SINCE START OF ACCIDENT, MIN Figure B-7: CELL PRESSURE 10/1(>/91

r CR3 BIACKOUT CONTAIN RESULTS DCH 1509.00 -

1359.00 -

1209.00 -

1059.00 -

8 m

909.00 -

N 759.00 -

m m

ed 0i 609.00 -

A

~

A h

459.00 -

309.00 -

159.00 -

9.00 i

i i

i i

i 480.00 482.00 484.00 486.00 488.00 490.00 49,2.00 494.00 496.00 498.00 500.00 TIME SINCE START OF ACCIDENT, MIN Figure B-8: CELL PRESSURE 10/16/91

~

,r.

'CR3 BLACKOUT CONTAIN RESULTS DCH 1509.00 -

1359.00 -

1209.00 -

1059.00 -

b m

b 909.00 -

Ed 759.00 -

m GrJ od A

609.00 -

A

~

J h

459.00 -

i 309.00 -

159.00 -

~

9.00 i

i i

i i

.- j 480.00 482.00 484.00 486.00 488.00 490.00 492.00 494.00 496.00 498.00 500.00 TIME SINCE START OF ACCIDENT, MIN Figure B-9: CELL PRESSURE 10/16/91'

P

~

CR3 BIACKOUT CONTAIN RESULTS DCH 31,,,_

4600.00 -

4100.00 -

~

3600.00 -

g N

g 3100.00 -

2600.00 -

8

'2100.00 -

NH

~

J 1600.00 -

,3 N

U 1100.00 -

i 600.00 -

100.00 q

483.00 485.00 487.00 489.00 491.00 493.00 495.00 497.00

' 499.00 501.00 503.00 TIME SINCE START OF ACCIDENT, MIN Figure B-10: CELL TEMPERATURE 10/17/91

.-r.

. ~

.m

m l

CR3 BIACKOUT CONTAIN RESULTS DCH 1.00 -

T 0.90 -

\\

\\

i ll \\

. i 0.80 -

1 i

\\

\\

e

's 0.70 -

'N

~~

'N N

,'N 2:

s O

0.60 -

'O g

j 0.50 -

,i 0.40 -

\\g g

.A 1

raa U

0.30 -

0.20 -

0.10 -

. _h.

i i

i i

i i

i 0.00 483.00 485.00 487.00 489.00 491.00 493.00 495.00 497.00 499.00 501.00 503.00 TIME 'SINCE START OF ACCIDENT, MIN Figure B-11: CELL 4 FRACTIONS 10/17/91

.m CR3 BIACKOUT CONTAIN RESULTS DCH 1.00-

~

,n.

0.90 -

- l 'g

~

ii

\\,

'g 0.se -.

\\

\\'

0.70 -

~ ~. -

ro 2:

0.60 -

o w

f 0.50 -

m i

0.40 -

N g

A WU 0.30 -

0.20 -

0.10 -

i

~

\\1 0.00 i

,i 483.00 485.00 487.00 489.00 491.00 493.00 495.00 497.00 499.00 501.00 503.00 TIME SINCE START OF ACCIDENT, MIN Figure B-12: CELL 5 FRACTIONS 10/17/91

CR3 BLACKOUT CONTAIN-RESULTS DCH 31,.

4600.00 -

4100.00 -

e 3600.00.-

N A

g 3100.00 -

~

b 2600.00 -

IW

}

k 2100.00 -

NH j

1600.00 -

N U

1100.00 -

600.00 -

100.00 i

i i

i i

i i

i 483.00 485.00 487.00 489.00 491.00 493.00 495.00 497.00 499.00 501.00 503.00 i

TIME SINCE START OF ACCIDENT, MIN Figure B-13: CELL TEMPERATURE 10/17/91 ;

.-~

p; CR3 BIACKOUT CONTAIN RESULTS DCH 51oo.co -

4600.co -

4100.00 -

~

3600.00 -

g W

A g

3100.00 -

H 2600.00 -

h 2100.00 -

WH J

1600.00 -

3 W

U 1100.00 -

600.00 -

I# #

i i

i i

i r- - '

r-7 483.00 485.00 487.00 489.00 491.00 493.00 495.00 497.00 499.00 501.00 503.00 TIME SINCE START OF ACCIDENT, MIN Figure B-14: CELL TEMPERATURE 10/17/91 m

a m-u

'+P

i

,;a -

APPENDIX C L

CONTAINMENT RESPONSE TO COMBUSTION OF HYDROGEN i

l l

i 1

58 534

C.1 Introduction The-temperature and pressure response to a postulated hydrogen burn in the containment was analyzed for the CR-3 plant. The amount of hydrogen burned was treated as a parameter in the form of an effective fraction of zirconium reacted. The core contains 50,100 lb of zirconium and a 100%

zirconium reaction would correspond to 1,100 lb-moles of hydrogen. Hydrogen generated from sources other than zirconium steam reactions can be converted to equivalent fractions of zirconium reacting and added in so that theoretically, more than 1,100 lb-moles of hydrogen could be available.

However,in practice, due to the inability to react all of the zirconium in the core and because early in the accident sequence other hydrogen sources are generally small, a 100% zirconium reaction is considered to be a very conservative upper limit, at least for considerations of early containment

. failure.

The hydrogen combustion reaction consumes 0.5 lb mole of oxygen for every 1.0 lb-mole of hydrogen and it liberates 104,000 BTU of exothermic energy for each lb-mole of hydrogen burned. Thus, the combustion of 1,100 lb-moles of hydrogen would consume 550 lb-moles of oxygen. The CR-3 containment volume is 2.04x10' cubic feet, and 550 lb. moles of oxygen corresponds to about 52%

of all the oxygen in the containment at 100 'F. Thus, there is sufficient oxygen to react all the hydrogen potentially generated at early times.

The heat of reaction of 104,000 Btullb-mole is smaller than the value of 124,000 Btu /lb-mole l

frequently seen in the literature. The former is called the lower heat of combustion and considers the i

reaction product (H O) as a vapor; the latter value is called the upper heat of combustion. It 2

considers the reaction product as liquid water and thus includes the heat of condensation. The reaction products immediately following a hydrogen burn will be in the vapor state and thus the correct heat balance to determine containment pressure is made with the lower heat of combustion.

The combustion of 1,100 lb-moles of hydrogen would release 114 million BTU of energy which corresponds to 49 full i: ore power seconds or the first 40 minutes of decay heat.

The effect of a hydrogen burn on the containment pressure and temperature is bounded by an adiabatic heatup of the containment gas mixture by the chemical energy of combustion. This is usually a reasonably good approximation, particularly for hydrogen concentrations of around 8% or more, because the burn completes in a few seconds and the energy loss to the containment structures

,_(

over the duration of the burn is small. However, it is important to account for the incompleteness

$35

I 1

of the burn (burn efficiency), and for the temperature-dependent specific heats of the constituent j

gases because of the high post-burn temperatures. For the subsequent calculations, temperature.

dependent internal energies of the constituent gases are taken from Reference C-1.

C.2 :

Global Ilydrogen Burning It is known from experimental data (Reference C-2) that for combustion of lean mixtures in a i

quiescent atmosphere, the pressure rise is significantly lower than predicted by an adiabatic complete i

burn model. This is in large part due to the fact that lean flammable mixtures do not burn completely.

For example, a 5.6% hydrogen concentration would only burn to about 50%

completeness, an 8% concentration would burn about 85% complete, and an 11% concentration I

i would burn about 95% complete (Reference C-2). In agitated or turbulent atmospheres, however, the data compiled in Reference C 2 indicate that significantly higher pressure rises can occur in the i

range of 4 to 8% hydrogen compared to quiescent atmospheres. In a large, dry containment with the debris heat source at the lowest elevation, it is expected that substantial mixing currents and convection will occur and that the containment atmosphere is agitated. As shown in Figure C 1 from a,

l Reference C-2, the data on burn pressure rise for agitated atmosphere tests are bounded by a f

pressure rise efficiency or burn efficiency which varies linearly from zero at 4% hydrogen to 100%

i at 8% hydrogen. His correlation has been incorporated into a constant volume, adiabatic burn j'

model with temperature-dependent specific heats from Reference C-1, which was used to calculate i

the post burn temperatures and pressures for different initial conditions. The results are shown in Figure C-2 in the form of a hydrogen burn nomogram for the CR-3 plant. Figure C 2 shows the results as a function of the pre burn containment atmosphere temperature, ranging from 100 *F to 350 'F, and of the amounts of zirconium reacted, ranging from 0 to the equivalent of 150% of the 4

core zirconium inventory. The results of Figure C 2 apply for a saturated containment atmosphere.

I In severe accidents, the containment atmosphere always has a significant steam content, typically in the range of 50% to 100% of saturated conditions at the containment temperature. Hydrogen burn

?

calculations were also performed for 50% and 75% saturation conditions, and the results are very i

similar to Figure C-2. Figure C-3 shows the maximum post-burn pressure (i.e., the highest pressure in the shaded area of Figure C 2) as a function of the steam saturation fraction. It is seen that the maximum pressure is nearly independent of the saturation fraction between 50% and 100%.

'Iherefore, Figure C-2'will be used together with the results from the CONTAIN analyses for l

hydrogen burns to estimate all hydrogen burn split fractions.

/

536 i

l 6

The lower portion of Figure C 2 shows the hydrogen concentration as a function of the fraction of

)

zirconium reacted at the six different preburn temperatures. He concentration of hydrogen in the

' containment depends only slightly on the normal temperature of the air. More importantly, it depends on the amount of steam in the air at the time when the burn initiates. At 100 F, there are initially 5.000 lb. moles of air in the CR 3 containment. If, after primary system blowdown, saturated steam and air equilibrate at a typical value of 200 *F, about 3,370 lb-moles of steam are mixed with the air. Thus the steam concentration would be about 40 percent.

At higher equilibrium tempe.atures, the amount of saturated steam increases further.

It can be seen in Figure C 2 that the global Hammability limit of 8% cannot be reached with a 100%

zirconium reaction fraction if the saturated steam atmosphere is at a temperature of 240 *F or higher.

f In dry air, an 8% hydrogen concentration would be reached when approximately 40% of the zirconium is reacted. The analysis assumes that the flammability limits are independent of the type of non-reacting diluent which is assumed to be nitrogen. However, steam is known to be a more l

cffective diluent than nitrogen and, thus, there would be a slight increase in the flammability limits for increasing steam concentrations, which is neglected in the analysis.

f

(

The post-burn pressure is shown in the upper portion of Figure C-2. De area within the hcavy lines indicates the area of interest for global burn conditions. This area is bounded by the lower Dammability limit at 4% on the left, by the curve for which the steam concentration does inert the atmosphere on the upper left, which corresponds to a saturation temperature of approximately 240

• F, and by the curve at the top which corresponds to 100% of the core zirconium reacted. A steam-hydrogen-air mixture is steam inerted if the steam concentration is about 56% (Reference C-2) irrespective of the hydrogen and oxygen concentration. A saturated atmosphere is steam inerted at a temperature of about 240 *F.

At 50% saturation, the atmosphere is inerted at a temperature of 300 *F. He shaded area in Figure C-2 indicates that the pressure in the CR 3 containment due to global hydrogen burning can not exceed 140 psia, even if hydrogen generation from as much as 100%

of the core zirconium is considered. Figure C-3 shows that this maximum hydrogen burn pressure is nearly independent of the steam saturation fraction. At 75% zirconium oxidation, the maximum i

hydrogen burn pressure is seen to be 107 psia and at 50% zirconium oxidation the maximum pressure is only 72 psia.

4 i-Figure 4.4 2 in Section 4.4 shows the probability of containment failure as a function of pressure.

,'.k It is noted that the probability of containment failure from hydrogen burning in the CR 3 537

containment could be as high as 0.53 from the limiting hydrogen burn in the shaded area of

~

Figure C 2. De adiabatic burn model may be conservative by a few psi. but the conclusion remains that a hydrogen burn could challenge the CR 3 containment, ne nomogram can be read either for a " percent zirconium reacted" value on the lower left axis, or from a " hydrogen concentration" value on the bottom axis. The line with arrows gives an example for reading the nomogram at 80% zirconium reacted, and a preburn temperature of 200 'F. Dis -

nomogram is used extensively to quantify the hydrogen burn and containment failure split fractions on the containment event tree as explained in Section 4.7.

C.3 Continuous Discharge Burning of Ilydrogen Hydrogen does not necessarily have to accumulate in the containment until a globally Hammable mixture is obtained. Hydrogen can also ignite at the point of release from the primary system or from an inerted region and burn continuously as a flame torch at the point of release. Discharge flames have not been studied as extensively in the past as global burns. For a fixed quantity of hydrogen burned, discharge burns can potentially yield higher pressures than global burns if the quantity of

('

hydrogen burned corresponds to less than an 8% concentration in the containment. His is due to the reduction in burn efficiency for hydrogen quantities that would yield a global hydrogen concentration of less than 8%. He global burn efficiency limit does not necessarily apply to a discharge burn. In fact, a highly concentrated discharge flame can burn completely, whereas the same quantity of hydrogen when distributed in the containment, may only yield a concentration ofless thca 4% and not burn at all. Figure C-4 shows the hydrogen burn map for discharge burns. It differs from Figure C 3 in that it does not include a reduction in the burn completeness below 8% hydrogen concentrations. His figure should be used for any situation where all the hydrogen burns. This can occur either as a discharge burn with a flammable mixing zone or in a hypergolic recombination.

A discharge burn can occur either at the location of a break in the primary sptem, such as a small or large LOCA break, or at the PORV drain tank relief valve. It can also occur at the point of release into the main containment volume if the containment sub-volume at the release location is inerted. Alternatively, for small LOCAs and transients, a discharge burn can occur at vessel failure -

as a result of the rapid discharge of the RCS steam and hydrogen inventory out of the reactor cavity region into the main containment volume. In order to assess the flammability of a hydrogen steam mixture discharging into an air steam mixture, use is made of the air-steam. hydrogen flammability 538

i diagram shown in Figure C 5 from Reference C-3.The containment atmosphere steam concentration

'is marked on the lower scale and the hydrogen concentration in the primary system steam is marked on the slanting scale. 'Ihe two points are connected to obtain the mixing line, which indicates the possible compositions of mixed gases. If the mixing line crosses into the flammable domain, a discharge burn is possible if ignited. By drawing the tangent to the flammability curve from the containment atmosphere composition point, it is possible to determine the minimum concentration of hydrogen required in the discharge mixture for flammability. This is shown to be 70% in the example in Figure C-5.

me, 539 s

N References for Appendix C

~

C1 Ashley, S. Campbell. Thermodynamic Analysis of Combustion Engines, John Wiley and Sons. New York,1979.

C2 Sherman, M. P., et al, "The Behavior of Hydrogen During Accidents in Light Water Reactors," USNRC Report NUREG/CR 1561 (SAND 80-1495), Sandia National Laboratories, August 1980.

-C-3 Pickard, Lowe and Garrick, Inc.,"Seabrook Station Probabilistic Safety Analysis," prepared i

for Public Service Company of New Hampshire and the Yankee Atomic Electric Company, PLG-0300, December 1983.

t l

l Figure C-1:

Experimentally Determined Pressure Rise From Hydrogen Burns (Data from Reference C-2) 1.0

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a to 12 HYOROGEN CONCENTRAfl0N M) 541

Figure C-2: Hydrogen Burn Map for Saturated Conditions - Global Burns i

200 180 160 i

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Figure C-4: Hydrogen Burn Map for Saturated Conditions - Locel Burns l

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Figure C-5: Flammability of Hydrogen-Steam Discharges into Air-Steam Mixtures (From Reference C-3) i i

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0 10 20 30 40 50

• 60 70 80 90 100 i

WATER VAPOR IN AIR (VOLUME. PERCENT) e STEAM CONCENTRATION IN CONTAINMENT: Xg O = 0.57 2

'

• MINIMUM FLAMMABLE HYDROGEN CONCENTRATION IN DISCHARGE:

MIN X H,D =0.7 545

APPENDIX D i

RCS FAILURE MODES AT HIGH PRESSURE l

e k..

546

D.1

' Introduction 1

l i

Accident sequences in which the RCS pressure remains high until vessel melt through can lead to natural convection heat transport within the reactor vessel and RCS. High wa!! temperatures and high pressures can lead to creep deformations and possibly creep rupture of the pressure boundary.

Therefore,it is necessary to address the vessel failure mode and the subsequent containment response to assess the risk associated with pressurized sequences.

Dree basic RCS failure modes have been identified for high RCS pressure conditions. All are related to the combination of high pressure and high temperatures that may occur in the RCS after core uncovery starts. He three potential failure location and failure modes are:

1. Thermal creep failure of the RCS hot leg or pressurizer surge line.

2. Thermal creep failure of the steam generator tubes.

3. Vessel melt-through failure by molten debris attack of the in-core instrument penetrations.

Thermal creep failure of the RCS hot leg and/or pressurizer surge line are a concern only for high l

(

RCS pressure conditions. A high RCS pressure promotes natural circulation in the vessel, hot leg, and steam generator when the temperature begins to increase after core uncovery. He transient i

analyses in Section 4.6 predict that the hot leg can reach temperature levels of 1250 'F before vesse!

breach. It is to be determined whether creep rupture could occur at the calculated temperature in

-l the time that the high temperature condition exists before vessel meit through.

The CR 3 hot legs and surge line are made of carbon steel, which has a lower temperature threshold q

{

for creep rupture than either stainless steel or Inconel. Creep rupture of the hot leg or surge line will be controlled by hoop stresses that are expected to yield a failure area of a size that results in rapid RCS depressurization. If depressurization does not occur rapidly due to the size of the initial

)

leak area, the flow of hot gases through the initial rupture would quickly cause local heating of the rupture area and further promote an increasing creep rupture area. Herefore, if hot leg or surge

]

line creep rupture is predicted to occur first, the RCS will depressurize quickly, resulting in low RCS 1

pressure at vessel melt-through.

l J

Creep rupture of one or more steam generator tubes is also a possibility at high RCS pressure.

Although less likely because the steam generator tubes are farther from the core, the tube geometry 547

I u less conductive to natural circulation, and the tube materials have a higher resistance to creep rupture than the hot leg piping. Rupture of steam generator tubes would be an undesirable event 4

because it 'would create a potential for a containment bypass release path if the RCS is not -

~

depressurized and if the faulted steam generator safety valve opened or if the atmospheric relief valve

+

j on the faulted steam generator is open.

1 s

' De RCS vessel is assumed to eventually fail by bottom head vessel mell-through'in all unmitigated i

core melt accidents. If this failure mode occurs at high pressure, there is a possibility of an early containment failure should the core debris be ejected from the vessel at high pressure, be finely fragmented, and cause direct heating of the containment atmosphere by the debris and exothermic s

chemical reactions. The driving force for debris dispersal and direct heating is reduced if one of the RCS failure modes discussed above occurs and depressurizes the RCS before the time of vessel

(

melt-through. Similarly, if the operator depressurizes the RCS by opening the PORVs, the RCS l

pressure at vessel melt-through would be so low that the concern about a steam generator tube creep rupture could be climinated and the concern about direct heating could be substantially reduced. -

l~

I D.2 Creep Rupture Failure Creep rupture of ductile materials is described by three parameters: time, temperature, and stress.

For PWR hot legs and steam generator tubes, creep rupture data were compiled and published in

{

the NUREG 1150 draft report (Reference D-1). Figure D-1 is reproduced from the NUREG 1150 draft document. It shows the creep rupture time for the carbon steel hot legs at Surry as a function of temperature and RCS pressure. The hot leg material at CR-3 is also carbon steel, and the l

thickness to diameter (t/D) ratio which governs the hoop stress is very similar. For the Surry hot l_

legs, t/D = 0.086, whereas for CR 3, t/D = 0.080. Herefore, the Surry curves in Figure D-1 can be used to evaluate hotleg creep rupture at CR-3.

i ne steam generator tubes at CR 3 are made of an Inconel type Ni-Cr Fe alloy. Figure D-2, also taken from Reference D 1, shows the creep rupture characteristics of Inconel, which can be used to

)

l

' evaluate creep rupture of the steam generator tubes at CR-3. It is seen from a comparison of Figures D-1 and D-2, that Inconel is considerably more resistant to creep rupture than carbon steel l

. In the CR-3 design, the steam generators form a cold well in the primary system loop, and therefore, the tubes will always be much colder than the hot legs. With lower temperatures and a higher creep

,,k rupture resistance, creep rupture of a steam generator tube in a B&W design will always be much 548

)

i

less likely than rupture of the hot leg or surge line. This conclusion was also reached in Reference D 2 for a Westinghouse design PWR, and it is even more strongly the case for a B&W PWR design.

Figures D 1 and D-2 will be used to evaluate creep rupture phenomena in Section 4.7.

D.3 Vessel Failure Mode with Depressurized RCS Depressurization of the RCS leads to a significant reduction in heat transport from the reactor vessel to the hot legs and steam generator. Therefore, the probability of hot leg failure would be significantly reduced. Since hot leg failure would be a desirable RCS failure mode and since at low pressure it would be less likely than vessel melt through, all sequences with the RCS depressurized are modelcd to proceed to vessel melt-through.

f l

e 0

1 s

e I

4 549

References for Appendix D D-1 Reactor Risk Reference Document" NUREG.1150, Volume 3, Appendix J, February 1987.

D 2_

" Risk Management Actions to Assure Containment Effectiveness at Seabrook Station",

~

PLG 0550, July 1987, 1'

4 4

1 4

4 a

550-

' A.

i j.

i u

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Figure D-1: Average wall temperature versus rupture time for hot leg nozzle (A-508, Class'2 carbon 7

L steel) (From Reference D-1) 4 1300 _

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Note' Phase change occurs at about 1005 K (1350 F). Calculated rupture times above this temperatu,e (indicated by dashed lines) may not be valid.

l 551 s

=..

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Figure D-2: Average wall temeprature versus rupture time for steam generator tube (Inconel 600) (Frcm Reference D-1) 1600 _

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i i iiiii ReI-Inconel 600 Technical Bulletin International Nickel Co. Inc 5

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1 i.

APPENDIX E i

i STRESS-STRENGTH INTERFERENCE ANALYSIS OF i

CONTAINMENT FAILURE PROBABILITY i

I i

l i

2 l

553

.. -. ~

~_

Containment failure probabilities are calculated using a stress strength interference method.

His method is illustrated in this appendix, using fictitious split fraction FSX as an' example. De

' uncertainty distribution for the containment pressure due to a direct containment heating event with the default hydrogen burn model from Appendix F is used to determine the split fraction value of FSX. De pressure load from a DCH event with the default hydrogen burn model (DEFHB) from Appendix F is used for the example. It is characterized by a median pressure of 82 psia and a 95 percentile pressure of 106 psia. His is interpreted as a lognormal distribution, which yields a lognormal standard deviation of 0.156. This distribution, which defines the containment load, is shown as the curve labelled " Load DEFHB"in Figure E 1.

De curve labelled " Capacity" in Figure E 1 is the composite (total) containment failure distribution curve in Figure 4.4-1. His curve defines the containment strength.

He probability of containment failure is defined as the probability that the containment load exceeds the containment strength. Since the uncertainty distribution for the containment load is independent of the distribution for the containment strength, the containment failure probability is defined by the stress-strength interference integral:

(

Pr (Containment Failure) =, " Pr(P, - p)

[ Pr(P, - p' jdp' dp where: P, is the peak containment pressure (load), and P,is the containment failure pressure (capacity).

De expression inside the brackets is the cumulative composite probability distribution for total containment failure CPr (P, < p), which is shown in Figure 4.4-2.

Performing the integration for the two curves in Figure E-1 yields the probability of containment failure, FSX = 0.018, which is given in Appendix F.

554

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APPENDIX F DIRECT CONTAINMENT HEATING AND 4

d CONTAINMENT INTEGRITY AT RCS FAILURE l

i t

J 1

d y

4 556 i

=-

F,1

Background

e

~ High pressure accident sequences can involve a series of physical prme<<ec at the time of vessel breach that may affect the magnitude of the containment pressure rise. In a large dry PWR, the blowdown of the primary system gases at high pressure and temperature can cause the containment pressure to increase to about 60 to 80 psia, without the action of any other mechanisms that can l

increase containment pressure. His pressure level is not sufficient to challenge the integrity of the containment. However, the integrity of the containment at the time of vessel breach can be

' challenged by a number of other physical prue<<en which may or may not occur simultaneously with the blowdown pressure rise and which may result in an incremental pressure increase whose magnitude can not be predicted with a high degree of certainty. Dree physical processes have been identified that can increase the pressure rise at vessel breach. These are:

l l

a)

A hydrogen burn b)

Direct containment heating c)

Rapid steam generation from debris quenching I

I De effect of hydrogen burns on *.he containment pressure is discussed in Appendix C. De effect of debris quenching in the containment is an integral part of the accident progression analysis, and therefore does not need to be addressed separately.

l Direct Containment Heating (DCH) postulates that the core debris is forcefully ejected from the reactor vessel due to the high RCS pressure and that the debris is finely fragmented allowing a significant portion of the sensible heat in the debris to be transferred directly to the containment atmosphere. It is further postulated that certain metals in the hot debris can be oxidized by the 3-

)

j.

steam and oxygen in the air releasing additional exothermic reaction energy, generating additional hydrogen and rendering certain fission products, which are more volatile in the oxidized form, airborne. Furthermore, the hot debris particles can act as a distributed ignition source, causing the hydrogen in the containment atmosphere to recombine at any concentration.

Conclusive results for assessing the impact of direct containment heating, which would generically resolve the issue, have not been published to date.' De Industry Degraded Core Rulemaking j

(IDCOR) program assessment of direct containment heating has concluded that there is no significant potential for direct transfer of debris sensible heat to the containment atmosphere. This is principally

$57

't i

i

because debris transport from the reactor cavity to the lower compartment is not believed to involve _

-l

)

a fragmentation and entrainment mechanism which would yield small diameter particles.

c t

l De most recent information from the NRC-sponsored research on DCH is contained in References b

F 1 and F 2. A semi mechanistic, dynamic model for direct heat transfer between the debris and the gas phase was incorporated into the CONTAIN code. His code is known as CONTAIN DCH, Version 1.0. In this model, some of the less well understood phenomena are treated parametrically.

j I

Reference F-1 includes some 50 separate sensitivity calculations, for a five node model of the Surry _

l l

i containment, which are documented in Tables 3.3 and 3.4 of Reference F-1. Dese results show a L

j significant reduction in the maximum containment pressure compared to the early estimates, which 7

l were based on ' nonmechanistic equilibrium calculations. De maximum containment pressure calculated in Reference F 1 is 178 psia, whereas the base case yielded a containment pressure of 102 i

psia.

l I

More recent calculations were performed in support of the revision to the NUREG-1150 analyses.

nese calculations also used the CONTAIN code, but with an 18-node containment model. Dese l

results were presented to the expert panels, and they are to be published as part of the

{

NUREG/CR 4551 reports which evaluate severe accident risks on potential risk reductions at several l

representative plants. Extensive comparisons and additionalinsights are not available at this time, except that the more detailed containment models result in a further decrease in the peak containment pressure by typically 5 to 15 psi.

I In a more recent report from the Brookhaven National Laboratory (Reference F 2), results of DCH calculations also using Version 1.0 of the CONTAIN DCH code for the Zion PWR are presented.

Dese analyses showed somewhat higher peak pressures compared to the results of Reference F 1.

Both sets of results are analyzed in the following sections. In the absence of plant-specific DCH 4

calculations for CR-3, the results of References F 1 and F 2 will be interpreted for CR 3 and applied.

a j

F.2 DCH Analysis De approach used to define the magnitude and uncertainties in the containment pressure at CR-3 after a high pressure vessel melt-through was to consider the most recent information publicly available from other sources and to interpret this information for the CR-3 containment t

5 558 1

F configuration. Equal weight (probability) was assigned to the calculations and to the cdaptation of -

the results from the NRC-sponsored research in References F-1 and F 2.

-He physical parameters which can be expected to have the most significant influence on the containment pressure after vessel breach are the free containment volume, the core mass and the

- RCS volume, whereby a large containment volume, a small core mass and a small RCS volume would be expected to reduce the DCH pressure peak. The Surry, Zion and CR-3 values for these parameters are listed below. It is seen that the values for CR-3 are intermediate to those for Surry -

and Zion. Therefore, at first glance, the DCH behavior at CR 3 could also be expected to be bracketed by the calculations for Surry and Zion.

Parameter Surry Zion CR-3 Containment Free Volume (Mcft) 1.8 2.6 2.0 Core UO2 Mass (1000 kg) 80 100 93 Core Zr Mass (1000 kg) 16.5 21.6 19.1 RCS Volume (Keft) 9.8 13.7 11.5 Both references consider cases where the CONTAIN default hydrogen burn model is used and cases where the unconditional hydrogen burn model is used. The default model requires that the ignition criteria with respect to the concentration of hydrogen, oxygen and inerting diluents like steam are satisfied before a hydrogen burn can occur. This hydrogen burn model is appropriate if the containment atmosphere temperature is below the auto-ignition temperature for hydrogen.

Hypergolic recombination, i.e., spontaneous recombination without regard to the concentrations of hydrogen, oxygen, and steam, can occur at temperatures above about 1,000 *F at low steam concentrations. The auto-ignition temperature increases to about 2,000 'F at very high steam concentrations. Due to the direct heating of the containment atmosphere by the debris in a DCH event, the temperature threshold for auto-ignition is readily exceeded for at least a portion of the containment atmosphere.

The unconditional hydrogen burn model assumes that there are no ignition limits, and that all the hydrogen is recombined, limited only by the availability of oxygen.

559

q Table F 1 lists all the ' full' debris". cases with the unconditional hydrogen burn model from References F 1 and F 2. Dere are 19 such cases for Surry and 14 cases for Zion. These 33 cases

.j were considered to establish a probability distribution for the peak containment pressure at CR-3 I

given a DCH event with an unconditional hydrogen burn. The columns from left to right identify l

the case number, the case identifier, the case probability weight, the calculated peak pressure in bar and psia, a case correction, the corrected peak pressure applied to CR-3 in psia, and the fraction of j

the core and other debris structures allowed to participate in the DCH event. He Surry cases from NUREG/CR-48% (Reference F-1) are identified as Surry Tx.y #z,where x.y is the table number and z is the case number from Reference F 1. The discrete probabilities are distributed to all the cases such that all the Surry cases add up to 0.5. The case correction is based on information contained in the Referen'ces. The results in NUREG/CR-4896 for Surry were adapted to CR-3 using the l

following corrections discussed in Reference F-1.

i Change in Peak Cause Pressure Rise i

E - Error Fix

-0.2 bar I

Modak Emissivities

-0.2 bar i

5 Node to 18-Node CONTAIN Model (Surry)

-0.5 bar 7 Node to 18 Node CONTAIN Model (Zion, estimate)

-0.4 bar he top two corrections were identified in Reference F-1. Rese corrections were applied to the appropriate cases in Tables 3.3 and 3.4 of Reference F 1. Care was taken not to double count corrected cases. De last correction has been applied to all cases. His resulted in the 19 corrected l

full debris cases with an unconditional hydrogen burn listed in the top portion of Table F 1. In these cases, the maximum amount of debris considered is ejected and participates in the DCH process. In the Surry calculations, this maximum amount of debris is 75% of a fully melted core and lower internals. De containment pressures calculated for these cases ranges from 119 psia to 174 psia.

Based on Reference F 1, Case 4 was identified as the base case, and it was assigned twice the weight

- of all the other cases. The total probability of all the Surry cases in Table F-1 is 0.5.

l s

1 560.

I

Fourteen Zion cases with an unconditional hydrogen burn and with the maximum amount of debris is involved in the DCH process are documented in NUREG/CR-5282. Dese cases are listed in the second block in Table F 1 (cases 20 to 33). They are identified as Zion A-#, where the # designates the case number from Reference F 2. In all the Zion cases,100% of the core and lower internals is assumed to participate in the DCH process. Only the last correction listed above was applied to the Zion cases, since the first two corrections had already been applied to all case calculation & The estimated correction for the limitations of a 7-Node CONTAIN model was based on the corresponding correction for Surry, because no Zion calculations with a detailed CONTAIN noding scheme were documented. De containment pressures for these Zion cases range from 136 psia to 207 psia. Based on Reference F-2, case 20 was identified as the base case and was assigned twice the weight of all the cnher cases. The total probability of all the Zion cases in Table F-1 is 0.5.

De amount of debris participation in the DCH process is one of the largest uncertainties regarding the effects of DCH. In each reference, some calculations were performed with only a fraction of the maximum debris amount participating in the DCH process. The results of these cases are shown in Figure F-1. The upper three curves represent unconditional hydrogen burn (UCHB) cases, and the lowest curve applies to default hydrogen burn (DEFHB) cases. Also marked are the only two Zion

{

cases for default hydrogen burns at 100% debris involvement. It is seen that when plotted against the debris mass fraction participating in the DCH process, the differences between the Suny calculation and the Zion calculation for the UCHB cases are not large, and for the default hydrogen burn model, the Surry curve extrapolates to in-between the two Zion cases at 100% debris participation. On the basis of Figure F-1, the behavior for CR 3 is represented as the average between the Surry and the Zion results. This is indicated by the curves labe!!ed CR-3 on Figure F-1.

The CR 3 UCHB curve is incorporated into the first three lines in the top header of Table F-1. The base pressure in the first line is the pressure with zero debris involvement in DCH. The second line represents the points on the CR-3 UCHB curve at 25%, 50%, 75% and 100% debris DCH involvement, as indicated on the third line. De results on the left side of Table F-1 were scaled to these four discrete DCH debris involvement fractions in order to cover more realistically the range of debris involvement given a DCH event. A zero debris fraction is not included, because that would be the "No DCH" case which is treated separately. The scaling was performed on the basis of the CR-3 UCHB curve in Figure F 1, and the results are shown on the right side of Table F-1 as the peak DCH pressure in psia. De resulting probability distribution is shown in Figure F-2 as the curve labelled " Unconditional Burn".

561 l

De exact same procedure is applied to develop the peak DCH pressura for the default hydrogen burn condition. The cases and results are shown in Table F-2. The range of containment pressures calculated by CONTAIN for the default burn cases ranges from 87 psia to 138 psia for the Surry cases and troni 107 psia to 113 psia for the two Zion cases. The scaling to the four discrete DCH

- deFris involvement fractions is based on the curve labelled CR-3-DEFHB in Figure F 1. He resulting probability distribution for the peak DCH pressure is shown in Figure F 2 as the curve labelled " Default Burn". He " Default Burn" curve shows a significantly lower peak pressure than the

" Unconditional Burn" curve, because at the time of vessel breach in a station blackout sequence, the 4

containment is steam inerted, and no hydrogen burn occurs with the " Default Burn" model.

F.3 Containment Failun Due to DCH The probability distribution for the containment pressure after vessel breach given a DCH event is defined by the two curves in Figure F-2.

These curves define the probability that the actual containment pressure following a DCH event at vessel breach in a station blackout accident sequence would be less than indicated by the X-axis pressure value. Rese curves are approximated by a log-normal distribution with the following characteristics:

(

Case Unconditional llydrogen Burn Default Hydrogen Burn Median pressure 116 psia 82 psia 5th percentile 97 psia 72 psia 95th percentile 153 psia 106 psia Range factor 1.26 1.21 i

i The probability that containment failure occurs due to DCH at the time of vessel breach is determined by the probability that the containment failure pressure is lower than the containment I

pressure load as described in' Appendix E.

Convolution of the curves in Figure F 2, with the composite containment failure distributions shown in Figure 4.4-5, according to the stress-strength interference integral yields the probability of containment failure given DCH as follows:

562 j

Case Failure Probability Containment failure given DCH, default hydrogen burn 0.018 Containment failure given DCH, unconditional hydrogen burn 0.252 These split fractions apply to high pressure vessel melt-through conditions, i.e.,

station blackout sequences without depressurization. In the CET quantification, these split fraction values have been used at the appropriate branching points. Rese analyses indicate that for high pressure core melt sequences where neither RCS depressurization nor hot leg creep rupture occurs before vessel breach, the Direct Containment Heating phenomena, if it occurs, would cause the containment to fail with a 25% probability if the containment temperature is sufficiently hot to support autoignition, and with a probability of 1.8% otherwise.

(.-

563

References for Appendix F F1 Williams, D. C., et al.," Containment Loads due to Direct Containment Heating and Associated Hydrogen Behavior: Analysis and Calculations with the CONTAIN Code," Sandia National Laboratory, NUREG/CR-48%,1987, i

F2 Tutu. N. K., et. al., " Estimation of Containment Pressure Loading Due to Direct Containment Heating for the Zion Plant", Brookhaven National Laboratory, NUREG/CR-5282, March 1991.

I i

i

(

564

^

p Table F-1: Containment Pressure Due to Direct Containment Heating with Unconditional Hydrogen Burn Base Pressure, Po (Psaal 84 84 84 84 P-CR3 Model(psi) 133 125.50 113 98.5 Debris Fraction 1

0.75 0.5 0.25 Probabehty 0.05 0.2 0.5 0.25 NUREG-4896 Surry case Corrected Esected Case Discrete NUREG 5282 Zion Correction Pressure Debns Peak DCH Pressure Probabihty (Bar)

(Psia)

(Bar)

(Barl Fraction (Psial 1

Surry T3.3 #8 0.025 9.8 142.1 0.9 8.9 0.75 137.19 129.05 115.48 99.74 2

Surry T3.3 #9 0.025 8.8 127.6 0.9 7.9 0.75 120.07 114.55 105.35 94.67 3

Surry T3.3 #10 0.025 8.4 121.8 0.9 7.5 0.75 113.22 108.75 101.30 92.65 4

Surry T3.3 #11 0.050 9.1 131.95 0.7 8.4 0.75 128.63 121.80 110.41 97.21 5

Surry T3.3 #12 0.025 9.3 134.85 0.9 8.4 0.75 128.63 121.80 110.41 97.21 6

Surry T3.3 #13 0.025 9.6 139.2 0.9 8.7 0.75 133.77 126.15 113.45 98.73 7

Surry T3.3 #15 0.025 9

130.5 0.9 8.1 0.75 123.50 117.45 107.37 95.69 8

Surry T3.3 #17 0.025 12.3 178.35 0.9 11.4 0.75 179.99 165.30 140.81 112.41 9

Surry T3.3 #19 0.025 11.6 168.2 0.9 10.7 0.75 168.01 155.15 133.72 108.86 10 Surry T3.3 #24 0.025 9.5 137.75 0.9 8.6 0.75 132.06 124.70 112.44 98.22 11 Surry T3.3 #25 0.025 9.1 131.95 0.9 8.2 0.75 125.21 118.90 108.39 96.19 12 Surry T3.4 #11 0.025 11.1 160.95 0.5 10.6 0.75 166.30 153.70 132.71 108.35 13 Surry T3.4 #12 0.025 8.9 129.05 0.5 8.4 0.75 128.63 121.80 110.41 97.21 14 Surry T3.4 #13 0.025 9.9 143.55 0.5 9.4 0.75 145.75 136.30 120.55 102.27 15 Surry T3.4 #14 0.025 10.5 152.25 0.5 10 0.75 156.02 145.00 126.63 105.31 16 Surry T3.4 #15 0.025 8.2 118.9 0.5 7.7 0.75 116.65 111.65 103.32 93.66 17 Surry T3.4 #16 0.025 8.7 126.15 0.5 8.2 0.75 125.21 118.90 108.39 96.19 18 Surry T3.4 #17 0.025 10.9 158.05 0.5 10.4 0.75 162.87 150.80 130.68 107.34 19 Surry T3.4 #18 0.025 9.1 131.95 0.5 8.6 0.75 132.06 124.70 112.44 98.22 20 Zson A-1 0.067 9.9 143.55 0.4 9.5 1

137.75 129.52 115.81 99.91 21 Zion A-2 0.033 11.3 163.85 0.4 10.9 1

158.05 146.72 127.83 105.91 22 Zion A-3 0.033 10.4 150.8 0.4 10 1

145.00 135.66 120.10 102.05 23 Zion A-6 0.033 9.7 140.65 0.4 9.3 1

134.85 127.07 114.09 99.05 24 Zion A-7 0.033 9.6 139.2 0.4 9.2 1

133.40 125.84 113.24 98.62 25 Zion A 8 0.033 10.1 146.45 0.4 9.7 1

140.65 131.98 117.53 100.76 26 Zion A-11 0.033 10.1 146.45 O.4 9.7 1

140.65 131.98 117.53 100.76 27 Zion A-13 0.033 14.3 207.35 0.4 13.9 1

201.55 183.56 153.57 118.79 28 Zion A-13e 0.033 10.8 156.6 0.4 10.4 1

150.80 140.58 123.53 103.77 29 Zson A-13b O.033 10.8 156.6 0.4 10.4 1

150.80 140.58 123.53 103.77 30 Zion A-14 0.033 13.7 198.65 0.4 13.3 1

192.85 176.19 148.42 116.21 31 Zion A-16 0.033 9.4 136.3 0.4 9

1 130.50 123.38 111.52 97.76 32 Zion A-20 0.033 12.9 187.05 0.4 12.5 1

181.25 166.36 141.56 112.78 l

33 Zion A-21 0.033 10.7 155.15 0.4 10.3 1

149.35 -

139.35 122.68 103.34 565

r.

Table F-2: Containment Pressure Due to Direct Containment Heating with Default Hydrogen Burn Model Base Presouro Po (Psee) 64.4 64.4 64.4 64.4 P-CR3 Model (pei) 101.8 91.50 81.4 72.7 Debris Fraction 1

0.75 0.5 0.25 Probabehty 0.05 0.2 0.5 0.25 NUREG-4896 Surry case Corrected Ejected Case Discrete NUREG-5282 Zion Correction Pressure Debns Peak DCH Pressure ProbatAty (Barl (Psial (Ber)

(Ber)

Fraction (Psial 34 Surry T3.3 #2 0.025 7.9 114.55 0.9 7

0.75 115.60 101.50 87.67 75.76 35 Surry T3.3 #3 0.025 6.7 97.15 0.9 5.8 0.75 91.59 84.10 76.76 70.43 36 Sorry T3.3 #4 0.025 6.4 92.8 0.9 5.5 0.75 85.58 79.75 74.03 69.10 37 Surry T3.3 #5 0.025 7.1 102.95 0.9 6.2 0.75 99.59 89.90 80.40 72.21 38 Surry T3.3 #6 0.050 7

101.5 0.7 6.3 0.75 101.59 91.35 81.31 72.65 39 Surry T3.3 #14 0.025 6.9 100.05 0.9 6

0.75 95.59 87.00 78.58 71.32 1

40 Sorry T3.3 #16 0.025 9.5 137.75 0.9 8.6 0.75 147.62 124.70 102.23 82.87 41 Surry T3.3 #18 0.025 8.8 127.6 0.9 7.9 0.75 133.61 114.55 95.86 79.76 42 Sorry T3.3 #20 0.025 7.2 104.4 0.9 6.3 0.75 101.59 91.35 81.31 72.65 43 Sorry T3.3 #21 0.025 7.9 114.55 0.9 7

0.75 115.60 101.50 87.67 75.76 44 Sorry T3.3 #22 0.025 7

101.5 0.9 6.1 0.75 97.5v 88.45 79.49 71.77 45 Surry T3.3 #23 0.025 7.4 107.3 0.9 6.5 0.75 105.60 94.25 83.13 73.54 46 Sorry T3.4 #3 0.025 7.2 104.4 0.5 6.7 0.75 109.60 97.15 84.94 74.43 47 Surry T3.4 #4 0.025 7.8 113.1 0.5 7.3 0.75 121.60 105.85 90.40 77.10 48 Surry T3.4 #5 0.025 6

87 0.5 5.5 0.75 85.58 79.75 74.03 69.10 49 Surry T3.4 #6 0.025 7.3 105.85 0.5 6.8 0.75 111.60 98.60 85.85 74.87 50 Surry T3.4 #7 0.025 8.4 121.8 0.5 7.9 0.75 133.61 114.55 95.86 79.76 i

51 Surry T3.4 #8 0.025 7.5 108.75 0.5 7

0.75 115.60 101.50 87.67 75.76 52 Surry T3.4 #9 0.025 8.4 121.8 0.5 7.9 0.75 133.61 114.55 95.86 79.76 53 Zion A-4 0.333 7.8 113.1 0.4 7.4 1

107.30 95.49 83.90 73.92 54 Zion A.5 0.167 7.4 107.3 0.4 7

1 101.50 91.28 81.26 72.63 i

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4 9

APPENDIX G CONTAINMENT FAILURE DUE TO LATE

~

HYDROGEN BURNING 569

i Dis appendix ' documents' the methodology for calculating the probability of containment failure due to late hydrogen burning. In the CPET. this is addressed in top event HD for late hydrogen burning and in top event FD for late containment failure. After vessel breach, the concentration of-I flammable gases increases due to the ex-vessel generation of hydrogen from the reaction of metallic

. elements in the debris and from' the decomposition of calcium carbonate in the concrete which i

releases carbon monoxide (CO).- According to the accident sequence analyses for accident sequences f

with continued basemat penetration, the concentration of carbon monoxide in the containment i

atmosphere steadily increases to levels comparable to the hydrogen concentration. Hydrogen and CO l

are very similar with respect to combustion effects. He molar heat of combustion of CO is about l

15% lower, and the ignition limits and the inerting behavior is similar for the two gases. Herefore,-

l to evaluate the effects of a late global burn CO will be treated as hydrogen on a molar basis, and consequently, hydrogen and CO are of comparable importance for late combustion events.

De assessment of late containment failure due to hydrogen burns is based on the notion that the l

calculations of long term concentrations of hydrogen and CO, and to a lesser degree the steam I

concentrations, are subject to significant uncertainties. High concentrations of hydrogen are possible.

For ' example, the concentration from the reaction of 100% zirconium in a saturated steam

[

containment atmosphere is 15% at 150 *F,12% at 200 *F, and 7.8% at 250 *F. At a temperature of 250 *F, a saturated containment atmosphere contains 57.5% steam, and therefore at higher

]

temperatures, the containment would be steam inerted.

Table G 1 illustrates the method for calculating the split fractions for late containment failure due to a late hydrogen burn. In the first column, the range of combined hydrogen and CO concentrations i

in the containment is discretized for the specific accident condition,in this case represented by split fraction HDA. His condition applies to station blackout sequences where there has been no prior i

i hydrogen burn. Dese sequences are without containment heat removal and the debris is not cooled, ne second column lists the equivalent fraction of zirconium reacted, estimated from the total i

number of moles of H and'CO in the containment atmosphere, at the point in the accident sequence 2

under consideration. His fraction can be greater than one due to the generation of CO. De third column lists the estimated probability based on the CONTAIN calculations that the long term E

hydrogen concentration would reach the indicated range in the absence of a hydrogen burn. For any given final concentration level, combustion can occur at the final level or at any of the preceding (lower) levels since the concentration has to build up through these levels to reach the final level

, _.(

. De fourth column lists the probability that the containment atmosphere is not steam inerted. De 570 4

a

i fifth column indicates the probability that combustion does occur at any concentration and the next l

live columns distribute the total burn probability to the lower concentration ranges. De difference between the total burn probability and the sum of the burn probabilities at lower concentrations is then the probability of a burn at the final concentration level. He probability of a hydrogen burn '

in a given concentration range is shown in the next column. It is calculated as the sum of all l

contributions to burns at this concentration level, including those which otherwise would proceed to a higher concentration. De containment temperature, before the burn which is appropriate to the accident condition, is indicated-in the column labelled 'T Before", ne post-burn pressure is determined from Figure 4.71 or Figure D 2 and is listed in the column labelled "P-After". - From t

Figures 4.4 2,4.4-4 or 4.4-6, the probability of containment failure, given that a burn occurs, is listed 5

in the next column. De figure which most closely corresponds to the preburn containment I

temperature listed in the column labelled "T-Before" is used. The second to-last column lists the conditional probability of containment failure due to a hydrogen burn. It is the product of the burn a

~

probability and the containment failure probability given that a burn occurs. The last column finally gives the unconditional probability of a containment failure due to a hydrogen burn.

a De different hydrogen concentration ranges are combined to determine the integral probabilities

(

listed at the bottom of the table. The first of the three rows at the bottom correspond to the split fraction value HDx that a late hydrogen burn occurs. The middle row gives the unconditional probability of a late containment failure due to a burn, and the last row gives the conditional probability of a containment failure as a result of a late burn, given that a burn occurs. In l

Section 4.7,-this methodology is applied for all split fractions for top event HD and for the corresponding split fractions in top event FD.

4 1

)

i i

4 l

1 I

571 1

A 1

Table G-1: Methodology for Computing Containment Failure Due to a Late Burn e

i SPUT FR ACTION:

HDA KPOS: K7D No p6er burn, deb 6s not cooled X(H2 + CO)

Equivet. %

Prob of Prob not PROBABluTY OF IGNITION Probability T-Before P-Atter Conteenment Fedure Conc (%)

Zr-Owidation X(H2 + CO) inerted Total

<4% 4-6%

6-8 % 8-10% 10-12%

of Burn (F)

(PSIA)

CFl Born C Fall

<4-

<52 O

O.3 0

0.132 546 75 0

0 4-6.

52-77 0

0.3 0.4 0.5 0.02895 546 90 0.020 5.8E-04 6-8 77-103 0.1 0.3 0.7 0.5 0.25 0.03195 546 135 0.523 1.7E-02 8-10 103-129 0.2 0.3 0.8 0.5 0.15 0.15 0.03585 546 174 0.953 3.4E 42 10-12 129-155 0.5 0.3 0.9 0.5 0.1 0.1 0.15 0.02625 546 197 0.996 2.6E42

>12

>155 0.2 0.3 1

0.5 ' O 05 0.1 0.1 0.1 0.009 546 220 1

9.0F-03 PROGABluTY OF LATE BURN =

0.2640

= HDA L

PROBABluTY OF CONTAINMENT FAILURE =

0.0866 PROBABillTY OF CONTAINMENT FAILURE I L ATE BURN =

0.3280 1

572 m.----

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+.

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APPENDIX H I

4 CONTAINMENT PHENOMENOLOGICAL EVENT

., (

TREE QUANTIFICATION d

i I

i 4

W 573 i

His appendix contains the details of the conditional CPET quantification in the form of a fully expanded CPET where the frequency at each branching point and at the end of each sequence is given. The CPETs are listed in the following order:

Figure H 1: CPET Quantification for KPDS K7D Figure H.2: CPET Quantification for KPDS K71H Figure H-3: CPET Quantification for KPDS K6BA Figure H 4:' CPET Quantification for KPDS K4K Figure H-5: CPET Quantification for KPDS K3BA The numbers listed on these figures are the cumulative conditional frequency of the sequence from the initiating event (KPDS) up to the point in the sequence where the frequency is given. He frequency of each entire sequence is listed in the column for top event BM. The quantification of each CPET is conditional, i.e., with a KPDS frequency of 1.0. In order to obtain the absolute frequency of each sequence, the conditional frequency for the sequence given in the column "BM" must be multiplied by the frequency of the KPDS.

574

~

~

Figure H-1: CR3 Containment Event Tree (CPET) for KPDS K7D Page 1of5 l -POS ' HL i 00 i HP i FP l DT i HS FS ' OC CS 6 HD I CA I FD i BM No. I R O.

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Page 2 of 5 j

Figure H.1

CR3 Containment-Event Tree (CPET) for KPDS K7D.

J

.\\NTtFICATION FOR XPDSr K70 l ' SOS I' HL DQ l HP

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Figure H 1: CR3 Containment Event Tree (CPET) for KPDS K7D Page 3 of 5 tNTWIC ATiON FOR KPDS:

K7D HS i FS DC 1

CS 1 HD i CA ! FD BM No. 1 R C.

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)

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l

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Figure H-1E CR3. Containment 5 dent Tree (CPET) for KPDS K7D

.Page 5 of 5

\\NTIFIC ATION FOR KPDS-K70' DT l HS FS 1

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FD i BM No-R.C.

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Figure H-2: CR3 Containment Event Tree (CPET) for KPDS K73H Page i of 5 1

UJTIFIC ATIOi4 FD*l KFDS:

K7JH

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Figure H.2: CR3 Containment Event Tree (CPET) for KPDS K7JH Page 2 of 5

\\NTtDC A TION FOR KPC$t K7)H

~ l DC 1 CS I

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CA 1

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Figure H-2: CR3 Containment Event Tree (CPET) for KPDS K7JH Page 3 of 5

\\NTIFIC ATION FO.7 KTOS-K7)H l n9CS I HL i DQ, HP FP DT HS FS OC CS HD i CA FD i BM No. I R.C.

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Figure H-2: CR3 Containment Event Tree (CPET) for KPDS K7JH Page 4 of 5 tNTIAC ATION FOR KPOS:

K7)H l =POS HL I

DQ l HP 1 FP ! DT i a5 6

FS I 00 CS CA i FD BM No. I R.C.

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Figure H-2: CR3 Containment Event Tree (CPET) for KPDS K7JH Page 5 of 5 SNTIFIC ATION FOR KPOS:

K7)H l JOS i HL I DO ! HP FP i OT ! HS 1

FS DC CS HD CA 6

FD l BM No. I R.C.

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Figura H-3: CR3 Containment Event Tree (CPET) for KPDS K6BA Page 1 of 5

%NTIFIC ATION FOR KPDS-K63A l /05 i HL 00 l HP i FP i DT i 'HS FS i DC CS 1

HD 1 CA FD 4 BM No-R.C-150E 1 k l

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Figure H-3: CR3 Containment Event. Tree (CPET) for KPDS K6BA. _.

Page 2 of 5 4

4 2

\\NTIFIC ATION FOR KPOS:

K68A FS I DC CS

(

HQ i CA I FD BM No-R.C. '

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__ i Figure H 38 CR3 Containment Event Tree (CPET) for KPDS K6BA Page 3 of 5

, tNTIFIC ATION FOR KPOS:

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HD i CA !

FD i BM No-l R.C.

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Figure H-3 L CR3 Containment Event Tree (CPET) for KPDS K6BA Page 5 of 5

' \\NTWIC ATION FOR KPOS:

K68A l =PDS i HL i - DQ l HP I FP i DT 1 HS I FS DC CS MD i CA I ~FD i

SM-No.

R.C.

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Figure H-4: CR3 Containment Event Tree (CPET) for KPDS K4K Page 1 of 5 l

tNTIFIC ATION FOR KPOS:

K4K l.s*DS i HL i 00 l HP l FP i OT i HS 1 FS CC i CS ! HO i CA l FD ! BM No. I R.C.

l 1 sof.1 l l

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Figure H-4 CR3 Containment Event Tree (CPET) for KPDS K4K Page 2 of 5

\\NTIFICATION FOR DOS?

K4K l, DOS I HL i DO 4 HP i FP i DT HS i FS CC CS i HD f CA i FD l BM l No. j R.C.

81 ALAQ l

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Figure H-4: CR3 Containment Event Tree (CPET) for KPDS K4K ~

Page 3 of 5 SNTIFIC ATION F021 KPOS:

K4K CC ! CS I HD l CA I FD l BM No. i R.C.

l IPOS t HL I DQ ' HP I

FP CT i HS I FS I

162 XDMS i

163 x6I5~

8 164 X6I'8~'

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i

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177~"~x0A0 l

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i i80 x6%D~

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7 91._.T~x,D AU X

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l Figure H-4: 'CR3 Containment Event Tree (CPET) for KPDS K4K Page 4 of 5 l

UJTIFIC ATION FOR KPCS:

K4K l aPOS i HL i DQ i MP i FP 1 DT I HS f FS f

CC CS HD CA i FD i BM No. I R.C.

I fil~i~~D A Sks I

i 243 A

6 37E 3 i

+

l 245~I XDMU l

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2E1,}~i6EU~

c fl0 t XN J

6 37E 3 5 74E 3 e i

l ffE*I~X6MU~

5 74E 3 6 5 74E 3 250 XDAV 6 37E 4

' 6 37E 4 2$i~~E6iU~

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6 37E 7 e

i 263 Xt55~

l 6 37E 7 i

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t 265~I~XDAQj I

i i

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8 f58 XN~

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ffe X6IU~

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Figure H-4:1CR3 Containment Event Tree (CPET) for KPDS K4K Page5of5 SNilFICATION FOR KPDS' K4K

~-. l. DOS I HL 1 00 i HP i' FP I-DT I HS ; FS i DC I CS l WD i CA i FD I

BM-No.

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l

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Figure H-5: CR3 Containment Event Tree (CPET) for KPDS K2SA Page 1 of 5 tNTirtC ATION FOR KPOS:

K3BA J

l./DS I HL I

DQ l HP FP t

DT HS FS l DC I CS HD CA 1 FD l BM No. I R.C.

i I

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2 53E 5 IT~I XOMS~~

i 5 06E 9 5 06E 9 12 XOAS X5'~~'

f XOAS 5 62E-6 562E6 13 5 57E 3 5 OtE 3 i 4 78E 3 i 2 39E-3 14 2 39E 3 1$~~I~I6M5~'

2 31E4 2 31E 4 18 XD $~

f 5 57E 4 5 57E-4 11 x6I$'~

l 6 25E 4 6 25E-6 ~5 62E-6 i 5 62E 6 281 E 6 18 XN

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3 75E 2 3 67E 2 184E-2 i 184E 2 i 8 64E 3 7 77E 3

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6 77E-3 fi~~I~f6M5~

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Figure H-5: CR3 Containment Event Tree (CPET) for KPDS K3B Page 3 of 5 tNTIACATION FOR KPOS:

K3BA 1 do$ I HL. I DQ l HP 1 FP I DT 4 HS I FS i DC !

CS 1 HD I CA I FD l

BM No. I R.C.

I 3 7tE 2

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Appendix 1: Disposition of Key Event Tree Considerations This appendix describes how each of key phenomenologicalissues defined in Tables 2.2 and A.5 of NUREG-1335 were dealt with in the CR3 IPE Level 2 model. Table 2.2 defines potential containment failure modes and mechanisms, whereas Table A.5 addresses parameters that are considered to have significant phenomenological uncertainties.

l.1 Containment Failure Modes and Mechanisms This section addresses the containment failure mode sensitivities identified in Table 2.2 of NUREG-1335.

l.1.1 Direct Containment Bypass Bypass containment failure modes were considered as separate initiating events in the Level 1 e

model and assigned to a specific group of bypass PDS identified by the last three columns of the PDS Matrix shown in Table 4.3-3 of the submittal.

l.1.2 Failure to isolate Containment isolation failures were explicitly considered in the Level 1 model. Sequences with small penetrations isolation failures (< 3 inches in diameter) were assigned to the small containment failure group identified by the letters "E, F, G or H" as the second character in the PDS designator. Sequences with large penetrations isolation failures (> 3 inches in diameter) were assigned to the large containment failure group identified by the letter "I or J" as the second character in the PDS designator. These two containment isolation failure groups were separately considered in the Level 2 analysis.

l.1.3 Vapor Explosions Vapor explosions which cause the reactor vessel head to become a missile that penetrates and fails the containment were considered by the NRC's steam explosion review group. The probability of this event occurring when the molten debris slumps to the bottom of the reactor vessel was concluded to be very low, on the order of 1E-4. If this event were postulated to occur at this probability, the frequency of the resulting early containment failure would be a factor of 1000 lower than the frequency of the dominant early containment failure sequence. Therefore, the CR3 IPE Level 2 results are not sensitive to assumptions about the probability of alpha-mode (steam explosion) containment failures, unless this probability is postulated to be on the order of 0.1 or larger.

3 1.1.4 Overpressurization Containment failure due to overpressurization both due to steam and non-condensable gases

~

is explicitly addressed in three distinct top events in the CPET, namely before vessel breach (FP), at vessel breach (FS) and late in the accident sequence (FD).

i i

1.1.5 Combustion Processes i

Hydrogen bums in the containment are explicitly considered in three distinct top events in the l

.CPET, namely before vessel breach (HP), at vessel breach (HS) and late in the accident i

sequence (HD).

i-1.1.6 Core Concrete interaction Containment failure due to basemat penetration from core concrete interaction is explicitly considered in CPET top event BM for late basemat meltthrough.

i j

'l.1.7 Blowdown Forces 5

-i l

Vessel thrust forces during a high pressure vessel meltthrough exert forces on the reactor vesselin the upward direction. If a vessel meltthrough failure is postulated to occur away from j

the vessel bottom area, and if it is postulated to propagate around the vessel circumference, then larger forces can be exerted on the vessel in the upward direction, up to a maximum force j

bounded by the vessel pressure multiplied by the inside cross-sectional area of the vessel.

j This maximum cross sectional area failure could be outside the vessel support skirt and the force would be resisted by the c.ot legs and cold legs penetrating through the biological shield.

l Such a failure mode can be postulated to occur due to a creep rupture of the vessel L

circumference at the top of the debris pool. If this failure mode is postulated to occur before vessel depressurization occurs, then an early containment failure would be possible, either by i

the vesselimpacting on the missile shield and containment dome.

in vessel designs with bottom instrument penetrations, like the CR3 design, vessel failure is

~

expected to occur at one or several of the instrument penetrations inside the vessel support f

skirt, and not by this circumferential failure mode. Therefore the circumferential failure mode i

was not considered.

i f.

1.1.8 Liner Meltthrough l

Liner meltthrough is a BWR issue, and it was not addressed in the CR3 IPE analysis.

l 2

i 1.1.9 Thermal Attack of Penetrations Local or leak before break failure modes, including thermal attack of penetrations has been analyzed in several other IPEs and PRAs for large dry PWRs. They generally show local failure modes to occur at 10 to 20 psi lower than the first gross failure mode, 4

j l

i and generally in the vicinity of 130 to 140 psid. Since the lowest structural containment failure mode for the CR3 containment was identified at a pressure of 140 psi (at 300*F) to 122 psig (at 800*F), it was conservatively assumed that all containment failures l

would occur by the leading structural failure mode at the basemat to cylindrical wall i

juncture. For this reason local or leak before break failure modes were not analyzed.

[

1.2 Accident Phenomenoloav and Parameter Sensitivity This section addresses the phenomenological and ;,arameter sensitivities identified in j

4 Table A.5 of NUREG-1335.

l 1.2.1 Containment Heat Removal The availability of containment heat removal is explicitly considered in the Containment Systems Event Tree (CSET) analysis. Four containment heat removal states are considered (containment heat removal and fission product scrubbing, containment heat removal only, containment fission product scrubbing only and no heat removal and scrubbing). Separate l

PDS columns were defined for each case and sequences are assigned to separate PDS bins for each containment heat removal state.

i 1.2.2 In-Vessel Phenomena - high vessel pressure Hydrogen production and combustion: The MARCH 3/CONTAIN calculation results were l

a.

considered as best estimates and uncertainties were explicitly considered in the CPET quantification of the three hydrogen bum top events.

g U-b.

Induced RCS failure: The probability of an induced hot leg creep rupture after core uncovery is addressed in CPET top event HL, but it was not found to be numerically significant for the OSTG design.

c.

Core relocation characteristics: These are considered in CR3 specific MARCH 3 model.

j However, no sensitivity analyses for different blockage assumptions were performed due

{

to the low importance of the hot leg and steam generator tube creep rupture issue.

I d.

Mode of reactor vessel failure: Downward vessel meltthrough at one or more instrument penetrations inside the vessel. support skirt is considered the expected failure mode for bottom instrument tube designs. Therefore, other failure modes were not considered to i

be significant.

I 1.2.3 In-vessel phenomena - low vessel pressure l

a.

Hydrogen production and combustion: The MARCH 3/CONTAIN calculation results were considered as best estimates and uncertainties were explicitly considered in the CPET i

quantification of the three hydrogen bum top events.

f i

____,.___.____,______________________,________._.______._______.______,______.__I

,_m,,

a

i l

b.

Core relocation characteristics: Same as item 1.2.2 (c) above.

c.

Fuel coolant interactions: The effect and sensitivity to in-vessel fuel coolant interactions is addressed in Sechon 1.1.3.

d.

Mode of reactor vessel failure: Same as item I.2.2 (d) above.

l.2.4 Ex-vessel phenomena - high vessel pressure Direct containment heating: The potential and the effects of direct containment heating a.

are explicitly considered in CPET top event DT, and the impact of DCH on containment l

integrity is explicitly considered in the quantification of top event FS.

b.

Eariy containment failure due to pressure loads: The probability of early containment failure due to the pressure transient resulting from an early hydrogen burn is explicitly addressed in top event FP. The probability of early containment failure due to the pressure transient resulting from a high pressure vessel meltthrough is explicitly addressed and modeled in CPET top event FP, and it considers the effects of blowdown forces, hydrogen bums and DCH at the time of vessel failure.

c.

Debris disposition and debris coolability: The CR3 IPE Level 2 analysis considered the debris disposition depending on the RCS pressure at vessel meltthrough. For high pressure sequences dispersal through overpressure failure of the two cavity access crawl tunnel doors and for low pressure sequences meltthrough of the doors by direct exposure to core debris and debris flow from the cavity through the cavity access crawl tunnel into the lower compartment was considered. In each case a debris bed of at least one foot in depth remaining in the reactor cavity was considered.

Debris 2

cootability, and the effects of the resulting core concrete interaction on containment failure and on the source term were explicitly modeled and assessed depending on the resulting debris configuration.

l.2.5 Ex-vessel phenomena - low vessel pressure t

l'ar1y containment failure due to liner meltthrough: Liner meltthrough after vessel a.

breach is a BWR issue, and it was not explicitly addressed.

b.

Long-term core-concrete interactions:

Core concrete interactions, non-condensable and flammable gas generation, and concrete penetration was explicitly considered in J

the CONTAIN model and calculations. These effects and the uncertainties associated with them on the time of containment failure and on the containment failure mode were explicitly considered in the quantification of split fractions for CPET top events DC, HD, FD and BM.

i 1

a 4

_, _ _ _ _ _ _ _, _ _ _ _ _ _ _, _ _ _ _ _ _, _