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Failure mechanisms of LWR steel containment buildings subject to severe accident loadings
Nuclear Engineering and Design 104 (1987) 371-379 North-Holland, Amsterdam
371
FAILURE MECHANISMS OF LWR STEEL CONTAINMENT BUILDINGS SUBJECT TO SEVERE ACCIDENT LOADINGS * D a v i d B. C L A U S S Containment Integrity Division 6442, Sandia National Laboratories, Albuquerque, N e w Mexico 87185, USA
Received 27 April 1987
The potential failure mechanisms in LWR steel containment buildings subject to quasi-static pressurization and elevated temperature are identified. For each failure mechanism, the relevant structural response measures are discussed. For mechanisms involving leakage, the importance of seal performance is also discussed. Criteria that can be used to evaluate threshold environments are presented for several failure mechanisms. Results of tests on scale models and seal tests that support the criteria are referenced.
1.
Introduction
During a severe (degraded core) accident, containment buildings may be subject to loads that significantly exceed their design basis loads. The performance of the containment building, which represents the last engineered barrier to release of radioactive material, affects the consequences of a severe accident. A partial list of the containment performance parameters that affect severe accident consequences follows: - Capacity: A r e the severe accident loads sufficient to cause a failure, i.e., a significant release of radioactive material? - Timing: In general, risk from an early failure is greater than that from a late failure, because early in an accident sequence radioactive materials are more likely to be airborne. - Location o f failure: Is the release direct to the environment or filtered through other structures, for example, the reactor auxiliary building? - L e a k a g e rate o f "'hole size": Is the leakage rate sufficient to preclude further pressurization of the containment building? Consequences are proportional to leakage rate at least over some intermediate range of values. Rupture could result in resuspension of radioactive aerosols. * This work supported by the US Nuclear Regulatory Commission and performed at Sandia National Laboratories, which is operated by the US Department of Energy under contract number DE-AC04-76DP00789.
It should be evident that the predicted failure mechanism will have a significant effect on risk estimates. In fact, failure mechanisms are containment specific as well as scenario dependent - for instance, for a given containment, the failure mode due to dynamic loading could be much different than that due to quasi-static pressurization. This paper will focus on the potential failure mechanisms in a steel LWR containment building subject to quasi-static pressurization and elevated temperature. The failure mechanisms considered in this paper are: Failure of the containment shell; Leakage past sealing surfaces (operable penetrations); Equipment Hatch sleeve ovalization; Pressure unseating equipment hatches and drywell heads; Personnel airlocks; - Leakage from electrical penetration assemblies; - Leakage through purge and vent valves; - Leakage due to failure of a bellows (expansion joint). For each potential failure mechanism, the important structural response measures and seal performance parameters are discussed. Preliminary criteria are presented for evaluating rupture of the shell and leakage from equipment hatches and drywell heads. The discussion is limited to the range of response wherein the containment building behaves like a freestanding steel shell. Most steel containment buildings are enclosed by a concrete shield wall. If large membrane plastic strains are developed in the containment shell before failure occurs, the containment and shield
0 0 2 9 - 5 4 9 3 / 8 7 / $ 0 3 . 5 0 © Elsevier Science Publishers B.V. ( N o r t h - H o l l a n d Physics P u b l i s h i n g Division)
372
D.B. Clauss / Failure mechanisms of L W R steel containment buildings
wall will come into contact. In BWR Mk I containment buildings, the gap between the steel shell and the shield wall corresponds to approximately 0.5% membrane strain; in other types of plants with freestanding steel containments the gap typically corresponds to 5% membrane strain. If contact occurs, the performance of the containment building will be affected. Different failure modes may come into play, or the shield wall may strengthen the containment building, or both. The effect of contact with concrete shield walls on the failure mechanisms of LWR containments are not considered in this paper.
2. Failure of the containment shell In this paper, the containment shell is defined as the pressure boundary of the containment, exclusive of those parts of the boundary made up of penetrations, including equipment hatches, personnel airlocks, piping, and electrical penetrations assemblies. A loss of integrity associated with the containment shell must involve a cohesive material failure, i.e., a through-wall crack or tear in the pressure boundary. If a through-wall crack or tear is developed, it is highly likely that rupture (unstable crack growth) will rapidly ensue. Based on results of tests on scale models of steel containment buildings that were pressurized to failure [1,2], a simple strain criterion appears to be adequate for estimating the point at which a through-wall crack or tear will develop. The steel most commonly used in US containments, A516 Gr 70, has exceptional ductility, and very large plastic strains may be developed prior to failure. The following criterion for evaluating rupture is proposed: • Rupture will occur if, at any point on the pressure boundary the equivalent strain exceeds the material's ultimate strain (the strain at maximum load as determined from a uniaxial tensile test). If the strain is due primarily to bending, somewhat higher strain may be allowable (to account for crack growth through the thickness). The above criterion should be used only with detailed three-dimensional finite element analyses, which would need to include at least major penetrations and stiffeners. Results can be very sensitive to design details such as stiffener patterns around penetrations [2,3]. Any failure criterion must be consistent with the analysis used to calculate the response measures used in the criterion. If axisymmetric analyses are used to predict rupture, the limiting value of strain used to predict
rupture must be decreased to account for the effects of penetrations and stiffeners. Many of the discontinuities in the containment shell, such as the springline and the base embedment, result in self-limiting states of stress, and consequently, a cohesive material failure is not likely to initiate in these areas. On the other hand, large penetrations cause local strain concentrations in the shell that reduce the capacity relative to an undisturbed shell. Stiffeners tend to increase the general yield pressure but they do not necessarily increase the ultimate strength by a like amount. There is a large body of evidence that indicates that if a through-wall crack or tear is developed a steel containment shell will rupture catastrophically. Kiefner [4] developed empirical equations based on a large number of tests on pressurized cylinders that show that rupture will occur unless the initial through-wall crack is the result of a deep surface flaw (depth of the surface flaw must be at least one-half the wall thickness). Four of the five steel containment models tested at Sandia ruptured; in the one exception, leakage from a through-wall crack was sufficient to depressurize the model. However, in that particular model, the cylindrical wall had been severely thinned by excessive grinding adjacent to a repaired weld, which was equivalent to a deep surface flaw [3]. At least qualitatively, the formulas developed by Kiefner explain why rupture did not occur in that model. A final example is the J-integral analysis of the Sequoyah containment conducted by Griemann [5]. Griemann's analysis showed that a through-wall crack originating in the thinner wall sections near the springline would not be arrested at the thicker wall sections used at lower elevations, and he concluded that unstable crack growth (rupture) would occur.
If a steel containment building is pressurized to the point where the capacity of the shell is the controlling failure mechanism (which may be precluded by leakage from penetrations or by contact with the shield wall at lower pressures), then it is highly likely that a rupture will occur. It is possible, though not likely, that stable, through-wall cracks could develop in the shell, in which case leakage would be proportional to the crack opening area. (Initial surface flaws are assumed to be less than 20% of the shell thickness.)
3. Leakage from operable penetrations Any penetration that is intended to be used for access to the containment is categorized as an operable
373
D.B. Clauss / Failure mechanisms of LWR steel containment buildings
penetration. In US steel containments, this would include equipment hatches, personnel aidocks, escape hatches, and BWR Mk I and Mk II drywell heads. The potential for leakage between the sealing surfaces of these types of penetrations is considered in this section. In general, leakage from operable penetrations depends on both displacement and performance of the seal. Tests have shown that if metal to metal contact is maintained between the sealing surfaces, significant leakage does not occur regardless of the condition of the seal material [6]. On the other hand, because the seals are compressed typically one-quarter inch, large separations of the sealing surfaces can arise without significant leakage if the performance of the seal is not compromised. Relative displacements of the sealing surfaces can be categorized as either separation (relative motion perpendicular to the seating surfaces) or sliding (relative motion in the plane of the surfaces). Tests on seal and gasket materials [61 suggest that there are at least two ways in which performance of the seal can be compromised: (1) radiation or thermal aging can result in a loss of resilency, and (2) high temperatures (500 to 650 o F, depending on the type of material, degrade the seal materials, that is, the materials outgas, then become dry and powdery). In special cases, leakage may be a function of performance of the seal only, or sealing surface deformation only. In personnel airlocks with inflatable seals, the door and bulkhead are not designed to be in intimate contact, and therefore, leakage will arise if the performance of the seal is compromised, irrespective of structural deformations. The performance of the seal may also be the only factor preventing leakage if there is significant out-of-flatness between the seating surfaces. This is a condition that would not be detected by Integrated Leak Rate Tests; at design levels, seal performance prevents leakage. However, in BWR Mk I and Mk II plants, severe accident temperatures are expected to be sufficient to degrade seals and out-of-flatness could result in leakage without additional separation or sliding of the seating surfaces. In cases where sliding or separation of the seating surfaces is very large, leakage can result even if the performance of the seal is not compromised.
DETAIL B
Fig. 1. Details of a pressure seating equipment hatch.
the containment shell. The horizontal diameter of the sleeve increases while the vertical diameter decreases by a like amount; the deformations can become quite large
,,
:A" I
~__{
a"
S~-IE DETAIL '~I~" G~I)VF- ~
•
3.1. Equipment hatch sleeve ovalization
Typical details of equipment hatches used in US steel containments are shown in figs. 1 and 2. In a cylindrical containment shell, an equipment hatch sleeve will deform into an oval shape due to interaction with
~u~l SLEEVE ~ ~0" ~ 9.' "/"
F C~'sKET
~E'rAl"
PRE~IU~ E L E VA,T|ON VIEW
Fig. 2. Details of a pressure unseating equipment hatch.
374
D.B. Clauss / Failure mechanisms of L W R steel containment buildings
(on the order of the sleeve wall thickness) after general yielding of the cylindrical containment shell. The equipment hatch cover and tensioning ring deform axisymmetrically, i.e., they retain their circular shape. Changes in the diameter of the tensioning ring are expected to be quite small (except possibly for pressure seating hatches that are susceptible to buckling). The mismatch caused by the sleeve sliding relative to the tensioning ring is greatest along the horizontal and vertical centerlines; as indicated in fig. 3a, eventually the sleeve and tensioning ring will lose contact at these points, resulting in leakage. Under these conditions, the
performance of the seal material does not affect the leakage potential. The deformation of the equipment hatch sleeve can be related to the average membrane strain in the containment shell in an approximate way. Ovalization of the equipment hatch sleeve is primarily an inextensional (no stretching of the midsurface) deformation, and consequently it is a relatively flexible mode of deformation. Thus, the equipment hatch sleeve basically conforms to the membrane displacements of the containment shell at their intersection. As illustrated in fig. 4, the increase in the horizontal radius of the sleeve, A r H, can be approximated as Ar H = o r ,
(1)
a r- . . . . I
where c is the average membrane circumferential strain in the shell at the elevation of the equipment hatch centerline, and r is the original sleeve radius. Assuming the cover and tensioning ring do not deform significantly, leakage due to ovalization is assured if
..,__._-.--------'-TENSIONING ~ - ~ RING
SLEEVE
~,~%,~
Ar > t,
(2)
where A r is the change in the sleeve radius and t is the sleeve wall thickness at the sealing surface. Combining eqs. (1) and (2), the approximate leakage criterion is c r / t > 1.
(3)
The average membrane strain in the shell at the elevation of an equipment hatch corresponding to leakage from ovalization, ~', can be calculated from eq. (3). For equipment hatches used in steel containments identified in a survey conducted by Argonne National Laboratory
b
[ ". . . .
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(r /
il'! Fig. 3. Leakage path created by ovalization of sleeve: (a) without seal degradation; (b) with seal degradation,
~
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UNDEFORMED SHAPE
R ~ CYLINDER RADIUS r ~ SLEEVE RADIUS I[ ~ "AVERAGE"MEMBRANE CIRCUMFERENTIALSTRAIN
SLEEVE AXIS
Fig. 4. Equipment hatch sleeve and containment shell interaction.
D.B. Clauss / Failure mechanisms of L WR steel containment buildings ( A N L ) [7,8], (' ranged from 2.5% to 7.3%, as indicated in table 1. Clearly, significant plastic membrane strains must be developed in the containment shell before ovalization of an equipment hatch sleeve will cause leakage. U n d e r certain conditions, leakage caused by ovalization may occur at lower values of average membrane strain. If the seal material is severely degraded, the grooves that hold the seals could act as a leakage path, as suggested by fig. 3b. For a pressure seating hatch, the grooves would be exposed to the inside of the containment at 3 and 9 o'clock and to the outside at 6 and 12 o'clock, yielding critical strains ((') from 1.7 to 4.5%, as indicated in table 1. However, in order for this leakage path to be realized, a substantial amount of the seal would probably have to be degraded and ejected from the seal grooves. The approximate method outlined above did prove accurate for the equipment hatches in the l : 8 - s c a l e model [2,3]. In particular, the change in the horizontal diameter calculated from eq. (1) was within 10% of the measured displacement. However, it is unknown how accurate this approximate method would be for different reinforcement designs and stiffener patterns. 3.2. Separation of pressure unseating hatches Schematics of a pressure-unseating equipment hatch and a B W R Mk I drywell head are depicted in figs. 2
375
Table 1. Critical strain for equipment hatch leakage (by ovalization) in steel containments Unit
t (in.)
r
(in.)
(,
a) (%)
(, b) (%)
Seal type
1
514
72
7.3
4.5
gumdrop
2
4½
72
6.3
4.2
gumdrop
3
4~2
84
5.4
3.9
gumdrop
4
4~
84
5.4
3.6
gumdrop
5 6
4~ 4~
126 126
3.6 3.6
2.3 2.3
gasket gasket
7 10
4~: 3
126 120
3.6 2.5
2.3 1.7
gasket gumdrop
~) Assuming seal degradation has not occurred. b) Assuming seal has been severely degraded and ejected.
and 5, respectively. In both cases the bolts are preloaded to maintain leak tightness during internal pressurization of the containment. Typically the bolts are tightened until metal to metal contact is achieved, and then a specified level of torque or load is applied. Internal pressure acts in the opposite direction of the bolt preload and tends to separate the sealing surfaces. The effects of temperature must also be considered; from a heat transfer viewpoint, the bolts, if they are exposed to air, will act as convective fins. Thus, in many
':;'°
I
• ~rr~.~.
Fig. 5. Details of a BWR Mk I Drywell Head.
~
•
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b" 1½"l
DETAIL C
,
,,
[
Fig. 6. Details of a personnel airlock with double dog-ear seals. zation is resisted primarily by bending action, and therefore, rather large out-of-plane deformations may arise for beyond design basis loads. Although the door is pressure seating, separation
may still be possible at certain points. For instance, in a simply supported rectangular plate subject to uniform pressure, significant reaction forces must occur at the comers; if the corners are not held down, the comers
R.
T...
PRESSURE;
I ii
t, ol DOO~R
"'1- I BULIC.HEAD ~LO" ~-I
NI:L ~TAB SEAL
r, LE
g
~," )IA.
2E'--
C)PENIMC¢
HINGE
,
Fig. 7. Details of a personnel airlock with inflatable seals.
DETAIL
377
D.B. Clauss / Failure mechanisms of LWR steel containment buildings
designs, the average temperature of the bolts will be less than that of the compression flanges. Differential thermal expansion (when the bolts expand less than the compression flanges) tends to maintain contact between the sealing surfaces, that is, it acts in the opposite sense of internal pressurization. The following criteria are proposed for evaluating the onset of leakage from pressure unseating hatches and heads: • If the seals are exposed to temperatures high enough to cause degradation (500 to 650°F, depending on the material), leakage will occur when the internal pressure is sufficient to cause separation of the sealing surfaces. • If the seals are not degraded, but at some time into the accident aging causes a loss of resiliency, a subsequent increase in separation will cause leakage. • If the seals are not degraded and do not lose resiliency, leakage will occur when the seal squeeze becomes equal to zero (when the separation is equal to the initial seal compression). Recent tests on silicone O rings suggest that even degraded seals may prevent leakage through small gaps, up to about 0.005 or 0.006 in.. If this is substantiated, the above criteria would have to be modified so that leakage for separations less than this amount is not allowed regardless of the condition of the seal. Assuming linear elastic behavior of the bolts and compression flanges, the internal pressure at which separation begins for pressure unseating hatches and heads, Ps, can be calculated as: Ps = [ ( k b + k f ) F i / / k f + k b ( ' T f - - CTb) L ] / ( 1 + 2 V k b / k f ) ~ r r 2.
(4)
The subscripts f and b denote the compression flanges and bolts, respectively. F i is the bolt preload, k is the axial stiffness, c T is the thermal strain, L is the bolt grip, v is Poisson's ratio, and r is the average radius of the equipment hatch sleeve or drywell head. Eq. (4) was determined by introducing terms representing the axial thermal expansion of the bolts and the compression flanges and the axial deflection of the compression flanges due to hoop stress to expressions in [9]. The resulting equation for the force in the compression flanges was set to zero and solved for p,. Additional pressure is carried entirely by the bolts, so that the separation between the sealing surfaces, s, is given by s = ~rr2(p - p s ) / k b
for p >Ps-
(5)
The separation pressure for drywell heads and pressure
unseating equipment hatches used in steel containments that were identified in the A N L survey of containment penetrations [7,8] is tabulated in table 2. The mean temperature of the compression flanges was taken to be 700°F and that in the bolts was taken to be 550°F (these temperatures are based on heat transfer calculations made for the drywell head region of a given containment). The results are very sensitive to the bolt preload, which is difficult to know accurately. Generally, preload is specified in terms of torque, and the axial load developed in the bolt depends on friction, which is difficult to characterize. The axial preload was calculated using the following expression from [9]: T = 0.2Fid,
(6)
where T is the specified bolt torque and d is the bolt diameter. The coefficient of 0.2 in eq. (6) corresponds to a friction coefficient of 0.15; lubricants may substantially reduce the friction coefficient, resulting in higher axial preloads for the same specified torque. On the other hand, relaxation, which reduces the axial load in the bolts and is also difficult to characterize, occurs with time. A survey of axial preload in bolts of pressure unseating hatches and heads actually obtained in the field would help quantify leakage potential for these types of penetrations. 3.3. Personnel airlocks
A personnel airlock with double dog-ear seals is shown in fig. 6. The inner door and bulkhead are the primary pressure boundary, however, if leakage were to occur past the inner door, there is a redundant outer door. The doors and bulkheads are essentially flat plates that are stiffened by webs and flanges. Internal pressuriTable 2 Separation pressure for BWR Mk I Drywell Heads Unit Drywell Heads BWR Mk I: 1 2-5 6 Equipment Hatches BWR Mk I: 1 6 BWR Mk III: 8-9
Ps kb (psig) (1061b/in)
s at Ps + 10 psi (in.)
135 127 119
199 410 278
0.0042 0.0033 0.0043
171 ? 34
91 140 194
0.0018 0.0012 a) 0.0023 b)
a) Specified torque not known. b) Mean temperature of bolts and compression flanges taken to be 350 and 400°F, respectively.
378
D.B. Clauss / Failure mechanisms of L WR steel containment buildings
will lift off. The door could be thought of as a plate that is simply supported by the bulkhead; separation of the sealing surfaces at the door corners may be possible. Also, if the stiffness of the door and bulkhead are not matched, their out-of-plane bending displacements may differ, especially near the horizontal stiffeners, resulting in separation of the sealing surfaces. If the performance of the seal material is compromised, small separations may result in leakage. Plans for testing a personnel airlock similar to that shown in fig. 6 are being formulated, as described in [6]. The personnel airlock will be subject to elevated temperature (700°F) and internal pressure; deformation of the sealing surfaces and leakage will be closely monitored. Finite element analyses will be conducted in support of this test. When the tests and analysis of this personnel airlock are completed, much more definitive statements about the leakage potential of personnel airlocks should be possible. Personnel airlocks with inflatable seals are used in a number of steel containments. As shown in fig. 7, there is a large gap between the door and the bulkhead that is blocked by the inflatable seals. The seals are pressurized and expand against the sealing surface; only seal performance prevents leakage. Inflatable seals are subject to degradation at elevated temperatures similar to other organic seals that have already been tested. In addition, high temperatures could reduce the material strength, and the seal could rupture due to its own internal pressure. In the absence of elevated temperature, it is suspected that inflatable seals would begin to leak when the internal pressure in the containment is approximately the same as the seal pressure. Since the seal pressure has a significant effect on leakage past inflatable seals, typical industry practice for pressurizing these types of seals will be surveyed. Tests are planned in the next year to determine the conditions under which the performance of inflatable seals is compromised.
4. Miscellaneous
Tests have been conducted on D.G. O'Brien and Westinghouse electrical penetration assemblies (EPAs) to determine their leakage characteristics during severe accidents [10]. Leakage was not detected during the testing of either EPA, although there was a small leak (0.13 cm3/s) from the D.G. O'Brien EPA during cooldown. The D.G. O'Brien and Westinghouse EPAs were tested to PWR and BWR Mk III severe accident environments, respectively. Both these environments have maximum temperatures below that required to degrade
the elastomer seals. A Conax EPA was tested in the more severe BWR Mk I and Mk II environment (maximum temperature of 700 F) in July 1986. It did not leak. Idaho National Engineering Laboratory conducted leak integrity tests on purge and vent valves subject to severe accident loads up to 350°F and 120 psig [11]. INEL's results seem to indicate that leakage during cooldown was possible under these conditions due to permanent set of the elastomer seal. Again there has not been testing in severe accident environments that have maximum temperatures capable of degrading the elastomer seals. There are also numerous mechanical and "dummy" penetrations that have flat plate covers. Although flat plates carry internal pressurization by primary bending, these covers generally appear to be designed for higher loads than the containment shell, and failure or leakage is unlikely to initiate at these points. The capacities of bellows (expansion joints) have not yet been documented in the open literature. Bellows are considered to be most susceptible to torsional and shear loads. Long range plans call for testing a bellows in severe accident environments to investigate failure modes and capacity.
5. Conclusions
There are a number of potential failure mechanisms in LWR steel containment buildings; leakage from operable penetrations or failure of the containment shell are most probable. If a through-wall crack or tear is developed in a steel containment shell, it is likely that rupture will rapidly ensue. However, there are a number of potential leakage paths in steel containment shells from which significant leakage could occur before a through-wall crack is developed. Leakage could be sufficient to depressurize the containment, thus precluding rupture. Although not considered explicitly in this paper, contact between the containment shell and the reactor building, which will occur if sufficiently large strains are developed, could also affect the failure mode and capacity of the containment building. For instance, if contact occurred, the likelihood of rupture might be reduced since redistribution of loads would then be possible. The leakage potential of penetrations and capacity of the containment shell are design dependent (containment specific), and need to be evaluated on a case by case basis. The leakage potential of mechanical and electrical penetrations does not appear to be significant. Operable
D.B. Clauss / Failure mechanisms of L W R steel containment buildings
penetrations appear to be the most likely source of significant leakage during a severe accident. Methods for evaluating the leakage potential of equipment hatches and drywell heads were presented in this paper. NRC plans for testing a personnel airlock, inflatable seals, and bellows will resolve many unanswered questions. Characterization of leakage from containment penetrations subject to severe accident conditions could be improved by undertaking the following efforts: - In-field survey of operable containment penetrations to measure (i) out-of-flatness of sealing surfaces, (ii) any decrease in the performance of seal materials (i.e., loss of resiliency) during normal operation, and (iii) actual bolt preloads in pressure unseating hatches and heads. - Quantification of leakage rates for irregular path geometries and also taking into account the effects of aerosols. - Experiments on seal and gasket materials to study the time dependence of performance degradation. The response and potential failure modes of steel containments subject to dynamic pressurization may also need to be studied more carefully. As mentioned in the Introduction to this paper, it is quite likely that, for a given containment, the failure mode due to dynamic loading is altogether different than that due to slow pressurization. Improvements to risk estimates could be obtained by an analytical or experimental investigation, or a combined approach, to study the response of containment buildings to dynamic loading.
Acknowledgement
The hypotheses and conclusions presented in this paper are based largely on experimental studies directed by L.N. Koenig, D.S. Horschel, and J.D. Keck from Sandia National Laboratories. Their efforts are gratefully acknowledged.
379
References
[1] D.S. Horschel, The Design, Fabrication, Testing and Analyses of Four 1:32-Scale Steel Containment Models, NUREG/CR-3902, SAND84-2153, Sandia National Laboratories, Albuquerque, NM (to be published). [2] D.B. Clauss, Comparison of Analytical Predictions and Experimental Results for a 1:8-Scale Steel Containment Model Pressurized to Failure, NUREG/CR-4209, SAND 85-0679, Sandia National Laboratories, Albuquerque, NM (July 1985). [3] D.B. Clauss, D.S. Horschel, and T.E. Blejwas, Insights Into the Behavior of LWR Steel Containments Buildings During Severe Accidents, Sandia National Laboratories, Albuquerque, NM, Nucl. Engrg. Des. (to be published). [4] J.F. Kiefner, et al., Failure Stress Levels of Flaws in Pressurized Cylinders, Progress in Flaw Growth and Fracture Toughness Testing, ASTM STP536 (American Society for Testing and Materials, 1973) pp. 461-481. [5] L. Griemann, F. Fanous, and D. Bluhm, Crack Propagation in High Strain Regions of Sequoyah Containment, NUREG/CR-4273, Ames Laboratory, Ames, Iowa (July 1985). [6] L.N. Koenig, Leakage Potential Through Mechanical Penetrations in a Severe Accident Environment, 3rd Workshop on Containment Integrity, Washington DC, NUREG/CP-0076 (May 21-23, 1986) pp. 557-568. [7] T.R. Bump, et al., Characterization of Nuclear Reactor Containment Penetrations - Preliminary Report, NUREG/CR-3855, SAND84-7139, Sandia National Laboratories, Albuquerque, NM (June 1984). [8] M.H. Shackelford, et al., Characterization of Nuclear Reactor Containment Penetrations - Final Report, NUREG/CR-3855, SAND84-7180, Sandia National Laboratories, Albuquerque, NM (February 1985). [9] J.E. Shigley, Mechanical Engineering Design (McGrawHill, New York, 1977) pp. 240-250. [10] J.D. Keck and F.V. Thome, Leak Behavior Thrgugh EPAs Under Severe Accident Conditions, 3rd Workshop on Containment Integrity, Washington DC (NUREG/CP0076 (May 21-23, 1986) pp. 569-580. [11] R. Steele Jr. and J.C. Watkins, Containment Purge and Vent Valve Test Program Final Report, NUREG/CR4141, EG&G Idaho, Inc., Idaho Falls, Idaho (September 1985).
371
FAILURE MECHANISMS OF LWR STEEL CONTAINMENT BUILDINGS SUBJECT TO SEVERE ACCIDENT LOADINGS * D a v i d B. C L A U S S Containment Integrity Division 6442, Sandia National Laboratories, Albuquerque, N e w Mexico 87185, USA
Received 27 April 1987
The potential failure mechanisms in LWR steel containment buildings subject to quasi-static pressurization and elevated temperature are identified. For each failure mechanism, the relevant structural response measures are discussed. For mechanisms involving leakage, the importance of seal performance is also discussed. Criteria that can be used to evaluate threshold environments are presented for several failure mechanisms. Results of tests on scale models and seal tests that support the criteria are referenced.
1.
Introduction
During a severe (degraded core) accident, containment buildings may be subject to loads that significantly exceed their design basis loads. The performance of the containment building, which represents the last engineered barrier to release of radioactive material, affects the consequences of a severe accident. A partial list of the containment performance parameters that affect severe accident consequences follows: - Capacity: A r e the severe accident loads sufficient to cause a failure, i.e., a significant release of radioactive material? - Timing: In general, risk from an early failure is greater than that from a late failure, because early in an accident sequence radioactive materials are more likely to be airborne. - Location o f failure: Is the release direct to the environment or filtered through other structures, for example, the reactor auxiliary building? - L e a k a g e rate o f "'hole size": Is the leakage rate sufficient to preclude further pressurization of the containment building? Consequences are proportional to leakage rate at least over some intermediate range of values. Rupture could result in resuspension of radioactive aerosols. * This work supported by the US Nuclear Regulatory Commission and performed at Sandia National Laboratories, which is operated by the US Department of Energy under contract number DE-AC04-76DP00789.
It should be evident that the predicted failure mechanism will have a significant effect on risk estimates. In fact, failure mechanisms are containment specific as well as scenario dependent - for instance, for a given containment, the failure mode due to dynamic loading could be much different than that due to quasi-static pressurization. This paper will focus on the potential failure mechanisms in a steel LWR containment building subject to quasi-static pressurization and elevated temperature. The failure mechanisms considered in this paper are: Failure of the containment shell; Leakage past sealing surfaces (operable penetrations); Equipment Hatch sleeve ovalization; Pressure unseating equipment hatches and drywell heads; Personnel airlocks; - Leakage from electrical penetration assemblies; - Leakage through purge and vent valves; - Leakage due to failure of a bellows (expansion joint). For each potential failure mechanism, the important structural response measures and seal performance parameters are discussed. Preliminary criteria are presented for evaluating rupture of the shell and leakage from equipment hatches and drywell heads. The discussion is limited to the range of response wherein the containment building behaves like a freestanding steel shell. Most steel containment buildings are enclosed by a concrete shield wall. If large membrane plastic strains are developed in the containment shell before failure occurs, the containment and shield
0 0 2 9 - 5 4 9 3 / 8 7 / $ 0 3 . 5 0 © Elsevier Science Publishers B.V. ( N o r t h - H o l l a n d Physics P u b l i s h i n g Division)
372
D.B. Clauss / Failure mechanisms of L W R steel containment buildings
wall will come into contact. In BWR Mk I containment buildings, the gap between the steel shell and the shield wall corresponds to approximately 0.5% membrane strain; in other types of plants with freestanding steel containments the gap typically corresponds to 5% membrane strain. If contact occurs, the performance of the containment building will be affected. Different failure modes may come into play, or the shield wall may strengthen the containment building, or both. The effect of contact with concrete shield walls on the failure mechanisms of LWR containments are not considered in this paper.
2. Failure of the containment shell In this paper, the containment shell is defined as the pressure boundary of the containment, exclusive of those parts of the boundary made up of penetrations, including equipment hatches, personnel airlocks, piping, and electrical penetrations assemblies. A loss of integrity associated with the containment shell must involve a cohesive material failure, i.e., a through-wall crack or tear in the pressure boundary. If a through-wall crack or tear is developed, it is highly likely that rupture (unstable crack growth) will rapidly ensue. Based on results of tests on scale models of steel containment buildings that were pressurized to failure [1,2], a simple strain criterion appears to be adequate for estimating the point at which a through-wall crack or tear will develop. The steel most commonly used in US containments, A516 Gr 70, has exceptional ductility, and very large plastic strains may be developed prior to failure. The following criterion for evaluating rupture is proposed: • Rupture will occur if, at any point on the pressure boundary the equivalent strain exceeds the material's ultimate strain (the strain at maximum load as determined from a uniaxial tensile test). If the strain is due primarily to bending, somewhat higher strain may be allowable (to account for crack growth through the thickness). The above criterion should be used only with detailed three-dimensional finite element analyses, which would need to include at least major penetrations and stiffeners. Results can be very sensitive to design details such as stiffener patterns around penetrations [2,3]. Any failure criterion must be consistent with the analysis used to calculate the response measures used in the criterion. If axisymmetric analyses are used to predict rupture, the limiting value of strain used to predict
rupture must be decreased to account for the effects of penetrations and stiffeners. Many of the discontinuities in the containment shell, such as the springline and the base embedment, result in self-limiting states of stress, and consequently, a cohesive material failure is not likely to initiate in these areas. On the other hand, large penetrations cause local strain concentrations in the shell that reduce the capacity relative to an undisturbed shell. Stiffeners tend to increase the general yield pressure but they do not necessarily increase the ultimate strength by a like amount. There is a large body of evidence that indicates that if a through-wall crack or tear is developed a steel containment shell will rupture catastrophically. Kiefner [4] developed empirical equations based on a large number of tests on pressurized cylinders that show that rupture will occur unless the initial through-wall crack is the result of a deep surface flaw (depth of the surface flaw must be at least one-half the wall thickness). Four of the five steel containment models tested at Sandia ruptured; in the one exception, leakage from a through-wall crack was sufficient to depressurize the model. However, in that particular model, the cylindrical wall had been severely thinned by excessive grinding adjacent to a repaired weld, which was equivalent to a deep surface flaw [3]. At least qualitatively, the formulas developed by Kiefner explain why rupture did not occur in that model. A final example is the J-integral analysis of the Sequoyah containment conducted by Griemann [5]. Griemann's analysis showed that a through-wall crack originating in the thinner wall sections near the springline would not be arrested at the thicker wall sections used at lower elevations, and he concluded that unstable crack growth (rupture) would occur.
If a steel containment building is pressurized to the point where the capacity of the shell is the controlling failure mechanism (which may be precluded by leakage from penetrations or by contact with the shield wall at lower pressures), then it is highly likely that a rupture will occur. It is possible, though not likely, that stable, through-wall cracks could develop in the shell, in which case leakage would be proportional to the crack opening area. (Initial surface flaws are assumed to be less than 20% of the shell thickness.)
3. Leakage from operable penetrations Any penetration that is intended to be used for access to the containment is categorized as an operable
373
D.B. Clauss / Failure mechanisms of LWR steel containment buildings
penetration. In US steel containments, this would include equipment hatches, personnel aidocks, escape hatches, and BWR Mk I and Mk II drywell heads. The potential for leakage between the sealing surfaces of these types of penetrations is considered in this section. In general, leakage from operable penetrations depends on both displacement and performance of the seal. Tests have shown that if metal to metal contact is maintained between the sealing surfaces, significant leakage does not occur regardless of the condition of the seal material [6]. On the other hand, because the seals are compressed typically one-quarter inch, large separations of the sealing surfaces can arise without significant leakage if the performance of the seal is not compromised. Relative displacements of the sealing surfaces can be categorized as either separation (relative motion perpendicular to the seating surfaces) or sliding (relative motion in the plane of the surfaces). Tests on seal and gasket materials [61 suggest that there are at least two ways in which performance of the seal can be compromised: (1) radiation or thermal aging can result in a loss of resilency, and (2) high temperatures (500 to 650 o F, depending on the type of material, degrade the seal materials, that is, the materials outgas, then become dry and powdery). In special cases, leakage may be a function of performance of the seal only, or sealing surface deformation only. In personnel airlocks with inflatable seals, the door and bulkhead are not designed to be in intimate contact, and therefore, leakage will arise if the performance of the seal is compromised, irrespective of structural deformations. The performance of the seal may also be the only factor preventing leakage if there is significant out-of-flatness between the seating surfaces. This is a condition that would not be detected by Integrated Leak Rate Tests; at design levels, seal performance prevents leakage. However, in BWR Mk I and Mk II plants, severe accident temperatures are expected to be sufficient to degrade seals and out-of-flatness could result in leakage without additional separation or sliding of the seating surfaces. In cases where sliding or separation of the seating surfaces is very large, leakage can result even if the performance of the seal is not compromised.
DETAIL B
Fig. 1. Details of a pressure seating equipment hatch.
the containment shell. The horizontal diameter of the sleeve increases while the vertical diameter decreases by a like amount; the deformations can become quite large
,,
:A" I
~__{
a"
S~-IE DETAIL '~I~" G~I)VF- ~
•
3.1. Equipment hatch sleeve ovalization
Typical details of equipment hatches used in US steel containments are shown in figs. 1 and 2. In a cylindrical containment shell, an equipment hatch sleeve will deform into an oval shape due to interaction with
~u~l SLEEVE ~ ~0" ~ 9.' "/"
F C~'sKET
~E'rAl"
PRE~IU~ E L E VA,T|ON VIEW
Fig. 2. Details of a pressure unseating equipment hatch.
374
D.B. Clauss / Failure mechanisms of L W R steel containment buildings
(on the order of the sleeve wall thickness) after general yielding of the cylindrical containment shell. The equipment hatch cover and tensioning ring deform axisymmetrically, i.e., they retain their circular shape. Changes in the diameter of the tensioning ring are expected to be quite small (except possibly for pressure seating hatches that are susceptible to buckling). The mismatch caused by the sleeve sliding relative to the tensioning ring is greatest along the horizontal and vertical centerlines; as indicated in fig. 3a, eventually the sleeve and tensioning ring will lose contact at these points, resulting in leakage. Under these conditions, the
performance of the seal material does not affect the leakage potential. The deformation of the equipment hatch sleeve can be related to the average membrane strain in the containment shell in an approximate way. Ovalization of the equipment hatch sleeve is primarily an inextensional (no stretching of the midsurface) deformation, and consequently it is a relatively flexible mode of deformation. Thus, the equipment hatch sleeve basically conforms to the membrane displacements of the containment shell at their intersection. As illustrated in fig. 4, the increase in the horizontal radius of the sleeve, A r H, can be approximated as Ar H = o r ,
(1)
a r- . . . . I
where c is the average membrane circumferential strain in the shell at the elevation of the equipment hatch centerline, and r is the original sleeve radius. Assuming the cover and tensioning ring do not deform significantly, leakage due to ovalization is assured if
..,__._-.--------'-TENSIONING ~ - ~ RING
SLEEVE
~,~%,~
Ar > t,
(2)
where A r is the change in the sleeve radius and t is the sleeve wall thickness at the sealing surface. Combining eqs. (1) and (2), the approximate leakage criterion is c r / t > 1.
(3)
The average membrane strain in the shell at the elevation of an equipment hatch corresponding to leakage from ovalization, ~', can be calculated from eq. (3). For equipment hatches used in steel containments identified in a survey conducted by Argonne National Laboratory
b
[ ". . . .
~--~...
~TENSIONING \
(r /
il'! Fig. 3. Leakage path created by ovalization of sleeve: (a) without seal degradation; (b) with seal degradation,
~
/
UNDEFORMED SHAPE
R ~ CYLINDER RADIUS r ~ SLEEVE RADIUS I[ ~ "AVERAGE"MEMBRANE CIRCUMFERENTIALSTRAIN
SLEEVE AXIS
Fig. 4. Equipment hatch sleeve and containment shell interaction.
D.B. Clauss / Failure mechanisms of L WR steel containment buildings ( A N L ) [7,8], (' ranged from 2.5% to 7.3%, as indicated in table 1. Clearly, significant plastic membrane strains must be developed in the containment shell before ovalization of an equipment hatch sleeve will cause leakage. U n d e r certain conditions, leakage caused by ovalization may occur at lower values of average membrane strain. If the seal material is severely degraded, the grooves that hold the seals could act as a leakage path, as suggested by fig. 3b. For a pressure seating hatch, the grooves would be exposed to the inside of the containment at 3 and 9 o'clock and to the outside at 6 and 12 o'clock, yielding critical strains ((') from 1.7 to 4.5%, as indicated in table 1. However, in order for this leakage path to be realized, a substantial amount of the seal would probably have to be degraded and ejected from the seal grooves. The approximate method outlined above did prove accurate for the equipment hatches in the l : 8 - s c a l e model [2,3]. In particular, the change in the horizontal diameter calculated from eq. (1) was within 10% of the measured displacement. However, it is unknown how accurate this approximate method would be for different reinforcement designs and stiffener patterns. 3.2. Separation of pressure unseating hatches Schematics of a pressure-unseating equipment hatch and a B W R Mk I drywell head are depicted in figs. 2
375
Table 1. Critical strain for equipment hatch leakage (by ovalization) in steel containments Unit
t (in.)
r
(in.)
(,
a) (%)
(, b) (%)
Seal type
1
514
72
7.3
4.5
gumdrop
2
4½
72
6.3
4.2
gumdrop
3
4~2
84
5.4
3.9
gumdrop
4
4~
84
5.4
3.6
gumdrop
5 6
4~ 4~
126 126
3.6 3.6
2.3 2.3
gasket gasket
7 10
4~: 3
126 120
3.6 2.5
2.3 1.7
gasket gumdrop
~) Assuming seal degradation has not occurred. b) Assuming seal has been severely degraded and ejected.
and 5, respectively. In both cases the bolts are preloaded to maintain leak tightness during internal pressurization of the containment. Typically the bolts are tightened until metal to metal contact is achieved, and then a specified level of torque or load is applied. Internal pressure acts in the opposite direction of the bolt preload and tends to separate the sealing surfaces. The effects of temperature must also be considered; from a heat transfer viewpoint, the bolts, if they are exposed to air, will act as convective fins. Thus, in many
':;'°
I
• ~rr~.~.
Fig. 5. Details of a BWR Mk I Drywell Head.
~
•
376
D.B. Clauss / Failure mechanisms of L WR steel containment buildings
/
,/X, I- o~
\i
'2
I
5-ZT-<,'~"
~
5-
,L_-ZLJ r ' I" l /
I
I pRES,;%~DE .
F~,SKET--~
DETAIL B
@loov
\\
1.344 W •
'-~.I Y.,"~'~ •
VIEW
I
A-A
" !,-I,.--~--~ ~.- ~ ,11',,'J
b" 1½"l
DETAIL C
,
,,
[
Fig. 6. Details of a personnel airlock with double dog-ear seals. zation is resisted primarily by bending action, and therefore, rather large out-of-plane deformations may arise for beyond design basis loads. Although the door is pressure seating, separation
may still be possible at certain points. For instance, in a simply supported rectangular plate subject to uniform pressure, significant reaction forces must occur at the comers; if the corners are not held down, the comers
R.
T...
PRESSURE;
I ii
t, ol DOO~R
"'1- I BULIC.HEAD ~LO" ~-I
NI:L ~TAB SEAL
r, LE
g
~," )IA.
2E'--
C)PENIMC¢
HINGE
,
Fig. 7. Details of a personnel airlock with inflatable seals.
DETAIL
377
D.B. Clauss / Failure mechanisms of LWR steel containment buildings
designs, the average temperature of the bolts will be less than that of the compression flanges. Differential thermal expansion (when the bolts expand less than the compression flanges) tends to maintain contact between the sealing surfaces, that is, it acts in the opposite sense of internal pressurization. The following criteria are proposed for evaluating the onset of leakage from pressure unseating hatches and heads: • If the seals are exposed to temperatures high enough to cause degradation (500 to 650°F, depending on the material), leakage will occur when the internal pressure is sufficient to cause separation of the sealing surfaces. • If the seals are not degraded, but at some time into the accident aging causes a loss of resiliency, a subsequent increase in separation will cause leakage. • If the seals are not degraded and do not lose resiliency, leakage will occur when the seal squeeze becomes equal to zero (when the separation is equal to the initial seal compression). Recent tests on silicone O rings suggest that even degraded seals may prevent leakage through small gaps, up to about 0.005 or 0.006 in.. If this is substantiated, the above criteria would have to be modified so that leakage for separations less than this amount is not allowed regardless of the condition of the seal. Assuming linear elastic behavior of the bolts and compression flanges, the internal pressure at which separation begins for pressure unseating hatches and heads, Ps, can be calculated as: Ps = [ ( k b + k f ) F i / / k f + k b ( ' T f - - CTb) L ] / ( 1 + 2 V k b / k f ) ~ r r 2.
(4)
The subscripts f and b denote the compression flanges and bolts, respectively. F i is the bolt preload, k is the axial stiffness, c T is the thermal strain, L is the bolt grip, v is Poisson's ratio, and r is the average radius of the equipment hatch sleeve or drywell head. Eq. (4) was determined by introducing terms representing the axial thermal expansion of the bolts and the compression flanges and the axial deflection of the compression flanges due to hoop stress to expressions in [9]. The resulting equation for the force in the compression flanges was set to zero and solved for p,. Additional pressure is carried entirely by the bolts, so that the separation between the sealing surfaces, s, is given by s = ~rr2(p - p s ) / k b
for p >Ps-
(5)
The separation pressure for drywell heads and pressure
unseating equipment hatches used in steel containments that were identified in the A N L survey of containment penetrations [7,8] is tabulated in table 2. The mean temperature of the compression flanges was taken to be 700°F and that in the bolts was taken to be 550°F (these temperatures are based on heat transfer calculations made for the drywell head region of a given containment). The results are very sensitive to the bolt preload, which is difficult to know accurately. Generally, preload is specified in terms of torque, and the axial load developed in the bolt depends on friction, which is difficult to characterize. The axial preload was calculated using the following expression from [9]: T = 0.2Fid,
(6)
where T is the specified bolt torque and d is the bolt diameter. The coefficient of 0.2 in eq. (6) corresponds to a friction coefficient of 0.15; lubricants may substantially reduce the friction coefficient, resulting in higher axial preloads for the same specified torque. On the other hand, relaxation, which reduces the axial load in the bolts and is also difficult to characterize, occurs with time. A survey of axial preload in bolts of pressure unseating hatches and heads actually obtained in the field would help quantify leakage potential for these types of penetrations. 3.3. Personnel airlocks
A personnel airlock with double dog-ear seals is shown in fig. 6. The inner door and bulkhead are the primary pressure boundary, however, if leakage were to occur past the inner door, there is a redundant outer door. The doors and bulkheads are essentially flat plates that are stiffened by webs and flanges. Internal pressuriTable 2 Separation pressure for BWR Mk I Drywell Heads Unit Drywell Heads BWR Mk I: 1 2-5 6 Equipment Hatches BWR Mk I: 1 6 BWR Mk III: 8-9
Ps kb (psig) (1061b/in)
s at Ps + 10 psi (in.)
135 127 119
199 410 278
0.0042 0.0033 0.0043
171 ? 34
91 140 194
0.0018 0.0012 a) 0.0023 b)
a) Specified torque not known. b) Mean temperature of bolts and compression flanges taken to be 350 and 400°F, respectively.
378
D.B. Clauss / Failure mechanisms of L WR steel containment buildings
will lift off. The door could be thought of as a plate that is simply supported by the bulkhead; separation of the sealing surfaces at the door corners may be possible. Also, if the stiffness of the door and bulkhead are not matched, their out-of-plane bending displacements may differ, especially near the horizontal stiffeners, resulting in separation of the sealing surfaces. If the performance of the seal material is compromised, small separations may result in leakage. Plans for testing a personnel airlock similar to that shown in fig. 6 are being formulated, as described in [6]. The personnel airlock will be subject to elevated temperature (700°F) and internal pressure; deformation of the sealing surfaces and leakage will be closely monitored. Finite element analyses will be conducted in support of this test. When the tests and analysis of this personnel airlock are completed, much more definitive statements about the leakage potential of personnel airlocks should be possible. Personnel airlocks with inflatable seals are used in a number of steel containments. As shown in fig. 7, there is a large gap between the door and the bulkhead that is blocked by the inflatable seals. The seals are pressurized and expand against the sealing surface; only seal performance prevents leakage. Inflatable seals are subject to degradation at elevated temperatures similar to other organic seals that have already been tested. In addition, high temperatures could reduce the material strength, and the seal could rupture due to its own internal pressure. In the absence of elevated temperature, it is suspected that inflatable seals would begin to leak when the internal pressure in the containment is approximately the same as the seal pressure. Since the seal pressure has a significant effect on leakage past inflatable seals, typical industry practice for pressurizing these types of seals will be surveyed. Tests are planned in the next year to determine the conditions under which the performance of inflatable seals is compromised.
4. Miscellaneous
Tests have been conducted on D.G. O'Brien and Westinghouse electrical penetration assemblies (EPAs) to determine their leakage characteristics during severe accidents [10]. Leakage was not detected during the testing of either EPA, although there was a small leak (0.13 cm3/s) from the D.G. O'Brien EPA during cooldown. The D.G. O'Brien and Westinghouse EPAs were tested to PWR and BWR Mk III severe accident environments, respectively. Both these environments have maximum temperatures below that required to degrade
the elastomer seals. A Conax EPA was tested in the more severe BWR Mk I and Mk II environment (maximum temperature of 700 F) in July 1986. It did not leak. Idaho National Engineering Laboratory conducted leak integrity tests on purge and vent valves subject to severe accident loads up to 350°F and 120 psig [11]. INEL's results seem to indicate that leakage during cooldown was possible under these conditions due to permanent set of the elastomer seal. Again there has not been testing in severe accident environments that have maximum temperatures capable of degrading the elastomer seals. There are also numerous mechanical and "dummy" penetrations that have flat plate covers. Although flat plates carry internal pressurization by primary bending, these covers generally appear to be designed for higher loads than the containment shell, and failure or leakage is unlikely to initiate at these points. The capacities of bellows (expansion joints) have not yet been documented in the open literature. Bellows are considered to be most susceptible to torsional and shear loads. Long range plans call for testing a bellows in severe accident environments to investigate failure modes and capacity.
5. Conclusions
There are a number of potential failure mechanisms in LWR steel containment buildings; leakage from operable penetrations or failure of the containment shell are most probable. If a through-wall crack or tear is developed in a steel containment shell, it is likely that rupture will rapidly ensue. However, there are a number of potential leakage paths in steel containment shells from which significant leakage could occur before a through-wall crack is developed. Leakage could be sufficient to depressurize the containment, thus precluding rupture. Although not considered explicitly in this paper, contact between the containment shell and the reactor building, which will occur if sufficiently large strains are developed, could also affect the failure mode and capacity of the containment building. For instance, if contact occurred, the likelihood of rupture might be reduced since redistribution of loads would then be possible. The leakage potential of penetrations and capacity of the containment shell are design dependent (containment specific), and need to be evaluated on a case by case basis. The leakage potential of mechanical and electrical penetrations does not appear to be significant. Operable
D.B. Clauss / Failure mechanisms of L W R steel containment buildings
penetrations appear to be the most likely source of significant leakage during a severe accident. Methods for evaluating the leakage potential of equipment hatches and drywell heads were presented in this paper. NRC plans for testing a personnel airlock, inflatable seals, and bellows will resolve many unanswered questions. Characterization of leakage from containment penetrations subject to severe accident conditions could be improved by undertaking the following efforts: - In-field survey of operable containment penetrations to measure (i) out-of-flatness of sealing surfaces, (ii) any decrease in the performance of seal materials (i.e., loss of resiliency) during normal operation, and (iii) actual bolt preloads in pressure unseating hatches and heads. - Quantification of leakage rates for irregular path geometries and also taking into account the effects of aerosols. - Experiments on seal and gasket materials to study the time dependence of performance degradation. The response and potential failure modes of steel containments subject to dynamic pressurization may also need to be studied more carefully. As mentioned in the Introduction to this paper, it is quite likely that, for a given containment, the failure mode due to dynamic loading is altogether different than that due to slow pressurization. Improvements to risk estimates could be obtained by an analytical or experimental investigation, or a combined approach, to study the response of containment buildings to dynamic loading.
Acknowledgement
The hypotheses and conclusions presented in this paper are based largely on experimental studies directed by L.N. Koenig, D.S. Horschel, and J.D. Keck from Sandia National Laboratories. Their efforts are gratefully acknowledged.
379
References
[1] D.S. Horschel, The Design, Fabrication, Testing and Analyses of Four 1:32-Scale Steel Containment Models, NUREG/CR-3902, SAND84-2153, Sandia National Laboratories, Albuquerque, NM (to be published). [2] D.B. Clauss, Comparison of Analytical Predictions and Experimental Results for a 1:8-Scale Steel Containment Model Pressurized to Failure, NUREG/CR-4209, SAND 85-0679, Sandia National Laboratories, Albuquerque, NM (July 1985). [3] D.B. Clauss, D.S. Horschel, and T.E. Blejwas, Insights Into the Behavior of LWR Steel Containments Buildings During Severe Accidents, Sandia National Laboratories, Albuquerque, NM, Nucl. Engrg. Des. (to be published). [4] J.F. Kiefner, et al., Failure Stress Levels of Flaws in Pressurized Cylinders, Progress in Flaw Growth and Fracture Toughness Testing, ASTM STP536 (American Society for Testing and Materials, 1973) pp. 461-481. [5] L. Griemann, F. Fanous, and D. Bluhm, Crack Propagation in High Strain Regions of Sequoyah Containment, NUREG/CR-4273, Ames Laboratory, Ames, Iowa (July 1985). [6] L.N. Koenig, Leakage Potential Through Mechanical Penetrations in a Severe Accident Environment, 3rd Workshop on Containment Integrity, Washington DC, NUREG/CP-0076 (May 21-23, 1986) pp. 557-568. [7] T.R. Bump, et al., Characterization of Nuclear Reactor Containment Penetrations - Preliminary Report, NUREG/CR-3855, SAND84-7139, Sandia National Laboratories, Albuquerque, NM (June 1984). [8] M.H. Shackelford, et al., Characterization of Nuclear Reactor Containment Penetrations - Final Report, NUREG/CR-3855, SAND84-7180, Sandia National Laboratories, Albuquerque, NM (February 1985). [9] J.E. Shigley, Mechanical Engineering Design (McGrawHill, New York, 1977) pp. 240-250. [10] J.D. Keck and F.V. Thome, Leak Behavior Thrgugh EPAs Under Severe Accident Conditions, 3rd Workshop on Containment Integrity, Washington DC (NUREG/CP0076 (May 21-23, 1986) pp. 569-580. [11] R. Steele Jr. and J.C. Watkins, Containment Purge and Vent Valve Test Program Final Report, NUREG/CR4141, EG&G Idaho, Inc., Idaho Falls, Idaho (September 1985).