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Proceedings of the OECD I CSNI specialist meeting on ~~u~~ ~EACTOR CONTAINMENT SAFETY

Ociober 17 - 21, 1983 Toronto, Ontawso

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CSNI Specialist Meeting on Water Reactor Containment Safety Toronto, Canada 17th-21st October 1983

ABSTRACT

Behaviour of Concrete Containment Structures Under Over-Pressure Conditions

by G.J.K. Asmis

Atomic Energy Control Board, Ottawa, KIP 5S9 )• 2

I' I This presentation reports on the most significant findings from a study to assess the response of prestressed concrete secondary contain- ment structures for nuclear reactors under the influence of high in- ternal overpressures. In 1975 the Atomic Energy Control Board of Canada began to sponsor a comprehensive study at the University of Alberta to determine the response of a containment structure to such overpressures. The length of the study was 5 years and was approximately equally divided between experimental and theoretical investigations. The principal investigators were professors J.G. MacGregor, D.W. Murray and S.H. Simmonds. A film was made of the test of the 1/14 scale model of containment. The response of the structure, in terms of crack progression, ductile expansion and finally the moment of rupture can be clearly seen. The reactors in a Canadian nuclear power plant of the Gentilly-2 type are enclosed in a cylindrial prestressed concrete structure. In the event of certain accidents in which pressurized gases or steam would be discharged, this concrete enclosure would act as a containment structure to prevent these products from escaping into the atmosphere. The largest internal pressures considered in design result from a complete rupture of a heat transport line. The Gentilly-2 plant is designed so that the increased pressure due to this release of steam would activate a water dousing system which would cool and condense the steam, limiting the maximum internal pressure to 124 kPa(g)(18 psiq). This is referred to as the design basis accident. As well, one of the design criteria for the containment is that there be no tensile stresses in the inner surfaces of the concrete under a proof pressure of 1.15 times this amount. In the unlikely event that a main steam line ruptures completely and simultaneously the dousing system fails to act, the internal pressure may reach several times the design pressure. As the internal pressure in a prestressed concrete containment structure increases, it is possible to postulate a series of stages of increasing damage to the structure which limit its usefulness. These conditions are referred to as "Limit States" and include such states as first surface cracking, first through-the-wall cracking, initial yielding of reinforcement, first yielding of prestressing tendons, and fracture of reinforcement and/or tendons. The objectives of the study were to develop a methodology that would permit the prediction of both the internal pressures and the locations in the containment where these limit states occur, and using this methodology to estimate the rate of leakaqe through the concrete portion of the containment at any given pressure. These objectives have been met and are reported in references 1 to 12. Procedures for evaluating the degree of cracking, the amount of deformation, the extent of yielding of the reinforcement, the rate of leakage associated with any given internal pressure, as well as the ultimate pressure to cause fracture of the tendons were developed. These procedures were verified by the testing to failure of a 1/14 scale prestressed concrete containment structure in the laboratory. Excellent agreement was obtained between the predicted and observed behaviour of this test structure. The leakage rates were not verified in this test. Details such as temporary openings during construction, air locks and I \ 3

other penetrations which could affect locally the behaviour of the structure were not considered. The test structure was patterned after the Gentilly-2 type containment and was designed to behave and fail in a similar manner. Practical considerations in the construction and testing of the model meant that the same scale could not be used for all structural elements. As a results while the overall behaviour of the model would be similar to the prototypes there would not be a one to one correspondence of pressure at each limit state between the structure tested and the prototype containment. For this reasons the test struc- ture should be viewed as a structure in its own right rather than as a model of a particular existing containment structure. Since the test structure was designed to indicate behaviour at high internal pressuress only those components of the prototype resisting internal pressure were included in the model. Hence the lower dome for the dousing system was not included in the model and the base was designed to be rigid rather than a slab on an elastic foundation. The connection between the base and the wall s howevers was designed to act as a hinge in a manner similar to the prototype. The test structure had a height above the base of 3.81 metres (12 I -6") and an outer wa 11 diameter of 3.2 metres (10'-6") correspondi no to an overall scale of 1/14 compared to the prototype. Reinforcement and prestressing details resembled those of the prototype and were proportioned to give the same sequence of behaviour as elastic analyses had predicted for the prototype. In order to maintain a constant pressure while readings were being taken and to obtain sufficient internal pressure to cause failures a flexible liner was used to prevent leakage. To prevent an explosive type failure the fluid used for loading was water. Extensive measurements were made during loading of the test structure. Quantities measured electronically included internal pressures def1ectionss steel and concrete strainss and those measured manually included concrete surface strains crack widths and meridona1 relations. Electronically read quantities were obtained from electric resistance strain gages and LVDT's (Linear Variable Differential Transformers) from which the voltage output was sampled and converted to digital form by various pieces of data logging equipment that were monitored and recorded using a NOVA 210/E digital computer. In general s readings were taken electronically at pressure intervals of 34.5 kPa (5 psi) and manual readings at intervals of 70.0 kPa (10 psi). The response of the structure to increasing pressure is documented in the movie. The sixth tests was to have been the test to failures however the test was stopped at 550 kPa(g) (80 psig) because of leakage of the liner and the difficulty in increasing and maintaining pressure. During this test extensive cracking of the structure was observed over much of the shell surface. A new liner was fabricated and inserted inside the patched initial liner. The final test to failure took two days to complete. No leakage through the liner was observed during this test until the final failure pressure was reached. Above 550 kPa(q) (80 psig)s the crack pattern over the wall and dome continued to develop and there was a strong correlation between the crack and tendon patterns. At higher 10adss these cracks became wider and new cracks deve10peds again reflecting the tendon and reinforcement locations. 4

The crack pattern in the wall at 930 kPa(g) (135 psig) is shown in the movie. Attention is drawn to the horizontal cracking existing alonq the face of the buttress. Also visible is the outward bulging of both the wall and buttress. At pressures higher than this, although no new cracks developed an increase in pressure caused a considerable increase in deformations attesting to the ductility of the test structure. r'twas difficult to increase pressure using the low volume hand pump. At a pressure of 1030 kPa(g) (149.5 psig) the loading was terminated after 16 hours of continuous testing. As a precaution, the load was reduced to 945 kPa(g) (137 psig) for overnight. When loading was resumed after an interval of 14 hours .using a truck mounted high capacity pump, the pressure had dropped only 13.8 kPa (2 psi). The gap between the two points after reloading to the previous pressures is an indication of the creep that occurred during this period. Thus, although the test structure was cracked extensively and the internal pressure was over 85% of that to cause failure when it was left over night, the test structure was able to maintain this pressure with very littl~ additional deformation. This was not altogether unexpected as the structure was essentially being held together by the prestressing tendons. Failure of the structure occurred at a pressure of 1,100 kPa(g) (159.5 psig) when the 7th horizontal tendon from the base and one vertical tendon fractured at mid height near the south-east buttress, permitting the liner to rupture. Immediately prior to failure the concrete cover over one of the horizontal tendon anchorages at this buttress was observed to pop outwards followed by rapid opening of the cracks in the immediate vicinity and spalling of the concrete. The dynamics of the failure have been caught on film. Although failure was due to fracture of the horizontal tendons, the buttresses in which the horizontal tendons were anchored showed distress at loads well below the failure load. At 552 kPa(g) (80 psig) well defined horizontal cracks were visible across the face of the abutment and each anchorage point had a crack immediately above and below the anchor plates. These cracks continued to open and near 896 kPa(g) (130 psig) at several anchorages the concrete over the anchorage was spalled outwards as the anchor plates pulled into the sides of the buttresses. The experimentally derived relationships were factored into the theoretical phase of this research work. The resulting expressions were applied to the Gentilly-2 secondary containment building. It was predicted that for this building, cracks would first penetrate throuoh the wall at 331 kPa(g) (48 psig) internal pressure, or 2.3 times the proof test pressure. After extensive inelastic deformation, failure would occur at 531 kPa(g) (77 psig). This assumes that the pressurizino medium could be supplied rapidly enough to maintain these pressures. Leakage calculations indicate that leakage through cracks would be negligible at pressures below 345 kPa(g) (50 psig), increasing exponentially as the pressure increases above that value. At 93 percent of the predicted failure load the calculations indicate a leakage rate approximately equal to the volume of the structure each second. 5

REFERENCES: 1) IIBehaviour of Prestressed Concrete Containment structures - A Summary of Findings" INFO-0031, Atomic Energy Control Board, KIP 5S9, Ottawa, Canada. 2) MacGregor, J.G. Sil11llonds,S.H., and Rizkalla, S.H. IICracking of ll Prestressed Concrete Containment Due to Internal Pressure , Paper J4/10, SMiRT 6, Paris, August 1981. 3) Simonds, S.H. MacGregor, J.G., and Rizkalla, S.H. IIResponse of ll Prestressed Concrete Containment Model in Post Cracking Region , Paper J4/11 SMiRT 6, Paris, August 1981. 4) Murray, D.W., Sil11llonds,S.H., and MacGregor, J.G. "predicting ll Behaviour of Gentilly-2 Containment , Paper 5/7, SMiRT 6, Paris, Augus t 1981. 5) Rizkalla, S.H., ~1acGregor, J.G., and Simmonds, S.H., IIAirLeakage ll Characteristics of Prestressed Concrete Containment , Paper J5/14, SMiRT 6, Paris, August 1981. 6) Murray, D.W., Chitnuyanondh, L., Rijub-Agha, K.Y., and Wong, C., ll IIConcrete Plasticity Theory for Biaxial Stress Analysis , Journal of the Engineering Mechanics Division, ASCE, Vol. 105, No. EM6, Dec. 1979, pp. 989-1006. . 7) Murray, D.W., IIAReview of Explosive Characteristics of Prestressed ll Secondary Containments Near the Ultimate Load , Nuclear Engineering and Design, Vol. 52, Amsterdam, 1979. 8) Chitnuyanondh, L, Rizkalla, S.H., Murray, D.W., and Machregor, J.G., IIEffective Tensile Stiffening in Prestressed Concrete Wall ll Segments , Paper J3/4, SMiRT5, Berlin, 1979. 9) Murray, D.W., Ch itnuyanondh, L., and Wong, C., IIMode11ing and Predicting Behaviour of Prestressed Concrete Secondary Containment Structures Using BOSOR5", Paper J3/5, SMiRT, Berlin, 1979. 10) Murray, D.W., Simmonds, S.H., MacGregor, J.G., IISomeAspects of Structural Response of Concrete Containment Buildings To Accident ll Condition , Meeting, Canadian Nuclear Association, Toronto, May, 1979. 11) Rizkalla, S., Sil11llonds,S.H., and MacGregor, J.G., IIATest of a Model of the Thin-Walled Prestressed Concrete Secondary Containment ll Structure , Paper J4/2, SMiRT5, Berlin, August 1979. 12) Rizkalla S.H., Lau B.L., Simmonds S.H. IILeakage of Pressurized ll Gases Through Unlined Concrete Containment Structures , Paper J1/5, SMiRT 7, Chicago, August 1983. A RATIONAL APPROACH TO ASSESS PIPEWHIP IMPACT LOCAL EFFECTS ON CONCRETE CONTAINMENTS

J.H.K. Tang. A. Danay. J.H.P. Lee Ontario Hydro. 700 University Avenue. Toronto. Ontario. canada

Presented at OECD/CSNI Containment Specialist Meeting

held 1n Toronto

on OCtober 17-20. 1983

l700L SllMMARY

The local response of a concrete containment due to pipewhip can be examined by a combination of rigorous analytical techniques. as well as by fundamental principles of mechanics and empirical formulas derived from experimental data available to-date.

The analysis involves the following procedures:

Firstly. interpretation of experimental data on missile tests and comments on empirical formulas for missile perforation; secondly. simple dynamic analysis of severed piping system subjected to jet impulse. coupled with more complex numerical approach using commercially available computer programs;

Thirdly. extensive study of dynamic stress wave propagation effects by the finite difference program PISCES which takes into account the non-linear characteristics of steel and concrete materials; and

Fourthly, evaluation of possible cracked width and resulting leakage path area by assessing the shear capacity after concrete cracking has occurred.

The results obtained from the above activities include the amount of energy absorbed by the concrete. depth of penetration and estimates of possible crdcked dred induced by the impdct due to a postulated pipe burst accident.

1700L 1.0 INTRODUCTION

As part of the safety assessment of the reactor design, the consequences of postulated breaks in the primary heat transport (PHT) piping. especially the effects on the containment have to be evaluated. For this purpose. a number of critical design basis pipe rupture cases would have to be hypothesized.

The forces which result from the impact of a ruptured pipe on the concrete structure are usually computed by non-linear dynamic analysis of the piping system in which the plastic moment and deformation in ovalization of a pipe are taken into account. The results include the velocity of the pipe at the instant of impact. the distribution of the contact forces and their individual time histories. The dynamic effects of these impact lOads are then examined to ensure that the reactor building can maintain its structural integrity for all postulated pipe break scenarios. In addition. local damages due to collision between the pipe as a missile and the concrete are studied using empirical equations such as Army corps of Engineers (ACE).,modified Petry and modified National Defense Research Committee (NDRC) formulas. Taking into account the energy absorption capability of the pipe, the amount of penetration into the concrete structure can be computed in order to assu:e that the containment would not be perforated.

However. in the region immediately surrounding a penetrating missile. the structure locally is obviously cracked and crushed under the severe impact. The extent of internal cracking cannot be easily determined from the previous analyses. Therefore. a more detailed study has to be carried out to address this problem in order to estimate any possible increase in leakage rate for the containment envelope resulting from this postulated accident condition. This paper suggests a rational approach to evaluate these local impact effects and describes some of the major results of a sample investigation using this methodology.

2.0 MISSILE IMPACT LOCAL RESPONSE

2.1 Description of Physical Phenomenon

Local impact phenomena for rigid missile impdct are schematically shown in Figure 1. The following terminology is commonly used to describe the damage characteristics: (a) Spalling is the ejection of pieces of concrete from the front face region surrounding the area of impact. thus leaving a front ctaterthat extends over a substantially greater area than the cross-sectional area of the striking missile.

(b) SCabbing means ejection of pieces of concrete from the rear surface as a result of the tensile effects of reflected compression and shear stress waves generated in the concrete barrier. The ShOCK wave effects form an almost conical plug of fractured concrete ahead of the missile. If the shear capacity Is exceeded, this plug would start to move and create tensile splitting at the back face reinforcing. Separation of the concrete cover occurs when the plug movement is excessive.

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Tests have shown that scabbing is always related to the rear face and the last layer of steel. and the zone of scabbing (and cracking) will generally be much wider. but not as deep. as the front face spall crater (Berriaud 1977).

(c) Penetration is originally defined as the depth to which the missile enters a massive concrete with "infinite" thickness. The concrete is assumed not to yield (scab) on the back face. However. recent interpretation of penetration (Rotz 1975. Nusbaum 1976. Stevenson 1977) often refers to the depth to solid plug as designated by "d" in Figure 1. This definition would include the possible movement of the plug after the occurrence of concrete "shear" failure, as discussed above.

(d) Perforation of the target will occur as the penetration cavity extends through to the scabbing crater and the projectile just passes completely through the slab. It must be noted that. in most cases, perforation is prevented only by the reinforcing mesh in arresting the plug movement by providing the "net" effect. More details with regard to this phenomenon are discussed later.

2.2 Empirical Concrete Impact Formulas The extent of local damages induced by a missile striking a reinforced concrete target depends highly on the impact parameters. The most importdnt pdCdmeters dce the missile chdcdcteristics, impdct conditions and target characteristics as listed in Table 1. A review of available experimental data for a range of impact parameters has found that the existing results are still rather limited (Sliter 1979, Godbout 1980). A summary of the range of parameters from impact tests on concrete slabs is presented in Table 2. Most of the tests are related to either solid or pipe missiles with the longitudinal axis of symmetry striking normal to the concrete target surface. Thus. these missiles were conside~ as nondeformable or intermediately deformable. As a result. almost all empirical formulas developed should only be applicable to these "hard" missiles (Kennedy 1976, Berriaud 1977 and 1979, McMahon 1977 and 1979. Degen 1980). Of all the formulas. the most commonly used in nuclear plant design to-date (ASCE 1980) is the so-called modified NDRC (National Defense Research Council) formula as given below:

(a) For penetration depth:

)1.8 x .. /4 KNWd ( V for ! ~ 2.0 (la) 1000 d d or

)1.8] x .. (KNW ( V + d for ! ~ 2.0 (lb) 1000 d d

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in which

x ". penetration depth in inches

"" missile weight in pounds

K concrete penetrability factor

180 where f~ ".concrete ultimate strength "" K N "" missile shape factor = 0.72

v ". velocity in ft/see

d ". missile diameter in inches

The validity of this formula was examined by comparing the penetration depths given by the formula with the results of 103 tests with solid missiles and 42 tests on pipe missiles with velocities ranging from 90 to 1000 fps as listed in Table 3 (Sliter 1979). It was found that for solid missiles. penetration depths greater than 0.6 of the missile diameter can be predicted by this formula to within !25 percent (see Figure 2). However. for smaller depths (x/d ~ 0.6). the NDRC formula predicted depths that were as much as eight times greater than those measured.

For the deformable pipe missiles. comparison between the measured and calculated penetration depths is made by adjusting the nose shape factor N and by using an equivalent diameter d in the formula. as shown in Figure 3. The best overall agreement is obtained using N ""0.72 and the outside diameter do of the pipe in the NDRC formula.

(b) For scabbing thickn~ss. ts' the modified NDRC scabbing formula gives the following relationships:

ts ""2.12 + 1.36 (!) for 0.65 ~!~11.75 (2a) d d d or

ts ""7.91 (!) _ 5.06 (!)2 for ! ~ 0.65 (2b) d . d d d For solid missiles. comparison with scabbing results has demonstrated that good agreement is achieved for small diameter missiles as shown in Figure 4(a). For large diameter solid missiles. scabbing occurs below the scabbing threshold predicted by the modified NDRC formula if the measured penetration depths are used as indicated in Figure 4(b). This means that for large diameter missiles. scabbing is produced by penetratJ.ons less than those predicted by Equation 2(b). However. if the predicted penetrations are used. there is apparent agreement between predicted and observed threshold of scabbing. as demonstrated in Figure 4(c). In other words. the disagreement

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• - 4 -

with test data for penetration and scabbing thickness fortuitously offset each other to give apparent agreement with observed values.

ror the deformable pipe missile. all the test data points confirming the appropriate applicability of the empirical NDRC curve. providing that the outer diameter do is used to calculate penetration and the equivalent diameter de is used to calculate scabbing thickness as shown in Figure 5. This figure also shows that use of de in both equations over-estimates the target thickness required to prevent scabbing.

(c) ror perforation thickness. tp:

~ = 1.32 + 1.24 (~) for 1.35 , ~ £ 13.5 (3a) d d d or

~ = 3.19 (!) - 0.718 (!)2 for ! ~ 1.35 (3b) d d d d For solid missiles. it was found that agreement. or near agreement is indicated for all test data except for large-diameter. heavy missile tests. For a large majority of these data. the modified NDRe formula over-estimates the wall thickness required to prevent perforation. Test data on perforation by deformable pipe missiles are not as much as data on scabbing. However. the NDRC formula using do in Equation 1 and de in Equation 3 gave results which are in general agreement with observed perforation or non-perforation. In summary. the general opinion is that. since perforated resistance is substantially affected by the finite thickness of a target (movement of shear plug) and is more dependent on the amount of reinforcing. the adequacy of the perforation formula for either the solid or pipe missiles should be reviewed when more relevant test data are available.

2.3 Highly Deformable Missiles It is well known that highly deformable missiles produce much less local damage than would be predicted by conventional formulas based on non-deformable missiles. The reason is that use of the formula implies that the resulting impact force is very much higher than the crushing strength of these types of missiles. Actually. during impact. the missile crushes significantly and the depth of penetration is reduced. This reduction in the depth of penetration. however. is expected to cause a lesser reduction in the required perforation and scabbing thickness.

Some of the highly deformable missiles considered in the design of nuclear power plant facilities are wooden poles. automobiles and aircrafts. The state-of-the-art procedures for design of barriers against these missiles relies first upon the accurate definitions of the interface forcing function by analytical methods. and then checking the punching shear capability of the concrete target (Riera 1968. McMahon 1977. Danay 1982 ASCE committee 1980).

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2.4 Dynamic Shear Capacity

The punching shear resistance mobilized by a concrete barrier due to impact of a missile was investigated by numerous researchers. It was generally found that the ultimate punching shear stress ranges between 9 .f'f::' and 12..fT? (ASCE-ACI Commit tee 1974. Albr itton 1969 and McMahon 1979).

Experimental evidence shows that severe local damage due to missile impact causes shear cracks to isolate an inner core of concrete. Subsequently. the shear transfer takes place primarily through aggregate interlock. It can be envisaged that. as the missile penetrates the concrete. the punching shear resistance offered by the core is reduced according to the following relationship: v .. vsb (d - x) (4) where

= shear resistance = ultimate punching shear stress :or circumference of a cylinder with diameter equal to diameter of missile plus slab thickness d total depth of slab x penetration One way to obtain a realistic estimate of the ultimate punching shear stress, vs' is to analyze test data. The report on full scale tornado missile impact tests prepared by the Sandia Laboratories provides a comprehensive set of material for this purpose. Major results for these tests were summarized in Tables 4, 5 and 6.

Typical missile decceleration of the pipe missile is shown in Figure 6. The maximum missile force can be calculated from the peak inertia value. Plots of shear resistance (v) against assumed depth of pene~tati?n (x) using Equation 4 are shown in F1.gure 7, with Vs equal t.o 10 fc'. It was found that for each test, the depth of penetration and the maximum missile force agrees very well with the observed damages. That is to say. if the peak missile force is less than the shear capacity corresponding to the depth of penetration, no shear failure occurs and the observed damage is confined only to cracking. If the peak missile ferce just exceeds the shear capacity. flaking occurs due to insignificant shear plug movement. However. if the peak missile force is very ~ch higher, concrete shear failure is experienced. resulting in large plug movement which is only stopped by the backface reinforcing and thus scabbing occurs.

In the pipewhip study, the section of the pipe impacting on the containment is "collapsible" and, therefore, is considered highly deformable. Following the methodology for the design of highly deformable missiles discussed in Section 2.3. the local effects on the concrete containment envelope due to a major pipe failure can be assessed by computing the peak impact load and checking with the dynamic shear capacity. The influence of the pipe deformability on the impact forcing function is demonstrated in the following sections. l700L - 6,-

3.0 ANALYTICAL CONSIDERATIONS FOR PIPEWHIP ANALYSIS

The overall kinetic energy acquired by the pipe as a result of the postulated break is shared- between a number of mechanisms which. most often that not. will require consideration of highly non_linear effects. These include both geometric conditions. like large displacements and rotations. and plastic range constitutive equations. The main energy absorbing mechanisms may be listed as:

(a) angular deformation of plastic hinges at various locations along the pipe. (b) ovalization of the pipe either at the impacted end or as a result of severe flexual deformations. (c) progressive crushing of the concrete under the large contact forces. causing some pipe penetration. (d) initiation of a compressive shock wave into the concrete. with additional effects such as spalling. scabbing and momentum transfer into the concrete barrier.

The simultaneous inclusion of all above mechanisms into the mathematical modelling. though desirable. may be highly impractical in the numerical solution of all but very simple structural components. Alternatively. since the forces during impact are mainly dependent on the pipe deformations. a good assessment of the contact forces time histories and of their global and local effects on the structure may be obtained by carrying out the analysis in two decoupled stages: non-linear dynamic analysis for pipe impact forces simulation of pipe-concrete interaction to determine local effects.

The first stageJwhich incorporates the mechanisms (a). (b) and (c) as described above. effectively result in the complete modelling of the pipe characteristics while the structure is represented only by an equivalent spring. On the other hand. the second stage. i.e .. pipe concrete interaction. will make use of the first stage results in order to simplify the pipe idealization while devoting most of the numerical effort to the solution of the local concrete response.

4.0 NON-LINEAR DYNAMIC ANALYSIS FOR PIPE IMPACT rORCES

4.1 Computer Analysis Since the forces during impact are mainly dependent on the stiffness of the pipe. the key to determine the realistic forcing function lies in the model of the pipe behaviour prior and during impact. Based on a number of tests on the distortion characteristics of several pipe sections loaded against a rigid surface (Peech 1977). the force-displacement relationships for the ovalization of an elbow and an open ended pipe could be established.

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The computer analysis may be carried out by using the following procedures:

(a) the pipe is represented by a finite element model;

(b) the jet thrust time history is included from the break occurrence:

(c) the concrete response during possible penetration is modelled by means of a non-linear spring with an assumed contact area;

(d) the non-linear pipe dynamic model makes allowance for the formation of one or more mid-span plastic hinges and consideration of ovalling effects at the impacted end;

(e) solution of equations of motion by numerical "techniques with the use of structural analysis computer programs to compute the dynamic response of the piping system and the dynamic reaction on the target structure.

Comput2r analysis of a sample case which involves a longitudinal break in a 22" diameter piping system was carried out by using a dynamic analysis computer program. The finite element model is shown in Figure 8. and the system results are summarized in Figure 9.

4.2 Simplified Method Verification In general. dynamic analyses involve numerical techniques. including non-linear material behaviour. large deformation and highly specialized skills on mathematical modelling. There are many computer programs with different numerical methods. which could be applied to the pipewhip problem. PIPERUP. ABAQUS and NONSAP are some of many available to-date for this purpose. In order to establish a "feel" for the validity of the results obtained from these programs. a simplified rigid pendulum approach can be used to study the computer results.

In this analysis. the pipe is modelled as a simple cantilever. with the sudden break at location B. as shown in Figure 10.

The pipe is assumed to rotate abcut a plastic hinge formed at some point A between the break location and the nearest anchor. The location of this plastic hinge can be determined by the simple relationship:

L=~ (5) F The maximum velocity at the tip of the pipe before impact is:

V=J6~Fa (6)

l700L - 8 - where

E'" jet thrust

Mp plastic moment L = distance from jet to plastic hinge

W = weight of pendulum

~ = angular rotation

V = velocity g gravity By using similar non-linear springs for the ovalization of the pipe and crushing of the concrete. the maximum impact load can be determined by equating work done in compressing the spring and plastic work at the hinge with the external work done by the jet thrust.

Table 7 compares the major results from the computer analyses with this simplified method. It can be seen that this simple pendulum method can be used to predict the peak impact velocity accurately. However. the peak impact force is very conservative because the secondary hinges are not accounted for in this simplified approach. With regard to local impact effect. this method can be employed to provide an effective mass of the rotating section by energy balance between the kinetic energy of the distributed mass whose velocity varies linearly along the length of the pipe and the kinetic energy of the effective concentrated mass at velocity V before impact. Thus.

Kequivalent = t ~ (7) With the use of this equivalent mass and velocity. the amount of energy imparting on the concrete structure can be estimated.

5.0 COMPUTER SIMULATION OF PIPE-eONCRETE INTERACTION

5.1 Dynamic Stress Waves Propagation It was discussed in Section 2.3 that. for highly deformable missiles. an estimate of damage 1n a concrete target subjected to missile impact can only be made by computer analysis to-date. The requirements for a computer program are that it 1s capable of modelling dynamic missile and target behaviour. large local deformations. as well as missile and target interaction effects. There are a number of computer programs which have the necessary features to analyze the local effects of impact on concrete targets (STEALTH 1976. PISCES 1975. Gupta 1977. Alderson 1977. Tu 1979). Since the loading is of short duration and the earliest part of the response is very important. an accurate and efficient numerical procedure to control the development of shock waves which occur due to the impact is necessary. Usually. finite difference techniques are preferred for the time integration.

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For the purpose of this study. the PISCES 2DELK computer code (Hancock 1976). available through Control Data Cybernet Services. was used. This program is capable of solving coupled equations of motion in a continuum by finite difference method using Euler-Lagrange formulation. It takes into account the laws of conservation of mass. momentum and energy allowing a material model relating stress to deformation in the non-linear range. Further details related to this program are contained in PISCES literature and. therefore. will not be repeated here.

5.2 Description of Mathematical Model

The dynamic analysis of the ruptured piping system for the longitudinal break has indicated that the pipe would slap almost flatly on the concrete vault ceiling across a length of over seven feet. As a result. it seems reasonable to use a two-dimensional model of the pipe-concrete by taking a unit width plane-strain condition. In addition. by taking advantage of symmetry. only one-half of the model is necessary. The steel pipe was modelled as shell elements (see Figure 11). The 60" thick concrete target was originally modelled as a Lagrange grid as shown in Figure 12. However. as the analysis proceeded further on. it was found necessary to extend the side boundary to account for the longer radiation and reflection of stress wave paths resulting in a model shown in Figure 13. The material properties used in the PISCES .analysis are as follows:

(a) Concrete

Compressive strength fc ' = 4 500 psi Young's modulus E = 3 800 ksi Bulk Modulus K '"' 2 110 ksi Shea r Modu 1us G '"' 1 580 ksi Sound speed C '"' 0.1371 x 106 in/see Density p '"' 150 Ib/cu ft Poisson's ratio v 0.2

(b) Steel Pipe

Yield stress fy 45 000 psi Ultimate stress fu 67 000 psi Young's modulus E = 26 470 ksi Bulk Modulus K = 22 058 ksi Shear Modulus G 10 180 ksi SOund speed C a .2194 x 106 in/see Density p 490 Ib/cu ft Poisson's Ratio IJ = 0.3 Total pipe weight w = 470 Ib/ft

5.3 Constitutive Models

Two types of constitutive laws are required for each material. They are (5) the equation-of-state to describe the hydrostatic pressure-volume change (or density change) relation and (b) the yield model to define the failure of material.

1700L - 10 -

For the steel pipe. the equation-of-state was considered a linear relation as shown in Figure 14(a). For yielding. the steel pipe was treated as elastic-linear strain hardening-perfeet plastic as shown in Figure 14(b). The constitutive model for concrete. is more difficult to define accurately. For this study, the equation-of-state is shown in Figure 15. The pressure at initial concrete crushing (Pcoll) and at completion of concrete particle compacting (Pcomp) were set at Pcoll = 4 fc' and Pcomp = 10 fc' respectively. The slope of line AS at compacting stage was assumed to be half of the concrete bulk modulus. This stepwise linear function was actually input using a polynomial function as required by PISCES. The yielding for concrete was based on the Mohr-Coulomb failure model as shown in Figure 16.

5.4 PISCES Results Two runs were made: (a) with initial pipe velocity equal to 300 ft/see. and (b) with pipe velocity of 73 ft/sec. Some seleeted results for velocity of 300 ft/see are shown in Figures 17 to 25.

Figures 17. 18 and 19 are stress plots illustrating the propagation of pressure waves as impact continues through snapshots taken at intermediate time frames. At cycle 800. it clearly identifies the formation of a shear cone behind the impact and concentration of tensile cracks within this region ("voids" appearing in the plots means material has exceeded the spall limits). FIgure 20 aemonstrates the sIgnIfIcant deformation occurring in the steel pipe. It has been practically flattened during the impact.

Figure 21 shows the interaction force resulting from the collision. This also represents the dynamic load applied on the concrete.

Figures 22 to 24 illustrate the plot of the transfer of energy and momentum between the steel pipe and the concrete target as a function of time. It can be seen that most of the kinetic energy in the pipe was spent in about four msec. The final sharing of energy is summarized in Tables 8 and 9 for the two velocities (300 ft/see and 73 ft/see).

Figure 25 identifies locations where tensile stress limit has been exceeded. It must be noted that these cracks are not formed simultaneously. because some stress reversals do occur. This figure is used merely to illustrate qualitatively that the three major tensile regions are generally similar to the damages experienced in some of the tests on missile impact. 6.0 EVALUTION OF THE CONCRETE LOCAL RESPONSE DUE TO PIPEWHIP

6.1 General Considerations Due to the complexity of the mechanisms involved in the evalution of the local structural response. it is apparent that no individual analytical method described therein is able to supply a definitive answer on its own merits only. Further analytical and experimental effocts are still required to expand the state of the art to more confident levels. Meanwhile. some judgement need be applied by using fundamental principles of mechanics and empirical data. as follows:

1700L - 11 -

6.2 Rigid Pipe Upper Bound Solution

It is obvious that the most conservative assumption in evaluating the local damage is to take the pipe as a rigid missile. Since this assumption implies that all the energy is used up in destroying the concrete. this definitely represents an upper bound solution.

From the simple pendulum analysis. the free impact length is 21 ft and. therefore. the effective weight is 6 570 lbs. Velocity at time of impact. as predicted by computer analysis. is 73 ft/sec.

Using similar assumptions regarding rigid pipe shape during penetration. the depth of penetration was computed using the NDRC formula to be 5.8 inches. Equivalent diameter at contact is 29.5 inches.

The modified NDRC formula predicts that the perforation thickness required is 46 inches. whereas the scabbing thickness is 70 inches. As a result. even with this very conservative hard missile assumption. the concrete would not be perforated. However. if the pipe is indeed rigid. scabbing would occur. possibly accompanied by large shear plug movement.

Since this rigid missile calculation is too severe. more realistic evaluation is deemed necessary. This may be achieved by making use of the PISCES analysis results.

6.3 Deformable Pipe Impact The non-linear dynamic analysis in Section3.l and the finite difference PISCES code both demonstrated that significant amount of energy would be spent in crushing the pipe at impact. Therefore. in order to provide a more realistic assessment of the local damage incurred on the concrete structure. this loss of energy in crushing cannot be ignored.

From the results of the PISCES analysis. it was found that approximately 69 percent energy was taken by the pipe. and only 31 percent energy was expended in the concrete. OHt of the 31 percent energy shared by the concrete panel. 27 percent was spent in crushing and penetrating into the concrete and 4 percent was left as kinetic energy in moving the centre plug of concrete.

If the PISCES analysis is to be continued. it would result in widening the major craCK extending roughly at 45° diagonally towards the back face. similar to that shown in Figure 25. This is because of the limitation of the concrete model which transcribes exceedance of principle tel~ile stress into cracks which lose all capability of carrying shear forces. However. it is well recognized that because of the size of the constituent particles and the roughness of the crack surfaces. the effect of aggregate interlocK can yield a significant capacity to carry shear stresses. Analysis has confirmed that the deflection for a target in which aggregate interlOCK has been ignored is excessively high when compared to experimental values obtained from a test (see Figure 26 extracted from Alderson 1977).

1700L - 12 -

The mechanism of shear transfer by aggregate interlock is also being studied by researchers in the reinforced concrete field. Tests have demonstrated that shear forces can be carried across precracked interfaces (renwick and Paulay 1968. Walraven 1981). Assuming. conservatively that the diagonal cracks are at 45°. the shear stress-shear displacement relationship is obtained as shown in r1gure 27.

From the peak interactive force computed by PISCES with impact velocity of 73 ft/sec. it has been found that the shear failure of the concrete is unlikely. The average peak force is 40 kips/inch. whereas the dynamic she~apacity 1s estimated to be approximately 84 kips/inch using 10 ~fc' as the limiting shear stress. In other words. it is not expected that the concrete plug would be dislodged.

However. if a plug movement 1s hypothesized. the maximum crack width could be evaluated by using the curve shown in rig. 27. Assuming conservatively that all the available kinetic energy left in the concrete is to be used up in shear movement. the maximum crack width was found to be approximately 0.02 inch and that the cracked area would be less than 1.8 sq. in. This. in turn. may be grossly overestimating the true leakage area. since. as illustrated in Figure 28. in order to develop the shear strength. the two concrete surfaces should be in contact with aggregate interlocking each other and thus making free passage of air almost impossible.

7.0 CONCLUSIONS It is expected that under a major pipe impact. localized cracking does occur in the concrete due to dynamic tensile stress wave reflected across the thickness of the concrete member. The actual crack size and pattern is difficult to evaluate. This is due to lack of experimental data on missile impact as well as insufficient information regarding dynamic strength of concrete. Therefore. the estimate of maximum probable cracked width could only be determined by using simplified. but cnservative. assumptions. It must be noted that. to our knowledge. the complex problem of assessing the leakage area has never been addressed by the nuclear industry to-date. The evaluations developed in this paper are considered reasonable. since a combination of rigorous computer analysis. empir1cal formulas. simple mechanics approach and engineering judgement are used.

1700L - 13 -

REFERENCES 1. Danay. A •. J.H.K. Tang. "The Validity and Conservatism of Applying a single Degree of Freedom Idealization for Impact Loads". OH Report ~78243. July 1979.

2. Kennedy. R.P .• "A Review of Procedures for the Analysis and Design of Concrete Structures to Resist Missile Impact Effects", Journal of Nuclear Engineering and Design, Vol. 37. NO.2, Hay 1976.

3. Nusbaum. M.S .• et at. "One Quarter Size Reinforced Concrete Missile Barrier Tests for Stone and Webster Engineering Corporation", lIT Research Institute Project J8184. May 1976.

4. Berriaud, C •• et aL "Local Behaviour of Reinforced Concrete Walls Under Hard Missile Impact". Fourth SMiRT Conference. paper J7/9. San Francisco. August 1977.

5. Alderson. M.A.H.G., r.L. Davis. R. Bartley, T.P. O'Brian, "Reinforced Concrete Behaviour due to Missile Impact". F'ourth SMiRT Conference. paper J7/7. San Francisco. August 1977.

6. Stephenson. A.E., "Full Scale Tomado Missile Tests", EPRI NP440, July 1977. 7. Sliter, G.E., "Assessment of E:mpirical Concrete Impact Formulas", ASCE Conference on Civil Engineering and Nuclear Power, Boston, Mass., April 2-3, 1979.

8. RoU, J.V .• "Results of Missile Impact Tests on Reinforced Concrete Panels", ASCE Specialty Conference on Structural Design of Nuclear Plant Facilities. New Orleans, December 1975.

9. McMahon, P.M., B.L. Meyers. K.P. Buchert, "The Behaviour of Reinforced Concrete-Barriers Subjected to the Impact of Tornado Generated Deformable Missiles". Fourth SMiRT. San Francisco. August 1977. 10. McMahon, P.M., S.K. sen. B.L. Meyers. "Behaviour of Reinforced Concrete Barriers Subject to the Impact of Turbine Missiles", Fifth SMiRT. Berlin. August 1979.

11. Berriaud. C •. et aL "Test and Calculation of the Local Behaviour of Concrete Structures Under Missile Impact", Fifth SHiRT. Badin. August 1979.

12. Degen. P .• "Perforation of Reinforced Concrete Slabs by Rigid Missiles". ASCE, Structural Division ST7. July 1980.

13. Godbout. P .. A. Brais. "A Methodology for Assessing Aircraft Crash Probabilities and severity as Related to the Safety Evaluation of Nuclear Power Stations - Phase III". Ecole Poly technique de Montreal, March 1980.

1700L - 14 -

14. Kar, A.N., "Barrier Design for Tornado-Generated Missiles", Fourth SMiRT, San Francisco. August 1977.

15. Gupta, Y.M .. L. seaman, "Local Response of reinforced Concrete to Missile Impacts", Fourth SMiRT, Vol. J, 1977.

16. Riera, J.D., "On the Stress Analysis of Structures Subjected to Aircraft Impact Forces", Nuclear Engineering and Design, No.8. 1968. 17. Danay, A., "Darlington GS A - Effects of Vehicle Impact on Steel Columns and Concrete Barrier ~all", O.H. Report No. 82411. July 1982. 18. ASCE Committee on Impulse and Impact Loads, "Preliminary Draft Report on Impactive and Impulsive Loads", Knoxville, Tennessee, 1980. 19. Peech, J.M., R.E. Roemer. S.D. Pirotin. G.H. East, N.A. Goldstein, "Local Crush Rigidity of Pipes and Elbows", Fourth SMiRT, San Francisco, August 1977.

20. "STEALTH", A Lagrange Explicit Finite Difference Code for Solids, Structural and Thermal Hydraulic Analysis". Report NP-176, SCience Applications, Inc., June 1976.

21. Abrams, J.I.. J.D. Stevenson, "proceedlngs of symposium on Structural Design of Nuclear Power Plant Facilities". University of Pittsburgh. December 1972. 22. Tu. O.K .• R.A. Larder. "Numerical Simulation of Tornado-Borne Missile Impact on Reinforced Concrete Targets", Lawrence Livermore Laboratory. UCRL-57607, February 1979.

23. "Dynamics of a Blunt Projectile Impacting Concrete at 1400 feet per second", Physics International Co .• San Leandro. Calif., April 1975. 24. Hancock, S., "Difference Equations for PISCES-ZDELK System of Computer Code Packages", Physics International Co., 1976.

25. Fenwick. R.C .• and Paulay. T., "Mechanism of Shear Resistance of Concrete Beams", J. of Ste. Div., ASCE. Vol. 94, No. STIO. Proc. Paper 2325, OCtober 1968, pp 2325-2350.

26. Walraven. J.C •• "The behaviour of cracks in plain and reinforced concrete subjected to shear, IABSE Colloguium on Advanced Mechanics of Reinforced Concrete. Deflt, 1981.

27. "The Shear Strength of Reinforced Concrete Members - Slabs", by the Joint ASCE - ACI Task Committee 426 on Shear and Diagonal Tension of the Committee on Masonry and Reinforced Cocrete of the Structural Division. Neil M. Hawkins, Chairman. Journal of the Structural Division. ASCE, Vol. 100, No. ST8. August 1974.

1700L - 15 -

28. Albritton. G.E •. "Deep Slab Subjected to Static and Blast Loading". Journal of the Structural Division. ASCE. vol. 95. No. STll. November 1969. 29. Physics International Co .. Internal Memo from S. Hancock to M. Cowler. March 21. 1981.

l700L TABLE l.--Major Missile Impact Parameters Affecting Local Response

Weight Size ~issile Characteristics Shape Oeformabil ity I I-' Velocity (along line of flight) ':J) Rotational velocity Impact Conditions Missile orientation at impact Obliquity (angle between line of flight and target surface)

Thickness Target Characteristics Unsupported span Concrete properties Amount of reinforcing TABLE l

SUMMARV OF IMPACT TESTS ON CONCRETE SLABS

mod e 1 5 C ale f u 1 1 - 5 C ale Range of Range of Range of Range of Range of \ Oataset velocities masses K diameters masses diameters I (m/s) (kg) ( an) (kg) (cm) actual aircraft 15-183 ---- 1 ---- 65-3050 20-140 u.s. DATA NDRC, 1946 280-305 NA NA NA 0.75-454 4-31 Sandia, 1976 28-132 ---- 1 ---- 3-337 2-33 Calspan, 1975 37-150 ---- 1 ---- 60-98 20 ~ -...J BRITISH DATA UKAEA, 1976 60-138 1.6 1/25 12 25,000 300 FRENCH DATA Preliminary, 1975 27-30 ------334-343 11-16 EDF, 1975 80-160 50-227 1/2 30.5 400-1,816 61 EDF, 1975 92-215 30-100 1/5 20-30 3,750-12,500 100-150 GERMAN DATA Part I, 1975-76 100-1.700 ---- 1/2 -1 ---- .020-.340 1.5-4.0 Part II, 1976 150-427 1.11,2.25 1/3.5 4.5 48, 97 15.75 In progress, 1977 to 300 ---- 1 ---- to 1,000 to 60 AUSTRALIAN DATA UofSydney, 1974 to 11 53-104 NA 7.5 ------TABLE 3.--Range of Parameters from Recent Impact Testing with Solid and Pipe Missiles

. Source Number Missile Missile Missile Target Target t/d of data of tests weight. ,~ diameter. d velocity. V thickness. t strength. fl in pounds in inches in feet per in inches in pounds c second per square inch (I) (2) (3) . (4) (5) (6) (7) (8) Solid Missl1es "

EOF/II (2) 4 750-756 4.3-6.1 89-96 6.9-13.8 5410-6190 1.1-2.7 ~ EOF/l1I (2) 32 352-661 1.9-12 213-567 15.1-23.6 4860-5800 1.3-3 ro CEA (2) 16 38.5-125.5 1.9-11. 8 324-1013 8.2-10.2 4930-6450 0.1-1.3 B(CIHEl (22) 7 214 8 122-371 12-24 4500-5770 1.5-3 S&'~(8) 7 12.5-23.5 3 89-205 4.5-6 3210-4400 1.5-2 OECIlTEl (1) 5 8 1 150-220 3-9 4810-5940 3-9 EPRI (18) 2 8 1 303-435 12-18 3550-3650 12-18 EHI (12) 20 2.5-5 1.8 495-1020 9 4730-1130 5.1 EHI (11) 10 0.24-1.97 0.8-1. 6 479-955 9.8-15.1 5140-6130 8.7-15 All sources 103 0.24-756 0.8-12 89-1020 3-24 3210-1130 0.7-18 Pipe Miss l1es

EPRI (18) 12 78-743 3.5-12.7 92-213 12-24 3340-4690 0.9-3.4 OECIHH (22) 9 132-210 8.6 135-475 12-24 4400-5770 1.4-2.8 S & W (8) 21 4.2-28.2 1.6-3.5 159-517 4.5-6 3210-4400 1.5-3.5 All sources 42 4.2-743 1.6-12.7 92-517 4.5-24 3210-5770 0.9-3.5

2 Note: 1 lb = 0.453 kg; 1 in. = 2.54 em; 1 fps = 0.305 m/s; 1 psi = 6.9 kN/m Table 4

SUMMARY Of' 'l'OHNAOO IMPACT TEST DATA

MAX. lUSSILE PANEL CONCRETE FRONTFACE BACKf'ACE TEST NO. MISSILE WEIGHT VELOCITY DEFOHHA'rION TIIICKNESS STRENGTH PENE1'RATION DAMAGE (lb. ) (fps) (in.) (in .1 (psi) (in. )

6 (a) Utility Pole 1500 205 50.5 18 3770 0 Cracks

7 Utility Pole 1470 204 69.4 12 3620 0 Cracks

1 1" Rebar 0 30) 0 18 )650 ).6 No Damage 13 1" Rebar 8 435 0 12 3545 5.8 Cracks

5 )" Pipe 78 212 ),9 12 ))40 4.6 Cracks I t-' W 8 12" Pipe 743 202 l.B 24 3795 6.8 Cracks

3 12" l'ipe 743 202 3.B 18 3350 7.0 Scabbing

4 12" Pipe 743 198 3.8 18 3560 6.8 Scabbing

12 12" Pipe 743 20] 5.0 18 4535 7.5 Scabbing

18(bl 12" Pipe 74] 213 2. :) 18 4420 9.1 Scabbing

19 (c) 12" Pipe 743 222 9. l 18 3900 4.3 Cracks

17 12" Pipe 74] 157 o. :!5 18 4070 4.1 Flaking

15 12" Pipe 743 152 O. l3 18 4205 5.) Flaking

9 12" Pipe 743 143 2. l 18 4325 5.0 Cracks

10 12" Pipe 74 ) 143 0 12 )(,90 12.0 Perforation

11 12" Pipe 743 90 0 12 )595 4.5 Scabbing

16 12" Pipe 74) 92 0 12 3350 3.5 Flaking

14(d) 12" Pipe 743 92 [) 12 3545 ).9 Flaking

(a) Ballast pipe attachment broke; full weight not effective. (h) Impact atorcuar intel"section (cl lmpilct 45 to panel face. (d) Impact near panel edfJc. Table 5 COMPARISON OF MEASURED AND COMPUTED PENETRATION DEPTIIS

MISSILE ClmRACT~RISTICS PANEL CHARACTERISTICS PENETRATION (in.) CONe. WEIGII'f VELOCITV TIIICKNESS STRENGTII . a PETRya NO. TVPE (lbs.) (tps) (In.) (psi) ACTUAL NDRC 1 I" rebar 8 ]0] 18 ]650 ].6 ].2 13.0 I 13 I" rebar -8 4]5 12 ]545 5.8 5.0 12 tv o 5 ]" pipe 78 212 12 ]]40 4.6 4.2 (6.2) ].7 (>12) ] 12" pipe 74] 202 18 3350 7.0 7.6 01.4) 2.4 (>18) 4 12" pipe 74] 198 18 ]560 6.8 7.3 (11.11 2.4 (>18) 12 12" pipe 743 203 18 45]5 7.5 7.0 lIO.7) 1.7 (>18) 8 12" pipe 743 202 24 3790 6.8 7.3 (ll.ll 2.2 (>24) 9 12" pipe 743 143 18 4320 5.0 5.2 (7.9) 1.1 (12.9) b b 10 12" pipe 743 143 12 3680 Perforated 5.4 (8.2 ) 1.2 (>12) 11 12" pipe 743 96 12 3595 4.5 3.9 (5.B) 0.5 (4.]) 14 12" pipe 743 92 12 3545 4.5 3.7 (5.6) 0.5 (4.5) 15 12" pipe 74] 152 18 4205 5.3 5.5 (8.4) 1.2 (>18) 16 12" pipe 743 92 12 ]]50 3.5 ].7 (5.6) 0.5 (4.6) 17 12" pipe 743 157 10 4255 4.1 5.6 (8.6) 1.] (>18) 18 12" pipe 74] 213 18 4690 9.1 7.2 (11.1) 2.0 (>18) c 7 Utility pole 1470 204 12 3770 0 9.4 4.0

aparentheses denote use of metal area (d = d ) in formulae. bNDHC pre dOlcts per foeoratlon. CUtility pole test NO.6 omitted (ballast attachment broken, effective weight indeternlinate). Table E COMPARISON OF MISSILE FORCE (INPUT) AND MEASURED REACTIONS (OUTPUT) FOR 12-IN. PIPE MISSILE

Slab Missile Penetration Max. Missile Sum of Test Thickness, t VelocitYI Depth, X -X Backface Force, F Reactions, R No. (in.) (ft./sec. ) (in.) t Damaqe (Kips) (Kips) R/F

B 24 202 6.0 0.263 Cracks 1045 420 0.41 N i--' 16 10 213 9.1 0.506 Scabbing 762 305 0.46

3 10 201 7.0 0.309 Scabbing 706 No edge beams -

4 16 190 6.0 0.378 Scabbing 694 354 0.51

17 18 157 4.1 0.228 Flaking 650 380 0.58

15 10 153 5.3 0.294 Flaking 645 355 0.55

9 10 143 5.0 0.278 Cracks 561 313 0.54

11 12 96 4.5 0.375 Scabbing 370 180 0.49

16 12 92 3.5 0.292 Flaking 203 176 0.62

0.52

Average Value TABLE 7: CXMPARISOO OF N(N-LINEAR ANALYSIS RESULTS

APPROXIMATE NCNSAP PIPERUP METHOO

T.iJre of nnpact (m sec) 99 86 91 .

N N Impact Velocity (fps) 207 197 196

.

Peak Impact Force (kips) 1,100 1,850 3,095 TABIE 8: STEEL PIPE - cetK::REI'E ENERGY FOR V=300 FT/SOC

Concrete Wall ('rarget) Steel Pipe (Missile)

Kinetic Energy Internal Energy Kinetic Energy Internal Energy (Strain Energy) (Strain Energy)

Due to Due to Due to Due to (Shear) Vollllle (Shear) Vollllle DeforTnation Change Defonnat ion Change

Initial ('1'=0) 0 100%

I 100% 0 0 N 0 0 0 W

36% 64% At tennination of Analysis 54% ('1'=0. 004472 32% second) 4% 29% 3% 10% 54% 0

Prediction at Full Stop or Rebound of Pipe 40% 60%

4% ]6% 0 60% 0 TABLE9 : STEELPIPE - a:::NCRE1'EEl'JffiGY FORV = 73 PI'/SOC

Concrete Wall (Target) Steel Pipe (Missile)

Kinetic Energy Interval Energy Kinetic Energy Internal Energy (Strain Energy) (Strain Energy) Due to Due to Due to Due to (Shear) Volurre (Shear) Volute Defonnation Change Defonnation Change . Initial 0 100% ('1'=0) IV 0 0 0 100% 0 0 "'"

At Termination 31% 69% of Analysis ('1'=0.0054 second) 4% 27% -2% 67%

18% 9% 67% 0 . Prediction at 31% 69% Full Stop or Reoound of Pipe 4% 27% 0 69% 0 - 25 -

'- f c !

+I 1 I a l' 1

I t

Front Fal.:2

Damage Char.act~ristics [ Symbol , I 1 Diameter or loose concrete at front surface. Neck diameter. I I Neck Depth. I I~I I d Depth to solid plug, e Diameter of loose concrete at back surface, I I f Major diameter of plug.

g Rear surface translation, \ h Rear rebar translation, I 9 Concrete plug angle, \,-

Fig. 1 Barrier DarrageCharacteristics. 9 0 a ~ EDF (2) ~ CEA(2) • Bechtel (22) 1 a Y Cl Stone &. Webster (8) o 6echtet (1) e ~ EMI (12) 8 A. EMI (11) (filled symbols 5 denote scabbing) XM:)RC -;( 4-

:3

2

1

0 2.0 0 0.5 1.0 1.5 xld

Fig. 2 Ccrnparison of calculated and rreasured penetration depths versus penetration for non-defomable cylindrical missiles. - ?o1

1.5 , _l1el ~.'2.1'Sil'I. 1.5 f.1Zitt. "~ t- 'Sin. '4 0 t-2 ..... --,~..." 1.;1 1'1J~.J88S,*

1.0 < ;; 0.5

~::~".:.':N -0.72. d. d1 I o.s ] 0.0 . 0 I I o.;z 1 I 0 Q 100 ::00 JOO '00 500 'ilro.l

'5 I I _<221 ~.&.1S21/'l. 1.1 o r.'2ift. o t.1...... __ I 1.4 • fe2."", \ .... I "'-0..7"2.4 -

'.0 < ;; 0 o.a I "- ••...0.:"2.1,J o.S I I Q.4

o.;z

a ~ Q '00 ::00 JOO 41 500 v,,=»

'.5 I I SIcN'~18) '.S ~.Jdl. .1-.5". • t-&'o1I\. I '.4 I~ ...... , -1 i - c..eu&at-e .. 1ft '.;I !": :..... ~ a-

Ul I !. !'\I-o.n..d-d. 1 I .. o.a

o.s J I 0.0 1 o.;z J I

0 I 0 '00 = JOO 500 'I~

Fig. 3 ~.easured and calculated (NDRC)~netrz.ticn de~tl;..s versus im?act velocity for steel 9iL:es. 28 -

_1- • ..L 1 0 I' T

75t'Vm 0_,,,o a..::n. (22) Q EPRt 1181 .:,5100 o ,:ftIeO~-..-..,

a a 1Q 12 III 18 :'0 Ud

CcIn;Jariso!1of NDRCscabbing fODnula with data on ncn-de:o~le cylindrical missiles wi~~ t/d ~ 3.

1.5 I -1-. 2>

Vd-J

1,0. o _l22l C .sra-. 'N.-w11l1 ~ _,Cl!Al2l • --_..-.01

a a

Fig. 4b Or:parison of -t-lDRC scabbing formula wi~" data on non-ucfornu1>lc cylint1ric.:ll m.i::;::;ile:::;L/d ~ 3.

,,5 I O~(271 C ~,weoaer 1,8' ~ ecF, CEAIZI --_...-.., 1.0.

0.5

a a 'IC

Fig. 4c Canparison of NDRCscabbing formula using calculated penetration depths fram data for non- deformable cylindrical missiles 'Nith t/d ~ 3. - 29-

I .;. 9"lII CSI:Ia ~181 ~.12.i'S'" 11\11.. JV"'IOOI_ denot. tc.IDOIt'OJ

~ .....I I !

lell

0.5 1 ~ 2.0 l.! 3.0

'..Ie!,

j I I .: SIaN' WeMtlf' aata Id' ~ _3.n. i mGeQ S)'fftCJOIS c:enoc. scu:t.,nqj J I j r- zoo I I I

i -- :-.e;::.:; :1 - :.; ,I --- 'jCRC .J. - ~,: I 'ell ~_i"!'!l:1'~ I •• -- .... Slon." ~.~ I I

o ! o :).5 1 a ::'0 0.0 1..5 .0

'J«ft~r: -.c;;c i // .~.-'J I I 1 I 4, • $.52 1". I -i \I\I' .. ~~ I /. oenot:.~J I -;/f I - i ;.1 ~L '--1 it -i II1/ .' 1 I I I: I -i J /1 / I~S4W : 1 -I I 1 -- ~~,~~-a,; : I --- ~CRC ,a, - -=.. i J _.-3lIcnt ...~1 • "I ,L ...... slone,.o'V.cst .... C'l"""lJt.1 o , 0 IS 30

fig. 5 Canparison of r->redictCYi and obser\le ..-~s~~3.bLi:lc

l i.~it.x.: rnissil.c :,.). - )0 -

1500.00

1000.000

(!) z z o ~ 500.000 a:: W ..J W (,) «(,)

o

-500.000 30.0 o 10.0 20.0 TIME IN MILLISECONDS

Fig. 6 Typical r1.issiles ceceleration - 31 -

\ \ 1500

'-\ \ \ II \ (/) \ -a. , \ ' ~ \\ \ ~\ >- \ I \ ~ \ o \ u \ \ ~ 1000 \ o IC \ <2: \ \ U ~ \o \ )I " a:: \ \ \ !/"TEST No 8(24 SLAB) <2: '\ \ w \ :I: (/) '\ W \ \. ~ \ i LLl '1 a:: U Z 500 8 TEST No 18(18 SLAB'j > ,,'~TEST N 3 (18" SLAB) Jt

TEST No 16(12"SLAB I o 5" 10" 15" 20" x PENETRATION (lNJ

Fig. 7 Shear Capa~ity versus Penetration. \51l

y

11

x

31

W N I

LEGEND I -NODE LOCATION ll 6 -MASSPOINT LOCATION r-r 2 rtfY1"-r f 6 -RUPTURE BREAK rrrti~ LOCATION

Fig. 8 Nathematieal t-'loclel - 33-

920 900 890

200

150

100

o 2095 216 220 196 210 213 190 200 m see

VELOCITY AT TIME OF IMPACT 73 ft /sec . I

1.87' 2.71' F2 Fl I 0.89 F3 FI F2 F3

~BREAK LOCATION

Fig. 9 Instantaneous Forces on Building During Impact - 34-

x

ANCHOR

~

A-LOCATION OF PLASTIC HINGE B -LOCATION OF POSTULATED PIPE BREAK L -LENGTH TO PLASTIC HINGE F -SLOWDOWN FORCE X -REFERENCE COORDINATE a-REFERENCE ANGLE

Fig. 10 Simple Pendulum System. - 35-

7', \ , ..) INTERNAL DIAMETER = 22 INCHES THICKNESS = 1.5 INCHES c:.

\ ' oJ

,.-..." t ~ J ,-; r ! .../ I ''\ \j 16 () 7 5

J

12

13

oJ I ... .:> FACE OF CONCRETE WALL (j

Fig. 11 Shell Sul:grids TRANSMITTIVE BOUNDARY " I'" 60 ,.1

I lJ 0\

o 10

lll...... ~,\ I

Fig. 1 3 Extended SlllxJrids. Fig. 12 Finite Difference Slll.xjrids. - 37 -

(hydrostatic pressure) 7 P KS = 2,2058 x 10

modulus)

(relativ~change of density)

Fig. 14 (a) Equation-of-State Model for Steel

0.00462 Mbar ( 67 KSI )

0,0031 Mbar (45 KSI)

6 0'17% TENSILE ST~AIN

Fig. 14 (b) Strain Hardening Yield ~.cdel for Steel - 33 -

p (hydrostatic pressure) C • I I P COMP r------E~8:

~SLOPE : I COMP I ___ A /: P COU lI \ / I I I / I

I / : I ~ I i ~K I / I " I / -1' IJ ( REL.-AT IVE ------.- CHANGE OF ~ COLL INCREASING DENSITY Ci::NSITY) ( DECREASING VOLUME)

Fig. 1 Equation-of-State Model for Concrete.

'( (yield stress)

5fc

4

'( max : 23,500 psi 2

P (hydrostatic pressure 5 6fc

_I I- SPALL LIMIT 500 psi

NOTE: fc = UNCONFINED COMPRESSIVE STRENGTH = 310 bar (4,500 psi)

Fig. 1 Mohr-Coulamb Yield Medel for Concrete. - 39

CYCLE 50

.....---_ ..

I .1 I .1

... _ .. -,- ~._'~..- ....--.-_.- CYCLE 100

Fig. 17 Initial Tensor Stress Vector Plots. - 40

.._-.----_ ...... •.••• ,~. I --:~::.::..•...... j

j ...."'lI

_ ..._ .... "., ..., ..,~ I ..... ~-""'-' . CYCLE 450

....,. ~ " " .

...... J I ! ~.~.-., I i ...... ---- ...... _ ...... , .,,-_ .., -1I .. u~ •••••... , l CYCLE 500 •. •••• --- ••.•.••••• "1 ~ ~ '•. ,. ''' __._ J 1- , ! ...... -...... _. ~ ....~..,...~<-...... "j ! ::~:::'_~~.-:~::~~~~:::~::)

Fig. 18 Inte.rrrediate Tensor Stress Vector Plots. _____ -1 _

I .... ~ I ...... --1

J"-._a --1 I '~"'."""""1 .-- , 1

.=~:::~::::::~~ CYCLE 750 . --_----.-•.•.••• _--l-i ::::=:: ::~:~~::::~::~~.1 ... _- ~~ _ . :::::?o:~~~.::j

Li;It:f?~~~;:i.__ • ...... ~_" J" -.... -. ::.~::--~. .~~ :; ~;;~;~.~:!~~~€f!f~ :::. <:::::::;<..,,-~:.: -

----.---- -_._._-j -..__ ---i _ ~ _-~ ._ , ii ."..."._ , 1 .- .:~:~~.~.~:~''''J CYCLE 800 -~::~::.~._.._ J ...... _c..._ 1 ~~~~:~~~:;~~::~:j f...... _.".~--- _...... ; ::;7::~ :7:.~--~.~-~~-.--..:.:.-~._~.- .""t,.~:~~~~~~:-:.-.• ~".':~~~:~~~.~ • •••...... - .- .•- -..... I ..~

Fig. 19 Final Tensor Stress Vector Plots. - ~2-

Fig. 20 Final Pipe Deformation. -43 -

F (\6':7)

100,000

50,000 TC see)

0.004 C005 a 0.001

Fig. 21 Pipe-Concrete Interactive Force. \>J( \6-it1)

35 ------5 3 x 10 ENERGY TRANSFEREO TO CONCRETE WALL 25 I .l'> .1.. 2 xl05

INTERNAL ENERGY STORED 1.5 .n PIPE DUE TO ITS DEFORMATION

I xI05 KINETIC ENERGY .5 REMAINED PIPE --nsec)

0007 0008 0009 0.01 o 0001 0002 0.003 0004 0.005 0006

Fig.22 System Energy Transfer. .'

'WcOb- i 1'\) COMPUTED TOTAL ENERGY TOTAL ENERGY IN CONCRETE WALL OVER ESTIMATION IN CONCRETE WALL OBTAINED BY SUBSTRACTING ENERGY OF ENERGY IN IN PIPE FROM INITIAL TOTAL ENERGY CONCRETE WALL 5 I-50 xlO i-----:--_.------1,25 \ ------

I-OC KINETIC ENERGY \_-----~-""'-::--.1 .".r- I ---- Ul INTERNAL (STRAIN) ENERGY "" ,(see) -25 0006 0009 0.01 0.004 0.005 0.006 0007 o 0.001 0-002 0.003

TIME ( SECONDS)

Fig. 23 Energy Absorbed by the Concrete Target. ~ ' I (lb. see) . 'II- ! 2.0 x 105

==----~ , ------I 1-4 I I MOMENTUM TRANSFERED 12 *"- TO CONCRETE WALL I a'I 10 !

0.8

0.6

04 I MOMENTUM 02 REMAINED IN PIPE I T(5€C) o 0.001 0002 0003 0.004 0.005 0006 0007 0008 0009 001

Fig. 24 M::IrentumTransfer. - I

- ~7-

Fig. 25 Tension Zones D1. Concrete. '-Cl1~- III r;0 !'i ~ .~ il_ tAl 8 I I I / .. -- I I

,. • Tarpt thiol

l: • Duration or Impact

o .5 2;0 o 0.'. 1.0 TOO: (t)

Fig. 26 I:A? f lection-Tirre Curves for Bending fvI£xleof r~sponse fran l\ircraft Impact Loading (l\LDERSCN et al (1977)) - 49 -

"WALRAVEN" 600

/ ~---/...... -In / ...... c. 500 ...... "FENWICK AND PAULAY" (/) ...... (/) 400 ( fc' = 3000 psi) W ...... a:: " ...... ~ (/) 300 "FENWICK 8 PAULAY" a::

0010 0020 0030 0.040 0.050 0060 1J.070 0.080 DISPLACEMENT (inch)

Fig. 27 Shear Stress - Displac~~t

l-GENERAL ORIENTATION OF CRACK I I I 1 1

t::. SHEAR DISPLACEMENT

CRACK WIDTH

Fig. 28 Displacerrent along a crad

CONTAINMENT INTEGRITY PROGRAMS*

Thomas E. Blejwas Walter A. von Riesemann Systems Safety Technology, Division 6442 Sandia National Laboratories Albuquerque, New Mexico 87185, U.S.A.

Introduction

The behavior of containment buildings of U.S. light water reactors (LWRs) during postulated severe accidents is being addressed in three coordinated programs at Sandia National Laboratories. The programs are being conducted for the U.S. Nuclear Regulatory Commission (NRC) by the Systems Safety Technology Division of the LWR Safety Department at Sandia. Through combined experimental and analytical efforts, the leakage and structural behavior of the containment building is being characterized ..

The containment building may leak during postulated accidents through a variety of mechanisms. The programs at Sandia are formulated to cover many of the most likely leakage paths. A program entitled Containment Safety Margins is addressing the overall structural and leakage behavior of the containment building, including the largest penetrations. The leakage potential of electrical penetrations is the focus of the Electrical Penetration Assemblies Program. Other penetrations, including personnel locks and pipe penetrations, will be included in the program entitled the Integrity of Containment Penetration under Severe Accident Loads that is presently being formulated.

Overall Containment Behavior

To assess and improve methods for predicting containment structural behavior during severe accidents, experiments and

This work performed at Sandia National Laboratories supported by the U.S. Department of Energy under contract number DE-AC04-760P00789 for the U.S. Nuclear Regulatory Commission, Office of Nuclear Regulatory Research. analyses are being conducted as part of the Containment Safety Margins Program. The experiments consist of tests of models of entire containment structures at reduced scale. Comparisons of analytical and test results are providing a basis for qualifying the analytical methods. Leakage measurements taken during pneumatic-pressurization experiments will provide qualitative information on containment leakage through structural tears and seals around equipment hatches.

Steel, reinforced concrete, and prestressed concrete containment models will be included in the analytical and experimental efforts. The loadings being considered in the program are static internal pressurization, dynamic internal pressurization and seismic loadings.

Of the three types of loadings being considered in the program, static internal pressurization is being addressed first. Because it is desirable to obtain data on the potential for leakage prior to catastrophic failure, pneumatic rather than hydraulic pressurization is being used. Safety concerns have dictated a remote outdoor testing site. The selected site is on Kirtland Air Force Base about five miles southeast of the Headquarters Building of Sandia National Laboratories, Albuquerque.

Program Objectives

The overall objective of the Containment Safety Margins Program is the qualification of methods for reliably predicting the capability of steel and concrete containment structures to function under loadings caused by severe accidents -and extreme environments. The suitability of existing analytical methods is being determined by comparisons with test results. Improvements to existing methods will be undertaken where deficiencies are found.

The end results of the program will be:

Benchmark data from tests of models of containment structures or features.

A set of qualified methods for estimating the structural capacity of containment structures.

Limited data on the leakage of containment models subjected to internal pressurization. This data can be combined with leakage data from the other NRC sponsored programs.

Near-Term Programs

The testing and analysis of steel containment models subjected to static internal pressure constitute the first phase of the

-2- program. The experimental set-up for testing prototypical models of about 1/30 the size of full-size BWR MARK III or PWR ice condenser steel containment structures has been completed. Pneumatic, quasi-static, internal-pressurization experiments with three model configurations (Figure 1) are nearly complete. Analyses of each experimental model is being completed before it is tested.

Design and fabrication of a large steel model of about 1/8th size (Figure 2) has been completed by Chicago Bridge & Iron Company. Delivery of the model is expected in early November 1983. Testing by quasi-static pneumatic pressurization will be started in March 1984. As with the smaller steel models, analyses will be completed prior to testing.

The second phase of the program is the testing and analysis of reinforced concrete containments subjected to internal pressurization. An assessment of the state-of-the-art of the analysis of concrete containments has begun; critiques and comparisons of sample analyses using various computer codes and techniques will be the end products. Work on the experimental modeling of concrete containments has also been initiated. Preliminary and final designs of a reinforced concrete containment model will be completed by the end of this year.

Results

A test to check out the experimental hardware and software for the small steel experiments was completed in December 1982. The model, which was a clean shell configuration (Figure 1), experienced significant plastic deformation. When the maximum circumferential strain reached about 20%, the model failed completely. Only limited data was successfully recorded during the experiment due to software problems.

A second model of a clean shell with full instrumentation was tested in April 1983. The model maintained structural integrity through gross membrane yielding. When the maximum global meridional strain was about 7%, a tear occurred near a repaired weld in the meridional seam. The model could not be further pressurized with the existing equipment. Data from this experiment are presently being compared with analytical results.

A model with ring-stiffeners was tested in July and August of this year. The first pressurization resulted in a tear due to improper grinding of a weld. The second pressurization after the model was repaired resulted in a complete failure. Data from the experiment indicates a significant change in material behavior between the two tests. Although plastic strains in the order of two percent were noted during the first testing, the complete

-3- ,,"

failure during the second testing occurred with very little measured plastic deformation. Strain aging may has caused a change in material properties; material tests to investigate this possibility are presently being conducted.

Leakage Through Electrical Penetration Assemblies

The purpose of the Electrical Penetration Assemblies Program is to assess the potential for containment leakage through electrical penetrations during severe accidents. A study of the numbers and types of' electrical penetrations has been completed. A limited number of tests of actual assemblies are planned. Combined pressure-temperature environments will be induced with a steam source.

Preliminary results from a study of the types of electrical penetration assemblies in use indicate the following. Many assemblies have both inner and outer seals; the outer seals will experience relatively low temperatures in many accident scenarios. Some assemblies have glass-to-metal seals that are likely to function at very high temperatures. The operational performance of electrical penetration assemblies during containment leak rate testing is statistically better than that of active penetrations (e.g. equipment hatches and personnel locks). However, some assemblies use sealing materials with service temperatures as low as 3500F. Testing in this program will be focused primarily on those types of assemblies that are available for testing and that have a high potential to leak because of loss of strength of the sealing materials due to high temperatures during postulated severe accidents.

Containment Penetration Integrity

The objective of the programs on the Integrity of Containment Penetration under Severe Accident Loads is to improve our understanding of the potential for penetrations (other than electrical penetration assemblies) to leak during severe accidents. Background studies for this program were performed by several U.S. national laboratories. Plans for the program are presently being formulated at Sandia in conjunction with the U.S. NRC.

-4- ..

HEMISPHERICAL DOME

THICKNESS 4.8mm

EQUIPMENT t- HATCH 5 PERSONNEL 1/ - LOCK E!l! (1 OF 2)_I Na: .,w "z'0 TYPICAL ::i PENETRATION u>- (1 OF 8) -J W Q STIFFENING -- o RINGS ~ C) z it 1.67 m ::l l- X o 0 ii:

:;

L1-1.17 m----J CLEAN SHELL RING REINFORCED SHELL WITH SHELL PENETRA TIONS 1/32 SCALE STEEL MODELS 1/8 SCALE STEEL MODEL

Figure 1 Figure 2 en m en en zo

-is:'" G.l~ IZ -i-i Zm mZ (J):t> (J)Z C> m o ", ~ A Experience with different containment leak detection methods and auxiliary system tests by European nuclear power plant operators.

by H.A. Maurer , CEC , Brussels P. Mostert , KEMA, Arnhem K.P. Termaat , KEMA, Arnhem

SUMMARY

Five different power plant operators of light water cooled reactors in five European countries reported in a study about their experience in aLJplying available methods and procedures in the periodic leak testing of containment systems and in particular on their experience in the application of available methods to extrapolate lea~ rates found at test pressures to leak rates at accident pressures. The main objective of the study was the containment shell itself but attention was also given to the isolation valve system, the heat removal systems and the ventilation and air filter systems. The study gives an inventory of national specifications, regulations and guidelines applied in the design, manufacture and testing of the different containment systems. The results of the study are analysed and particular attention is given ~o con~inment leakage rate determination.

Contair~ent leakage rate determination

The L~itial leakage rate tests of the reactor containments are in general performed at design pressure immediately after the pressure test, with penetrations and locks provisionally closed. Besides the inital test a final overall leakage rate test is performed when all penetrations and equipments are installed. The final leakage rate test and the periodical tests after a certain period of operation (so-called type A tests) are generally performed at reduced pressure because of possible damage to some of the installed equipment and instrumentation.

Divergent views were expressed in the study concerning the necessary pressure for the periodic leak tests. Some experts believe that it is still possible to obtain a sufficiently accurate representation of the leak at a test pressure of 0,5 atm. overpressure since, assuming turbulent flows, the extrapolation to the design pressure seems to produce reliable results.

Others are convL~ced that ~~e generally applied test pressure for period~c leakage rate tests of 0,5 atm. is far too low for measurement reproducibility and want a pressurization of the containment also for periodic leakage rate tests at half the accident pressure or at the design pressure. Operators of prestressed concrete containments believe that the periodic leak rate test at design pressure is the best way to represent the accident conditions with satisfactory precision.

The test periods of the leakage rate ~ests are generally fixed by the responsible national authorities in a case by case decision based on the results of the foregoing tests. -2-

The absolute method, and in some cases for cocparative reasons also the reference method, is applied to verify the leakage rate.

The paper refers also to local tests for detection and measurement of leaks from certain special components, the so-called type B tests (to measure leakage through containment penetrations) and type C tests (to measure leakage through the isolating valves).

Auxiliary systems

The ventilation and air filter systems are regularly inspected by the shift personnel (once every month or once every 3 conths). In all the nuclear power plants considered the particle filters which are installed to reduce the airborne activities in the building are tested by the DO? (0ioctyl-Phtalate) method or other methods. Active charcoal filters are nor-~lly installed for iodine retention; they are tested by special procedures, as e.g. Freon-12 sample tests (once every year) and functional tests (once every 2 years).

The study included investigations on heat removal systems installed within (and outside) the containcent because they are of a quite different design. The tests performed are functional tests rather than leak rate tests.

Conclusion

Even for technically most developed systems and structures as containments and related auxiliary equipment important divergencies still exist which are partially based on the difference in national specifications and guidelines. Some of the most important divergencies can only be avoided by coordination of a stepwise harmonisation of design, manufacture and testL~g procedures.

1. Introduction

Because of the importance of containment integrity in reducing the hazards associated with a nuclear power plant and because of the possibility that deteriorations of certain parts of the containment system may occur over long periods of time it was considered of high priority that the working group no. 1 of the European Community on "Methodology, Criteria Codes and Standards" - WG 1 - investigated the problems associated with the containment leakage rate determinations. When the activities of this working group started (in 1975) there were no generally applicable requirements for integrated leakage rate testing in European countries. Some plants have already had retesting experience and have established individual methods and procedures on the basis of this experience and on the basis of already existing US requirements.

As an input to the activities of WG 1 the European Community asked five European nuclear power plant operations with experience in the area of integrated leakage rate testing to -3-

participate in an European comparative study. The present paper reports on the results of this study and on the envisaged future activities in the framework of the European Community working group no. 1.

2. Leakage rate determination

The methods of leakage rate testing in the five European countries under investigation were all derived from codes originating in the USA. This is typical for the whole LWR industry, since this technique carne to Europe together with the design of the first power plants.

This becomes evident from the appendix where the codes and standards used in the design, construction and testing of containment, containment isolation, containment ventilation and post accident cooling are given for the-different countries. Also it is conspicuous in this appendix that the larger industrialized countries like Germany and France have already worked on a development of national codes, while the nations with smaller national efforts in the nuclear area use the codes of the manufacturer. This is particularly clear for countries like Belgium and Holland who have used American, French and German codes dependent on the manufacturer of the plant.

For the containment testing the applicable code in the USA is Appendix J of 10CFR50. All countries involved have used this as a basis, but as a result of their experience interesting deviations have developed. Since the comparison is done on the older plants of the European countries, the origin of these deviations can still be traced.

After a short description of the containment designs we will compare the following points: - the pre-operational testing, test pressures, instrumentation - the periodic testing, applied methods, test pressures and intervals, instrumentation and extrapolation to the calculated accident condition.

3. Containment designs

In essence, there are no different designs used in this study. The single containment consisting of a prestressed concrete cylinder clad with steel on the inside, like it is used in France, and the double containment in use in the other countries. Of these double containments two variations are in use, namely the steel sphere surrounded by a prestressed concrete cylinder with an airgap in between and the prestressed concrete cylinder with a metal cladding on the inside with a second concrete cylinder around it with an airgap in between. The latter design is used in Belgium. The airgaps are maintained at subatmospheric pressure. -4-

- 4. Choice of testing method

The two main testing methods are the absolute method and the reference method. It is remarkable that with the growing experience in the different countries the reference method is loosing ground in favour of the absolute method. Only in Belgium the reference method is still recommended as a cross-check of the re~ults of the absolute measurement. In Italy the authorities do not express a preference for one or the other method.

As a draw-back of the reference method the German study mentions the irreproducibility, while the French study points out the princiral difficulties as follows: the reference method is used to measure the differential pressure between the volume of the containment and that of one or more leaktight reference containers, placed within- the containment and initially having the sane pressure. This reference containmer is a gas thermometer, the purpose of which is to make automatic corrections for thermal effects. This method, although it obviates the necessity of taking very precise pressure measurements and a large number of temperature measurements, suffers from two major disadvantages: - in a complex containment structure with little temperature homogeneity, it is very difficult to set up a reference container system, providing correct temperature weighting, moreover the heat capacity of the tubes used for this purpose is much greater than that of the resistance thermometer; - since the volume of this containmer is much smaller than that of the containment, its leaktightness must be of an exceptionally high level; creating and verifying such leaktightness are very e.ifficul t.

The result is that the reference method is no longer used in France, Germany and The Netherlands. This of course does noe mean that the absolute method has no problems. In essence this method consists of measuring the variations in pressure of dry air in the containment structure and the variation in mean temperature. Since the pressure P of dry air is equal to the absolute total pressure in the containment minus the partial pressure of water vapour, it is in ae.dition indispensable to know the mean humidity, obtained by vol~~e weighting the local humidity measurements. Likewise the mean temperature results from appropriate weighting of a sufficient number of local measurements of the t~perature of local volumes.

5. Pre-operational testing and overall tests (tvpe A tests)

As is to be expected the period of the acceptance tests for the containment, performed before the installations are taken into service, is also used for the first measurement of the leakage rate of the containment.

In a number of countries these pre-operational leakage rate tests are performed at the calculated accident pressure,Pa,the design pressure, and at a reduced pressure. The idea of a measure~ent at reduced pressure is to have a reference measurement so that the periodic testing later in reactor life can be done at this reduced pressure only. -5-

There is very little uniformity in the use of this reduced pressure between the different countries. In the French reactors this measurement at reduced pressure is abolished entirely. The reason given for this is that the risk of deterioration of the static parts of the contaimuent structure can only have two causes: design or construction defects in the penetrations,. which are revealed after several months of operation and structural fatigue, which is a factor to be considered in the long term. As a consequence the three-yearly tests at reduced pressure are abolished in favour of tests at full pressure Pa, during the first refueling and every ten years there after.

In the Belgium practice the reduced pressure is taken "not less than 0,5 Pa". Three leakage rate tests are performed in a 10 years period.

In the Italian standard the reduced pressure is not fixed in advance. It is taken as close as possible to the design pressure, compatible with the capability of the equipment in the containment to withstand the test pressure ar.d at any rate not less than 1 kg/cm2 above atmospheric. The leak tests are performed at the first refueling and subsequently once every 5 years.

The reduced pressure used at Obrigheim was 401 mm WG. However, this measurement turned out not to be reproducible. Therefore, this pressure was abandoned and the periodic tests were performed at 5 m h'G. The tests are repeated every 4 years.

In the Netherlands (at Borssele) a reduced pressure of 0,48 kp/cm2 overpressure was used. This same overpressure is used for the periodic tests. These tests are performed every 4 years.

Extrapolation from measurement conditions to calculated accident conditions is done by assuming rough surface leaks with turbulant flow.

5. Overview of testing intervals 2nd pressures

A comparison of the different testing pressures and intervals for the different countries is given in Table I.

7. Local leaktightness tests

The components involved in the type B tests are electrical penetrations, door seats of the personnel air locks, penetrations in the personnel air locks, the plug seats of the equipment entrance ways, the seal of the transfer tube and any resilient jOints in the penetrations. The type C tests are concerned with the isolation valves. The testing frequencies for these components vary widely from country to country. Table I - Overview of the testing pressures and intervals as applied in five European countries.

Country Leakage Hate Design Pressure Preop. Test Per iod ic 'I'est Periodic 'rest Pressure Pressures Intervals

Belgium Lt=O,lO%/day >Pa Pa Pt = ~ pa ~ 3a to 0,20%/day and La=0,25%/day ~ Pa

Italy La=U,25%/day >Pa l':Jpa not fixed 1st refueling but at least 'r.V.=l,O kg/cm2 and ~l,O kg/cm2 o.p. o.p. ~5a Netherlands Ld=0,25%/day 3,9 ato (>Pa) 3,36 kp/cm2 Pt~1.5 kg/cm2 4a (Borssele) and (Lm~U,002%/day) 2 I 0,43 kp/cm o.p. 0\ I France La=0 ,25%/day Pa Pa Pa 1st refueling ,(W 900) and lOa

Germany La=0,25%/day Pa 3,05 ato 5 rn \'1G 4a (Obrigheim) and 401 nun l'1G

Ld design leakage rate; La = allowable leakage rate; Pa = accident pressure -7-

Generally these components can be tested during normal operation, but in some countries most of the tests are performed during the refueling shutdown. Sometimes these variations are due to the component design (e.g. some newer penetrations designs have a possibility for continuous monitoring). But in other components no apparent reason can be found why in one country the testing is done every month and in another every year. . Especially in this area international exchange of experience might prove to be useful.

8. Auxiliary systems

Ventilation and air filter systems

Tests on punps and valves of the ventilation and air filter systems are functional tests rather than leaktightness tests. Test frequencies vary between one month and several months. Test frequencies and methods are prescribed in the technical specifications of the operating license. Air filters are checked on p=oper installation and integrity. The adopted method to verify iodine retention capability of active carbon filters is the Freon-12 test. Particle filters are checked for proper operation by means of the OOP test or the Uranima test method. Filter testing frequencies vary between one and two years. Heat removal svstems

In the event of an accident, two independent systems of different design can provide containment cooling and pressure reduction. These two systems are internal ventilation cooling batteries and/~r a containment spray systems. Test on the heat removal systems are functional tests and inspection. They are performed on a periodic basis as prescribed in the technical specifications of the power plant. Test interval vary between one and three months.

9. Envisaged future activities of working group no. 1 (WG 1)

As it was shown by the study there are still divergent views of the importance of possible influences on the results of leakage rate measurements. The most important factors affecting the leakage rate results seem t; be: - the measurement inaccuracies of instrumenst used at the test - the temperature fluctuations and gradients within the containment vessel - the variations of water-vapor pressure and air in the vessel during the testing interval - the temperature changes of the ambient air which surrounds the containment - the changes in free volume of the containment due to large temperature change. -8-

It is the intention that these problems be further discussed within a special ad hoc subgroup of WG 1 of the European Community in order to overcome at least some of the present difficulties arising in leakage rate measurements, their results and the interpretation of leakage rate behaviour. The main aim of this discussion is to arrive in a stepwise approach to a kind of harmonisation of presently applied leakage rate testing practices, the interpretation of test results and the prediction of leakage behaviour in European countries.

10.Nationallyapplied regulations in the European countries

In Appenix I of this presentation a summary is given of the regulations, guidelines and specifications for the construction and/or testing of the containment and auxiliary systems for nuclear power plants in five European countries. Some general remarks for these five cases are as follows.

Belgium:

In general the American codes, standards and guidelines are applied in Belgium for both the construction and operation of a nuclear power plant. Those American rules which apply exclusively to the nuclear field are mandatory, however'rules of general application may be replaced by equivalent rules reflecting Belgian or European practice. Design consultants acting on behalf of the electricity producers have together drawn up documents of their own with the aim of facilitating and/or supplementing use of the ASHE code: the "CCA 's" ("Clauses Complementaires d'Application") and the "Transpositions" (implementing rules according to the context of the Belgian legal system). Since all the American rules are applicabl~ it was felt improfitable to list in the appendix those rules which are either too general (e.q. the general design criteria) or not specific to containments or the auxiliary systems (e.g. the IEEE IS) •

Italy:

Work is underway to set up national specifications for the containment system. To date ENEL has drawn up the Technical Specifications covering the specifications, regulations and guidelines to be met in the construction of the Italian nuclear power stations. These technical specifications partly reflect the procedures in the USA and on top of that require that the reactor and related primary cooling system be housed in a double containment system that with the help of the isolation, heat removal, emergency air ventilation and filtration systems, minimizes accidental radioactive releases to the atmosphere. -9.

The Netherlands:

The basic provisions---governing nuclear installations in The Nethe=lands are set out in the Nuclear Energy Act (Kernenergiewet) which became effective in 1969. Neither of the two nuclear power plants in The Netherlands (Borssele and Dodewaard) were entirely treated in accordance with this Act because of its late appearance. In fact te ~Dutch Rules for Pressure Vessels" have been applied and as these rules are not specifically made for nuclear power plants, they are supplemented by rules used in the USA e.g. the appendices of 10CFR50, parts III and XI of the ASHE Boiler and Pressure Vessel Code and the Regulatory Guides of USNRC. At the moment the Safety Criteria for Nuclear Light Water Power Plants are set up by the Safety Board of the Ministry of Social Affairs. The parts 84, 85 and 86 of these criteria apply respectively to the containment, the containment cooling system and the ventilation systems.

France

The list of documents given in the appendix for France applies only to the acceptance tests and periodic tests for the prestressed concrete containment sturctures with leaktight metal liner for PWR's. France has developed its own documents for these tests. The philosophy differs from the American practice mainly by replacing P/2 tests to be executed every three years with full P tests during the initial refuelL~g shutdown and then every ten years.

Germany:

The list of documents given in the appendix for Germany applies only to the Obrigheirn nuclear power plant. At the time of the planning of Obrigheirn in 1964 no special technical guidelines or regulations governing the construction of nuclear power stations in the Federal Republic of Germany were available. Therefore reference was made to the standard regulations for construction and testing of the Obrigheim containment. Examination and clearance of this specification, containing the listed reguirements, fell to the Technical Inspection Authority (TUV). For the present situation newer regulations have been developed e.g. Kerntechnischer Ausschuss 3405, Kerntechnischer Ausschuss 3401. I -.

jJ jJ jJ jJ l=: .J l=: c:: l=: c:: 0 c:: Regulations, Guidelines and Specifications QJ QJ QJ c:: QJ -rl QJ fl C a jJ applicable to the containment system . C-S ~ -~ .S~jJ .~ ~ .~ .[ (construction and testing): c:: t1l t1l t1l o jJ jJM ~'j jJM .~ ~ U c:: c:: 0 c:: c:: c:: 0 H 0 QJ 0 o Ul o QJ o 0 BELGIUM Po< u (J) U U .rl U :> U U

v USNRC 10CFR50 (appendix A and J) v v v v

USNRC Regulatory Guide no. 1 10/19/90 142 11 7152 1 6.2.5 6.2.6 6.2.3 6.2.4 6.2.2 USNRC Standard Review Plan, paragraph no.: 6.5.1 101 1 45.4 271 ANSI. Standards, N no.: 509/510 III,/XI III /XI III /XI ASME Boiler and Pressure Vessel Code, section no.: III 1,2 +Tr CCA +Tr7 CCA +TrlccA* TRABEL Documents (incorporations of ASME-code 4 sections into Belgian provisions)

Belgian Design Standards based on the IBN v v

International Recommendations of FIP-CEB v v

USNRC Branch Tecllnical position BTP /CSB no.: 6.3 6.4

ACI 349 Code Requirement v

AFNOR-NR-X-44-011 v

ASHRAE-52 v

~ '0 '0 ro :J p. ~...... I a jJ Regulations, Guidelines and Specifications rl QJUl1 ~ • Ul ~ III ~ rl ~ Ul Ul ~ :> ~~ t:>~ of the containwent system (construction and ~ o ~Ql 'H" Ul. 111 Ql Ql Ql 0 Ql ~ o >, -rl L ~ a H rl H -.-I a u ~ Ql~ oliik testing) o III ::J ~ ~ -~ Ql III + • 0 I ~ III • tJ> P -rl ~ ..... ~ .~ B ~ ~ 111 .ri ~ H a Ql .-t ~ a . U ~kP~ ~ tJ> 111 U H U '0 111 111 Ql tJ> -rl Ql U -rl Ql ::J ~ III QJ -oj ill III ITALY ~ ::J ~rl ~ ~ ~. P jJ ~ H ~ H ::J ~ 0 III Ul Ql k Ul ~n.~:>rl' ~ ~ 0 Ul Ql >, ~Ul Ul .001 3 ~ o ~ o a Ql >. U III HUl'H U ..... a: Ul o~ ilIaUl~ c....'O W :> III U -ri [; ill -rl ...

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ACI 318-71 Building Code Requirements v - ASME Boiler and Pressure Vessel Code, section no.; III - AISC Specification of Structural Steel v

ANSI N271 "Containment Isolation Provisions" -+-v

Royal Decree of May 12, 1927 • CSEIN regulations v

UNICEN standard (UNI-)768) "Leak Tests" v

APED 4470 "Extrapolation of Contairullcnt Vessel Leak 'rests" v

Operating License issued by Ministero Industria v v

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Dutch Stoomwezen "Rules for pressure vessels" v

USNRC lOCFR50 applicable appendices v

ASME Boiler and Pressure Vessel Code, section III and XI v

v v USNRC Regulatory Guides v

Preliminary Safety Criteria for Nuclear Power Plants, part. no.: 84 85 86

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USNRC 10CFR50 Appendix J v v

Note REAM "Palier W900 - Auscultation des biitiments" v

Note SEPTEN "Surete des enceintes de confinement en beton precontraint" v

v ORNL, NSIC 26 by P.C. ZAPP "Testing of containment systems"

ANSI N45.4 1972 "Leakage rate testing of containments" v

Bureau Veritas "Mesure des taw< de fuite des enceintes" v

Power Engineering, 1976 "Containment leak rate testing" v

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Pressure Vessel Specifications UVV 17, 18 and 19 v Construction Specifications for tanks and vessels DIN 4119 v VdTUV notices concerning pressure vessels J09 and DIN standards v Constructional Requirements (materials, tolerances etc.) v ASME Boiler and Pressure Vessel Code, section III, par. 4 v v AD Merkblatt II v DIN 0560 BJ v AUS 7 v VdTUV J09 v Additional special requirements (welding plans etc.) v DIN 17 135 v DIN 17 155 v AD Merkblatter v Additional special specifications (e.g. leaktightness tests) v MAINTENANCE OF LEAK TIGHTNESS

OF CANDU 600 MWe CONTAINMENT

C. Harvey*, R. Beaudoin**, S. Alikhan*,R. Blanchette**

OECD NUCLEAR ENERGY AGENCY

CSNI SPECIALIST MEETING ON

WATER REACTOR CONTAINMENT SAFETY TORONTO, CANADA

17 - 21 October 1983

* New Brunswick Electric Power Commission

** Hydro Quebec MAINTENANCE OF LEAK TIGHTNESS OF CANDU 600 MWe CONTAINMENT

C. Harvey., R. Beaudoin •• , S. Alikhan., R. Blanchette ••

OECO NUCLEAR ENERGY AGENCY CSNI SPECIALIST MEETING ON WATER REACTOR CONTAINMENT SAFETY TORONTO, CANADA 17 - 21 October 1983

• New Brunswick Electric Power Commission •• Hydro Quebec Page 1 of 14

TABLE OF CONTENTS

1.0 INTRODUCTION

2.0 CONTAINMENT-TESTING PRIOR TO INSERVICE

2.1 Building Leak Rate Tests 2.2 Proof Pressure Tests 2.3 Visual Inspections

3.0 CONCLUSIONS OF TESTING PRIOR TO INSERVICE

4.0 CONTAINMENT - IN SERVICE TESTING

4.1 Testing the Major Extensions of Contaiment 4.1.1 Airlock Tests 4.1.2 Containment Isolation Valve Tests 4.1.3 Containment Door Tests

4.2 Building Leak Rate Test Procedure 4.3 Visual Inspection 4.4 Liner Sample Tests 4.5 Leak Tests on Typical Penetrations

5.0 CONCLUSIONS Page 2 of 14

LIST OF FIGURES

FIGURE 1 Simplified Diagram of Containment Envelope FIGURE 2 Perimeter Wall Penetrations FIGURE 3 Leak Rate Formula FIGURE 4 Building Leak Rate FIGURE 5 Soap Bubble Test Equipment FIGURE 6 Ultrasonic Test Equipment FIGURE 7 Reactor Building Pressure Test FIGURE 8 Strain Gauge Results FIGURE 9 Tests for Airlocks FIGURE 10 Typical Leak Rate for Interseal Space FIGURE 11 Tests for Containment Isolation FIGURE 12 Tests for Containment Door FIGURE 13 Full Pressure Test Procedure FIGURE 14 Building Leak Rate Trend Prediction FIGURE 15 Plot of Building Leak Rate at 20% and 2% Increases FIGURE 16 Typical Embedded Part Page 3 of 14 ! •

1.0 INTRODUCTION The Containment System of a CANDU 600 nuclear power plant is one of the four special safety systems . designed to restrict the release of radioactivity to the environment within the maximum permissible limits, both under normal operation and accident conditions. This is achieved by limiting the integrated overpressure transient following a loss of coolant accident and by ensuring that the containment remains leak tight to the specified requirements. The containment structure is designed to withstand an internal pressure of 124.0 kPa(g). The accepted standard for containment leakage is 0.5% of the total enclosed free building volume over a twenty four hour period. Although conservative, 0.5% is achievable, and has been used for accident analysis purposes in the Safety Analysis Reports. If required, further analysis may prove that this acceptance value can be relaxed. To satisfy the licensing requirements, it is required to demonstrate the strength of the structure by performing a proof pressure test at 115% of the design pressure, and to demonstrate leak tightness of the building to the CSA Standard N287.6 prior to critical- ity as well as during in-service on a periodic basis. Figure 1 shows a schematic diagram of the containment system, which is comprised of: • a cylindrical reinforced post-tensioned concrete reactor building, 41.5 m in diameter with wall thickness of 1.07 m and 53 m in height enveloping a total free volume of 48480 m3• Because concrete is not leak tight, the interior surfaces of the reactor building are coated with an epoxy lining • • two Airlocks (an Equipment and Personnel), forty three (43) Electrical Penetrations, 100 Mechanical Penetrations (process piping and instrument tubing), and a Spent Fuel Discharge System. Figure 2 shows all perimeter wall penetrations and their locations • • a Containment Isolation System, which isolates major process piping emerging from the reactor building automatically on sensing high pressure or high radioactivity levels in the Reactor Buidling • • a Dousing System, provided to condense steam and thus reduce containment pressure following a loss of coolant accident. Up to 2090 m3 of water, stored in a tank formed by the inner and outer building domes is sprayed automatically if the containment pressure exceeds 14 kPa(g). Page 4 of 14

1.0 INTRODUCTION (Cont'd) • Local Air Coolers assist the Dousing system in their function of preventing overpressure of containment. They start automatically when Reactor Building . pressure exceeds 3.45 kPa(g). • a Vapour Recovery system and a particulate/charcoal filter assembly to allow depressurization of the containment envelope in a controlled manner following an accident •. 2.0 CONTAINMENT - TESTING PRIOR TO INSERVICE 2.1 Building Leak Rate Tests These tests are performed prior to criticality to demonstrate that leakage through the entire containment boundary does not exceed the specified limit. With the building pressurized to design pressure . :' (124 kPa{g», readings of the Reactor Building absolute pressure, air temperature and humidity are taken on ten minute intervals. This data is used to calculate the leakage rate. The data can be manually or I ' ' automatically entered into the plant computer which is progr ammed to: compute the fractional leakage for each successive set of readings using the formula given in Figure 3. - compute the actual leakage rate by linear regression analysis for the fractional leakages over a 24 hour interval. If an automatic system of data collection is used, there should be in addition, manual data collection to provide backup. Figure 4 shows the plot of the building leak rate for a test conducted at Gentilly-2. It has been experienced that a 6 to 12 hour stabilizing period is required following pressurizing of the building, before consistent data is obtained. The results of the building leak rate measurement tests to date have proven that the total leakage from containment is less than 0.5%. The leak rate measured for Gentilly-2 was 0.156% and for Point Lepreau it was 0.255%. Page 5 of 14

2.0 CONTAINMENT - TESTING PRIOR TO INSERVICE (Cont'd) 2.1 Building Leakage Rate Tests (Cont'd) During building pressurizing and at scheduled stops, leak searches were carried out. The two methods of leak detection used were: - Soap bubble test on 100% of the external surface of the Reactor Building - Ultrasonic scans on the inside of the building. The soap bubble test was carried out using ftLeakTechft fluid. About 1400 litres were required. Figure 5 shows a picture of a garden type spray that was used to apply the fluid. Many of the leaks detected using the method were evident at building pressures of 30 kPa(g) or less, however as the pressure was raised, other leaks developed.

; l The ultrasonic scans were conducted at hold points, when personnel access inside the building is permitted. Figure 6 shows the Ultrasonic Detector used. It was a Hewlett Packard Model 4905 Ultrasonic Translator Detector with a Model 18043A Ultrasonic .Reflector. It proved to be a very useful tool in pinpointing any defects in the liner, which by other means were extremely difficult to locate. The areas where leakage were found, many of which were common to both sites, were: - Bulkhead Swagelock fittings on seal plates - Pyrotenax cable fittings - Electrical cable glands - Spent Fuel Discharge Bay walls - Around major embedded parts (Main Steam and Equipment Airlock) - Temporary construction openings The following information has been gained from conducting previous presure tests and should be kept in mind for future tests: • Pressure testing should be kept to a minimum, as repetitive stressing may eventually cause problems in maintaining a leak tigh structure. • Experience has shown that reduced pressure building leak rate tests are not correlatable with the design pressure test and are not recommended. Page 6 of 14

2.0 CONTAINMENT - TESTING PRIOR TO IN-SERVICE (Cont'd) 2.1 Building Leakage Rate Tests (Cont'd) • The depressurizing rate should not exceed 5 kPa/hour. Exceeding this rate could result in extensive damage to the epoxy liner • • An extensive study will be required prior to in-service tests to extablish the state of each system consistant with the requirement to maintain safe operation of the plant • • All staff involved with the test should be on 12 hour shifts. This helps eliminate problems associated with shift changes, particularly with regard to entrances to the Reactor Building while it is under pressure • • Adequate numbers of staff of each discipline, i.e~ operators, electricians and maintenance should be qualified to work in a compressed air environment. This is recommended because the working time for individual in compressed air is limited, and rejection of a significant number of candidates may be expected • • A miminum of 1200 litres per second total compressor capacity. This will permit pressurizing at a rate of about 8 kPa/hour • • Sealing of liner and caulked joint defects can be achieved with the building at pressure. Thick coats of fast curing epoxy (Nickel Epoxy) or Thiokol, when applied over the defect, will be entruded into, and effectively seal the leak path • • Electrical cable glands can be repaired with the building pressurized. Extreme care must be exercised to secure the cable before the gland nut is removed. Movement could result in shielding sand being blown into the gland, preventing re-sealing • • It has been demonstrated at Gentilly-2 that it is permissible to conduct a pressure test under peak summer temperature conditions, however it is advisable to schedule testing for the cooler months, since stresses induced by heating on the external surfaces of the structure are additive to those induced by internal building pressure. Page 7 of 14

2.0 CONTAINMENT - TESTING PRIOR TO IN-SERVICE (Cont'd) 2.1 Building Leakage Rate Tests (Cont'd) • A weighted average of building temperatures obtained from a large number of existing RTD's will provide good results for leak rate calculations. The plant computers can be programmed to print out the temperatures and averages as required • • No entries into the Reactor are permitted during the leak rate period. If entries are made, the test must be res tarted • • When evaluating any leak rate test results, comparison to previous test results must be made to evaluate the condition, or changes in the condition of. the components being tested •. 2.2 Proof Pressure Test The Proof test, carried out in parallel with the initial Building Leakage Rate Test, demonstrates the structural soundness of containment. The proof pressure is 143 kPa(g), 115% of design pressure. Figure 7 shows the history of pressurization and a brief description of the major activities for the Proof and Building Leakage Test at Point Lepreau. Note that the Design Pressure Leak Rate Period 6 is the period of time shown in Figure 4. Tests and observations are used to verify that the containment boundary is structurally sound. The building must hold proof pressure for at least one hour. The pressure is monitored during the Proof Test to verify that a major leak has not developed. Figure 8 is a plot of building deformation as measured using strain gauges. The predicted values are plotted for comparison purposes. Strain gauges, installed at strategic locations to measure building deformation, were monitored at zero pressure, each hold point, and at proof pressure. The data was used to determine the behaviour of the building under pressure. Results of studies for both stations verified that the buildings performed as per design and that deformations were in close agreement with the predicted behaviour. Page 8 of 14

2.0 CONTAINMENT - TESTING PRIOR TO IN-SERVICE (Cont'd) 2.2 Proof Pressure Test (Cont'd) Hairline cracks were observed on the inside walls of the reactor building at both sites. The cracks were concentrated in the lower section of the perimeter wall. Several studies were conducted to determine the nature, cause and effects, as well as corrective action. Dye was injected into the cracks and core samples taken, and the cracks were monitored for length and width during pressure tests. An extensive study was carried out on the concrete itself. The results of these studies concluded that: A maximum width of 008 mm and depth of 13 cm ~ere noted • • The cracks were the result of concrete shrinkage, and developed as the concrete cured • • The cracks were of superficial nature and did not affect the strength of the structure. During building pressurizing the cracks opened slightly. A maximum crack opening of 0.026 mm was noted. On depressurizing, the cracks returned to normal width. The local movement at a crack dictated a need for a flexible coating. An elastomer based sealant was tested and found to be a suitable material to be used on the containment boundary and to accommodate movement. The sealant was applied over all cracks and was found to remain intact during subsequent building pressure tests. 2.3 Visual Inspections The readily accessible areas of the containment boundary, at both Point Lepreau and Gentilly-2 were given a general visual inspecton prior to going critical to ensure no damage had been done between the last pressure test and criticality. 3.0 CONCLUSIONS OF TESTING PRIOR TO IN-SERVICE The results of all tests conducted to date have shown the entire containment boundary and its various components perform better than the design requirements. Extensive efforts undertaken during the construction and startup phases of both Gentilly-2 and Point Lepreau have reduced leakage from containment to the lowest leakage rates possibly achievable. Page 9 of 14

4.0 CONTAINMENT - IN SERVICE TESTING Containment is the final barrier which restricts release of radioactive material to the environment. In order to ensure that hazardous releases of radiation will not occur, the leak tightness of containment must be ensured. This can be accomplished by testing the entire containment boundary as well as individual components. The following tests are conducted to ensure leak tightness and reliability. 4.1 Testing of The Major Extensions of Containment j; 4. 1 • 1 Airlock Tests Figure 9 lists the tests and the frequency at which they are conducted for the Airlocks. Alarm Test, Equipment and Personnel Airlocks (26 weeks). Twice a year, each parameter monitored by alarm circuits are tested by simulating alarm conditions. Proper response of each alarm are verified at the local, main control room, and secondary control room panels. - Interseal Space Test (4 weeks) This test measures the pressure decay of the interseal space. The pressure decay is compared to a standard curve to evaluate the performance of the seals. (Figure 10) - Air System Pressure Decay (26 weeks) The air system is capable of providing 24 hour air supply to the seals and two full operations of the personnel doors following loss of makeup air. The air system is tested to ensure it will still meet this requirement. - Electrical and Mechanical Interlocks (26 weeks) These tests verify the operation of the interlocks which prevent possible breaches of containment. Page 10 of 14

4.0 CONTAINMENT - IN SERVICE TESTING (Cont'd) 4.1 Testing The Major Extensions of Containment (Cont'd) - Airlock Leak Test (52 weeks) The airlock Leak Tests ensure that the leakage from either airlock does not exceed its design leak rate. The results of each leak test are compared to previous test results and any sources of leakage found are repaired. The tests to date at both Gentilly-2 and Point Lepreau have shown airlock leakage to be below the design limits. - Perimeter Seal Leak Test (52 weeks) The perimeter seals, which form part of the Equipment Airlock pressure boundary are leak tested annually, to ensure that there has been no damage or degradation which would affect airlock leakage. 4.1.2 Contaiment Isolation Valve Tests Piping used to transport low pressure air present leakage paths for radioactive materials under accident conditions. These lines, ie., ventilation systems, are provided with pairs of Containment Isolation Valves. Periodic tests are required on these valves to ensure their continued leak tightness and reliability. Figure 11 shows these tests and their frequency. Containment Isolation Valve Operation Control Test (4 weeks) The control logic circuit and valve operation are tested each month to ensure that the valve will operate properly when called upon to do so. The valve closing time is another test that is required to ensure that the valve will meet its design requirements. The 30. valves must be fully closed in 500 milliseconds and all others in less than 2 seconds. - Containment Pressure Monitoring Loop Tests (4 weeks) High Reactor Building pressure is one of the two parameters which will cause the containment isolation valves to close. Trips on two-out-of-three channels on high pressure initiate a button-up of containment. At a frequency not exceeding one month each of the three channels of the Reactor Building Pressure monitoring loops are tested for proper response by simulating high reactor building pressure. Page 11 of 14

4.0 CONTAINMENT - IN SERVICE TESTING (Cont'd)

4.1 Testing The Major Extensions of Containment (Cont'd) Containment Activity Monitoring Loop Tests (4 weeks) High radioactivity detected by two-out-of~three monitors located in the Vapour Recovery and Reactor Building Ventilation Systems is the other parameter that will initiate a Containment Buttonup. Monthly, each activity monitoring loop is tested by feeding a test signal into the circuitry and verifying the circuit trips and alarms properly. 4.1.2 Contaiment Isolation Valve Tests - Intervalve Space Leak Test (13 weeks) To ensure that each containment isolation valve is leak tight, every three months the valves are leak tested. This is accomplished by presurizing the interspace between each pair of valves to 124 kPa(g) and measuring the presure decay over a period of time. The acceptable pressure drop for each pair of valves will be different because of the differences in the interspace volume and valve sizes. In order to facilitate testing procedures, both utilities have adopted a standard conservative acceptance value for all valve pairs. If the pressure decay is less than 7 kPa per hour, the valve pair is considered leak tight. Between 7 and 15 kPa per hour the valve is flagged for repair. If immediate repair is not achieved, the repair will be scheduled for the next shutdown. If the pressure decay exceeds 15 kPa per hour, one valve remains closed until the other is repaired or the acceptance value re-assessed, on an individual basis. The test results to date show the leakage rate to vary between 0 to 4 kPa per hour. Eight cases of non compliance have been detected to date. Three of these required seal replacement, and 5 were corrected by adjustment of the close stop limits. Page 12 of 14

4.0 CONTAINMENT - IN SERVICE-TESTING (Cont'd)

4.1 Testing The Major Extensions of Containment (Cont'd)

Verification of Closing of Tritium in air Sampling Solenoid Valves.

These valves are tested for proper Qperation annually to ensure that they are closing when selected closed. All other valves have limit switches -that operate when the valve reaches its closed position, which is not the case with these solenoid valves. The test verifies the valve closes when it is de-energized.

4.1.3 Containment Door Test (52 weeks)

The access route to the Spent Fuel Discharge Bay is sealed by the Containment Door. Under certain conditions the Containment Door can form part of the containment boundary. Because of this, the door must be tested to ensure it remains leak tight. Figure 12 lists the tests carried out for the Containment Door. - Air Supply Reserve Test

The reserve tanks are required to maintain the door seals inflated for nineteen hours at a pressure above 140 kPa(g). This test is performed annually.

- Breach of Containment Test

The pressure switches which alarm on containment door seal failure are calibrated and tested annually. The alarms which are brought in by these switches are checked as part of the test procedure.

4.2 Building Leak Rate Test Procedure

The only method of confirming the leak tightness of the entire containment boundary is by conducting a design pressure leak rate test. However, because of several disadvantages the frequency of these tests must be kept as low as possible. Tests and other measures must be used to compensate for the reduced frequency of Building Leak Rate Tests.

The in-service building leak rate tests will be similar to the preoperational tests. Figure 13 shows a plot of Reactor Building pressure against time for what may be a typical inservice building leak rate test. Refering to Figure 13, the building will be pressurized to 124 kPa(g), followed by a three hour stabilizing period. Page 13 of 14

4.0 CONTAINMENT - IN SERVICE-TESTING (Cont'd) 4.2 Building Leak Rate Test Procedure (Cont'd) During this time, leak searches wil be conducted, but no repairs will be attempted. This will allow determination of the leak rate value before the test is started. Following the three hour stabilizing period, a fifteen hour leak rate test will be conducted. When the leak rate is determined, a decision will be made to repair or use the measured leak rate. If the leak rate exceeds the previous test results by more than roughly 20%, then repairs will probably be made. If.the leak rate exceeds the design value of 0.5%, then repairs must be made. Repairs will be made at pressure if possible. If any repairs are made, the new leak rate must be established. Figure 14 is a hypothetical plot of what may be expected over the life of a plant. Point (A) is the expected value of the next leak rate test. If the next leak rate in fact proves to be higher than this, Point (B), repairs may be made so that the leak rate is reduced to Point (C). Figure 15 shows two different plots of leak rate increases. Curve A shows a 20% increase per test and Curve B a 2% increase. Currently the test frequency is about 5 years. The results of future tests may help in extending the interval between tests. 4.3 Visual Inspections Concrete is not leak tight because of its porosity. The leak tightness of the Reactor Building is achieved by lining the inside with an air tight liner. Any deterioration of this liner will result in an increase in the leak rate, therefore the liner must be well maintained. A detailed visual inspection of all of the liner will be conducted between each building leak rate test. This respresents an area of approximately 20% of the total surface area each year. Other areas of concern, for example, high traffic and material handling areas 4r~ included in the Annual Shutdown inspection. Inspections to date have revealed a substantial amount of minor damage caused by movement of material and personnel. There was no conclusive evidence of downgrading of containment leak tightness due to aging or deterioration, however, the damages found in the following two areas may be dUe to aging. Page 14 of 14

4.0 CONTAINMENT - IN SERVICE TESTING (Cont'd) 4.3 Liner Smaple Tests (Cont'd) On caulked construction joints, the epoxy paint-~over a few of the joints was cracked. In some cases, the crack was observed to travel into the caulking material (thoikol), however, these were the exception. - Fiberglass liner on the inside of the Spent Fuel Discharge Bay (Rl-001) was cracked on inside corners. It is probable that the damage resulted from an impact load. The extent of damages found on the base slab, roof of the Spent Fuel Discharge Bay and in the hoist way indicates a need for protection and inspection on a yearly or more frequent intervals. During these inspections, all defects will be repaired. 4.4 Liner Sample Tests During the course of the visual inspection, if a defect in the liner becomes apparent for no obvious reasons, extensive tests may be carried out on a sample of the liner. Some of these tests are hardness tests, ! elasticity tests, effects of radiation and aging i~ effects.

! , 4.5 Leak Test on Typical Penetrations Some of the embedded parts have sealed plates on both ! I sides of the reactor building wall. (Figure 16). It is possible to perform local pressure tests on these penetrations if the leak tightness was suspected for any reason. 5.0 CONCLUSIION ( Maintenance of Containment leak tightness of the Candu 600 MWe is one means of protecting the public. Cooperative exchange of information and experience will improve our ability to maintain and improve the leak tightness.of containment...... :---, ,._--~, _.] ...... /._ ..... j. j- . ..' l " .. : l. , ,

DOUSING TANI{ TOTAL H 0 HOLDUP 2 STRUCTURE

w~ >u

C> .' i=;! 'l> ~(f) [QOUSING' TANK I t> ~ti EN-TiL-Al-'IO~ D," ." IRIB AOUOOlEAS] I>b FILTER ....4 (HEPA 8 CIIARCOAL) 8LDG. ACCESS ~ .. (> c:: AIRLOCK [AIRLOCKS) -_If=-_ FREE VOLUME 40:l00m! O' 1>''- .(> 2 0 0 VAPOUR RECOVERY SURFACE AREA 2230010 2 c ;" .~. DRYERS ~, DESIGN PRESSURE 124 t(Pa(g) ,,: .: : .. PROOF PUESSlJIlE 142.7 tfPa(g) o..~ CONTAINMENl] ...-.,..' '.~'. ISOLAIl.QN " '..:." 1>'" -- .....

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FIGURE 2 PERIMETER WALL PENETRATIONS FIGURE 3 : LEAKAGE RATE FORMULA

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----NOTES: I-PRESSURIZE I COMPRESSOR CAPACITY 850 LIS \(2) PRESSURIZING RATE 4k~a/HOUR OQ---IOOkPa(g)

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- ..... - _._ .._. MAINTENANCE OF LEAK-TIGHTNESS FOR A TYPICAL ONTARIO HYDRO MULTI-UNIT NUCLEAR STATION CONTAINMENT ENVELOPE

PAPER PRESENTED AT THE OECD/CSNI CONTAINMENT SPECIALIST MEETING (OCTOBER 17-21, 1983)

R.I. Chen and M. Mazza Civil Design Department, Ontario Hydro Toronto, Canada

1695L SUMMARY

This paper describes a typical Ontario Hydro Multi-Unit Nuclear Station Containment Envelope (i.e., Bruce NGS A) and the potential leakpaths which must be monitored and maintained in order to continuously assure the desired leak-tightness of the envelope. Also discussed are the types of tests used to determine the leakage rate (or leak-tightness) of the containment envelope, methods used to isolate and individually test certain penetrations without pressurizing the whole containment envelope, and.commissioning and in-service experience with leakage of containment.

l695L TABLE OF CONTENTS

1.0 BRUCE NGS A CONTAINMENT ENVELOPE 1

1.1 General 1 1.2 Containment Envelope Structures (Description) 1

2.0 POTENTIAL LEAKPATHS THROUGH THE CONTAINMENT ENVELOPE 4

2.1 Penetrations 4 2.2 Joints in the Concrete Structures 8 2.3 Cracks and "Pores" in the Concrete Containment Shell 8 2.4 Vacuum Building Roof Seal . 9 2.5 Vacuum Ducts 9

3.0 TYPES OF LEAKAGE RATE TESTS 10

3.1 Commissioning Leakage Rate Tests 10 3.2 Continuous Leakage Rate Monitoring 10 3.3 Periodic In-Service Tests 10

4.0 ISOLATION AND INDIVIDUAL TESTING OF PENETRATIONS 11

4.1 Electrical Penetrations 11 4.2 Shield Tank Seal 11 4.3 Airlocks and Transfer Chambers 11 4.4 Button-up Dampers or Valves 12

5.0 COMMISSIONING AND IN-SERVICE LEAKAGE EXPERIENCE 12

5.1 Commissioning and In-Service Overall Leakage Rate Test Results 12 5.2 Commissioning and In-Service Test Experience for Penetration Seals 12 5.3 Commissioning and In-Service Inspection Experience for Expansion and Construction Joints 13 5.4 Commissioning Test Experience for Cracks in Concrete Containment Structures 13 5.5 Commissioning and In-Service Experience for Roof Seal 14

1695L 1.0 BRUCE NGS A CONTAINMENT ENVELOPE

1.1 General

This report deals with the leakpaths, leakage testing, and leakage experience for the Bruce NGS A containment envelope.

The basic Bruce NGS A containment envelope consists of the following reinforced concrete structures, as shown in Figure 1:

(a) Four reactor vaults:

(b) Fuelling duct;

(c) Cent ral fuelling area;

(d) Two pressure relief ducts;

(e) Pressure relief valve manifold (including vacuum ducts); and

(f) Vacuum building.

This basic envelope has been extended to cover any other system where failure of a pipe, duct, valve or damper would result in an opening in the envelope. These pathways through the containment boundary (except dampers) are beyond the scope of this report and thus will not be dealt with here.

To maintain the integrity and leak-tightness of the containment envelope and to prevent loss or downgrading of heavy water, the inside surface of the concrete is lined with steel in most areas (i.e., the reactor vaults, the central fuelling area, and the fuelling duct throughout with the exception of the ceiling and the upper two-thirds of the walls under the Reactor Auxiliary Bay).

1.2 Containment Envelope Structures (Description)

The following is a brief description of each of the containment envelope structures mentioned above, with particular emphasis on the general location and accessibility of various system components, penetrations, etc. Any structural features which can affect leak-tightness are also identified (e.g., construction and expansion joints).

1.2.1 Reactor Vaults

The reactor vaults are rectangular, reinforced concrete structures (Figure 2). The inner surface of the concrete is lined with carbon steel to control leakage following an accident, and to prevent D20 migration through the concrete. Each vault is located over, and is connected to, the fuelling duct.

The reactor vault houses the calandria, and all associated high temperature, high pressure, heavy water systems (Figure 2). Most of the mechanical equipment associated with the primary heat transport system is accessible from outside of the reactor vault. This includes the steam generators (or boilers), heat transport pumps and reactivity mechanisms.

1695L - 2 -

The preheaters, however, are located inside containment, in a sealed steel enclosure which protrudes through the roof of the vault. The shield tank (upper extension) also protrudes through the roof of the vault. It is sealed at its periphery to the reactor vault ceiling by three reinforced, elastomeric seals.

The minimization of the reactor vault volume, and the accessibility of major components, mean that there are a large number of penetrations through the vault structure (i.e., approximately 325 penetrations through each reactor vault, including airlock, transfer chamber, mechanical equipment, electrical, instrumentation, etc., penetrations). The reactor vault is connected to the fuelling duct through two floor openings, each 25 feet by 30 feet in size. The openings are at either end of the calandria, and provide access for the fuelling machines. The reactor vault can be isolated from the rest of the containment envelope by blocking these openings with bulkheads, and closing connecting drains and ventilation ducts.

1.2.2 Fuelling Duct

The fuelling duct, which extends over the full length of the station, is a rectangular reinforced concrete structure with walls 5 feet thick and internal dimensions of 25 feet wide by 30 feet high (Figure 3). The floor of the duct is below grade. The sections of the fuelling duct which are located below the reactor building are steel-lined throughout for leak-tightness. Only the floor and lower eight feet of the walls are steel-lined at section of the fuelling duct located under the Reactor Auxiliary Bay. The steel lining here extends 3.25 feet above the water level of the lake, and is intended to prevent the influx or out flux of water through any cracks which may form in the duct walls. The ceiling and upper two-thirds of the walls are lined with a high quality vinyl paint.

With respect to the containment envelope, the purpose of the fuelling duct is to connect the reactor vaults to the pressure relief ducts. The connection is made at two "Tee" sections, which are located on either side of the Service Building, between units 2 and 3. Special expansion joints are used at the connections between the fuelling duct and the reactor vaults, central fuelling area and pressure relief ducts (15 joints in total). There are also four special expansion jOints in each of the intermediate fuelling ducts (east and west). Since these expansion joints breach the containment boundary, they are sealed with a metal strip which is seal welded to an embedded part(s) and/or to the steel liner in areas where it exists.

There are few penetrations through the fuelling duct structure (13 in total). Four of these penetrations are for airlocks while the others are piping penetrations for the emergency injection system and D20 vapour recovery system.

1.2.3 Central Fuelling Area

The central fuelling area (CFA) is located in the Service Building. It includes the portion of the fuelling duct which runs through the service Building, four fuelling machine rooms located above, and open to, this portion of the fuelling duct, and the transfer chambers connecting the four fuelling machine rooms to the Service Building. l695L - 3 -

There are approximately 192 penetrations through the CFA structure. These include penetrations for system piping, transfer chamber door locking and control mechanisms, air supply lines, etc., electrical cables, transfer chamber door frames and spent fuel and new fuel ports.

1.2.4 Pressure Relief Ducts

The. two pressure relief ducts (PRO's) are reinforced concrete structures with walls and roof 4 feet thick and floor 5 feet thick. The sole purpose, with respect to the containment system, of the PRO's is to connect the fuelling duct to the pressure relief valve manifold.

The interior surface of each duct is unlined. There are only two penetrations through the PRO structure. These are required for pipes.

1.2.5 Pressure Relief Valve Manifold

The pressure relief valve manifold is a rectangular, reinforced concrete structure which runs around most of the outer perimeter of the vacuum building at grade level (Figure 4).

It houses the containment pressure relief system. The system comprises 16 main pressure relief valves (PRV's), four auxiliary pressure relief valves (APRV's) and two reverse flow valves.

The main PRV's and the reverse flow valves are connected from the manifold to the vacuum building by vacuum ducts. Each vacuum duct passes from the manifold, through the vacuum building containment floor. The portion of the vacuum ducts which go through the basement are considered as part of the containment envelope.

The manifold concrete structure has 12 expansion joints. They are sealed by a special metal strip which is seal welded to an embedded part on each side.

Penetrations through the manifold structure include the PRV and reverse flow valve penetrations, APRV penetrations, APRV air test penetrations, vent duct penetrations, cable penetrations, instrument, service and breathing air penetrations and airlock and transfer chamber penetrations.

All of the above penetrations, except those for the airlock and transfer chambers, pass through the manifold floor.

The ceiling, walls and floor of the manifold are 3.25 feet thick. The interior surface of the ceiling and walls is coated with latex paint. The floor is coated with epoxy floor paint.

1.2.6 Vacuum Building

The vacuum building is an upright, cylindrical, reinforced concrete structure (perimeter wall is post-tensioned), consisting of a basement, main chamber and an upper chamber (Figure 4). The basement is not part of the containment envelope. The perimeter wall of the building is 3.6

1695L - 4 - feet thick, the containment floor is 4.5 feet thick, and the roof is 3 feet thiCK. The reinforced concrete roof slab is connected to the perimeter wall by an elastomeric ring seal.

The containment floor and roof slabs have construction joints. A special joint is used between the perimeter wall and base slab.

There are several penetrations through the vacuum building structure (50 in total). These include penetrations for vacuum ducts, system piping and I&C lines, access hatches and roof closures.

Most of these penetrations are in the containment floor and main roof slabs.

It is important to note that the inside of the vacuum building is not accessible while any of the units are in operation.

2.0 POTENTIAL LEAKPATHS THROUGH THE CONTAINMENT ENVELOPE

As mentioned above, the containment envelope is basically a reinforced concrete shell with a few steel sections (e.g., the preheater enclosures and hatches) and some steel-lined areas. The primary function of the concrete shell is the retention of the containment atmosphere (to within acceptable limits) in the event of the accidental release of radioactive contaminants from the reactor system. The measure of effectiveness -for this primary function is the degree of leak-tightness achieved by the containment shell. Absolute leak-tightness is not practically attainable in large containment structures because of the existance of a variety of potential leakpaths. Potential leakpaths through the Bruce NGS A containment envelope include:

(a) Penetrations required for piping, airlocks, mechanical equipment, etc.;

(b) Construction joints and expansion jOints (in concrete components) ;

(c) Cracks and "pores" in the concrete;

(d) Vacuum building roof seal; and

(e) Vacuum ducts.

In the discussion which follows, each of these potential leakpaths are described.

2.1 PENETRATIONS

Penetrations are one of the major sources of leakage through the containment envelope.

There are a large number of penetrations, especially in the reactor vaults, central fuelling area and vacuum building, which are required for a variety of components such as system pipes, electrical cables,

1695L - 5 -

instrumentation lines, mechanical equipment (e.g., boilers, heat transport pumps, etc.), airlocks, transfer chambers, and other system components.

Several types of penetration seals are used. These include:

(a) Welded Penetration Seals:

(b) Cable Penetration Seal:

(c) Instrumentation Line Penetration Seal;

(d) Metal Bellows Seals;

(e) Replaceable seals.;

(f) Airlock and Transfer Chamber Door Seals;

(g) Structural Seals:

(j) Preheater Enclosure Seal.

Table 1 shows the quantities for each type of seal for penetrations through the containment envelope.

2.1.1 Welded Penetration Seals

This category includes piping penetrations which are sealed by welding the pipe to a seal plate. The seal plate itself, is welded to an embedded part in the containment structure (e.g., pipe sleeve). Figure 5 shows details of a typical welded penetration seal for Bruce NGS A and also leakage paths such as cracks and weld "pores".

2.1.2 Cable Penetration Seal

This category includes seals for electrical and instrument cable penetrations. Sealing of the cable at the containment penetration cannot be done by welding as in the case of piping penetrations. Instead, sealing is accomplished by use of an electrical cable fitting as shown in Figure 6. The fitting consists of a steel gland, silicone rubber gasket (with gasket holder) and a gland nut.

When tightened, the gland nut compresses the rubber gasket and a seal is produced between the gland and cable. The effectiveness of this seal depends upon the resiliency of the rubber gasket. The bottom of the gland is welded to a seal plate which is welded to an embedded part.

* Airlock and transfer chamber door seals are not included in this category, even though they are replaceable. These seals have been placed in a separate category.

l695L - 6 -

There are three possible ways by which leakage can occur:

(a) Through the cable itself

(b) At the cable (i.e., at the interface of the gland and cable sealed by the gasket); and

(c). At the seal welds.

The latter two leakpaths are indicated on Figure 6.

2.1.3 Instrumentation Line Penetration Seal

Instrumentation tubes are generally sealed at the containment penetration by a swagelok fitting which is welded to a seal plate. The seal plate itself is welded to an embedded part. The swagelock is designed to produce a leak-tight seal at the interface of the male connector and tube by tightening the swagelok nut.

Leakage of this seal can occur either at the interface of the male connector and tube or through the seal welds.

2.1.4 Metal Bellows Seals

Metal bellows seals are normally used wherever displacements due to thermal expansion and contraction of penetration components (e.g., large piping and mechanical equipment for heat transport system) must be accommodated by means of a flexible design arrangement so that no undue strains are imposed upon the penetration or piping. A typical metal bellows seal for piping is shown in Figure 7. The bellows assembly is seal welded to the pipe and embedded part. A shroud (pipe sleeve) is used to protect the metal bellows.

The metal bellows seals for the heat transport pumps and boilers are shown in Figure 8 and 9, respectively. Welded construction is used to seal the bellows to the containment structure and the boilers. A reinforced asbestos gasket is used to provide a seal between the bellows and the top of the heat transport pump. This gasket is bolted to the top casing of the pump.

Potential leakage paths in metal bellows seals are through cracks in the bellows membrane, and at bellows connections (i.e., seal welds and asbestos gasket). These are illustrated in Figures 7, 8, and 9.

2.1.5 Replaceable Seals

This category includes elastomeric o-ring and/or gasket seals for door handwheel shafts, valve spindles, door locking devices, vault viewing ports (periscope), boiler support rod covers, pressure relief valves (seat), access hatches, etc. Also included is the special organic seal for the shield tank. Figures 10 and 11 show the o-ring seal for the pressure relief valves and Figure 12 shows the special seal for the shield tank.

1695L -7~

The shield tank seal consists of three reinforced, elastomeric sealing elements. The inner seal (closest to the reactor vault atmosphere) is inflatable and can be used as a temporary backup to the other two, as well as for periodic leak testing or replacement of the latter. The two outer seals are of the diaphragm type, with sufficient flexibility to accommodate the thermal and other movements of the shield tank. The seals are capable of withstanding pressures up to +20 psig. Only one of the two inner seals needs to be intact in order to maintain the containment envelope.

An important point to note abut replaceable seals is that provlslon has been made for replacing them in case deterioration or aging negates their sealing capability.

Leakage of replaceable seals can occur at the seal interface, through cracks (or splits) in the sealing element, or through seal splices (Figures 10, 11 and 12).

2.1.6 Airlock and Transfer Chamber Door Seals

Airlock (AL) and transfer chamber (TC) door seals are essentially replaceable seals, however, due to their importance as leakage sources they are treated separately in this report.

Three types of sealing arrangements are used.

These include:

(a) An o-ring (rubber) seal between the door and frame:

(b) A flat gasket (rubber) seal between the door and frame: and

(c) An elastomeric, dual inflatable seal between the door and frame (Figure 13).

The first two arrangements rely on the compression of an o-ring or gasket to produce enough contact at the interfaces of the door and o-ring and frame and o-ring in order that a seal is created. Their application is limited to AL's and TC's which are infrequently used. The dual inflatable seal is quite different. It consists of two special rubber sealing elements which can be inflated with air just like a tire tube (see Figure 13). Only one of these sealing elements is necessary to maintain the seal. Each seal is inflated to the position shown in Figure 13 (Detail 1). The inflated seal compresses against the AL or TC frame and a leak-tight seal is produced. The bottom half of each seal is placed in a special metal retainer. The dual inflatable seal is used wherever frequent AL or TC operation is essential. Over 70 percent of the AL's and TC's in the Bruce NGS A containment envelope utilize this type of seal.

Leakage of AL and TC seals can occur at the seal interface, through cracks (or splits) in the sealing element, or through seal splices (see Figure 13).

l695L - 8 -

2.1. 7 Structural Seals

A structural seal is always used for penetrations located in areas of the containment envelope where there is no steel liner. A puddle flange(s) (or steel water bar) welded to the embedded part reduces or prevents leakage at the interface of the embedded part and hardened concrete and provides anchorage as well (Figure 14(a». Caulking is sometimes also used to provide a seal at this interface. It should be noted that all seal places, cover plates, etc., are either seal welded to the embedded part or are sealed by a gasket placed between the plate and embedded part. Pipes, tubes, cables, etc., which go through the seal plate are sealed by welds, swageloks, pressure fittings or gland fittings (Figure 14 (a» .

Leakage of these seals can occur at the interface of the embedded part and concrete and at seal welds, swageloks, gaskets, gland fittings, etc., (Figure 14).

2.1.8 Preheater Enclosure seal

The preheater enclosure seal consists of a steel angle or plate which is seal welded to a steel enclosure and an embedded part (Figure 15).

Potential leakpaths for this seal are through the seal welds or between the embedded part and concrete. These leakpaths are shown in Figure 15.

2.2 JOINTS IN THE CONCRETE STRUCTURES

Expansion joints and construction joints, especially in the vacuum building, pressure relief valve manifold, pressure relief duct and sections of the fuelling duct which are not steel lined, are a potential source of containment leakage. The amount of leakage depends predominantly on the type of joint seal detail employed.

A special seal detail is used for some expansion joints. This seal consists of a metal strip or plate which is seal welded to an embedded part(s), and/or to the steel liner in areas where it exists (Figure 16). Caulking is applied at the interface of the concrete and embedded part. A polyvinyl chloride (PVC) waterstop is also used as shown in Figure 16.

All other expansion joints are the same as construction jOints (see below) except there is a 1 inch gap between the concrete surfaces which is filled with styrofoam (Figure 17(a».

A typical construction joint detail is shown on Figure 17(b). These joints are sealed using a sealant and a PVC waterstop.

Leakage sources include the concrete/embedded part interface, cracked seal welds, sealant/concrete interface and through the sealant itself.

2.3 CRACKS AND "PORES" IN THE CONCRETE CONTAINMENT SHELL

Cracks and "pores" in the concrete walls, slabs, etc., of the containment envelope are also a potential leakpath. Cracks in concrete arise if the tensile stress exceeds the ultimate tensile strength. This may be caused by drying shrinkage, cement hydration, changes in temperature (i.e., decrease in ambient temperature), excessive creep or live loading.

l695L - 9 -

Permeation of air through the concrete shell occurs through an interconnected system of microscopic pores (capillary pores and gel pores), randomly distributed throughout the concrete (i.e., cement paste). The capillary pores, which are three orders of magnitude larger than gel pores, are mainly responsible for the permeability of concrete.

2.4 VACUUM BUILDING ROOF SEAL

The elastomeric ring seal* connecting the perimeter wall to the roof slab (Figures 4 and 18) is potential leakpath through the containment envelope.

The roof seal is held in place by a post-tensioning system consisting of steel tendons (Figure 18) which are secured to special anchor blocks. This system ensures that there is intimate contact, and therefore a leak-tight seal between the roof seal and its upper and lower supports.

Potential leakage sources are shown on Figure 18.

2.5 VACUUM DOCTS

There are 18 vacuum ducts which connect the valve manifold to the vacuum building. The portion of these ducts which go through the basement of the vacuum building are part of the containment envelope, and are therefore designed for leak-tightness to prevent leakage during a LOCA.

Each vacuum duct consists of several pieces of carbon steel pipe (87 inches in diameter and 3/4 inches thick) joined together with seal welds (Figure 19).

Each vacuum duct has two expansion jOints in order to accommodate movements due to thermal expansion and contraction (Figure 19). Two types of expansion joints are used. One type is a rubber expansion joint. It is comprised of a rubber-neoprene-cot ton duck "laminate", reinforced with steel rings. This expansion joint is tightly secured to the vacuum duct or embedded part with bolts, to produce a seal. The other type is a metal expansion joint (Figure 20). It consists of a steel bellows assembly which is seal welded to the vacuum duct and embedded part. Out of a total of 36 expansion joints, 28 are of the rubber type and 8 are of the metal type.

Leakage through the vacuum ducts can occur at the interface of the rubber expansion joint and vacuum duct or embedded part, through cracks, tears and voids in the rubber and through cracks in the steel bellows and seal welds.

* Fabric reinforced, ethyl propylene diene monomer (EPDM) is the material used for the roof seal.

l695L - 10 -

3.0 TYPES OF LEAKAGE RATE TESTS

Overall testing of the entire containment envelope (or each containment structure) is relied upon heavily in Ontario Hydro's current test program. There are basically three types of overall tests now in use - Commissioning (or Pre-operational) Tests, Continuous Monitoring, and Periodic In-Service Tests. The results obtained from these tests vary in quality and range, .and they represent the summation of leakage through the containment envelope (e.g., penetrations, construction joints, cracks in concrete, concrete permeation, etc.).

3.1 Commissioning Leakage Rate Tests

Commissioning leakage rate tests are performed Defore the station is placed in service. These tests serve a number of important roles. First, they are used to demonstrate that the magnitude of leakage is acceptable. Second, they demonstrate the relationship between leakage rate and pressure. This is necessary to confirm safety analysis and design assumptions and to form a basis for the correlation of future in-service test results. In order to achieve this objective, the measurements are taken at intervals over the complete positive and negative pressure range that containment is expected to operate. Lastly, they should demonstrate that the leakpath geometry is not significantly affected by pressure. A laminar, constant geometry relationship is assumed and should be demonstrated in the tests.

3.2 Continuous Leakage Rate Monitoring

The leak-tightness of the containment envelope (except vacuum building) is continually confirmed during station operation by keeping containment subatmospheric with the exhaust fans for the normal reactor building ventilation systems. Containment pressure, temperature, humidity, air inflow (instrument air, breathing air, etc.), etc., are monitored. Any change in pressure which cannot be accounted for by changes in known variables such as temperature, humidity, air inflow. etc., would indicate a possible deterioration in leak-tightness of the containment boundary.

The leak-tightness of the vacuum building is demonstrated continuously while it is under normal vacuum by noting the number of hours each month that the vacuum pumps run. Any significant increase in running hours could be an indication of decreased leak-tightness.

3.3 Periodic In-Service Tests

Complete containment volume leakage rate tests (excluding the vacuum building) are conducted on a regular' basis by lowering the pressure inside containment 7 to 28 kPa below atmospheric pressure and noting the pressure recovery rate. An approximate leakage rate is calculated based on pressure. temperature and humidity measurements and known/inferred air inflow into containment. The leakage rate determined from this data is compared to commissioning test results and previous in-service test results, and trends in the leakage rate are determined. A trend indicating a continuously increasing leakage rate would indicate decreasing leak-tightness (or leakpath development) due to aging. deterioration, etc.

l695L - 11 -

The vacuum building leak-tightness is also checked on a periodic basis by pumping down the particular chamber being tested to the lower operating pressure limit (i.e .• -14 psig) and then shutting off the vacuum pumps and buttoning up. The vacuum decay is noted and together with temperature. pressure. and humidity data. an approximate ieakage rate is determined. This is compared to previous test results to determine whether the leak-tightness of the vacuum building has changed.

Certain containment appurtenances such as electrical penetrations. airlocks. transfer chambers. dampers. reactivity mechanism deck seals. etc .. are leak rate tested separately on a periodic basis as part of the Safety System Test Program for the station.

4.0 ISOLATION AND INDIVIDUAL TESTING OF PENETRATIONS

Certain types of penetrations in the Bruce NGS A containment envelope can be isolated and tested individually. These include electrical penetrations. the reactivity mechanism deck. airlocks and transfer chambers. and button-up valves and dampers.

The method of isolating and testing these penetrations is summarized in this section.

4.1 Electrical Penetrations

Electrical Penetrations have a secondary seal on the containment side .(Figure 6) to allow the performance of leakage rate tests on each penetration without pressurizing the whole containment building. These penetrations are periodically tested by pressurizing the interspace between the primary and secondary seals to 70 kPa(g) and recording the pressure decay for hour. A leakage rate of 35 kPa/hr or less is considered acceptable. Unacceptable leakage rates are corrected by determining the leak site(s) with a soap bubble solution. and taking the appropriate remedial action (i.e .• sealing procedure).

4.2 Shield Tank Seal

As mentioned previously. the shield tank seal consists of three reinforced. elastomeric sealing elements (Figure 12). The interspaces between the sealing elements can be pressurized and leak tested without pressurizing the whole containment building.

The outer interspace is tested by inflating the outer seal. pressurizing the interspace to 138 kPa(g). and recording the leakage rate. The inner interspace is tested by pressurizing it to 138 kPa(g) and recording the leakage rate. A leakage rate of 138 kPa per half hour is considered acceptable. Unacceptable leakage rates are corrected by determining the leak site(s) with a soap bubble solution. and sealing the leak(s) with the appropriate sealant/material.

4.3 Airlocks and Transfer Chambers

The airlock and transfer chamber door seals have been described in section 2.1.6. The airlock is leak tested by inflating the dual inflatable seals on the inner and outer airlock doors. then pressurizing

1695L - 12 - the interspace to 48 kPa(g), and recording the leakage rate. Acceptable leakage rates vary with the airlock volume. The bigger the airlock the lower the acceptable leakage rate.

4.4 Button-up Dampers or Valves

Dampers (Figure 21) are used at penetrations for the Reactor Vault and Fuelling Duct Ventilation System. There are 2 dampers (or valves) per penetration. In order to test the dampers, they are closed and the interspace is pressurized to a specific pressure. The leakage rate is then monitored and recorded.

5.0 COMMISSIONING AND IN-SERVICE LEAKAGE EXPERIENCE

5.1 Commissioning and In-Service Overall Leakage Rate Test Results

Commissioning leakage rates for various sections of containment ranged from 0.17 (for the PRVM and PRD's) to 0.36 (for RV's) percent of the contained air mass per hour at positive design pressure (+10 psig).

Periodic in-service leakage rate testing of the Bruce NGS A Containment Envelope shows that the leakage rate has increased, but it is still below the maximum allowable leakage rate of 2 percent of the contained air mass per hour at positive design pressure. Much of the increase is attributable to replaceable penetration seals. The rest of the increase seems to be due to the development or deterioration of various leakpaths in the unlined concrete sections of containment caused by ageing.

5.2 Commissioning and In-Service Test Experience for Penetration Seals

Commissioning and in-service test experience provides a good indication of which types of penetration seals are susceptible to leakage (or most likely to lose their leak-tightness). A review of this experience reveals that:

(a) Several of the penetration seals which utilize fittings (i.e., cable penetration seal, instrumentation line seal and pressure fitting seal) have leaked due to loose fittings. The cause of the loose fittings was probably improper tightening or poor installat ion:

(b) Numerous airlock and transfer chamber inflatable door seals have leaked due to pinholes in the sealing element or abrasion of the sealing element. Seal replacement is quite common. To date, most door seals have been replaced about once a year:

(c) Several gasket and o-ring seals have leaked, the gasket leaks being mostly a result of loose hold down bolts: and

(d) There have been numerous leaks between the caulking and embedded part or concrete and enilieddedpart (i.e., for structural seals)*.

* These leaks were detected during commissioning tests.

1695L - 13 -

It can be inferred from the above observations that replaceable seals (including AL and TC door seals), structural seals and seals which utilize fittings are susceptible to leakage.

5.3 Commissioning and In-Service Inspection Experience for Expansion and Construction Joints

Most experience to .date has been from commissioning tests (i.e., construction proof tests and leak rate tests). In-service inspection experience has been limited.

Commissioning test experience is as follows:

(a) There were several leaks between the caulking and embedded part for expansion joints (see Figure 16, Detail 5).

(b) There were numerous leaks between the caulking and concrete (for both construction and expansion joints);

The pressure at which most of the joint leakages were detected (i.e., first noticed) ranged from 2 to 4 psig. Except for a major leak in the expansion joint between the east pressure relief duct and valve manifold, most of the leaks detected were minor. It should also be noted-that all leaks were at expansion and construction joints located in areas of the containment envelope which are not steel lined.

5.4 Commissioning Test Experience for Cracks in Concrete Containment Structures

Commissioning and in-service test experience with cracking is as follows:

(a) Numerous cracks in the vicinity of penetrations, especially in the north wall of each reactor vault;

(b) Numerous cracks (i.e., too numerous to map), several fully penetrating, in the area of the construction opening for the west end closure of the fuelling duct. This closure is fully lined with steel;

(c) Several long cracKs (up to eight feet in length) on the outside face of the south wall of the East Intermediate Fuelling Duct. Crack widths ranged from .002 to .005 inches.

(d) A few cracks in the walls and floors of the pressure relief duct and valve manifold.

(e) Several hairline cracks in the vacuum building roof. All these cracks leaked. Leakage was first noticed at VB pressures of 2 psig.

(f) A few hairline cracks in the vacuum building containment floor. No leaks; and

1695L - 14 -

(g) Several cracks and water leaks developed in the vacuum building containment floor during the floor loading test*.

It is important to note that most of the cracks indicated above are long in relation to their width.

5.5 Commissioning and In-Service Experience for Roof Seal

The performance of the roof seal has generally been quite good to date.

During the commissioning pressure testing of the Vacuum Building, only minor leakage occurred at the roof seal. Most of these leaks were between the roof seal and its supports (Figure 18).

In-service experience shows that the roof seal has performed quite well. The roof seals' continuous good performance has been demonstrated by the fact that the vacuum in the VB has been maintained without problems since the VB came into operation.

* The containment floor is flooded with 18 feet of water and vacuum building is subjected to +8 psig, simultaneously, for a period of 72 hours.

1695L TABLE 1

BRUCE NGS A - QUANTITIES OF EACH TYPE OF PENETRATION SEAL

Unit Seal 0* 1 2 3 4 Total Type (Containment)

Welded Penetration 73 61 61 61 61 317 Seal

Cable Penetration 47 63 63 63 63 299 seal

Inst rument Line 19 31 31 31 31 143 Penetration seal

Metal Bellow 0 30 30 30 30 120 seal

Replaceable Seals 28 119 119 119 119 504

Airlock and 21 11 11 11 11 65 Transfer Chamber Door Seals

Structural Seal 193 193

Preheater 4 4 4 4 16 Enclosure Seal

* Unit 0 includes the Central Fuelling Area, Pressure Relief Duct, Fuelling Duct, Valve Manifold and Vacuum Building.

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REFERENCE- BRUCE NGS A SAFETY REPORT (FIGURE, 4.1-1)

FIGURE 2 PRIMARY HEAT TRANSPORT SYSTEM AND CONTA~NMENT 21.33000-6 REV.31975 5'-0" 251-0" 51-0" , ' ~i

I CONCRETE. :....:.... : [.:'.,.' ----- S.V '. @ ~ 0,,' .",,;., ',0. = l 9 ! 91 :.: '.() ~1 I

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STEEL LINER o. 0 00' I

. " . . '.~'i.' , . , .. ..,..0 .~. . y .. ..~., O~II' .v . '0" . ':~ ..

FIGURE 3 - CROSS - SECTION OF FUELING DUCT. 1 INTERNAL STRUCTURE 9 PERSONNEL AIRLOCK 2 EMERGENCY WATER STORAGE TANK 10 EQUIPMENT AIRLOCK 3 DISTRIBUTION AND SPRAY HEADERS 11 SERVICE TUNNEL 4 VACUUM DUCT 12 CATCH BASIN 5 VALVE MANIFOLD 13 ROOF/WALL SEAL 6 PRESSURERELIEF VALVE 14 BASEMENT 7 PRESSURERELIEF DUCT lS SUCTION PIPES 8 MONORAIL AND HOIST 16 DIFFUSING SCREENS 17 VACUUM DUCT COVERPLATE

FIGURE 4 - CUTAWAY VIEW OF VACUUM BUILDING AND VALVE MANIFOLD CONTAINMENT SIDE

1/411 THICK C,S. PLATE (TYP.) ---. -/~~---_/:_,- SEAL WELD ~(r...... " v''00

(TYP.l l '.0 o' . ~o. . O~. .,•. °0 • 'of : ' NO POISON PLATE REQID. \ '. ~ .0' FOR C.s. PIPE. 1/411 THK. FILLER MAT t,-. il' 0'" '. ; ~'d6 ... ' S. ST. POISON PLATE (TYP) lLLMENITE ~ Cl .. : 0 (FOR S S. PIPE ONLY) SAND :~ .. 1/411 THK. BRASS STRIP (TYP.) (FOR COPPER TUBE ONLY)

,,"/

SEAL TO ... ,,-,v.q ....;;O() .. v .i>.~ .• ~.o.. SUIT Q .~ .... '(7. , . 0 . q. . o. " ., .. (THIOKOL OR • . . .: .v. O' WELD (TYP.) EQUAL) 0:

" '. 1" . . . 6 ./y--_ .. : . ' ---//. ~CONTAINMENT VAULT LINING

C9 DENOTES MAJOR POINTS OF POTENTIAL LEAK DEVELOPMENT (TYP.)

FIGURE 5

TYPICAL WELDED PENETRATION SEAL...... ~~ COAT THREADS WITH .~ Q ~ .Oro .: LUBRIPLATE 630-2- O' STOCK CODE No. 892N- o /.) ". . .., 0055 ~ 0. 0 'f ' . ., . ~. , o • of Ii o.o..~: .....

CABLE SEAL (GLAND) SEE DETAIL BELOW ...... rfl. . LEAKAGE THROUGH o' ...... 6 ~ 'V.: THE CABLE (SEE • <>.'. d~.. :'.~ 0°0'. F..IGURE 10) 00. (9: "t. REF. .. /~ :., " .. '. 0. ,. L3 0 o • --_ OUTSIDE WALL OF

0. • .. 0 .. o.

o cP" -.1 CONTAINMENT '. 0':' 0. :'0: . /lcP ~.~. GLAND ~ \" WAVE SPRING 0) 0 0 .c4 "'t GASKET- .. ' -:J NUT '" \WASHER .'1. :0 SILICONE .'" RUBBER <{Q'oO CD DENOTES MAJOR '000 '4 . 0 POINTS OF POTENTIAL ... '1o~ LEAK DEVELOPMENT . (TYPICAL>

FIGURE 6

DETAIL OF EPOXY (L GLAND .MOULD CABLE PENETRATION SEAL ASSEMBLY ITYP'> GASKET VAULT LINER INSIDE ~I REACTOR VAULT WALL CONTAINMENT. SHROUD END RINGl ,- FASTENING PLATE

I END RING --- i---, ,~~.I, I SHROUD (SLEEVE) I \ ~:~. ~...' SHROU D PLATE RI SITE WELD N? ,,')'0 '. 21\ ./ "~~ I I ,./ SEAL BELLOWS I \, . '" ,.C) " ~ // ,DETAIL A , :\-: .,." " iMIN.I./ -- --" -. '-- ,,," SITE WELD " DETAIL B 111-----11 T ---I

o r-

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SEAL BELLOWS

END RING FIll.ET

DETAIL A DETAIL B

FIGURE 7 TYPICAL METAL BELLOWS SEAL OUTSIDE . . CONTAINMENT. ~_I 5'-6" DIA.(REFJ. 1 .. - 4 -5" DIA.(REF.) SEAL~ SEAL:JI WELD WELD 7 JO , 3/8' >:X 0 r!. 4-2 1/2" DIA.(REFj~ I : !J:lO~I?,-C-,-DIAJBE:f) i CD--: COVER PLATE <: iD 1 i I 'I 3 -81/4" DIA.(REF.) I i" , - EMBEDDED --l UPPER FLANGE . , , I PART. , I EL. 1 HOLD ~/ -----_.-673 -1" DOWN -~ (REF,) FRAME

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(b) DENOTES MAJOR " POINTS OF POTENTIAL LEAK DEVELOPMENT. INSIDE CONTAINMENT SECTION

I 1 11 i ..(UPPER FLANGE 5 -6 DIA. ! \ I, ! II 2 II R... n ./ L1 l ~/4" I I = i C0 OUTSIDE ~! CONTAINMENT I ~! - I

in! 001 GASKET- KLINGERIT- 1000 I I REINFORCED COMPRESSED .050" 1 11 i ASBESTOS JOINTING AS -l' JHICK 4 -81/8 DIJ\, rMANUF. BY THE ANCHOR , il cg !(f" 4'-5" D1A. iJ,L i PACKING CO. LTD. OR EQUIV. 1/1611 THK.x 31-9 11 x :i 3/8 1.0. 1 3/8 t-f-~ // - L --- ;t"- ii 31-11 7/8" O.D.

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I1 LOCK WASHER 112" DIA. SEAL BELLOWS ASSY. BOLT-1/2 , MEDIUM SERIES DIA. ! DETAIL-I '---PUMP FIGURE 8

HEAT TRANSPORT PUMP-METAL BELLOWS SEAL EL.6751-6t1 BOILER I -_._---+------_;'1'-10" dIA.(:~.._._._--- I (REF. ) 11 -i 71~101/4 i DIA.~REF'> I I i ,I . 'TOP i

SEE DETAIL I ACCESS DOOR BELOW COVER ROTATED '# .' " FOR CLARITY. . s .

.' "j INSPECTION PORTS .. ' :7, ~. ... ROTATED FOR . 1 CLARITY. o . ~. ~ . j .. ~.BOTTOM

if •. SEAL .... waD C0

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\.!::)o "-~ . DETAIL I L l -B

(lJDENOTES MAJOR POINTS '6F POTENTIAL LEAK DEVELOPMENT. (TYP')

FIGURE 9

B01LER METAL BELLOWS SEAL EXPANSION JOINT

VACUUM DUCT

CD DENOTES MAJOR POINTS OF POTENTIAL LEAK DEVELOPMENT

FIGURE 10: PRESSURE RELIEF VALVE INSIDE CONTAINMENT (VALVE MANIFOLD)

PRESSURE , : RELIEF VALVE: : STRUT ~,- - t-rr- PRY SEAT' ~ I' 11:II

~~~:OLDCS\t I. =~.I I'j I :; FLOO~_l-. _ L n II ~I $11 L,", RUBBER . ""'1.r,- i,? I, )lff'/ f r" .' . ii : ,I O-RING . . Q ...... r,A,"... ";.t" /: .f.l //'-,..:' ;:,!,""<1" '. ~ .. ~' ;" ",'," ;/ /,< .:.' , .' ~ ' ... : I~' . ..:;- './ .',/ .:'J .. .~ . "..' , • • I ,'J' "",', INSIDE ,d., .D;.: .;. • :QTI'I:II:rr-~01.', ... '.:" . "0 CONTAINMENT "/i,l:ii .' ." " .,. . ,r.'1 '\ .. .' . c. ((~1\,VACUUM . ' .. ~. a VACUUM :: I :: DUCT RING...... v. >J 1+ , .' DUCT. PLATE ' ".: . . ' 'n ~ , 0 •. <.,( ~ O 0 . '. ,"...... "~:';'

OUTSIDE CONTAINMENT

@ DENOTES MAJOR POINTS OF POTENTIAL LEAK DEVELOPMENT

FIGURE II

DETAIL OF O-RING SEAL (PRESSURE RELIEF VALVE) TY.P > ~ ji) C9 318 / 0~ / ~/ .

. / • :'. t • /~ . .1 >- / INSTALLED POSITION

PRESSURIZED 0 - POSITION --=rrrt' II ~Ii~...J_

20 O' tHIS OUTSIDE DIR~CT-I CO\ITAI N MENT. ION. i ON~Y \. , i) (Q DENOTES MAJOR POINTS OF POTENTIAL LEAK ~. I DEVELOPMENT (TYP.) TYp.C9. ,J

1~ .." ,'\ . ) ,yo. .' ' o~ ~~ h . SHIELD TANK EXTENSION . ) ". . . ~ ~.'. . " (1. , •.

PARTS KEY I UPPER DIAPHRAM MOUNTING 10. COVER 18. LOWER DIAPHRAM MOUNTING 2 RUBBER CEMENT II. INNER DIAPHRAM 19. INFLATER CONNECTION 3. CAP SCREW(BUTTON HEAD) 12. OUTER DIAPHRAM 20. WASHER 4. CORNER DIAPHRAM CLAMP 13. BELLVILLE WASHER 21. WASHER 5. DIAPHRAM CLAMP 14. FERRY CAP"COUNTER22. NPT JAM NUT 6. DIAPHRAM CLAMP BOREij SCREW 23. STEEL COUPLING 7. DIAPHRAM CLAMP 15. DIAPHRAM CLAMP 24. STEEL PLAIN NIPPLE a DIAPHRAM CLAMP 16. INFLATABLE SEAL 25 VALVED NIPPLE 9. DIAPHRAM CLAMP 17. BACK UP BAR

FIGURE /2

REPLACEABLE SEAL (SHIELD TANK SEAL) 0_ IJ • ' 7- I -_-:c'~_JGROUT p, I j

I J / I I I - \-y I I I ------'-- - 1- - ~ _ _ ---l riJIL _ ---J \ I DOOR \\- SEE DETAIL "01" I JIll PLAN u- I I I -t~JLL "~'- INFLATED I (0DENOTES MAJOR POINTS OF POTENTIAL rt- , LEAK DEVELOPMENT I III: DEFLATED -,L-I~ "Ir 'I I I I I I l-- -- .1 I ...,I----$EALI D I (SEAL DETAIL)

.... '"J • . o I', '-,- II- ~GROUT FIGURE /3 I ,0", ~.lt I I ' -- ,-_., " 1--

TYPICAL DUAL INFLATABLE DOOR SEAL SECTION SI GASKET (S) l\

(FLANGE ;' I SEAL r-pLATE CONCRETE

SURFACE! ! WELD !

",(..p .~. ;,~' '.' ~ " PUDDLE .0 •. ' , . , ~. PUDDLE1 " J FLANGE t ,,£ FLANGE' ,'1; '.

( B)

~SWAGELOK TUBE //' (S) PIPE TUBE SEAL PLATE

,', -;.,.. It ~ .' ~ OJ .., ".V'" '.' .: ... -..;.----- _. . c ~ •• J .'" .. (C) ( D)

C0 DENOTES MAJOR POINTS OF POTENTIAL LEAK DEVELOPMENT ( TYPJ

FIGURE 14

STRUCTURAL SEAL IN COMBINATION WITH OTHER SEAL TYPES L THROUGH STEEL ENCLOSURE WELDS ETC.

CONTAINMENT SIDE

SEAUNG ANGLE WELDED TO STEEL ENCLOSURE a EP l

'>-:J .... .l: ..,' 'Of ~', . ' '. : • ...6, .. :' (;?".,':'" '. :,7..~,' ',~G." '~. . c/C?" :' ,CP~" ~ • ['7 , ,0 ' C9-' • Q \) .. EMBEDDED' PART' ,~' () p. '. ~ :',. I I,.... • , . '1 ' o , ,'. G . 'l' ft t) I I". d. . /J , ' ~ 0 , O/) 1:;6 . • ., I I . LF . ':V.

( I I EMBEDDED CONCRETE MAIN GIRDER? I r-----ANCHOR BOoLT.

I I 0 I I @

C0 DENOTES MAJOR POINTS OF POTENTIAL LEAK DEVELOPMENT

FIGURE 15

PREHEATER ENCLOSURE SEAL ( ,

II tI . _OUTSIDE FACE OF tI BAR 4 x I/S ~ I 3/S I" ~ I 31 "LINER PLAT.E . WALL 00 ".,.'~'~.It'J2V2~1~.~I~PAN~!Qti--_. :". '-,L:----t...--=-- --.-.--..- ~ L ~~IUNG~..PANE'L". ~9'INT. -c$. -E.P. 1712. L~ 1\ ~I V VS . ~ ., .' JO 0. 0" '. .~. '0' '. . ': ....J- l ". I. : 4.... '. V •.•. -:: ...•: .. 04 ..•. ~. :'Pi1.'~."'W~ ..,

I' '•... g.~.'(1 ...... '. ~-, \~.~ • 't ." 'y-.. .:.. , '0. ., OUTSIDE FACE \ ""'GRIND CORNER OF CHANNEL , " ',~' OF LINER PLATE \\AS REQUIRED (~R~ll~g~o~~-~1Z~~~~IO~__JOINT ~.\..BAR 4". 1/8" BASE SLAB) \ .~ - PVC WATERSTOP "'-" . NOTE: PVC WATERSTOP IS ALSO !fJ5TtFfoLU-h~W.R l PRECAST FLR. PANEL 00 WALL PANE~ ~~DS~NO£~:AILS 2,3,4 a 5 BUT

RE~~t~AuLT . (TYPICAL SEALING OF LINER PANELS @ EXP. JOINT.)

GRIND CORNER OF 1...III'i~8~~SION __- CHANNEL TO SUITnI/21/l/\I __ 'h} I ~ :r1 EPI713 OR EPI71 . /::/

.J ~1:Y8"tl" l~~ ~17avl L BAR 4"xVg~ lIS. . ~DDITIONAL 1/4"x I 1/2"x 1511 --50~ff LINER (ONLY IN LINED AREAS OF FUELLING STRAP ANCHORS AT 31-0" c/c. BY CONST. DETAIL 3 DETAIL 4 DETAIL 5 FUELLING DUCT iti~~~~lgJV~v~~lM~ .~~ ... ~IPRE.I..~Ef'L. ULErL'Nb ~..!.!1 Yl EEl~SSEC~Cf' .. 'JIS. t: 8NTFOLD' (0DENOT6S MAJOR POINTS OF POTENTIAL LEAK DEVELOPMENT FIGURE 16 SPECIAL EXPANSION JOINT - TYPICA.L DETAILS CAULK

11 I 1/2 CAULKING DEPTH I t:~':~P0LYET~-ri.ENE ~~ ~o, "~0? • I :.~ ~E~UL~A~~. ~PHA~T~ ". ,.' • ,~. i 1-, --~ 'b-~ I " \1 • Co" ,4 PVC WI\:TERSTOP ." ,[) ,., i ' ,-t, --WHERE"~SHOWN + b . r- ~~::~AM. 0' °/1 ~.oO ' . . ~ ~j':'"P9LYETH'f~NE TA/lo" ".... •. ," ."".. ." <;)

\ 1/211 CAULKING "DEPTH "---- CAU LK . NOTE: THIS TYPE OF DETAIL IS USED FOR ALL EXPANSION JOINTS EXCEPT THOSE REQUIRING THE SPECIAL JOINT DETAIL SHOWN IN FIGURE 29 (SEE SECTION 6.0 OF THIS REPORTJ

(A) TYPICAL EXPANSION JOINT DETAIL

I 1/211x 45° CAULK I ') " , . ' , 'Ocr ,'(j, ~., , ',0',Oa~1' ~ I ',~,Q,PVC WATERSTOP" ,OJ '~, '<' ' • , ,J (j I ' WHERE SHOWN ... t'., ' , ... ~ . , ,~. ON DWGS. ,0 '. , 1/3 THICKNESS', 4- OR AS SHCM'N , ! '0 ~~~~NC. DETAI~ . -.." I CO', ! .. "'0-' i .... g:) 0, . I I ! ;,'", 0 j ! I i ' r ~ VARIES " i,0 ,9.L( \, , ; _ 0',., • _ "0 I - 0.' 0,;, . . u! , ,D , 'c ~ ., " -:..9

• , ' j'I ..... " .. I °a .. 0" 61 D

(8) TYPICAL CONSTRUCTION JOINT DETAIL.

FIGURE 17 BRUCE NGS A - TYPICAL EXPANSION AND. CONSTRUTION JOINTS TIGHTEN NUT SO THAT ROOF SEAL r WILL BE HELD ~L SECURLY IN PLACE I/ZV \ UPPER t POST TENSIONING ROOF SEAL I TENDON (TYP.) SUPPORT,' (CG -25125-1002) (BG-25159-i 1010) ! \ sO 11 745'-111/8 i I TOP POST TEN- ./

SIONING TENDON 1 11 T 745 - 6 518 i BEARING fL. ---- ! (DG- 25159-1011)

+1 SEAL I (CG-25125-1001) 1 I 80 - 3 R ! I • . BOT. POST TEN - \ ROOF SLAB ! SIONING TENDON \ BEARING ~. ~\ LOWER ROOF

FIGURE 18 ROOF SEAL DETAILS _____-628'-0"L

VACUUM DUCT

VALVE MANIFOLD !-FLcx:REL. 602'- 0"

I VACUUM BUILDING 7 r- CCNTAINMENT FLOOR 1 I EL. 595 - 9" I-i_ 1 t _f_590 -0". :! I EXPANS~ I FIRST JOINTS \S)-' RISER:\ \

BASE SECTION L THROUGH DUCT STEEL FIELD JOINT

@ DENOTES MAJOR POINTS OF POTENTIAL LEAK DEVELOPMENT

FIGURE 19

SECTION SHOWING VACUUM DUCT 8 LOCATION OF EXPANSION JOINTS......

.J c:r - z <;- ! ~ o::E - z 3/16 I -C\I t 1/2" X 1/2"x III LG. STOP SA (6-REQ1D EQUAL SPACED) 3116 FLEX ISLE SEAL DIM. TO ACCOMODATE "3/1611 VERTICAL MOVEMENT MATERIAL: ASTM A240, 31-71/211 RAD. TYPE 304 OR 304 L, . ----'-----~/(/ 18 GAUGE 4'-1 1/2" RAD. 1 81-3" CIA. BOLT 21RCLE .

J lit.. I :Y411 HOLES I EXISTING

C0 DENOTES MAJOR POINTS OF POTENTIAL LEAK DEVELOPMENT.

FIGURE 20

SECTION OF STEEL EXPANSION JNT. (BELLOWS) FOR VACUUM DUCTS ACTUATOR CARTRIDGE TYPE MOUNTING FLANGE O-RING SHAFT SEAl.: '------:AND COVER PLATE

TWO-WAY THRUST BEARING VALVE BODY SHAFT BEARING

DISC/DR IVE SHAFT VALVE ASSEMBLY DISC

SEAT MATING :SURFACE

FIGURE 21 TYPICAL DAMPER DETAIL CONTAINMENT, LEAK TESTING AND REPAIR, DOUGLAS POINT NUCLEAR GENERATING STATION

by D.F. Cooke - Ontario Hydro; Bruce Nuclear Generating Station C.A. Simmonds - J.M. Hopwood Atomic Energy of Canada Limited CANDU Operations R.K. Jaitly -

PRESENTED AT: CSNI Specialist Meeting on Water Reactor Containment Safety Toronto, Canada. October 17-21, 1983

SPONSORED BY: OECD Nuclear Energy Agency Committee on the Safety of Nuclear Installations

HOSTED BY: Atomic Energy of Canada Limited, CANDU Operations CONTAINMENT p LEAl< TESTING AND REPAIR p DOUGLAS POINT Abstract

Douglas Point Generating Station is the first full-scale prototype of the CANDU series of nuclear generating stations, and has been in operation for nearly two decades.

In 1980, a program of containment testing and repair was undertaken during a station outage for other reasons. At the request of the Atomic Energy Control Board of Canada (AECB), testing was carried out at the design overpressure of 41.4 kPa(g) (6 psig). The initial test at this pressure in April, 1980 yielded a leak rate of 59 dm3/s (125 scfm), 3.6 times the design maximum. This prompted an extensive repair and testing program. In all, five series of leak rate measurements were performed. Subsequent to each test series, efforts were directed towards repairing the leaks. The leak rate at 41.4 kPa(g) (6 psig) was eventually reduced to as low as 23.6 dm3/s (50 scfm) after extensive efforts at refurbishing the reactor building containment. Containment repairs consisted mainly of replacing deteriorated caulking material and sealing a large portion of the interior of the perimeter wall with a butyl rubber coating.

The key safety criteria for nuclear plants in Canada are the reliability and overall response of the plant safety systems in limiting the consequences of major accidents. The Douglas Point containment assessment was based on these criteria. Radiation dose to the public was the measure used. Hence, containment effectiveness was assessed as a component part of the overall plant response in mitigating public dose, or as an intermediate measure, plant fission product release for specific accidents. w~ile an increase in the tested containment leak rate at a nuclear power station may be a deviation from the original plant design, the fission product release In specific accidents may be very little affected, and the overall plant accident response acceptable. This is the case at Douglas Point. It was found that a nearly ten-fold increase in leak rate is required to double a predicted release for a given source term for likely worst accidents for example, a large loss of coolant accident (LOCA), based on 1-131 and assuming release of building pressure through the Filtered Air Discharge System three hours after the accident. TABLE OF CONTENTS

1.0 INTRODUCTION

2.0 REACTOR BUILDING DESCRIPTION

3.0 OBJECTIVE

4.0 TEST PROCEDURE 4.1 Preparations 4.2 Test Methods 4.3 Measurement Techniques

5.0 SEQUENCE OF TESTS AND REPAIRS 5.1 Series 1 5.2 Series 2 5.3 Series 3 5.4 Series 4 5.5 Series 5

6.0 IMPACT OF CONTAINMENT LEAK RATE

7.0 LEAK TESTING OF CANDU 600 STATIONS

8.0 CONCLUSIONS

LIST OF TABLES:

TABLE 1: 1980 REACTOR BUILDING LEAKRATE TEST RESULTS TABLE 2: RELEASES FROM CONTAINMENT; ORIGINAL & REVISED ASSESSMENTS TABLE OF CONTENTS (Cont'd)

LIST OF FIGURES:

FIG. 111: REACTOR BUILDING SECTION - DOUGLAS POINT NUCLEAR GENERATING STATION FIG. 112 : REACTOR BUILDING SECTION - DOUGLAS POINT NUCLEAR GENERATING STATION FIG. 113: SECTION OF VERTICAL CONSTRUCTION JOINT - DOUGLAS POINT NUCLEAR GENERATING STATION FIG. 114: REACTOR BUILDING PLAN - DOUGLAS POINT NUCLEAR GENERATING STATION FIG. 115: TYPICAL LEAK PATHS - DOUGLAS POINT NUCLEAR GENERATING STATION FIG. 116: REACTOR BUILDING JOINT AT DOME SPRINGLINE - DOUGLAS POINT NUCLEAR GENERATING STATIoN FIG. 117 : TYPICAL CANDU 600 GENERATING STATION REACTOR BUILDING

APPENDIX A HISTORY OF LEAK RATE TESTS 1.0 INTRODUCTION

Douglas Point Generating Station is the first full-scale prototype of the CANDU series of nuclear generating stations and has been in operation for nearly two decades.

In 1980, a program of containment testing and repair was undertaken during a lengthy station outage for other reasons. At the request of the Atomic Energy Control Board of Canada, testing was carried out at the design overpressure 41.4 kPa(g) (6 psig). This was the first time the reactor building had been subjected to this pressure since the construction proof test in 1965, prior to equipment installation. All intervening tests had been performed at a pressure of 13.8 kPa(g) (2 psig), the most recent of these in 1973. The design maximum leak rate was 0.1% of building volume per hour, at an overpressure of 41.4 kPa(g) (6 psig).

2.0 REACTOR BUILDING DESCRIPTION

The building is cylindrical in form, comprising a reinforced concrete outer wall, roofed by a hemispherical steel dome, with reinforced concrete and steel internal structures. (See Figure 1)

The outer wall is 39.62 m (130 feet) internal diameter and 32.31 m (106 feet) high above grade; 23.16 m (76 feet) of this is 1.22 m (4 feet) in thickness and the upper 9.1 m (30 feet) (parapet wall above dome spring line) is 0.61 m (2 feet) thick. The wall extends below grade at 1.22 m (4 feet) thickness to form the outer boundary of the various basements. The depth of the wall below grade varies from 3.35 m (11 feet) to 9.45 m (31 feet). The self weight of the wall exceeds considerably the upward force on the dome from the max. internal overpressure. The dome is hemispherical and 19.81 m (65 feet) internal radius, constructed by welding 12.7 mm (1/2") thick formed steel plates. An annular parapet roof of light weight concete is provided between the top of the parapet wall and the dome.

- 1 - The reactor building wall is constructed of ordinary reinforced concrete, constructed in segments with smaller filler segments between them, cast afterwards (See Figs. # 2 & 4). Containment relies on caulking of the construction joints as well as the vinyl coating on the interior of the perimeter wall to provide a seal.

The reactor building is designed to withstand a uniform internal overpressure of 41.4 kPa(g) (6 psig). At this pressure, maximum allowable design leak rate is conservatively specified to be 0.1% of the building volume per hour.

3.0 OBJECTIVE

Including the series of five Reactor Building Leak Rate Tests that were carried out on the Douglas Point Reactor building in 1980, a total of 11 such tests had been carried out until then. As indicated in Appendix A, a number of leak tests were carried out at 13.8 kPa(g) (2 psig) only, and results were extrapolated for 41.4 kPa(g) (6 psig) design overpressure.

The objective of the containment testing/repair program undertaken in 1980 was threefold: a) Establish the relationship between reactor building leak rate at 13.8 kPa(g) (2 psig) and 41.4 kPa(g) (6 psig). b) Verify that no irreversible damage is done to containment at 41.4 kPa(g) (6 psig) overpressure; for example to any elastomeric seals. c) Carry out repairs as necessary to achieve target leak rate.

The test sequence was designed to achieve the first two objectives. Leak rate measurements were performed, while pressurizing the building at 13.8 kPa(g) and 41.4 kPa(g) (2 and 6 psig). The reactor building was then depressurized and measurements taken at 27.6 kPa(g) and 13.8 kPa(g) (4 and 2 psig).

- 2 - 4.0 TEST PROCEDURE

4.1 Preparations

Building pressure had only previously been raised to 41.4 kPa(g) (6 psig) in 1965 prior to equipment installation. It was therefore deemed prudent first to subject representative samples of components to a pressure test of 41.4 kPa(g) (6 psig) in the main airlock. No ill effects were observed.

Prior to pressurizing the reactor building all interior doors were left open and all tanks were vented. Various isolations were performed to prevent operation of safety systems - dousingt emergency injectiont etc. As a general rulet systems - particularly those with containment penetrations - were left in a state as close as possible to normal so that the test results would be more meaningful. All non-essential air supplies were isolated.

4.2 Test Methods

For each pressure testt the leak rate from the containment building was measured in -two ways; by pressure run-up tests and by controlled leakage tests. The sequence was as follows:

Pressure run-up to 13.8 kPa(g) (2 psig) Controlled leakage test at 13.8 kPa(g) (2 psig) for 2 hours;

Pressure run-up to 41.4 kPa(g) (6 psig) Controlled leakage test at 41.4 kPa(g) (6 psig) for 2 hours;

Depressurize to 27.6 kPa(g) (4 psig) Controlled leakage test at 27.6 kPa(g) (4 psig) for 2 hours;

Depressurize to 13.8 kPa(g) (2 psig) Controlled leakage test at 13.8 kPa(g) (2 psig) for 2 hours.

- 3 - 4.3 Measurement Techniques:

Leak rate tests were performed in two ways: i) Pressure-Run-Up Test: Following isolation of the reactor building. it was pressurized initially by breathing air system. at a rate of 3.5 kPa (0.5 psi) per hour. At a pressure slightly below the test pressure. all air supplies except the low pressure instrument air system were isolated. The building pressure continued to rise slowly because of the maintained low pressure instrument air flow. The rate of supply of instrument air was measured. The leak rate was equal to the difference between the rate of supply and the rate of increase of building air content (calculated by increase in pressure). ii) Controlled Leakage Test: The building was pressurized as above and a controlled leak rate was superimposed on to the actual leakrate to stabilize the reactor building pressure. The leak rate was equal to the difference between the rate of supply and the controlled leak rate.

Inlet and outlet air flows were measured using Rootes gas flow meters. The times at which the meters turned over were manually recorded, along with inlet and outlet pressures and temperatures.

Meanwhile building pressure, vapour pressure and temperature were recorded along with atmospheric pressure at times corresponding to inlet Rootes gas flow meter turnovers to allow for corrections to standard conditions.

The data gathered was indexed to the inlet Rootes gas flow meter turnovers. Time intervals between data sets were of the order of 3 - 4 minutes. An array of 30 - 40 data sets constituted a full test at a given building pressure.

- 4 - The Ontario Hydro computer program "PRESTEST" used to calculate the leak rate was originally operated in a mode which calculated individual leak rates over the 3 - 4 minutes time intervals and averaged the resulting array. Because of the relatively small time intervals, this generated noise in the array and greater uncertainty in the final result. A smoothing technique was applied later which significantly reduced the scatter. This consisted of performing the individual leak rate calculations between data sets separated by N time intervals and stepping this through the array of raw data. Through this supporting procedure, the standard deviation in the final leak rate measurement was reduced to typically 2.36 dm3 (5 scfm). (See Table 1.)

- 5 - 5.U SEQUENCE OF TESTS AND REPAIRS

5.1 Series 1

At pressures above 37.23 kPa(g) (5.4 psig), air was observed bubbling in significant quantities from the irradiated fuel transfer mechanism in the spent fuel bay because an isolation valve was leaking. The head of water in the bay is " not adequate by itself to prevent leakage at 41.4 kPa(g) (6 psig). Starting the shuttle transport pump increased the head sufficiently to stop this leakage and allow the test to continue.

Pipe and cable penetrations for the new Emergency Coolant Injection System (ECIS) accounted for some of the leak rate. Other sources of leaks included the following: (See Fig. #5 and Fig. #6)

main airlock inner pressure equalizing damper.

uncapped redundant boiler vent line at main steam line penetration.

cracked weld on redundant vault drain line.

12 cable penetrations.

The rest of the leaks observed during Series 1 tests were associated with the concrete reactor building structure:

expansion joint in the fuel transfer trench where it penetrates the reactor building wall.

reactor building wall vertical construction joints.

the dome seal at the top of the RIB wall.

The pipe and cable penetration leaks were easily eliminated using conventional sealing techniques. The defective caulking in the fuel transfer trench expansion joint was removed and replaced.

- 6 - 5.2 Series 2

The repairs described above were effective in reducing the leak rate at 41.4 kPa(g) (6 psig) from 65 dm3/s to 37 dm3/s (138 scfm to 78 scfm). the rate measured in series 2 tests. While pressurized during the series 2 tests. very slight leaks were observed. at 5 cable penetrations and at the upper ECIS pipe penetration. The bulk of the leakage was attributed to the reactor building structure itself. Leaks were detected at the vertical wall joints. the dome seal area and 1n the ceiling of the fuel transfer trench near the previously repaired leak.

At this point, an investigation of the nature of the concrete structure revealed the probable root of the problem. The reactor building is constructed of individually poured concrete slabs with smaller filler segments between them cast afterwards. Inter-slab joints run the full height of the building. Figure 3 illustrates the section of a vertical joint. The joint caulking and interior paint form the primary containment barrier. The vertical joint caulking was generally in good condition, its deterioration could not account for the multitude of leaks seen externally at the vertical joints.

There were two most-probable paths for leakage past the primary barrier into the construction joints:

defective caulking at wall-floor joints which intersect vertical joints.

hairline cracks in the concrete which run into the joints.

A simple test proved that the latter provided an effective leak path to the joint system. A pipe flange was glued to the interior wall directly over a suspect hairline crack. This allowed a 2S mm (one-inch) length of the crack to be pressurized via an air line from a tank with a pressure gauge. The rate of pressure drop in the tank through the 2S mm (one-inch) length of crack indicated a leak rate of 0.19 dm3/s (0.4 scfm). Migration of air in the crack was confirmed with "leak tek" (soap bubbles) as far as one meter (3.28 ft) away from the flange - on the far side of a caulked joint.

- 7 - I _ The interspaces between interior vertical joint caulking and the vinyl waterstops are all interconnected via the horizontal wall - floor joints. Therefore, a leak at some point on the interior could follow a tortuous path of least resistance to the exterior. In other words, leaks observed in the fuel transfer trench or at the dome seal could easily have originated at a remote point in some defective floor - wall joint.

Repairs subsequent to Series 2 tests consisted mainly of removal and replacement of defective caulking at floor - wall and ceiling - wall joints. In addition, the expansion joint between the uppermost floor and the perimeter wall was refurbished. A few areas where there were significant defects (spalling, large cracks) in the concrete were sealed by coating with a urethane membrane.

5.3 Series 3

The repairs following Series 2 improved the 41.4 kPa(g) (6 psig) leak rate by about 5 dm3/s (10 scfm). Other than the continuing problems with the reactor building structure and the main airlock equalizing dampers, there were no other leaks.

In an effort to stop the leaks through the structure, the upper half of the interior perimeter wall was sealed with a butyl rubber membrane applied by airless spray in emulsion form. This area contained the largest concentration of concrete shrinkage cracks. In addition, the caulking at the underside of the uppermost floor slab was replaced with great difficulty. Other caulking at vertical joints in interior crosswalks was also replaced. A urethane rubber membrane (Gaco UWM-28) was applied at the exterior of the dome seal.

Because of the still relatively high leak rate, a complete review was undertaken of all system states and isolations set up for the tests as well as a check of instrumentation and leak rate calculation methods. No oversights were discovered.

5.4 Series 4

The repair effort following Series 3 reduced the leak rate at 41.4 kPa(g) (6 psig) to as low as 28 dm3/s (59 scfm). During Series 4, the pressure was

- 8 - held at 13.8 kPa(g) (2 psig) for 24 hours and at 41.4 kPa(g) (6 psig) for about 20 hours while exhaustive leak searches were carried out.

Approximately 10 leaking spots were founa in the dome seal area as compared to about 160 leaks found originally in this area. Also. several large blisters had developed in the rubber membrane in areas of imperfect adhesion to the concrete. Leaks were still apparent at the exterior of the vertical wall joints. although these were judged to have diminished somewhat.

A phenomenon was observed during Series 4 tests which had not been seen in earlier tests. In an exterior reactor building sump. air was bubbling from the box drains system header. This system is a network of channels under the lowest reactor building floor. Leakage into the construction joint grid was reaching the exterior via this route. This had not been seen in previous tests. even though the sump had been regularly checked. When there was sufficient head of water in the sump. the bubbling stopped and the measured leak rate diminished by about 5 dm3/s (10 scfm).

Diagnostic work following Series 4 tests confirmed the interconnection of the box drains with the construction joint system. roe box drains header was pressurized with air and "reverse" leaks were found at an elevated floor to wall joint where caulking had been removed but not yet replaced.

Between Series 4 and 5 tests. more repair work was carried out. All remaining accessible caulking at floor and ceiling joints was replaced - no matter what its condition. Caulking in all floor joints in the lowest floor was replaced and this floor surface was sealed with epoxy paint. Most of the remaining interior perimeter walls were sealed with Decadex Butyl Emulsion Membrane.

5.5 Series 5

The Series 5 test was designated as the final "as left" test. The leak rate at 41.4 kPa(g) (6 psig) was 29 dm3/s (61 scfm). No special measures were taken to seal the airlock dampers. However. the airlocks had been individually tested prior to Series 5. and met leakage specifications. Some bubbling was observed in the external drain sump while the building was under pressure. It should be noted that the leak rate of 29 dm3/s (61 scfm) is most likely somewhat higher

- 9 - than the actual leak rate as it was discovered that a defective pressure switch had been in place, which presented a leak path. The switch has since been replaced.

Subsequent to the final test. the reactor building underdrainage system was pressurized with helium gas. which. in turn. pressurized the construction joint system. A leak search ws performed inside the reactor building with 'leak tek' (soap bubbles). but no bubbling could be found. However, the presence of helium could be detected, using a Matheson gas leak detector in many. seemingly unlikely. locations on the reactor building perimeter walls, such as:

in most locations where bolts are anchored into the wall.

in some areas where the vinyl paint was cracked.

in places where interior walls intersected the perimeter wall and other places where caulking was inaccessible.

additionally. and most significantly. a strong indication of helium was detected where the dousing tank support legs go through the boiler room floor. Since these legs rest on concrete pads on bed-rock. it is apparent that this and probably other similar areas. such as beneath the calandria vault and others represent leak paths.

- 10 - 6.0 IMPACT OF CONTAINMENT LEAK RATE

The effect of containment leak rate on plant accident response was assessed. The key safety criteria for nuclear plants in Canada are the reliability and overall response of the plant safety systems in limiting the consequences of major accidents. To this end, containment effectiveness is assessed as a component part of overall plant response in mitigating public dose, or as an intermediate measure, plant fission product release for specific accidents.

The leak rate testing also coincided with the commissioning of a redesigned Emergency Coolant Injection System (ECIS). With the new system used to restrict fuel damage in the event of a Loss of Coolant Accident (LOCA) predicted fission product releases into containment are lower.

For a normal containment response, the fission products inside containment can escape initially by distributed leakage through the building walls and subsequently by planned use of the Filtered Air Discharge system. In predicting accident consequences, early use of the Filtered Air Discharge system, i.e., three hours following the accident, and pessimistically low charcoal filter efficiency (96% for iodine removal) were assumed. Given these assumptions,. the major part of the predicted release would be through the Filtered Air Discharge system and a very small part through the distributed leakage from the building structure. Increase in the leak rate will therefore only minutely affect predicted release. For example (See Table 2), for a postulated LOCA with ECIS assumed to be unavailable, 50% increase in leak rate contributes only 7% increase in predicted release. In fact it is found that nearly a tenfold increase in leak rate is required to double a predicted release for a given source term.

From Table 2, it can be also seen that for the LOCA accident analysed in accordance with the latest methodology, (e.g. Application of Computer Program FIREBIRD-III) and the effect of the redesigned ECIS in reducing the source term, considerably outweigh the effect of increased leak rate. This release is in fact substantially lower than the original predictions.

- 11 - Operator response to the use of the Filtered Air Discharge system at three hours is intended to bring down containment pressure to atmospheric. After this time, the heat sinks in containment can cope with any remaining heat sources, so that no containment repressurization would occur.

7.0 LEAK TESTING OF CANDU 600 MW STATIONS

The reactor building utilized for the four recently commissioned CANDU 600 stations is a pre-stressed concrete containment structure; has a double dome; a cylindrical perimeter wall; and a flat base slab. The perimeter wall is slip- formed; concrete for the dome is placed continuously and completion of the structure is followed by post-tensioning. The internal concrete surfaces of the containment are lined with epoxy and fiberglass epoxy. The various internal walls and floors are structurally separate from the perimeter wall (see Figure #7 and Figure #2 for comparison with Douglas Point).

These features combine to provide a high standard of leak tightness with no deterioration expected in lined pre-stressed concrete.

The acceptance value of leakage rate for CANDU 600 stations is 0.5ioof building volume per day at a design overpressure of 124 kPa(g) (18 psig). Values of about 0.3% were achieved in tests carried out at various stations.

8.0 CONCLUSIONS

At Douglas Point, the leak rate at 41.4 kPa(g) (6 psig) was reduced by a factor of 2 after exhaustive efforts. The equivalent hole size corresponding to the aggregate final leaks is approximately one sq.cm (0.16 sq.in). The residual leak rate is the net result of many small leaks past the primary containment barrier into the network of construction joints. Following conclusions are drawn from the 1980 testing and repair work program:

1. The leak rate measurement is probably higher than the actual leak rate at present since a defective pressure switch had been in place during the final test which presented a leak path. The switch has since been replaced.

- 12 - 2. All of the accessible caulking at interior floor and wall to perimeter wall joints has been replaced and is now in good condition.

3. All visible hairline concrete shrinkage cracks have been sealed with a flexible membrane.

4. Sealing of the outside of the building (top of RIB wall at dome seal) achieved marginal success.

5. All pipe and cable penetrations are sealed.

6. Any remaining air leakage from the building originates from small imperfections in the concrete surfaces within the building, both in the perimeter wall and interior floors and walls, and migrates to the outer surface of the building through the network of construction joints to the path of least resistance, i.e. the vertical construction joints, box drains. system and dome seal.

7. Increase in leak rate results in only marginal increase in release of radiation dose following a tOGA. A ten-fold increase in leak rate is required to double the release of radiation dose, based on the methods used and the conservative assumptions made, i.e. based on 1-131 and use of the Filtered Air Discharge System three hours after the accident.

- 13 - TABLE1 1980 REACTORBUILDINGLEAKRATERESULTS

Measured Leak Rate Test Measured Adjusted Standard Pressure Leak Rate To 6 Psig Deviation Series Type of Test kPa(g)(Psig) dm3/5 (scfm) dm3/5 (scfm) dm3/5 (scfm)

1 Pressure Run-up 14 (2.03) 24.4 (49.56) 79 (167.34) 7.9 (16.24) Controlled Leakage 15 (2.17) 16.4 (34.83) 51.8 (109.69) 4.0 (8.47) Pressure Run-up 41.2 (5.98) 58.2(123.32) 59.3 (125.68) 3.4 (7.28) Controlled Leakage 28 •84( 4 •18) 42.4 (89.91) 65.1 (137.86) 10.3 (21.75) Controlled Leakage 14.1 (2.05) 22.7 (48.08) 76.68(160 .35) 5.4 (11.45)

2 Pressure Run-up 14.3 (2.08) 12.1 (25.55) 39.5 (83.68) 2.1 (4.48) Controlled Leakage 14.7 (2.13) 16.1 (34.18) 51. 6 (109.39)11.8 (25.08) Pressure Run-up 41.2 (5.98) 37.0 (78.47) 38.0 (80 .46) 1.4 (2.95) Controlled Leakage 41.4 (6.00) 36.1 (76.39) 36.9 (78.14) 0.93 (1.97) Controlled Leakage 28.3 (4.11) 21.9 (46.31) 34.3 (72.74) 1.5 (3.24) Controlled Leakage 14 (2.03) 7.3 (15.36) 24.5 (51.90) 0.43 (0.92)

3 Pressure Run-up 13.6 (1. 97) 10.5 (22.19) 36.7 (77.85) 2.4 (5.10) Controlled Leakage 13.9 (2.02) 11.2 (23.72) 38.2 (80.85) 2.6 (5.65) Controlled Leakage 14 (2.03) 13.6 (28.89) 46.3 (98.12) 4.5 (9.63) Pressure Run-up 41. (5.94) 30.6 (64.88) 31.5 (66.74) 2.8 (5.96) Controlled Leakage 40.8 (5.92) 31.1 (65.89) 32.0 (67.84) 3.0 (6.38) Controlled Leakage 28.1 (4.07) 21.7 (46.00) 34.3 (72.71) 2.8 (5.93) Controlled Leakage 14.27(2.07) 14.2 (30.09) 47.1 (99.78) 5.1 (10.86)

4 Pressure Run-up 14.5 (2.11) 8.1 (17.11) 26.3 (55.64) 3.5 (7.34) Controlled Leakage 14.9 (2.16) 5.5 (11.69) 17.4 (36.90) 5.5 (11.63) Controlled Leakage 13.9 (2.01) 7.0 (14.76) 23.8 (50.42) 2.9 (6.22) Controlled Leakage 41.7 (6.0 5) 23.6 (49.95) 23.6 (50 .05) 2.4 (5.19) Controlled Leakage 42.3 (6.13) 28.0 (59.26) 27.6 (58.58) 5.5 (11.68)

5 Controlled Leakage 41.2 (5.97) 28.1 (59.54) 28.9 (61.34) 1.3 (2.70)

- 14 - TABLE 2 RELEASES FROM CONTAINMENT ORIGINAL AND REVISED ASSESSMENTS

Original Assessment Revised Assessment

Bases: Design leak rate Leak rate 1.5 x design

Large LOCA Worst case LOCA

Latest analysis methodology

Loss of Coolant Accident (LOCA) 10 Ci 1-131 1.8 Ci 1-131

LOCA with ECIS unavailable 670 Ci 1-131 720 Ci 1-131

- 15 - 1 CALANDRIA 12 PRIMARY HEAT TRANSPORT PUMPS 30 BUILDING VENTiLATION 3 REACTOR SHIELDS 17 EMERGENCY WATER STORAGE TANK 31 FUELLING MACHINE ACCESS DOOR 5 FIJELLING MACHINE 18 SPRAY TANK PIPES 32 DOUSING WATER PASSAGES 7 BOOSTER AND ABSORBER RODS 21 FUELLING MACHINE VAULT 33 DECONTAMINATION TANK 8 PRIMARY HEAT TRANSPORT 23 AIR LOCK ENCLOSURE ~ STEAM GENERATOR INSULATION REACTOR OUTLET HEADER 26 THERMAL SHIELD COOLING DUCTS 35 HELIUM STORAGE TANK 9 PRIMARY HEAT TRANSPORT 28 PRESSURE WALL 3ll SHIELD TANK REACTOR INLET HEADER 29 DUMP TANK 37 ION CHAMBERS

FIGURE 1 REACTOR BUILDING SECTION DOUGLAS POINT NUCLEAR GENERATING STATION REINFORCED CONCRETE

rI"" _ •••

TYPICAL CONSTRUCTION JOINTS

FIGURE 2 REACTOR BUILDING SECTION DOUGLAS POINT NUCLEAR GENERATING STATION I / , I " WATERSTOP JI

INTERIOR

"-d

KEY CAULKING

SHRINKAGE CRACK

4

EXTERIOR

SLAB

FIGURE3 SECTION OF VERTICAL CONSTRUCTION JOINT DOUGLAS POINT NUCLEAR GENERATING STATION ~ c h <= C ~ ~ U

r"G ~ Gor 0 05 oL -.J 0 o( )

III1 ~~ 0 ~" 0 "0

TYPICAL CONSTRUCTION JOINTS

FIGURE 4 REACTOR BUILDING PLAN DOUGLAS POINT NUCLEAR GENERATING STATION DOME/PERIMETER WALL JOINT

DEFECTIVE CAULKING AT FLOOR/PERIMETER WALL JOINT

FLOOR

MAIN AIRLOCK PRESSURE EQUALIZER DAMPER I CIRCUMFERENTIAL i AND LONGITUDINAL I: SHUTTLE TUBE. JOINTS r LEAKING VALVE IN THE BAY

EXPANSION JOINT IN THE SPENT FUEL TRANSFER TRENCH

FIGURE 5 TYPICAL LEAK PATHS DOUGLAS POINT NUCLEAR GENERATING STATION DOME STIFFENER

OUTSIDE INSIDE I

SEALING RING

\ ADDITIONAL SEAL APPLIED . .... THIOKOL I ~:;. CAULKING COMPOUND ----~~~~~-::~ BASE PLATE

",:.:" -...... , .....:•...• ::t...... ~~, ..• t!-..I. . , '" -' ',~ .: .. ~~. . '.-: " . .', ...;~~.!"'." .~:::~"" ", . ~;~.:. " .. .;." .. ~r . ~. CORBEL . ,/ .~.. : " GROUT " ~ .'

FIGURE 6 REACTOR BUILDING JOINT AT DOME SPRINGLINE DOUGLAS POINT NUCLEAR GENERATING STATION CONSTRUCTION RE.INFORCED JOINT ..~.... :--' ...... ',. ..' ..... '." ~ CONCRETE

:,','

i:

.:...• f_ t. ~'. 1-.- ': .'.~. f:- ....'- II i.:" . < " ------.L t: r;~'~ I.: -0 ,' r; -, '-.

PRE.STRESSED CONSTRUCTION SEPARATION MEMBRANE CONCRETE JOINT

FIGURE 7 TYPICAL CANDU 600 GENERATING STATION REACTOR BUILDING APPENDIX A

HISTORY OF LEAK RATE TESTS

DATE LEAK RATE @ LEAK RATE ~ LEAK RATE 13.8 kPa(g)(2 psig) 41.4 kPa(g)(6 psig) % of R/B dm 3/s (scfm) dm 3/s (scfm) VOLUME PER HOUR @ 41.4 kPa(g) (6 psig)

TARGET 4.7 (10.0) 16.8 (35.7) 0.1 1965 - 1l.2 (23.8) 0.07 1968 July 20.3 (43) 68.4 (145*) 0.40 1968 Oct. 5.7 (12) 18.9 (40*) 0.12 1969 Nov. 15.6 (33) 52.9 (112*) 0.33 1970 Nov. 7.6 (16) 25.5 (54.1*) 0.15 1973 Sept. 3.1 (6.6) 9.4 (20.S*) 0.06 1980 Apr. 16.5 (35) 59.5 (126) 0.36 1980 May 12.3 (26) 36.8 (78) 0.22 lY~O July 10.4 (22) 31.6 (67) 0.19 1980 Sept. 7.1 .(lS) 23.0 (50) 0.14 1980 Oct. - 28.8 (01) 0.18

* 41.4 kPa(g) (6 psig) leak rate extrapolated from value measured at 13.8 kPa(g) (2 psig)

A-1 CSNI SPECIALIST MEETING ON WATER REACTOR CONTAINMENT SAFETY TORONTO, CANADA 17th - 21th Oct. 1983

SIGNIFICANT ASPECTS ASPETTI SIGNIFICATIVI RELATI- CONCERNING THE LEAKAGE VI ALLA MISURA DEL TASSO DI TESTS OF THE SAFETY PERDITA DEI CONTENITORI DI SI CONTAINMENTS: CUREZZA: VALUTAZIONE DELLA OPERATIONAL EXPERIENCE ESPERIENZA OPERATIVA. EVALUATION

by: G. Grimaldi

Direzione Centrale Sicurezza

Nucleare e Protezione Sanitaria ENE A Via Vitaliano Brancati n. 48 00144 Rome - Italy SIGNIFICANT ASPECTS CONCERNING THE LEAKAGE TESTS OF THE SAFETY CONTAINMENTS OPERATIONAL EXPERIENCE EVALUATION

ABSTRACT

Our experience during inspection of leakage tests on nuclear plants con- tainments focalized two significant aspects: 1) confidence on results of overall leakage tests and relative inspec- tions; 2) confidence level on leak tightness capability conservation between two subsequent tests.

With reference to the firs point ENEA/DISP developped a computer code for independent evaluations of leakage rate (CA.TA.PE.COS). Statistic values of interest are computed together with the upper confidence li- mit and the evaluation of leakage rate after rejection of experimental spurious data. Recentely we introduced in the test program for the evaluation of the leakage rate a test without cont4inment pressurization (zero differen- tial pressure test), preliminary to the overall leakage measurement at test pressure, with the following objectives: a) verify the effectiveness of the actions intended to minimize thermal gradients and their consequences on the representative errors; b) verify if conditions exist to mask an effective leak, such as compre~ sed air or water reentry; c) verify, before pressurizing test volume, if some gross leak exists. Fist evaluations, still in progress, confirmed the effectiveness of the method with reference to points a) and c) above; we are confident on c~ pacity of the "zero differential pressure test" also with reference to point b).

Referring to point 2) above our present activity is in revising local tests results with regard to: d) optimization of maintenance program, particularly with reference to inspection frequencies; e) individuation of critical leakage paths with reference to valves ty- pes and related operating conditions; 2.

f) referring to the preeceding point e), eventual substitution of existing valves with others better quality types and/or revision of valves ope- rating conditions. Following these lines we intend to perform periodic overall leakage tests right after plant shut-down, aimed to verify the effectiveness of the en-" tire maintenance and testing program with reference to containment capa- bility to performe its function between two subsequent tests. 3.

ASPETTI SIGNIFICATIVI RELATIVI ALLA

MISURA DEL TASSO DI PERDITA DEI CONTENITORI DI SICUREZZA

VALUTAZIONE DELL'ESPERIENZA OPERATIVA

SOMMARIO

La nostra esperienza nella verifica delle prove di tenuta dei contenito- ri per impianti nucleari ha evidenziato due aspetti significativi: 1) garanzia dei risultati nella valutazione delle prove di tenuta globa- Ie e verifiche relative; 2) garanzia di conservazione delIa capacita di tenuta fra due prove suc- cessive.

Con riferimento al primo punto l'ENEA/DISP ha sviluppato un cod ice di calcolo per la verifica indipendente del tasso di perdita (CA.TA.PE. COS). Vengono calcolate Ie grandezze statistiche di interesse, compreso il limite di confidenza superiore e la valutazione del tasso di perdi- ta dopo scarto dei punti anomali. Recentemente e stata introdotta nel programma di prove per la valutazio- ne del tasso di perdita una prova senza pressurizzazione del contenito- re (prova a pressione differenziale zero), preliminare alIa prova di te- nuta globale in pressione, con i seguenti obiettivi: a) verificare che Ie azioni intese a minimizzare i gradienti termici e Ie loro conseguenze sugli errori di rappresentativita siano efficaci; b) verificare se esistono condizioni in grado di mascherare un reale ta! so di perdita, come rientro di aria compressa 0 di acqua; c) verificare, ancor prima delIa pressurizzazione del volume di prova se esiste una perdita consistente. Le prime valutazioni, tuttora in corso, hanno confermato l'efficacia del metodo con riferimento ai punti a) e c); si e fiduciosi sulla efficacia delIa prova a pressione differenziale zero anche relativamente al punto b).

Con riferimento al punto 2) si stanno riesaminando i risultati delle pr£ ve di tenuta locali con l'intento di: d) ottimizzare il programma di manutenzione con riferimento aIle freque~ ze di verifica; e) individuare Ie vie critiche di fuga con riferimento ai tipi di valvo- le e aIle condizioni operative relative; f) con riferimento al punto e) precedente, predisporre eventualmente la sostituzione di valvole esistenti con altre di migliore qualita e/o rivederne Ie condizioni di impiego. 4.

In questa linea si e orientati a richiedere l'esecuzione delle prove di tenuta globale subito dopo la fermata dell'impianto, allo scopo di verificare l'efficacia dell'intero programma di manutenzione e di prova nel garantire la capacita del contenimento a svolgere la sua funzione nell'intervallo tra due prove successive. 5.

INTRODUCTION

Scope of the containment building in nuclear plants is to limit fission products external release in case of accidents with fuel damages. Containment capability is verified by periodic tests of containment buil ding and related safeguards. An overall leakage test is performed as well, with plant in shut-down con dition and the containment building at the peak pressure calculated for the design reference accidents. Leakage rate measurement is performed by application of the "perfect gas law" to the containment atmosphere and by evaluation of presure decay versus time. Instantaneous air mass in the containment volume at time ti is:

yi = kPi/Ti (1)

with: k: scale factor Pi: absolute pressure at time ti Ti: average absolute temperature at time ti (on spatial basis)

Corresponding relative leakage in a fixed time interval between t1 and t2 is:

Lr = 1 - « P1 T2)/(P2 Tl». (2)

Allowable leakage rate values are normally very low (few per mille of the air mass in the free volume of the containment). Consequently temperatures and pressures in eq. (2) must be determined with great precision and resolution. Furthermore statistical evaluation of leakage rate is necessary, based on different determinations of eq. (2), due to inherent statistical variations of phisical measurements. As an example leakage rate evaluation can be based on a "plot" of istantaneous masses in eq. (1) versus time and a linear regression of experimental points. On this basis leakage rate is evaluated as: b K . t (3) a where: L leakage rate r K scale factor b slope of least squares line a intercept t time interval in hours 6.

1. MAJOR DIFFICULTIES IN LEAKAGE TESTS

One of major errors in overall leakage test measurement is determined by "rapresentativity errors" associated with the difficulty to evalua- te physical quantities of interest as accurate as necessary. Specifical ly thermal gradients due to internal (residual heat) and external (solar irradiation) sources make difficult to evaluate the average instantaneous temperature of test volume. For this reason containment volume is divided into several subvolumes each containing one temperature measurement sen- sor. Consequentely global instantaneous temperature is calculated as:

T = VI T1 + V2 T2 + ••• + Vn Tn (3) wi th VI, ... , Vn subvolumes in % of total Tl, ... , Tn related temperatures

"Vi" coefficients in eq (3) are calculated on volumetric basis,conside- ring the influence on each temperature sensor essentially by the surroun ding air.

The smaller subvolumes and thermal gradients are the more exact is the value of the average air temperature. Due to this fact, sometimes at the end of an overall leakage test the "mass plot" do not result linear. In such a case leakage rate evaluation becomes an hard task and the test has to be repeated with a different and a more accurate choice of subvo- lumes on a physical rather than a geometrical basis. A second difficulty in overall leakage measurement is related to test verification. For example a straight line as a result of "mass plot" ap- plication is not sufficient to assure the correctness of the calculated leakage rate. Experimental points in fact could be distributed on a strai ght line, but not on the right one. The superimposed leak verification test, as per ANSI/ANS- 56.8-1981, is generally a valid method to verify the leak test. However some problems still remain: - the superimposed leakage evaluation is based on total free volume, wh£ se calculation is affected by large uncertainty; the verification criterion is based on a difference and then is little sensitive to systematic errors, such as 'representative errors' (tem- perature measurements or leakages due to compressed air leak or water reentry, of opposite sign respect to the real leakage). Particularly a water reentry ooth in pressure test and in verification test can be masked in such a verification. A third group of difficulties is related to the not easy repeatibili- ty of the local tests. 7.

2. SCOPE

Scope of this paper is to refer ENEA/DISP experience in leakage tests verification, and consequent improvements. In particular we focalized two significant aspects: 1) confidence on results of the evaluation of overall leakage tests and relative inspections; 2) confidence level on leak tightness capability conservation between two subsequent tests. With reference to point 1) a code was developped for an evaluation of leakage tests independent from plant utilities (CA.TA.PE.COS). Recent~ ly we introduced in the test program for the evaluation of the leakage rate a test without containment pressurization (zero differential press~ re test) with the aim to improve the" verification tests". Referring to point 2) above our present activity is in revising the local test results with the aim to optimize the maintenance program, to individuate critical leakage paths, to substitute if necessary critical valves with better quality ones.

3. "CA.TA.PE.COS" CODE

CA.TA.PE.COS code is normally used by ENEA/DISP for inspection verifi- cation during leakage tests of containment buildings in Italy (nuclear power reactors and research plants). Basically input data (temperatures, pressures, vapore pressures) are e- laborated with the "Mass plot" method; the "reference chambers" method is available as well but not more used. Statistical evaluation of the data is based on a "Student" distribution A data rejection criterion is used, based on the following formula:

labs V(i)/ - Fs • SO> a (4) where labs V(i)/: absolute value of the difference between the generic istantaneous mass y(i) and the calculated mass on linear regresion basis yc(i)

student coefficient based on the number of experimental points and the assumed confidence level

standard deviation on "V" distribution

An alternative method to "mass plot" based on "travelling windows" is available in the cases where it is necessary to minimize the effect of periodic disturbance such as temperature sistematic errors with co- stant period (24 hours). Basically this last method computes leak rates on time intervals of 24 8.

hours (disturbance period) with a progressive shift equal to the time between two subsequent data acquisition. The disadvantage of this method is the necessity for test duration grea- ter than 24 hours; however the method can results useful for metallic containments without isolation from external conditions.

~ A SIGNIFICANT TEST: GARIGLIANO 1977

On June 1977 a leakage test was executed on the plenum between the two equipments a:cesses of Garigl iano power stat ion, at a pressure of 1.11 kg/cm2• The tested volume is an horizontal cilinder of about mc. 110. The con- ditions of such volume are representative of the whole containment buil ding (a metallic spheroid of about mc. 48300) referring to atmospheric conditions. The test intended to locate the causes of the poor statistic significa~ ce of the tests performed on the Garigliano containment in the previous years. The test volume was divided in two subvolumes, each containing two tem- perature sensors and two humidity sensors, only one absolute pressure. sensor was used. Experimental data were elaborated as below: 1. "Mass plot" without spurious data rejection (Graf. 1) Input data (instantaneous pressures,vapore pressures.and temperatures) and elaborated data (instataneous masses, calculated masses on best fitting basis) are reported versus time. It can be Qbserved a large data distribution around the interpolating straight line, expecially in the daily hours when the solar irradia- tion determines large thermal transients on the test volume.

2. "Mass plot" of the same data with spurious data rejection application (Graf. 2). It can be noticed (see table 1) a better statistical distribution of instantaneous masses (Standard deviation less than in the first case). The code rejects all experimental data in the day time.

3. "24 hours travelling windows" (Graf. 3). The intent was to sinchroni ze the disturbance due to the thermal gradient in the test volume. The diagram reports leakage rates calculated on daily basis every hour and the average leakage rate. Single leakage rates are still scatte- red but the periodical characteristic of the disturbance is elimina- ted. 9.

4. "12 hours travelling windows" (Graf. 4). Periodic disturbance is computed alternatively with direct and oppo- site phases, therefore its effect increases. The experience has showed the necessity: - to reduce the influence of the solar irradiation and other thermal sources as well, that can introduce thermal gradients in the test volume; to monitor the test volume temperature with sufficient accuracy aimed to well represent the real temperature in any time.

On this basis an experimental program was started on the Garigliano sphere containment (~8300 m3) aimed to determine the better practical ac- tions to reach the preceding "goal points". Basically this program provided an accurate examination of external and internal heat sources and their effects on transients inside the test volume. The test report [1], issued on July 79 reported experimented techniques intended to reduce external thermal gradients (temporary containment i~ sulation by water spraying on the metallic sphere during leakage test) and to correctly monitorate the test volume (for a good test it is ne- cessary to use several temperature sensors, about fifty, distributed in the free volume with care of the vertical and radial thermal gradients and moreover a good air omogeneization).

5. BASIS CONSIDERATIONS ON THE ZERO DIFFERENTIAL PRESSURE TEST.

Due to the difficulties about the superimposed leak test verification, recentely we introduced in the test program for the evaluation of the leakage rate a test without containment pressurization (zero differen- tial pressure test). A similar approach is performed in France by "Bu reau Veritas DJ This verification test, preliminary to the overall leakage test has the following objectives: a) ~erify the efficacy of the actions intended to minimize thermal gr~ dients and their consequencies on the representative errors, b) verify if conditions exist to mask an effective leak, such as com- pressed air or water reentry; c) verify, before pressuring test volume, if some gross leak exists.

The principal consideration for this verification test is that no lea- kage exists if no differential pressure exists between theinside and the outside atmosphere. 10.

A normalisolatilln of the containment system is performed after instru- mentation installation without pressurization. Dataco1lection and eva- luation are performed as for a normal pressure test with the "mass plot" techni que. This plot is expected to be linear and zero sloped if no changements exist in the test atmosphere and the used instrumentation and codes work properly. In effect during the test the temperature variations in the containment can produce some pressure increase or decrease and men produce a real leakage with free mass modification. Ussually it is not expected a significative leak rate due to this cause. The problem is still under evaluation to define. practical acceptance criteria.

6. CAORSO NOV. 1980 TEST

Caorso power plant is a BWR with a nominal power of 860 MWe. The containment building is a Mark II with a free volume of about 8000mc. Maximum allowable leakage rate is 0.5% at a pressure of 3.07 kg/cm2• In Nov. 1980 the first leakage rate test was performed after commercial operation date. On the basis of the good results of the preoperational test, when perfect straight lines were obtained as a result of the mass plot both at full pressure and at reduced pressure, test volume was divided in 7 subvolumes both for vapor pressure and temperature measurements. Preliminary to the pressure test a "zero differential pressure test" was executed for a ti- me period of 12 hours. Graf. 5 reports absolute pressure, average vapor pressure, dry air pressure, average instantaneous temperature and calcu- lated free air mass in the test period. Note that primary water temperature was slightly increasing during the first part of the test due to residual heat. After nine hours from the beginning of the test the RHR*system was actuated to reduce primary wa- ter increasing temperature. Correspondentely the mass plot lost the regular horizontal aspect obser- ved till that moment (see fig. 5). On this consideration during the pressure test the RHR system was not u- sed, but .the necessity of limiting primary water temperature brought to use the some system, at full power, right before the beginning of the pressure test. That caused larg therma) gradients in the test volume during the pressu- re test. In fact at the end of the test the mass plot showed a non linear shape (Graf. 6).

* RHR: Residual Heat Removal 11.

An average leakage rate widely below the maximum allowed was estimated but with a low confidence level. In conclusion the agreement was reached with the utility to repeat the leakage test at the next refuelling.

7. CAORSO MARCH 1983 TEST

On March 1983 the second operational test was performed on Caorso-con- tainment building. For this test specific actions [31 were introduced able to reduce ther- mal gradients in the test volume and to better minitorate temperatures. The leakage test was programmed at the end of the refuelling period to reduce the influence of the residual heat. A contin~ys use of the RHR system was provided during the test but only 1 Division and parziali- zed with the aim to remove exactly the residual heat produced and then with primary water and metal vessel temperatures as constant as possi- ble. The atmosphere between primary and secondary containment was sligh- tly warmed to reduce thermal exchange from the primary containment. A fine suddivision of the test volume was performed with 19 subvolumes each with two temperature sensors. Finally, during the test, action w~ re performed by reactor operators such to mantain primary water level as constant as possible (in a range of about 1 cm). Graf. 7 shows the mass plot of the test volume during the zero pressu- re test. Two phasis can be noted. In the first one an excessive positive (from outside to inside) leakage rate appears due to drain pipes of steam II nes inadvertentely opened. Note that these drain pipes communicate with the turbine building at a pressure greater than the extablished pressu- re in the primary containment for the test. At the end of the first phasis, when the drain pipes were closed the mass plot suddenly assumes a much more smooter slope but non even hori zontal. At this time the pressure gage used for the zero pressure test was discovered defective. It was substituted with a new tested one_but the zero pressure test was not repeated. During the pressure test a linear mass plot was obtained; a leakage ra- te widely below the allowed one was computed; the verification test with a superimposed leak confirmed the measured leakage rate. Fig. 8 shows in the firs part the mass plot during the pressure test, in the second part the mass plot with the superimposed leak. Test results confirmed the correctness of the actions introduced for performing the test. 12.

The adopted test system resulted so sensitive to clearly reveal a varia- tion of about 200 It. in the test volume (1 cm. variation in the reactor water level).

8. "ZERO PRESSURE TEST" ON ESSOR PLANT

Essor is a research plant in Ispra (Italy) whose operation is under Eu- ropean Economic Comittee responsability. Its containment building is a metallic cylinder with an emispheric dome on the top with a thermal in sulation from the external atmosphere. Free volume is about 52000 mc. An annual" leakage rate test is programmed at 0.1 kg/cm2 with an allow~ ble leak rate of l%/Day. Storically the first leakage rate tests were performed with 51 subvolu- mes (original test). In 1971 a reduced method was introduced with only 8 subvolumes based on the good correspondence between the original and the reduced test results,obcained in such occasion. In 1983 ENEA/DISP requested a repetition of the test verification, no more repeated after 1971. Several "zero pressure" tests were performed with the subsequent more significant result. The first test on April, 29-1983 showed a large positive (from outside to inside) computed leakage rate (+ .422 %/Day). The test was stopped and a gross leak was discovered due to an open pipe through the con- tainment. It was found that a pipe (15 mm in diameter) in a system recentely di- sactivated had been inadvertentely Left passing the containment. Such pipe should have been blank flanged. After the necessary corrective actions another "zero pressure" test was performed on May, 14 - 1983. The mass plot showed a positive calculated leakage rate of .487/Day. A compressed air reentry was supposed and the test suspendend for neces- sary verifications. The compressed air circuit was modified for a better isolation of the containmente building. After that a further test was performed on June, 30 - 1983.Graf. 9 reports the significant data collected during the test and the calculated leakage rates. It can be noted an apparent leakage rate in the first 16 hours and a ze ro leakage rate in the last 8 hours when thermal gradients were nearly absent. On these considerations an insufficient representativity was supposed in the first 16 hours and a prevision was made for next tests for more detailed volume subdivisions and/or omogeneization of the free air in the test volume. 13.

9. CONFIDENCE LEVEL ON LEAK TIGHTNESS CAPABILITY CONSERVATION.

Preliminary to the overall leakage test on Caorso plant, local tests were performed in March 1983. According to ANSI the utility was requested to report measured values before (as found) and after (as left) maintenance. A first examination of the collected data was executed even before the final utility test report. Subsequent significant problems were observed: a) P42 Closed cooling water system (Penetration T23-GG062 - FF309 and FF326 butterfly valves, 8" pipe.::.diameter). A widely large leakage rate was found before maintenance ( .....0,313 Nmc/h) • The measured leakage rate after maintenance was about 0,033 Nm3/h. During maintenance consistent oxide fragments were observed inside the valves. b) B21 Feedwater system (Penetration T23-GG0516 F032A, E51 - F013 and G33 - F039 valves, on 20" diameter pipes). An excessive not measured leak was .discovered on F039 valve. A not complete valve closure was found. A significant leak rate reduction was measured after valve disc lap- ping but still too large. An acceptable leak rate was measured only after a complete revision of the F032A valve (0,938 Nm3/h). c) T48 Ventilation system. (Penetration T23-GG017 FF 306 (20" diameter). FF307 (6" diameter) and FF308 (20" diameter) valves). An excessive leak was found before maintenance (8,36 Nm3/h). The measured leakage rate after maintenance was 0 Nm3/h. d) T48 Ventilation system (.FF310 and FF311 valves, 20" diameter). Several excessive leakage rates were measured after repeated maint! nance. Only after various subsequent open-closure cycles of the FF311 valve an acceptable leakage rate was measured (0,075 Nm3/h). Incoherent oxide fragments were observed inside all T48 system pipes and valves (with teflon seals). From these experiences it is necessity to take in a greater conside- ration local leak rate tests, with the aim to improve the confidence level on the leak tightness capability conservation in the time bet- ween two subsequent overall tests. The possible future action on this problem are: optimize the maintenance program, particularly to inspection fre- quencies; - individuate critical paths with reference to valve types and rela- ted operating conditions; - provide eventually substitution of existing valves with other bet- ter uqality types and/or revise operating valves conditions. 14.

Today experience on containment leak tests bring to perform the overall leakage tests execution right after plant shut-down, aimed to verify the effectiveness of the entire maintenance and testing program with re ference to the containmente capability to perform its function betw~en two subsequent tests. 15.

This paper was prepared utilizing the experiences of the Nuclear Reactor Safety Division of ENEA/DISP during the inspection activity. The author is grateful to Mr. P. Bastianini of Ispra C.C.R. Euratom for ESSOR data collection and evaluation and to Mr. A. Linari by ENEL and to Mr. R. Romanacci by ENEA/DISP for their precious collaboration. 16.

TAB LEI

Garigliano plant leakage test on equipment acces on June 1977. Comparison of the various methods results:

1. "Mass plot" without data rejection. 2. "Mass Plot" with data rejection 3. "24 hours Travelling Windows" without data rejection. 4. "24 hours Travelling Windows" with data rejection. 5. "12 hours Travelling Windows" No data rejected

Upper (1) N. Leak Rate Standard N. Confidence Rejected O/oo/Day Deviation Interval Points

1 1.67 .21 .341 0 2 1.52 .06 .099 57

3 1.55 .94 .174 0

4 1.47 .76 .147 7

5 1.83 6.85 1.111 0

(1) U.C.!. 24000 to 95 SA/B SA: Standard deviation on slope of least squares line B : Intercept of least squares line 17.

References:

[1] CRUCIANI, GRIMALDI, LINARI, ROMANO - Presupposti Specifica Tecnica di Prova - Centrale ENEL Garigliano - Luglio 1979. [2] D. HURE' - Statistical Analysis of the Leak Rate Measurements of Reactor Containments.

[3] LINARI, RAVERA - Procedura di Sorveglianza 4.7.A.1.jA Sistema di Contenimento Primario Misura del Tasso di Perdita Globale - ENEL Caorso [4] 10 CF 50 App. J (5J ANSI.ANS 56.8 - 1981 ( R.TR.PE.CO.S - CRlCOlO Dll IRSS~ 01 '(RDIIR DR UN CONllN! !O~( 01 UCURtllR C(N I RRI t DEl Cfl~ DLlIINO rRO~R lENUIII 'ORIR INDRtSSO R"ARlCrtIIRIUR[ 25/6/11 SCRRID rUMI' RNe All '"'[0110 •••••••• MSi 'lOl DR" ~f1ODD IISSOl 10

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CONSIDERATIONS WITH RESPECT TO

PROCEDURES AND MEASURING METHODS

FOR CONTAINMENT LEAK RATE TESTS.

BY

W.van DOMSELAAR.

H. PRUIMBOOM.

R.V.van WIERINGEN.

This paper is presented at the CSNI Specialist Meeting on Water Reactor Containment Safety. Toronto, Canada 17th - 21 st October 1983. -1~

Considerations with respect to procedures and measu- ring methods for containment leak rate tests

Authors:

W. van Domselaar Gemeenschappelijke Kernenergie- centrale Nederland, Dodewaard, Netherlands.

H. Pruimboom Netherlands Energy Research Foundation, Petten, Netherlands.

R.V. van Wieringen Ministry of Social Affairs and Employment, Nuclear Department, Voorburg, Netherlands.

Preface

The information presented in this paper is for a great deal the result of discussions and advices of a working group (Working group Safety Containment) . Besides the authors of this paper the other permanent members of this working group are:

J. Bornkamp Nuclear Power Station Borssele, Borssele, Netherlands.

R.J. van Santen Dutch Authority for Boiler and Pressure Vessels, The Hague, Netherlands.

J.W. de Vries Interuniversity Reactor Institute, Delft, Netherlands.

The authors wish to thank the other members of the working group for their contributions to the various items of this paper. -2-

1. Introduction

It was already in the seventies that it became apparent that uniform test methods and interpre- tation of test results were necessary for con- tainment leak rate measurements. After discussions between the regulatory body and the Netherlands Energy Research Foundation (ECN) it was decided that a proposal should be made, by the Research Foundation, for a direc- tive for containment leak rate tests. This study resulted in a draft guide line, ECN 82-036/037. Based on practical experiences, acquired during leak rate tests on research reactors and nuclear power stations, the regulatory bodies proposed to establish a working group composed of representa- tives from the regulatory bodies, the research reactor institutes and the utilities. This working group was established in November 1980 and operates under the auspices of the Committee for Reactor Safety (CRV). The principal objective for the working group is to establish standard procedures and measurement methods for the leak tightness testing of safety containments. After a careful study the working group decided that the report ECN 82-036/037 will be a good basis for a guideline on containment leak rate measure- ments. However, modifications to this report are necessary because of the experiences acquired during contain- ment leak rate tests and the discussions in the working group. Special attention should be paid to disturbing in- fluences on leak rate measurements. These influences are: - the diffusion of gas into concrete and certain insulation materials. - changes in ambient conditions during measurements. - non homogeneous and time varying distribution of vapor pressure and temperature distribution inside the containment during the measurements. It has also been found necessary to pay attention to the following subjects: - the possibilities for quantitatively testing of penetrations. - the use of the perfect gas law compared with the "van der Waals" gas law. -3-

- deterioration of the containment structure and the penetrations during normal reactor operation. - relation between local and integrated leak rate tests including the test frequency. - verification of test procedures. - the assurance that the containment does not ex- hibit any major leak path at the beginning of an operational period by means of a so called "blundertest". - statistical treatment of measured data.

It should be mentioned that the basis for the leak rate tests in the Netherlands is the American Standard ANSI N45.4-1972. However, because of ex- periences acquired during leak rate tests, there are some deviations from ANSI 45.4-1972, and also from ANSI/ANS 56.8-1981. The objective of the working group is to produce a final directive for the containment leak rate tests in the Netherlands. If, in the future, new nuclear power installations or special laboratories will be erected in the Netherlands then this directive will be applicable. The working group has the opinion that there should be an annex to the directive giving an explanation and motivation concerning the methods and procedures prescribed. This annex should prevent discussions about interpretation of the directive. The directive will be related to ANSI/ANS 56.8-1981 so that personnel familiar with this Standard will not have to much difficulties in using this final directive. -4-

2. Relation between local and integrated test, ac- counting for deterioration

The main objective of the leak rate tests is to assure the leak tightness of containment structures including the related systems. The specified limit for the leak rate, being a fixed datum, is important for the integrated leak rate in the design phase. The integrated leak rate is the sum of the local leak rates of the pene- trations and other unknown flowpathes. This means that in the design phase each penetration or group of similar penetrations ha~e been assigned a maximum allowable leak rate. In the construction phase the first integrated leak rate test can be performed. The results of these initial tests can be compared with the specified design limits. Performing these tests implicates that the contain- ment, the penetrations and containment systems should be in a state equal to the operational condition. The results of the tests will be a reference for the tests in the future. Before commercial operation an integrated and local leak rate test at the periodical test pressure should be performed. The results of these tests should be compared with the previous tests. During the first operational period no repairs on the containment or the containment systems will be allowed as long as the leak rate will be lower then specified. Also in this operational period the local tests will be performed more frequently on those penetrations and systems which can be tested during normal operation. After the first operational period, but before refue- ling, the integrated leak rate test and the local tests should be performed. Comparing the results of these tests with the previous tests gives an indication of the behaviour of the containment and the containment systems over the first normal operation period. After some time data has been collected on the leak rate of the penetrations or group of penetrations. According to ANSI/ANS 56.8-1981 the integrated leak rate test shall be performed 3 times in 10 years. This can be translated in a sequence of 3-3-4 years. The local tests do have a test frequency depending on their construction and use in an operational period with a minimum of once a year. The schedule for the tests is shown in fig. 2.1 and is in accordance with ANSI/ANS 56.8-1981. -5-

A test schedule for the penetrations can be made effective after the first operational period, but it is possible that, after some operational periods, it is found necessary to change this test schedule or to alter the test method~ It can also be possible that a modification of a penetration is a necessity in order to improve the leak tightness. A careful recording of the results of the tests and correlating these results to the containment total leak rate, enables to identify a trend in the leak rate of the containment, accounting for the known leak pathes. It is a common assumption, when a dete- rioration tendency is not yet known, to use a figure of 10%/year. In a normal operation period a course in the con- tainment leak rate can be represented graphically as is shown in fig. 2.2.For simplicity all the contain- ment leak pathes have been divided into some groups. When a local test has been performed then a new value for the containment leak rate can be calculated. Further, when it turns out that a repair on an pene- tration is necessary, the local leak rate after the repair is used for a new calculated containment leak rate. Keeping a containment under surveillance in this way can give a reliable impression of the behaviour of the containment and an assurance that the con- tainment leak rate remains below the permissible limit. To assure that the containment leak rate is below the specified limit between 2 integrated tests, the local tests are of great importance. The local tests will give more information about the behaviour of the con- tainment and when sufficient data are available the 10%/year tendency can be adjusted. To obtain data from the local tests it will be neces- sary to design the penetrations in such a way, that quantitative measuring methods can be used. Also it is obvious that the local and integrated test results shall be comparable with regard to the test pressure and temperature. A proper administration of test results will give a complete history of the containment and can prevent expensive, unexpected repairs. I--

I -6-

3. Blundertest

After refueling, before operation, it will be good practice to assure that the containment is ready to fulfil its function. When an integrated leak rate test has been performed and this test has been successful, then the licensee can be certain that the containment is in an opera- tional condition. If no integrated leak rate test has been scheduled it will be necessary to do a quick test. Such a quick test, called the "blundertest", can be per- formed as a short integrated pressure test and it prevents that major leak pathes will be unnoticed before operation starts. -7-

4. Practical experiences with integrated leak rate tests, performed at the High Flux Reactor (H.F.R.) at Petten (Netherlands)

4.1 General

The H.F.R. is a 45 MW materials testing reactor owned by the European Community, represented by the Commission of the European Communities. The reactor is situated at the establishment of the Joint Research Centre (J.R.C.) at Petten, while the manage- ment of the reactor operation is delegated to the Netherlands Energy Research Foundation (E.C.N.). The reactor installation is contained in a domed cylinder of steel with a free volume of about 10.000 m3 (see fig. 4.1.1). The inside of the steel wall is covered with an insulation layer of 4 cm. The basement is embedded in the soil with a free volume of about 2000 m3. The maximum alowable leakage rate of the containment is 0.1 wt%/24 hrs at an internal overpressure of 0.5 bar. Integrated leakage rate tests are performed each year at low pressure (0.2 bar) and each fourth year at 0.5 bar. For the tests the absolute method is applied. The measuring equipment consists of: - 21 thermocouples for the air temperature, - 12 thermocouples for the wall temperature, 1 thermocouple for the ambient temperature, 1 absolute pressure device and 4 dewpoint temperature detectors. The 39 electrical device signals are connected to a data logger which in turn is coupled to a computer. Some characteristics of the applied computer program are: . Erroneous signals are automatically detected and replaced. The substitute signal is derived from surrounding detector signals . . The total volume is divided into six subvolumes. . The thermal expansion of the dome wall is taken into account in the calculations . . For each new scan the total time leakage rate, the point to point leakage rate and the linear regression line with standard deviation is calculated . . For the calculation of the mass of air the "van der Waals" state equation is applied. . The reference scan can be changed during the measurements. -8-

• Calibration tests can be handled. • On line plotting of mass-time and leak rate- time curves.

From earlier integrated leakage rate test results of the HFR-containment a periodic disturbance can be observed (see fig. 4.2.1). This systematic dis- turbance with a period of about 24 hours shows a ~ clear correlation with the varying ambient condi- tions (day-night cycle). Due to the presence of this more or less systematic error the total measuring time for obtaining accep- table confidence levels amounts to 3 to 4 days. Moreover a multiple of 24 hour periods is required for the total measuring time, in connection with the periodic nature of this disturbance. In order to exclude the influence of the ambient conditions on the measuring results it was decided to install a water spray system on the outside of the containment wall. This system provides the outer wall surface with a thin water layer of constant temperature. Measuring results obtained with the spray system in operation show a remarkable decrease of the periodic disturbance as compared with earlier testresults (see fig. 4.2.1). As a result of this spray system the total measuring time could be reduced to 1 to 2 days. A possible explanation for this periodic disturbance might be the diffusion of gas in the concrete struc- tures within the reactor containment. By accident a secondary application of this spray system is the temperature conditioning of the con- tainment atmosphere during warm summer days.

In general integrated leakage rate tests are perfor- med using the absolute method. This implies that the mass of air inside the containment is derived from spatial measurements with regard to temperature, total pressure and vapour pressure. Commonly the mass of air is then calculated using the Boyle-Gay-Lussac state equation:

P.V n =: (4.3.1) R.T -9-

However, this equation is valid for a perfect gas, which means that the volume of the gasmolecules as well as the binding forces between the gasmolecules are neglected. One of the state equations which accounts for the above mentioned effects is the "van der Waals" equation: 2 a.n1 (P+ 2 (v-n1 .b) V (4.3.2) R.T

2 a. n 1 The terms (2 and (n1.b) in this equation V represent the cohesian forces between the gasmole- cules and the volume of the gas respectively. The coefficients (a) and (b) are dependent on the type of gas concerned. The leakage rate which has to be determined is defined as:

dn 1 (4.3.3) L = !h . dt

The relative influence of the applied gas state equation on the final leakage rate can be expressed as:

L - L 1 D =--- (4.3.4) L L

in which (L'l)and (L) refer to "van der Waals" and "Boyle-Gay-Lussac" respectively. Substituting eq. (4.3.1), (4.3.2) and (4.3.3) into eq. (4.3.4) yields the following approximation for D : L

dP dT (A-B) - (2 A-B) P T D (4.3.5) = dP dT dV L - + P if' V

in which: 2 a.n b. n 1 1 and B = (4.3.6) A = 2 v . P V -10-

Substitution of the parameter values for a represen- tative H.F.R. case as presented in tabel 4.3.1, results in:

D = 3.5% L

If the variations in temperature and pressure during ,I dT dP the measurements amount to ~ = 0.05 and p- = 0.049 respectively, the error increases to:

D = 17.0 % (see fig. 4.3.1). L

In general it can be concluded that in the framework of leakage rate tests the "van der Waals" forces cannot be neglected. The final influence on the mea- suring results is strongly dependent on the occuring dT dP ~ and p- and furthermore on the leak rate to be measured.

4.4 ~E~E~~E~~~1_~Y~!~~E~2~_2~_~~~~~E~~S_~~E~_~~~1~~~~s calibration test

During the performance of an integrated test the leak rate can change as a consequence of a leak repair or the introduction of a calibration flow. The total collection of measuring points could be divided into two or three parts with a constant leak rate each (see fig. 4.4.1). In cornmon practice the leaka:jerate is determined for each part separately. However in this way discon- tinuities in the gas mass exist in the combined repre- sentation of the total measuring period (see fig. 4.4.1). These physical impossible discontinuities can be eliminated by the application of an integral data evaluation for the total measuring period. Figure 4.4.2 gives an illustration of this integral lineair regression model for a data collection with one change in the leakage rate. By solving the coef- ficients A, B, and C according to the cornmon proce- dures, (least squares method with three degrees of freedom), the leakage rate values for both regions can be determined. ~------

-11-

The advantages of this integrated statistical eva- luation are twofold: - the accuracy of the obtained results is better as compared with the separate treatment procedure, because of a better correspondance of the model with physical reability. - the standard deviations for ~he leaka~e rate values are smaller, since more data points are involved (total data collection is applied for determination of B, C and D) .

This method can also be used in case of a step cali- bration flow. In this case a known amount of air is released is a relative short time period. After this step release it is preferable to add a short undis- turbed measuring period. (see fig. 4.4.3). The released amount of gas can be derived from the known time interval (tc - tk) and the calculated leakage rate (C).

Commonly the leakage rate of a containment is defined as:

dm 1 (4.5.1) L = m(o) dt

In order to ascertain the mass of air as a function of time the equation of "Boyle-Gay-Lussac" is commonly applied:

m(t) (4.5.2)

Fig. 4.5.1 gives an example of the measured m-values versus time. Due to random errors in the measuring equipment and in the method, the mass values show more or less scattering. Assuming the leakage rate is constant during the measuring period, a statisti- cal average leakage rate can be derived from the mass points in time. This can be accomplished by a linear regression ana- lysis (see fig. 4.5.1).

m(t) = A + B.t (4.5.3) -12-

Using the least squares method the coefficients A and B and their corresponding standard deviations sA and sB can be solved. From eqs. (4.5.1) and (4.5.3) the leakage rate can be expressed as:

B L = (4.5.4) A while the standard deviation is:

(4.5.5)

The final measuring result can be presented as:

(4.5.6) in which the factor (K) is defined by the required confidence level and the number of mass points (see table 4.5.1). For intermediate judgement of the measuring results the values of (L) and (sL) are recalculated after each new measured mass point. A typical plot of (L) and (sL) versus time is presented in fig. 4.5.2. The decision wether to terminate the measurements or not is a.o. based on the required upper confi- dence limit (L + K.s ) and on certain stability L condi tions for (L) (see fig. 4.5.2).

The procedure described above is generally applied for integrated leakage rate tests and should give reliable results, provided that the assumptions made for this procedure are satisfied. One of the assumptions is that the leakage rate is constant in time. However, possible dis.turbances in the assumed linear mass-time relation can be caused by: - stabilization effects - diffusion of gas into concrete - external day-night influences - internal disturbances (inflow from pressurized gas systems etc.) - systematic errors in the measuring equipment or in the method. -13-

The presence of any systematic disturbance can be demonstrated by a plot of the deviations of the mass points with respect to the regression line (d.) (see fig. 4.5.3). If1no systematic effect is present a random scatte- ring pattern is found. A better way of presenting this error distribution is shown in fig. 4.5.3. In this figure the error density distribution is compared with a "Gaussian distribution". Any signi- ficant deviation indicates a systematic effect and an investigation on the possible cause should be started. Summarizing, it can be mentioned that for a proper statistical treatment of the data points, attention should be paid to the following aspects: - the data collection applied for statistical eva- luation shall not show any significant systematic deviation from the assumed linear mass-time rela- tion. - the total measuring time shall be long enough to detect any long term systematic effect. The significance of this statement can be demonstra- ted by assuming a 24 hr. periodic disturbance. (see fig. 4.5.4). From the data points in the choosen 8 hr. measuring period no clear indication of the presence of this periodic disturbance can be derived. -14-

5. Merits and shortcomings of various calibration methods

In order to demonstrate that the whole measuring set-up, including measuring equipment, computer- codes etc. yields reliable leakage rate results, a calibration test is performed. In practice three calibration methods are applied: 1. calibration by a known (measured) in- or outflow, with the same order of magnitude as the actual leakage rate (see fig. 5.1.a). 2. calibration by a stepwise release of a known (measured) amount of air (see fig. 5.1.b). 3. calibration by creating zero leakage conditions. For this method the containment atmosphere is put in equilibrium with the ambient conditions (see fig. 5.1.c).

In the following the merits and shortcomings of the above mentioned methods will be summarized:

With this method the response of the measuring system to a known change in the actual leakage rate can be determined. The advantage of this method is that the measuring procedure is cali- brated in a representative way. (small leakage rate change, long measuring period) . The disadvantages of this method are: - additional calibration time is long, with corresponding cost increase. - no check on unknown inleakages. no check on certain systematic errors in measuring equipment. - interpretation of the calibration results is complex.

With this method a relative large amount of containment air (1.% to 5%) is released in a short time (0.5 to 2.0 hr). The amount of released air is measured with an accurate integrating flowmeter. The advantages of this method are: required time is short, small cost increase. - no interference with long term systematic effects. -15-

- interpretation of the calibration results is easy • The shortcomings of this method can be summa- rized as follows: - calibration method is not representative (large change in mass of air, in a short measuring period) . - no check on unknown inleakages. no check on long term systematic errors in measuring equipment. - disturbing effect on stabilized conditions in the containment. (diffusion of gas out of concrete structures, temperature effects etc. ).

With this method the driving force for any out leakage is eliminated by creating a zero pressure difference condition. When the equi- librium condition is reached the containment is closed and the gas content as a function of time is measured along the common procedures. This zero pressure test should be performed preferably before the actual test is started, in connection with diffusion effects in concrete. The merits of this method are: - check on unknown inleakages. check on long term systematic errors in measuring equipment. - the test procedure is representative.

However, due to the unconditioned environment (ambient pressure, temperature, humidity) there are some serious drawbacks for this method in practice: - long measuring time, increasing costs. - deviations from zero pressure difference condi- tions caused by internal heat sources or changes in the ambient conditions.

Concluding it can be said that none of the three cali- bration methods is adequate for an absolute check on the measuring set-up. Only the last method (3) gives reliable results provided that the ambient conditions are very stable. (cloudy weather) .. In practice the shortcomings of the first two methods (1 and 2) can be eliminated by separate checks on the various parts of the measuring system. (calibration of devices, testing of programs, etc). -16-

6. Practical experiences with integrated leak rate test performed at the nuclear power station "GKN" Dodewaard in the Netherlands

6.1 General

The Dodewaard nuclear power station is a BWR of General Electric design, with a power output of 54 MW (electrical). The plant, built in cooperation with the Dutch industry, was the first nuclear power plant in the Netherlands. The plant (fig. 6.1.1 and 6.1.2) is situated along the river "Waal" (fig. 6.1.3). Since 1969 the power plant is in continuous operation with an accumulated availability of 84%.

The reactor has been installed in a Mark I contain- men t (fig. 6.1 .4) . This containment consists of a reactor chamber (drywell) and two pressure suppression vessels (wetwells) connected by vent pipes. The design accident is a main stearn line rupture (LOCA) • The criteria are given in the next table.

)esign Reactor Suppression Vent :::riteria chamber vessels pipes

)ressure 5 bar 3.15 bar 5 bar ,-emp. 150°C 90°C 135°C humidity 100% 100% 100% ~ree air volume 296.1 m3 2x 141.5 m3 2x 35.8 m3

3 Total free air volume is 650.9 m • The maximum allowable leak rate is 0.5 wt%/24 hrs. defined at a test pressure difference of 1 bar.

During normal operation there is no admittance to the containment. The.' personnel lock, hood flange, penetrations and isolation valves do all have the possibility to perform local tests, most of them in a quantitative way, during the refueling period.

The integrated leak rate tests are performed by N.V. KEMA in accordance with ANSI 45-4. -17-

Local tests are performed every year on all openings such as personnel lock, hood flange and manholes, after every use.

Yearly local tests are performed by means of a helium test during refueling. The maximum test pressure difference is 0.3 bar and the helium concentration is 1.0%.

The intermediate local tests are performed by means of measuring the time required for a pressure drop from 0.25 to 0 bar. The pressure is applied between the O-rings or gaskets of the openings.

During normal operation both doors of the personnel lock are closed and are part of the containment boundary.

The measuring equipment consists of: 1 absolute pressure device for the reactor chamber and pressure suppression vessels. 1 pressure device for the pressure difference between reactor chamber and personnel lock. 16 thermocouples for the air temperature inside the, containment. 3 thermocouples in the personnel lock. 2 platinum resistance thermometers for the water- temperature. 4 thermocouples for the ambient temperature. 2 dewpoint detectors.

All the electrical devices are connected to a scanner which is connected to the data logger/com- puter.

Some characteristics of the applied programme are: - the total volume is divided into 19 subvolumes. In case of failing sensors adjustment of the weight factor is possible. - thermal expansion of the containment wall is taken into account. - for each new scan the leak rate from the reference point to the actual scan point, and the linear regression with standard deviation is calculated according to the least square method. - pressure, average temperature, humidity, and leak rate values are plotted. - the reference scan can be changed during the measu- rements, and since all the measured data are col- -18-

lected on disks, this can also be done afterwards. - automatic correction of all sensors for zero drift.

During each cycle of 120 seconds aIle sensors are measured once. For the accuracy of the devices see next table.

6.2.1 Integrated leak rate test 1979.

The containment was pressurized to 1 bar overpres- sure. The previous measurements (1976) were taken at 20 minutes intervals~ In order to improve the accuracy and to reduce the measuring time the intervals were reduced to 120 seconds. After 8 hours of measuring the required confidence level was reached. However, the overall test time was 64 hours, due to unexpected disturbances. (See fig. 6.2.1.1). Forced air circulation inside the containment impro- ved the homogenity of the temperature but did disturb the pressure measurements considerably.

6.2.2 Leak rate measurement 1980.

A publication of mr. M. Engel from "Technischer Oberwachungsverein Bayern" and French sources called out attention to the phenomena of gasdiffusion into the concrete and insulation materials inside our containment.

This explained, in our opinion, a slow decrease in the measured leak rate during the test. In the past we have terminated the measurements when the leak rate was just below the maximum allowable value and stable.

According to the above mentioned publication a tem- porary pressure rise in the beginning will shorten the saturation time. Therefore the containment was pressurized to 1.2 bar. This pressure was maintained during 2 hours. While decreasing the pressure to the test pressure followed. by a stabilisation time of 2 hours we used forced circulation to obtain an homogenuous air mixture inside the containment. (See fig. 6.2.2.1).

After acceptance of the results by the inspecting authorities we continued the measurements to obtain further information concerning temperature distri- / Measuring devices

Application Device description Range Accuracy

- pressure inside containment high precision quartz 0-2 bar resolution pressure sensor - pressure difference between high precision quartz 0-2 bar 0.25% of 2 bar containment and personnel pressure sensor lock - pressure outside containmen~ anroide barometer 710-810 rnmHg :!: 0.5 rnmHg - temperature in-and outside chromel-alurnel thermo- 50°C reproducibility containment couples 0.01°C - dewpoint temperature inside hygrometer model 600 -20°C to reproducibility containment includinq diqital +30°C 0.1°C vol tayer.,eter 100 millivolt resol~tion~) l:::~ .' ~-_. - ". ,". _. ,'; .~'_.-;.'~.'~' ....I 1.0 I -20-

bution inside the containment and the gasdiffusion. The forced circulation showed: pressure measurement was influenced by the air circulation. - we had to few temperature sensors or, - the applied weighing factors for the temperature sensors were not correct. For the next measurement an adjustment of the weighing factors is foreseen. Furthermore no forced air circu- lation will be applied (see fig. 6.2.2.2).

After depressurization the containment was isolated again to demonstrate the release of gas from the concrete. During this measurement a pressure increase was found and since inleakages from pressurized gassystems inside the containment can be excluded this must be caused by the gas release from the concrete. After 5! hours the measurements were terminated. From these measurements a saturation time can be derived of approximately 8 hours. See fig. 6.2.2.3. This experiment shows that a temporary overpressure of 0.2 bar during 4 hours instead of 2 hours seems to be the optimum. -21-

6.2.3 Leak rate measurement 1983.

With the experiences of 1980 we improved our proce- dures concerning the pressurization for the leak rate measurements. As explained before the pressure of 1.2 bar was applied for 4 hours. A stabilisation time of 2 hours was applied after depressurization to test pressure. The total measuring period was 24 hours with a cal- culated leak rate of 0.21wt% t 0.09%/24 hours, which was well below the specified limits. The leak rate was stable during the last 2 hours of the measurement. (See fig. 6.2.3.1).

In advance special attention was paid to the helium local tests. The first series was tested directly in the beginning of the outage. Some valve glands needed repair, and leaking elec- trical penetrations were under high pressure in- jected with epoxy resin. The second series on the end of the outage showed a large improvement. To avoid the day and night disturbing influence the concrete lids (see fig. 6.1.4) were placed. To minimise the influence of changing humidity the down comer pipes in the pressure suppression vessels were closed for 98% with aluminium lids. However during this measurement both humidity sensors showed stepwise oscilations. Therefore the correspon- ding weight factors were set equal to zero during the measurement. The leak rate became immediately stable. Afterwards the calculation was performed with the correct weight factors. The results of both calcu- lations were exactly the same. The humidity proved to be stable. -22-

Some German and French sources have indicated an influence of the diffusion of gas into the concrete on the measurement of the leak rate. As little was known about this phenomenon an ex- tensive literature study has been performed. The results of this literature survey give the indication that a saturation time of about 2 days is necessary for the concrete. However because no experimental data were available it was decided to set up an experimental program to study the influenc~ of the variations of the physical state of the con- crete. In due time results of the experiments will be available. The experiments are aimed on obtaining data in order to get insight in: - time constant for the gasdiffusion - possibilities for influencing the saturation, time. - penetration depth into the concrete. - influence of concrete surface protection (paint, a.s.o.) . -23-

8. Conclusions

The containment, being an important safety part of the nuclear reactor installation, should be cared for with special attention. It should be clearly recognized that the containment is the last barrier to the environment. Maintaining the containment in a proper condi- tion will assure an acceptable minimum for the uncontrolled release of radioactive materials to the environment, during accident conditions. Some important aspects of the total leak tightness quarantee program are: - local leak rate tests are of the same order of importance as the integrated leak rate tests. recording and evaluating the test results in a proper way is important for developing an ade- quate test and maintenance schedule. - penetrations should be designed in such a way, that intermediate local leak tests can be per- formed in a quantitive way. - performing a simple "blundertest" before opera- tion excludes the existence of any major leak path, caused by human failure. - proper interpretation of the intermediate local test results can yield a reliable impression of the containment total leak rate tendency during normal reactor operation. - stable ambient conditions during integrated leak rate tests are essential for obtaining reliable results within acceptable measuring periods. - gasdiffusion cannot be neglected, so sufficient stabilizing time should be maintained. A tempo- rary overpressure can reduce the required sta- bilization period. - application of the perfect gas law does not always yields reliable results and the use of more rea- listic gas state equations like "van der Waals" should be considered. - an integral statistical data evaluation procedure should be applied for the actual leak rate test and subsequent verification test for obtaining reliable results. - one should be aware of the shortcomings of the various calibration methods and proper measures should be taken to compensate for these short- comings. -24-

- the presence of systematic disturbances in the assumed linear mass-time relation has a negative influence on the reliability of the final results. - forced air circulation during a leak ra~e test can have a disturbing influence on the pressure measure- ments. TESTSCHEDULE

B B IB B B B IB B B II B B III B B I I I I I I I I I I I I I I I I I I I I I I I I I I I I I I Ic Ip BpI I I Ip I I Ip I I I Ip I I Ip I I Ip I I Ip I I I I I I I I I I I I I I , , , , , t , t t YJ t • t " 1 1 r I 1 I I I 1 ~ I II II II II II II II II II II II II II II II II II " II 10 II ..

Lc p p p p l :~ JP L I I 1 1 I I I 1 1 1 1 I I t IL I J~I I I J~I I I J~I I I t : I: I I I I I I I I : I: I I I I I I I I I I - I I I I I I I I I I I I I I I I I I I I I I I I --,...... ~ ~ ~ Lo La Lo Lo ~. I I I I I I I I I I I I I I I I I I I I LYEA~S.

B = BLUNDERTES T INDICES: c = construction

I = INTEGRAL TEST 0= normal operation

L = LOCAL TEST p = periodical fig. 2.1 LEAK RATE fig. 2.2 L = LEAK RATE AT PERIODIC TEST PRESSURE

INDICES: a =plplng penetrations group 1 b =piping penet rations group 2 e =electrical penetrations I =locks t =total 0/0 u =unknown flow pathes v =ventilation t LEAK RATE LIMIT 100 I , 1

I 90 I I Intf gral tes / 80 / I . ~ ----- ~ ~ ~ V ~ 1- ...--~ ------70 ------

g 60 ~ 0:: W LI- c:> 50 =.-oJ I W .-J

~ 40

Le I 30 ----

20 Lv .. -,.- + + + + +- Lu ~------.. - 10 ------_ ... La ~ - . - - - - """.~ --- -- LI - ... /. -- - - . "'"' - -.- - ~ ~~- - - Lb ..-f-- - -- i 1 2 3 4 5 6 7 8 9 10 11 12 1 . ~ • MONTHS'" FIG 4.1.1 H.F.R. PETTEN. ISOMETRIC DRAWl NG OF THE REACTOR BUILDING. r--~r--...,--...,.~--r-----.....--"-,""'---.--v-.--_-.....r--"""i,....~_ 0.4 I I -. {Q. ): I 0.3 1 I WITHOUT SPRAY SYSEM I 0.2

0.1 " ,,' .. , " I 1 ...... , ,. " . , , 0.0 .. "'. ;.- ' '. '.. ., •. .. ,.I 0.1 , 'f. • 1 .. I , . , . 'llo ~. , I ". . t. , ~.,.

0.2 I

IJII ,0.3 I a:

< ,0.4 u. o III , , III < o 10 20 30 40 60 :E w TIME IN HOURS AFTER FIRST MEASUREMENT X ~ Z . . . w 0.4 . Z "< (b) X o 0.3 . w > WITH SPRAY SYSTEM ~ 0.2 < ...I W a: 0.1

.. " " 0.0 ...... '. ;, ',' ", ,.'" , ','. "." ,,~ " " , . , , .' .." .' .. "." .. I - 0.1 'to, " ".•..... J ."1 •

. 0.2

-0.3

-0.4 I I I I ! I J l . I I o 10 20 JO 40 50 60

FIG 4,2.1 H FR LEAKAGE RATE TEST RESULTS WITH AND WITHOUT WAT ER SPRAY SYSTEM, Table 4.3.1 H~F.R. conditions with regard to leakage rate tests.

Variable Value IHmension Remarks

I) 3 m volume R.C. 2 P 150.000 N/m abs. press.of gas "- _ .. -- .-_. --_.- T 300 Kelvin abs. temp. pf gas ------,. -~_._._._-_._--_...__ .__.- --- R 8314.4 Joule/gr.kilomol Mol.Gasconstant. ~_ .._-_. ---_ _" __ .__ .------_._------.--- _-----_.__ . dP 9.10-3 - In 24 hQurs P -- ._--_ ....._- dT In 24 hours T a -1":~~-~~5-----''-';~-m4ikilomo12. 80% N ; 20'%02 2 ------._------_ ...._. _ ..._--_ ...- --...- ...__ ...... _---- ..__ ....._- . 3 b 0.03767 m /kilomol 80% N2 + 20% 02 ---.-- ..- e------_ ....- 2 P.V n 6.01366 ]0 kilomol n - R. T --_._-_. __ .....__ .- ...._-_ ..__ .. .• __ -.J. .• ~.__ ._ ._. ._' 2 6.02034 10 kilomol Ce'l" 4.3.2 )

dV o expans~on of dome V

1) R.C. = Reactorcontainment ~ 'c%) n .02. I ! I I ' I . -j-- ---1- ~-i--i--- ii I I I I I . , I . I .. _.._- j ------.- --- ., ------", : VAn d~ .. WAA\lll : " Boyle-~Ay_luflS (,

.01 ...----.----.-,..- --- .------,.-----...-.-.-.

I

_..._.__._. ..__1~ .'~ __ 4 __

. , i L -.. - ..- - - _.~--.- ..._-- --~---_.._~------.... ! I I --+------4P . dT ( --_ .... ~m> _ t:t - ) 'P T I .00 -~--- _ ...... -

o 4 s 6 .. FIG..4.~.1 rNF'LUENCE Of "VA N 1)tR WAA\..S EFFECTS

DN J\1E RE'SULi~ OF LEA KACaE' RATE TESTS.

------._. -_._------'LOr 1 UIl.~U~.U~ 1Ut;::;"Q ~ U,"J JUtt-IUti'L3t t;m< UI~Ut VLI( ~.~ I . III l' ~

I M ( t!A~ A~ !Q)

M=A2+C(T-TK)

0 t . G . . • . • :-, . )' . ..:.' . )4-A3+D(T-TC) . . • • • ,. I I • • • ~~ :\i=A1+B(T-T1) \ ~. :D 1S,C.ONTI!l UI,. f.s ~..,

~ ~ H ~. ~ I I •

I

Ll ~K RE'PAIR ~ALIBR ATION i'LOW

... r (TIM])

T1 TK TC TN FIG.4.4.1 SEPARATE TREATMENT OF VARIOUS LEAK RATE REGIONS Wl l U<;l.'u ... <;l £u~ ~ .. U, 1<;10,) ~-£UV'I.CJg, LN\ U£ ...... L.lI ,Colt 0."

I M ( ~~)

I M=A+B(TK-Ti)+C(T-TK) • I I ~ ~ • ~ 0 • 14- A+B(TK-n)+C(TC-TK)+D(T-TC) -:-L. ~ l · _ 0 II I ! • / • " ~ M=A+B(T-Ti) •,...... /

~~ ~ o I ~ eo ~. ~ I •

I I$AK HEPAIR :;ALIBR A.TION n,OW I I

r (TIM])

Tl TK TC TN FIG.4.4.2 INTEGRAL TREATMENT OF VARIOUS LEAK RATE REGIONS \ I

m lC3ASmQSS) 1 i I I I .------,i----l~-: r-!----1------1- __ I I ,! .. I! i: -I . I i I .1. ' .. _ I I •••• ~-'r'--''- r •• "'1 • ", .• • , •• i. ; • '" • '" • , • ._---'.-;. - j " ...... 'r m.. A + B (I:kl- l:I ) + f (l:- l:~) I C,IlLlG RJ\TE t) 1 Go" ~ ~eL \t'A~E'; i -+ m. A + a (t-It.a) I Ac.rUAl. t.~A~_ 1\AT£" ._ ...... -. -- . ------..-.-- ..J- I ------~------j-- I i !

. .. . ." ,1__."..".-'. ; .. --- ... - --- ...-t-. ------j------••• ••.• - i" • ""'. 1 • .. iiiIl I... ;'."1 • . I

- - --- ...... _----

O\::r A ...n (ek- t:a) .. C. (-l:.t,-I:lc ) + ~(t. \:;~)

J\C.TUAL LEAK ~ATE":

___ ..... Iilo... t '(I:i11ft c.. ) ~ + ~ !

~a tk l:c. tN

FIG 4.4.3. l.NT£&RAL. tAlC.UlATION MODEL 'TOR 1.,l.R,T. lNC.l.UDII\!6 C.ALII3RI\TION -rEs-T. I ...... _- _..,.... -_. ~- .. __ ... _- _0"- .. i------.- ...- --_ .. -.-l.~-.- I......

'"E 2 .u () '-oJ 2!e .u - .AD .....~ au OC ! . I Lu I r (13\< t-- 1 , B Cfl -1 VI ..J" 4Ct L L I «. ..: LU .~ ~ ...J- ,..-.. •• ..u,- .1J C[ .u Cl! ~ 0 LA. (!) + J- II-- o -< V) II &AI o cr E ce :J 0 o ell I- """'VI tf) lfI II( ~ I.U E -r V7

G oft ~ -. ~ .. ::r ;.1~--- oM I .&J- "LL.- N TWO SIDED .. + 5% 2% 1% 0, 1%

I 12,71 31 ,82 63,66 636,6 2 4,303 6,965 9,925 31 ,60 3 3,182 4,541 5,841 12,92 4 2,776 3,747 4,604 8,610 5 2,571 3,365 4,032 6,869 6 2,447 3,143 3,707 5,959 7 2,365 2,998 3,499 5,408 8 2,306 2,896 3,355 5,041 9 2,262 2,821 3,250 4,781 10 2,228 2,764 3,169 4,587 11 2,201 2,718 3, 106 4,437 12 2,179 2,681 3,055 4,318 13 2,160 2,650 3,012 4,221 14 2,145 2,624 2,977 . 4,140 15 2,131 2,602 2,947 4,073 16 2,120 2,583 2,921 4,015 17 2,110 2,567 2,898 3,965 18 2, 101 2,552 2,878 3,922 19 2,093 2,539 2,861 3,883 20 2,086 2,528 2,845 3,850 21 2,080 2,518 2,831 3,819 22 2,074 2,508 2,819 3,792 23 2,069 2,500 2,807 3,.767 24 2,064 2,492 2,797 3,745 25 2,060 2,485 - 2,787 3,725 26 2,056 2,~79 2,779 3,707 27 2,052 2,473 2,771 3,690 28 2,048 2,467 2,763 3,674 29 2,045 2,462 2,756 3,659 30 2,042 2,457 2,750 3,646 40 2,021 2,423 2,704 3,551 50 2,009 2,403 2,678 3,495 60 2,000 2,390 2,660 3,460 80 1,990 2,374 2,639 3,415 100 I,984 2,365 2,626 3,389 200 I,972 2,345 2,601 3,339 500 1,965 2,334 2,586 3,310 '" 1,960 2,326 2,576 3,291

t I 2,5% 1% 0,5% 0,05% N ONE SIDED

jA6L.F ~.S".I SlUOENT'.s Te-ST. I

, j< .- ;------T .. , ~ . - 1 ...... -,..---+--.;.---....1

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TIME.

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FrG. ~.J;;.2.. TYPICAL PLOT Of' LeA1

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":£GP.ES$(C~ LI~~ FOQ" -1I ; I .tlt HOUR. - ""l'eA\oQ i . . • . . .. - . • •• 0 • • • tJ ..

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---_~ It, (~ift\c..): o .24 h...

F\Gr. itS."'. Lf'A\~~GrE 'RATE Tf~T ReSULT6 WITH r»eRlO'bic. 't)'~TlJRBA"'c.e. M (moss)

'1. Q,l,r. ... t:.AA, f \&\0\,) . ~ .-. ~.~......

I

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m (mQs~) , (b) . .... ------'. . -- -- ..-....-....------,.- .. - . ct. ... __ ." • • . 0.- .... •• ..;. • . .• . . . ~.L. r . d..lor, ... c-..l. ,--.s r ,..., ... se. . I . I

Q.LY" •. . . ----po t (tiVW' .. )

(t.. ) c..L.,. . • ...... '...... ,. . '. . . ,

V :":"ro to t"'''1 P"-& ~ s. \) yo', z. CL \: lD V\ ...... st.bili::ua.'\lh\...... ' ......

FIG.. S".I TY PI CA L ?'l..OTS OF lEA 1<'-' es.E RATE lEST ~ES U LT S FoR VA~,ouS ...._.,_.R~... ~~----"'I_~_...-s.-,-'" ": "-.-.~~

-..~.~

2'-~.

Fi~ 61.1 G.K.N. Do dew a a r d N

,.. .J-':: 1_ II 1, 1

1\

\ ,. '---.-/ o o

1817• 11ll11t .. . ' ..

lL...._ . _.' . '-'-{H:.~eBm~I : ; 17""_ i I I J f-- i i) '. -:"; -:' I---.- 0 ~ : ~ / t}'~V:.:---/ =~ O. -:- L.:".1 {. ~ , , 0" , 1" ! I _-- _ plillll. : I ~~~ I 0 j, o ,0 , I i J 000000 .IL.... -- E.. F ~ I~.

_ ..-J1fI'-- ~;; J~- : .'~ ,.,.'"

MAX. 0flLI(, 5 /IIlA. NAJt TE~ 152°C NORM. TEM.'WC

,...-..,...... ~ FIG. 6.1.4. R...... -.- """"'.p..k

DIl\.IIl flU. CM5C~JV'" WUZ....eEH I I I I I

I I ( ,! ( /

\ I / ... ,- _ ..

'" 1 I --,--.'"" \ , " '-..... - ./'

(

, • .t

FIG.613 TEMPERATURE CC

WITHOUT WITH CIRCULATION CIRCULATION

29.2 27,7

29,1 27,9

27,B 27,2

27,5 27,B

28,0 27,7

27,6 27,1

FIG. 6.2.2.2 ", FIG, 6211, .

1.0 I \ - STABILIZA ION :MEASUR ~MENT 0.5 \

o o 10 20 30 50 \ 40 60

FIG 5.2.2.1.

I10 ~ , 0:: (j) (j) W 0:: Cl. 0:: 0 W 0 10 20 30 40 SO > 60 0 ~ (j) lJJ ~ FIG. 6.2 3.1

-

10 n 1 - STAB. ME.lSUREM NT

0.5

o o 10 20 30 50 60 TIME IN HOURS ,. MEASURED INLEAK AFTER DEPRESSURI ZATION. CAUSED BY GAS DIFFUSION IN CONCRETE.

1.0

08

0~ a:

UJ ~ / Z .... / ~

TOTAL VOLlJvIE CONCRETE 26 M3

MEASURED FREE VOLUME 8 %

SURVEILLANCE AND CONTROL OF CONTAINr1ENT BY MEANS OF RADIOACTIVE MEASUREMENTS

H. ROCHE*, J.J. SEVEON*, L. ROUSSEAU*, J. DELALANDE**

CNSI Specialist meeting on Water Reactor Containment Safety

Toronto, CANADA - 17th 21st October 1983

* Commissariat a 1 'Energie Atomique - Institut de Protection et de Saret~ Nucl~aire - B.P. n° 6 - F - 92260 FONTENAY-AUX-ROSES (FRANCE) ** Electricit~ de France - Service Etudes et Projets Thermiques et Nucl~aires - Cedex n° 8 - F-92080 PARIS-LA DEFENSE (FRANCE) Meaning of french system abreviations used ~n this paper

BAN Nuclear auxiliary building BAS Safeguard auxiliary building BK : Fuel building EAS Containment spray system EBA Containment sweeping ventilation system EDE Containment annulus ventilation system EPP Containment leakoff monitoring system ETY Containment atmosphere monitoring system EVF Containment cleanup system KRT Plant radiation monitoring system RCP Reactor coolant system ReV Chemical and volume control system REN Nuclear sampling system RIS Safety injection system RPE Nuclear island vent and drain system RRI Component cooling system SVA Auxiliary steam distribution system TEG Gaseous waste treatment sys.tem TEP Boron recycle system TES Solid waste treatment system TED Liquid waste treatment system 2

I - INTRODUCTION The containment function concerns all measures in a plant intended to ensure that the radioactive substances produced are maintained 'tJithin a given space (passive containment) or that their spreading to other areas lS limited (dynamic containment). In the case of PWR's, the confinement involves setting up a succession of several barriers, the tightness and other properties of which have been designed to accomodate all the operating conditions studied. Ihe purpose of these barrlers and of the usual and safety systems with which they are associated (purification, cooling, ventilation, filtering, ... ), is to isolate hazardous materials, so as to prevent plant personnel and members of the public from radiation exposure and environmental pollution. These barriers have to assume their role under normal or accidental operating conditions. A fundamental rule of containment safety is the evaluation of the degree of containment available at a given time, that is the knowledge of radioactive transfer possibilities across the different barriers. For this purpose, the radioactive measurements are of greatest interest : on one hand, they are very sens itive measurement methods, on the other hand they allow to have a rather immediate idea of the potential hazard which has to be evaluated. Although it is customary to consider that the containment function is normally fulfilled by a succession of three barriers (cladding, primary cooling system, reactor containment) there are exceptions to the three-barrier-sequence rule. These give rise to special measures, notably as concerns devices for isolating penetration of the reactor containment designed to restore its status as an leaktight barrier. Nevertheless, there are normal or accidental operating situations in which the containment cannot completely fulfill its third barrier role; then, we generally consider the auxiliary buildings as an extension of the third barrier. In this paper, we present the radioactive measurements participating in the surveillance and control of the reactor containment as 'tIelI as the possible procedures or operating rules related to, especially the uitimate procedures which could be implemented in case of a beyond of design accident. However, we first give an overall view of the plant radiation monitoring system installed on the French plants. If necessary, differences between 900 MW and 13UO MW units will be emphasized. II CONTROL OF RADIOACTIVE TRANSF~RS IN FRENCH PWR'S BY MEANS OF RADIOACTIVE MEASUREMENTS In this section, all the radioactive measurements installed on 90G MW French units are briefly described in order to give an overall view of the plant radiation monitoring system (french system KRI )(1). 3

This system is much the same on 900 MW and 1300 MW units in spite of differences in the two unit designs (simple or double containment, nuclear auxiliary building associated with two or one units). Therefore the whole system installed on 1300 MW units will not be described, only thos~ of the radioactive measurement involved in their containment surveii1ance are detailed in section III. In French 900 ~1W PWR's, about 75 radioactive measurements (for the nuclear island including 2 units) ensure the surveillance and control of radioactive transfers as well in normal operating conditions as in accidental situation. Figure 1 shows in a schematic way thelr location and table 1 gives their main characteristics. We may distinguish the three following complementary functions involved in the radiation monitoring system - surveillance of barrier tightness, surveillance of ambiant conditions, in relation to the prevention of the plant personnel radiation exposure, - surveillance of radioactive releases, in relation to the prevention of environmental pollution and population radiation exposure. The system is briefly described according to this classification, which is somewhat arbitrary, one measurement allowing to serve more than one purpose. 11- 1 - SURVEILLANCE MEASUREMENTS OF BARRIER TIGHTNESS 11- 1 - 1- First barrier: cladding0; In normal operation, the primary coolant activity is continuously monitored on the ReV letdown line (chemical and volume control system).In accidental conditions, this system is isolated and the primary activity is then measured on the REN sample line (sample system). At last, in case of REN dewatering, the activity monitoring can still be achieved by putting a mobile measurement near the RIS (safety injection system) or EAS (containment spray system) recirculation lines. 11- 1 - 2- Second barrier: primary circuit Four measurements are devoted to the primary-secondary leakage surveillance : one CV measures the radioactivity of gases extracted from the condenser, the others G) are installed on the blowdown 1ines of each steam generator. Two other systems are controlled : RRI (component cool ing ~vater system) by@and SVA (auxiliary steam system) by@ lhe atmospheric contamination in the reactor containment is detected at any time by a counting channel including a detector for aerosol sand iodine and an other one for gasesQV. In accidental operation the dose rate within the containment is measured by two redundant counting channels with large measuring range. ~--- . ------~ 4

11- 1 - 3- Third barrier (see II - 3)

11- 2- Ambiant environmental survei I lance measurements Solid waste drumming operations are controlled by four dose rate measurements@. The radioactivity of RCV (chemical and volume control system) and TEP (boron recycle system) head-filterstD~ is overseen so as to change them before the radiological standards (personnel exposure and radioactive material transportation) are no longer respected. If an atmospheric contaminati~ of the site occurs, the control room inhabitability is controlled by QY, which measures the radioactivity of supplied air. In case of going out-af-limits condition, there is a self-actuated switching of the ventilation through iodine filters. 11- 3- Surveillance and control measurements of radio~ctive releases II 3 - 1- Limitation of liguid wastes releases Liquid wastes have not to be released in environment if thelr activity concentration, measured on the discharge canal ~, exceeds 2.10-3 Ci/m3• Beyond this value, the release valve is automatically closed. II 3 - 2- Limitation of airborne effluent releases The main surveillance system of airborne effluent releases is located at the st)~f of the nuclear island and it includes two redundant counting channels~, each of them including itself: - an aerosol and iodine activity measurement - a low - range ~-gas measurement operating during normal operation 6 (10- 1O-~ Ci/m3) - a high-range ~-gas measurement implemented in accidental conditions (10-3 - 10+2 Ci/m3) Furthermore, other measurements taking place in reactor building, fuel building and nuclear auxiliary building are intented to limit airborne effluent releases. Fuel building: Above a given threshold, the two measurements @ set up above the pool allow to stop the normal fuel building ventilation and start a restricted flowrate exhaust air system fitted with iodine filters. Reactor containment: In normal operating and out-of-l imits condition, measurement @) leads to isolation of ETY* (containment atmosphere monotoring system) and RPE (nuclear island vent and drain system) penetrations. \~hen the reactor is shutdown - automatic isolation of containme~ ~ then inhibited-, above a certain threshold the t\vO measurements Qj) Q.9) actuate EBA (containment sweeping ventilation system) isolation ; ~ is a dose rate measurement above the reactor pool, ~ is the activity measurement of exhaust air ESA ventilation system * On French P\~R's plants, ETY system assumes two functions: first, a containment sweeping one in normal operating and in certain pressure, activity conditions, secondly a hydrogen recombination function in accidental conditions. 5

Nuclear auxiliary building: Up-stream exhaust fans, t\'~o detectors @, one tor each unit, measure radioactivity in 14 ventllation ducts by scanning each of them, that are 10 ventilation ducts of the nuclear auxi1 iary building and these of the two fuel buildings and the t'tJO control buildings. As far as they are concerned with contamination, it is so possible to pOlnt out a room contamination; then a remote control in the control room allows the operator to switch the ventilation through iodine filters so as to limit iodine releases. For 1300 MW units, an additional measurement makes this operation automatic.

At last, it must be noticed that additional dose rate measurements with high measuring Aange taking place in sump rooms of ~ nuclear auxil iary building Q]} (10.f - 105 Rad/h) and fuel building Qj) (1 - 10. Rad/h) wi1 be soon installed. The paragraph II- 3 - 2 "Reinjection of high-level radioactive waste into the containment" deals with the 1300 t.1W unit corresponding measurements.

III SURVEILLANCE AND CONTROL OF CONTAINMENT BY MEANS OF RADIOACTIVE MEASUREMENTS

There is a major difference in the design of the 900 MWand 1300 MW containments : one has a simple concrete containment with an internal steel liner while the other one is made of a double concrete containment without any steel liner. An other difference between the two plants designs is the presence in a 1300 MW plant of one nuclear auxiliary building for each unit instead of one for two 900 MWunits.

It sti1 I remains that the basic design of containment surveillance by means of radioactive measurements is much the same on the two types of reactor. Therefore this section deals only with 1300 MWcontainment(2).

French 1300 ~lW units are not yet in operation, the first unit core loading (Pa1ue1 1) is to be achieved on november of this year. Previously the standard safety report has to come under examlnation by safety organisms. This is the reason why informations given in this section should be considered with some caution.

Survel11ance and control of containment by means of radiation monitoring system will be described according to the three following situations :

1- normal operating reactor 2- cold shutdown reactor 3- accidental conditions for which we sha1 I distinguish:

- radloactivity measurements involved under this sltuation ; - reinjection into the containment of high-level radioactive liquid waste in case of auxiliary building contamination; - ultimate procedure U2 : response to a containment building isolation failure; - ultimate procedure US : fi Itered venting of the containment to avoid uncontrolled, delayed, loss ot containment. 6

Figure 2 shows all the radioactive measurements, listed with their main characteristics on table II, that ensure radiological surveillance of containment and systems associated with it. The reader may compare it with figure 3 deal ing ~~ith 900 MW units and see a great 1 ikeness between the two. III - 1 - Surveillance and control of containment, normal operating reactor In normal operating conditions, a continuous air sampling is carried out by EPP (containment leakoff monitoring system) -- manual valves 18 VA open, 17 VA closed, see figure 2 - which supplies four on-line measuring devices : - one measures activity concentration of aerosols@ trapped on a paper filter, which is automatically changed either because of the signal saturation or by a deliberate actuation; its processing (derived signal) allow to know activity concentration without delay; - the same is likewise true of iodine 131 ltrapped on activated charcoal filters) activity concenJ[ation measurement~. - the two 1ast ones W@ give a redundant measure of the acti vity concentration of noble gases within the containment. Above a threshold to be defined, @ and @ measurements actuate some of containment isolation devices concerning (table III, figure 2) : - RPE lines (Nuclear island vent and draw system) - ETY system (containment monitoring system system) which involves two functions. One is the first containment sweeping used either in case of an important nitrogen leakage or to reduce noble gases activity concentration. rhe second one implemented under accidental conditions is the hydrogen recombination function which this paper does not deal with; it has not been representated on figure 2, either. III - 2 - Surveillance and control of containment, cold shutdown reactor In cold shutdown operating conditions, EBA (conta1nment sweeping ventilation system), which is an open system, operates as an air conditioning system, so as to keep an acceptable ambiant temperature with1n the containment (one a1r renewal per hour). Given that safety injection signal 1S then inhibited, there can not be self-actuated containment isolation. Therefore Electricite de France in accordance with safety authorities has set up the following procedure, that is isolation of EBA and RPE (nuclear island vent and drain systemJ ~netrations by going out-of-limits conditions (only measurements Q)~). The above mentionned measurements~ (section III-I) analyse air containment radioactivity through the EBA ventilation ducts (valve 17 VA open, valve 18 VA open, see figure 2, table III) instead of EPP ducts (conta inment leakoff monitoring system) in normal operating reactor conditions. An additionnal measurement, with two redundant monitoring devices @ CD, is devoted to the surveillance of the reactor pool so as to warn operators of any failure of this biological protection (pool draining, fuel hanling accident, clad failure) and if so to evacuate building 7

reactor and close EBA penetrations (containment sweeping ventilation system) . All these dispositions ensure a quick containment isolation in case of a contamination within the building reactor w~n the safety injection signal is inhibited (cold shutdown reactor). It must be noticed that, in these condltions, the operator may place an internal purification system into operatlon. This one, EVF (containment clean up system), is a closed ventilation system with iodine filters, the efficiency of whiCh depends on temperature and relative humidity within the containment.

III 3 Surveillance and control of containment by radioactive measurements in accidental condition 111- 3 - 1 - Radioactive measurements

~ecause of the containment isolation signal phase I, measurements~ CD@ are no longer supplied, and the operator has no more any activity concentration measurement ot containment atmosphere but two high-range dose rate measurements @ (1- 10+7 Rad/h) give him an idea of it and allow to follow every change in the course of the accident.

Moreover, the primary coolant radioactivlty may be known, thanks to an other dose rate measurement

Nevertheless, -the principle not to transport high-level radioactive wastes outside the containment has been recognized (see next paragraph), the sample llnes penetrations are consequently closed by a high dose rate signal of@measurement (see table IiI and figure 2). The threshold value has not been yet defined.

III - 3 - 2 - f

Following the Three t'liles Island accident, according to the prescription of the Safety Authorities, Electrlcite de France has engaged certain studies and assessments regarding radioactive transfers inside the plant, which have notably led to identify all the systems and rooms allowable to be contamined in accidental situation and to provide solutions in order to insure that the containment level of radioactive substances remains acceptable opposite to occupationnal exposure and releases into the environment.

In case of an accident, will be used the tollowing systems: RIS (safety reinjection system), EAS (containment spray system), ETY (containment atmosphere monotoring system) and REN (nuclear sampling system) which go out of the containment and which are then allowable to highly contaminate auxil iary buildings. The possible leakages, "normal" ones (valves stuffing box, pump packings, condensate produced by ETY (containment atmosphere monitoring system) )as well as accidental ones are 8

recovered by RPE sys tem (nuc 1ea r is1and vent and draw sys tem) wh ich is normally provlded to transport them towards waste treatment systems. However, as these latter are not designed to receive, storage and treat such highly radloactive waste, Electncite de France with Satety Authorities agreement decided to forbid transportation of these effluents towards liquid waste and airborne eftluent treatment systems and to reinject them into the containment which remains the best way to confine radioactive substances. This objectlVe aims at the non dispersion of effluents and the limitation of contamination within spaces regarding the accident. ~o as to overtake this obJective, the following provisions have been made : - ReV (chemical and volume control system) is not used in accidental conditions, consequently there is not any waste producti on and transfer towards TEP (boron recycle system) and TEG (Gaseous waste treatment system) ; - all the sumps allowable to be contaminated (process drain or tloor drain) \vhich are fitted with a high range measurement (10-"- 10. Rad/h) which delivers, first an alarm (threshold 1) and secondly an automatic isolation signal (threshOld 2) of waste transfer towards TEU (liquid waste treatment system) ; the two thresholds values are not yet defined; - reinjection of wastes into the containment is actuated from the control room by the operator who opens the remote control valves of the sumps- bui Iding reactor lines; after sump draining, isolation valves are closed again. Reinjection of wastes from safeguard auxiliary building is made with EAS recirculation lines (containment spray system). Reinjection of wastes from fuel building, nuclear auxlliary building and service building is made through a specific penetration which can also be used for REN sampling waste recycle when normal recycle lines are no more available. Figure 4 shows the reinjection of high-level radioactive waste into the containment. Dnly pumps allowable to receive such etfluents are represented on it, with their radioactlVe measurements and associated active organs. All the necessary actions i~volved during reinjection operation, automatic actions actuated by radioactive measurements and voluntary actions by the operator from the control room, have been noted on table IV. III 3 - 3- U2 procedure - Response to a containment building isolation fai 1ure In case of an accident precludlng access to the reactor building, possible defects in the penetrations and in the isolation systems outside the containment can be searched out and el iminated. They may concern electrical and mechanical penetrations, personnel air locks, equipment hatch and fuel transfer tube as \vell as isolation valves. This is the purpose of the U2 procedure, based on the flowchart in figure 5 (3). Except if contalnment failure is due to a tailure in isolation valves shutting, this objective requires the knowledge of all possible ways of radioactive releases, thus involves the nearly whole radiation monitoring 9

system. That also lmplies that available lnstrumentation allow to find out the leakage quickly so as to limit radiological consequences.

Amo ng a I 1 the passib i1it ies 0f 10 ca1isat ion 0 f suchac 0 ntam inat ion source, it was originally expected to use a mobile measuring mean to be connected on plugs of main ventilation d~cts. In fact, difficulties encountered by such a method led Electricite de France to improve radiation monitoring system by setting up a measuring device which implements an automatic scanning of each of the five main ventilation ducts (4 in the nuclear auxiliary buildlng and lin the fuel building). In short, U2 procedure, the operating rule of which is currently available for reactors in operation, involved the following steps: - verification of isolation valve shutting - surveillance ot radioactive measurements and alarms so as to detect: leaking systems, contaminated ventilation systems and rooms, leaking organs. - eliminatlon of leakage. III - 3 - 4- US procedure. Filtered venting of the containment to avoid uncontrolled loss of containment (3) (4) . Dependi ng on the assumpti ons made about the eros ion speed of the concrete and the corresponding gas production, tne pressurization of the containment atmosphere up to the design pressure requires a lapse of time between 12 hours and a tew days. The actual loss of integrity of the containment would occur after still longer intervals. They would allow the operator to implement a system which would make possible a controlled and filtered venting of the containment. The US procedure, which would make use of this system, aims at the following objectives - avoid uncontrolled loss of containment; - reduce by a significant factor the release of radioactive materials in aerosol form; - conduct the filtered gas to the stack and use accident monitoring equipment to measure the amount of radioactivity released. This system is expected to be installed on the top of fuel buildlng (for 900 t~w, on the top of the nuclear auxiliary building, Hie filter being common to two units). A programme of studies and tests is under way at CEN-Cadarache and aims at determining technical specifications for the filtration system designed. Current results show that a filtration factor of 10 can be obta ined (5). IV - CONCLUSION Originally, radiation monitoring systems installed on PWR plants were essentially designed as passive systems ot surveillance, measuring means of radioactivity levels of some of the circuits within some of the nuclear 10

island areas considered as the most representative ones for all the normal operating situations and conventional design accidents. The assessment of the Three ~liles Island accident which has been undertaken in every concerned country has led the plant operators as well as the safety authorities, among the lessons drawn from it, to consider again the role of radiation monitonng system it might have as concerns the detection, the diagnosis and the tollow-up of any abnormal operating situation even those unforeseen within the framework of design studies. In France, these considerations enhanced by both Electricite de France and Satety organisms have led the former to increase the number of radioactive measurements and to make them have an active role in different operating conditions such as normal operating, shutdown reactor, accidental sltuations including beyond of design ones. So, there will be about 75 measurements for a pair of 900 MW units against about 45 previously, and about 50 for a 1300 MW unit. Ihese modifications in the design of radiation monitoring system enhance the greater role of the containment. as the last confinement barrier of radioactive substances. This statement is expected to be well illustrated by the descriptlon of this system and espec1ally the procedures such as the containment isolation in shutdown reactor condition, the reinjection of high-level radioactive liquid waste into the containment, and the U2 one, response to a containment bui Iding isolation failure. 11

REFERENCES

(1) Rapport 'standard de sGret~ - Centrales nucl~aires du palier 900 MW. (2) Rapport standard de suret~ Centrales nucl~aires au palier 1300 MW train P4. (3) J. PELCE, C. OEVILLERS Consid~ration of severe accidents procedures and design implications ACRS/GPR Meeting, 24-25th march 1983. (4) R. CAPEL R~acteurs a eau pressurlsee 900MWe. Mesures d'urgence en cas d'accident. Proc~dure et dispositifs correspondants. Note EOF ~44-83/1330 du 7/3/83 (5) A. L'HOMME, J. PELCE French regulatory requirements concerning severe accidents in PWRs and assoclated research programme. International meeting on light-water reactor severe accident evaluation Cambridge, 29 th aug.-2nd sept. 1983 Fig. 1 - RADIOACTIVE MEASUREMENTS INSTALLED ON 900 MWe UNITS

-uV v

MAIN (ONTROl ROOM ~:~:~ --~ S.lil ---eb-- - Auw"ad .S.li blQwdQwn5 _).._JI I INTAKE (ANAL (....) I ...\....r:!~ 6ifAlfiil RAW WATER L~~~fR n.. OISCIlARGE ~~-= (ANAL r ::-:~j=~::~.-~o; 'v no;.a- ';~ ~ ,,.\ J- -.-= .'~ ~~_.~~ ~ _ .. - ._....~.~~-.-.---_.__ .-----_.- , SEWERF : IAN B I (4 TPF1' -. . I - ROOMS [SUJ ++F~ 6l0WOOWNS STEAM f L,~:'"'' .,...... GUll drlin. liENERA T OR _. I I [-. SYMBOLISM lEU HAll o IURfllNE IEr=:'jl,G ._. ~--'EP'~-_iTffi~-;Bfs-S . o RADIOAClIVf I1fASUIlfllfNIS lEAKAGES Y OROG£NAlEO ~ _,j(mg!L'@(v ~'~-j~~)~f;~~~--J ~~ ----= AIR OR VAPOR (IRCUlI IUSUALl Y lOW lfVfl AClivi .!ff~lJ~N!L ____ V.;!'d~fVAPOOAI~ mG~J~ 1Co ') ~SIORAlif:, .o~ ... _ r'" _--+). l ------PRIHARY SYSlEH ! ------~-~ J -- L__ (I) __ ••• _ ••• _ VfRY AOIVf ORCUIl IN ACCllIfNIAl CDNDlIIONS -~'. @ 8 A . NG ~~~~ USUAllY lOW lfVfl AC1IVIlY CIRCUlI NU(LEAR AUXILIARY BUllOI Fig. 2 - SURVEILLANCE OF CONTAINMENT BY MEANS OF RADIOACTIVE MEASUREMENTS (1300 MWe UNITS)

Slack fOE @

R(P RPE lEG fly Primary ettluents tank RPE lEP ------14 VP POOL \ Loop fly ::001 2 tt I ____ t:.. __ J RRA 273 VP Loop EPP 4 3VA RRA EPP Vapor phase Ipre:;:;urizer I 4 VA Liquid phue Ipre:;surizerl EBA 15 VA OVA RPE

16 VA 14 VA Proce:;:; drain:; RPE SYMBOLISM EBA 1 VA lVA floor Radioactive measurements - o drains 2 VA 4 VA fan <9 ~ l1D, Iodine filler Fig. 3 - SURVEILLANCE OF CONTAINMENT BY MEANS OF RADIOACTIVE MEASUREMENTS (900 MWe UNITS)

Stack r/ @)

R(P RPE lEG fly -- "- ... ,...... - Primary effluents 3 VY tank RPE lEP - 17 VP 18 P - (--POol--~' ..~ J • n J • n 00 :iH - - II :r I":' .:£ I 'i" 6 VA 4 VA II --T-) 102 VP I EPP RRA ~ 43 VA I I Vapor phase (pressurizer I EPP ~"" I RRA 44 VA- I liquid phase ------i4- II mvp (pressurizer I I ~"'" I 1BVP E8A '---" 114 VA 11VA .. I I IT W II RPE ~ 16VA II 15 VA I Process . (I J • LI VI" II 28 VP drains RPE SYMBOLISM lVA ~ o EBA Radioactive measurement5 floor ~ drains 4 VA 1VA @fan (f[] Iodine filter -- -- Fig. 4 - REINJECTION OF HIGH LEVEL RADIOACTIVE EFFLUENTS INTO THE REACTOR BUILDING

Disconnect able coupling t lEU REN 580 VP Rev 601 VP

1,01, VP L/l 1,02 VP :z :;:{ 1,03 VP IX CI

VI VI 1,59 VP 1,70 VP W l.J o IX r-:--Jr-- a. t 1,1,1,VP

L/l et. u.J GJCD I QJ0 111,VEl! LA 365 I LA 362 JJ-.------~ 161 BA I I TEU ~ I I I , 1,51 VB 11,53VB ~VI et. lJ.-r __- IX CI

IX o o @., -'

"'~-'et. SYMBOLISM al.- =>:z 00 CI l.J o radioactive measurements normal lines ~} rein jectlon lines FIGURE 5 D' 2------?!tOaDURE :LOW CEAR'!

d,U t:om.u: ic shut:ing n.o' I at iso lat:ion valves " Shut~ing- pO'ssible :=om 0.0' yes cout:=ol room

yes Sh.u:~ing possible :rom elect:=ical auxi 1ia.ry rool!LS . . yes Manuel shu:=i:lg 0 possib 1e

Sh.u: aan valve down sC'eam [''"

On-5~Ot: verifica:ion of valve shu::ing

'triggering of radioactivicy ala~ Isolate successively all Water ventila:ions concerned from ehe outside ~cwards reac:or building Identify valve I concerned ,y _

Leakage Locali:ed at

Force valve shut: a penecatiCtl. an isO'LaciCtl. syscem or Shut nex: valve dO'wn sC'eam !sO'la:e Force '73.1veshut ~he area or Sh.ut:'J.e:t~ vaivl; down sc=ea.m or :soiate :he a.rea TAE.LE I

PI.'R900 :-!WeUNITS - ~IOACTI\TE ~SURE~NT CHARACTERISTICS

~umber of :1easuring range 3ererence energy Radioactive ~easurements aleasuremencs

137 CJ>(primary coolanc - RCV Iluni c 0,15 - 1250 :nrad/h 't Cs 137 1 primary coolanc - REN I/unic 0, I - 105 Rad/h 'f Cs 6 3 3 85 Q) Condensor I/uni t 10- 10- Ci/m ? Kr 6 _? 3 137 Q) Steam generacor blolJdowns I/steam generacor 10- - 10 - Ci/m ~ cs 6 2 3 137 8 III 2/unit 10- - 10- Ci/m ) Cs 6 2 3 137 CD Auxiliary vapor I12 uni tS 10- - 10- Cilm '5 Cs 6 1 3 85 (]){ Concainmenc (radioactive gas) I/unic 10- - 10- Ci/m ~ Kr 7 4 60 Concainment (aerosols) 1/ unic 10- - 10- Ci ~ Co 7 IJ3 CD Concainment dose race 2/unic - 10 Rad/h Y Xe 137 room TES 1/2 unics \03 mRad/h i Cs \37 8 used resins TES 0.15 1250 mRad/h ( Cs 137 I{~"'I evaporator concentrates - TES 0,15 1250 mRad/h ;( Cs 137 I drums room TES 0,15 - 1250 mRad/h .~ C5 10 RC'l Filcer 1I unit 0,15 - 1250 mRad/h '3 i~ICS 137 i@ TEP Filcer 1I uni c 0.15 - 1250 mRad/h '6 Cs 2 85 I@ Main control room I/unic 10-3 - 10 Rad/h ~ Kr 6 -? 3 137 ;@ Liquid releases Iluni c 10- - 10 - Ci/m 5 Cs ! 3 85 (gas) Ilunic 10-6 10-1 Ci/m ~ Kr I f SO«k 7 -4 60 i@ Scack (aerosols) I I unic 10- - 10 Ci ? Co 3 3 35 ,I , Scack (gas - high range) I/unic 10- - 10"2 Cilm p Kr 3 137. \9 Spent fuel pool 2/unic 10 ::lRad/h .t Cs 3 137 f@ Reac tor pool 2/unit 5 \0 :DRadih '5 C5 6 1 3 !@ ::BA ventilacion ducc II uni c 10- - 10- Ci/~ ? 35:<:r 6 -1 3 35., :@ 'lAN, 3K, 31.'"en ci lacion ducts 14/2 unics 10- \0 Ci/~ ? "r 5 137" :@ BA.'! sumps 2/2 Imics i),J - 10 Rad/h 3 <...s 5 137 @ SK sumps 3/unic - 10 Rad/h f C3 TABLE II

PWR 1300 MWe UNITS - RADIOACTIVE MEASUREMENT CHARACTERISTICS

! Radioactive measurements Number of Measuring Reference measurements range energy (per unit)

10-2 _ 104 MPD (1) 60 0 Con tainmen t (aerosols) I ~ co (2) 0 Containment (131 iodine) I 1 - 1000 MPC 131I 6 1 85 Containement (gas) 2 10- _10- Cilm3 Kr ~ ~

-3 +? 137 0

5 137 0 BK sumps 1 0, I - 10 Rad/h ';{ cs

5 137 0 BW sumps 1 0,1 - 10 Rad/h 't Cs

5 137 0GG0 BAS sumps 4 1 - 10 Rad/h i Cs 5 137 @0 BAN sumps 2 0,1 - 10 Rad/h '6 Cs

(I) MPD Maximum permissible dose

(2) MPC Maximum permissible concentration TABLE III

Penetrations isolation by radioactive measurements

p gas measurement in atmospher containment

(0 train A 0 train B Threshold I - shutting EBA VA Threshold I - shutting EBA 3 VA EBA 2 VA EBA 4 VA EBA IS VA EBA 13 VA EBA 16 VA EBA 14 VA

Threshold 2 - shutting RPE 28 VY Threshold 2 - shutting RPE 27 VY RPE 74 VP RPE 73 VP RPE 107 VP RPE 106 VP RPE 308 VP RPE 307 VP ETY II VA ETY 12 VA ETY 161 VA ETY 162 VA

manUaltETY 22 VA manual { ETY 21 VA opening ETY 152 VA opening ETY IS! VA

RPE 192 VP

Dose rate measurements on the top of reactor pool

0 train A 0 train B Threshold - shutting EBA VA Threshold 1 - shutting : EBA 3 VA EBA 2 VA EBA 4 VA EBA IS VA EBA 13 VA EBA 16 VA EBA 14 VA

Dose rate measurement on sampling line (REN system)

Threshold 1 - shutting REN 274 VP REN 294 VP REN 314 VP TABLE IV

Necessary actions to reinject high-level radioactive wastes into the containment

, ! Automatic isolation of high-level radio- Voluntary action of active wastes towards treatment system operator in main

Sump by radioactive measurements control room !

I Radioactive ." Shutting Tr~p Opening measuremen t

BAS - LA 355 421 VP 321 PO 423 VP 0 322 PO 113 VB

BAN-A - NA 365 431 VP 431 PO 433 VP 0 432 PO 401 VP

BAS - LA 365 459 VP 141 PO 443 VP G 113 VB

BAS - LB 358 451 VP 351 PO 453 VP 0 352 PO 114 VB

BK - 151 BA 461 VP 152 PO 463 VP 0 153 PO 401 VP

BAS - LB 362 470 VP 142 PO 444 VP 0 114 VB

BW - 161 BA 189 VP 171 PO 483 VP 0 401 VP

BAN-A 8\ BA 491 VP 191 PO 493 VP I 0 192 PO 401 VP

Reinjection of nuclear sampling rate by following voluntary actions

- shutting ReV 601 VP valve

- putting a disconnectable coupling down stream REN 580 VP valve and up stream RPE 407 VP valve

- Opening RPE 401 VP valve RECENT DEVELOPMENTS CONCERNING DEGRADED CORE HYDROGEN CONTROL FOR LWR PLANTS

by W. R. Butler, C. G. Tinkler~ and R. L. Palla, Jr.

1. Introducti on The need to deal with hydrogen generation has been generally recognized since the early days of water moderated reactors. The potential sources of hydrogen during reactor accidents include: 1) metal-water reaction involving the fuel element cladding; 2) the radiolytic decomposition of the water in the reactor core and the containment sump; 3) the corrosion of certain metals in contain- ment due to the action of spray solutions; and 4) possible synergistic effects of chemical, thermal and radiation by products of accident sequences on protec- tive coatings and electric cable insulation. These potential sources were debated during licensing proceedings for many years. While the various phenomena associated with these potential sources are relatively well understood, the actual course taken during accident conditions in nuclear plants cannot be accurately predicted. Design margins are, therefore, used to deal with the existing un- certai nties.

The NRC's licensing requirements in the past have been confined to those asso- ciated with design basis accidents (DBA). These requirements are detailed in 10 CFR 50.44, with guidance provided in Regulatory Guide 1.7 (Revision 2). With regard to metal-water reactions, these DBA considerations involve metal- water reactions of no more than 5% of the fuel cladding.

The TMI-2 accident in March 1979, was an accident more severe than our licensing basis DBA's. It produced about 990 pounds of hydrogen, which is equivalent to

-1- a metal-water reaction of about 45% of the fuel cladding with water. The com- bustion of this amount of hydrogen produced a 28 psi pressure spike.

Based on the TMI-2 experience, the NRC is re-examining its overall licensing requirements for dealing with accidents that are more severe than the DBA's. New licensing requirements for degraded core accidents (DCA) and core melt accidents (CMA) are under consideration. The DCA's are those accidents that are more severe than the DBA's but are successfully arrested prior to core-melt. Research efforts to develop the data base for supporting these new requirements are underway. Associated with these efforts are: 1) the Commission's Policy Statement on Safety Goals for the Operation of Nuclear Power Plants (48 FR 10772, March 14, 1983); and 2) the Proposed Commission Policy Statement on Severe Acci- dents and Related Views on Nuclear Reactor Regulation (48 FR 16014, April 13, 1983).

The licensing requirements that might stem from the consideration of core melt sequences are too far in the future to even speculate on at this time. In this paper, I plan to focus on recent developments associated with hydrogen control during postulated degraded core accidents. The requirements for dealing with hydrogen releases during DCA's are interim requirements, pending completion of our longer term efforts for dealing with CMA's.

II. Hydrogen Control Requirements for DCA's

Several licensing requirements have already been issued and others are in the final stages of preparation for dealing with hydrogen control during postulated degraded core accidents. It should be recognized that none of the prior requirements for hydrogen control during postulated design basis accidents has been deleted.

-2- A. Small Containments

The interim requirements related to the Mark I and II containments were issued in the form of a final rule (46 FR 58484) on December 1, 1981. These requirements were intended to address several concerns over hydrogen control in a post-accident environment.

It is the staff's interpretation that the requirements of this rule would address the hydrogen control capability needed for both the degraded core accidents involving a significant fuel cladding water reaction and those accidents considered as design basis accidents with a relatively small cladding reaction as considered in 10 CFR 50.46 and modified by 10 CFR 50.44(d). The requirement to inert the smaller pressure suppression containments was instituted because of the limited ability of these designs to tolerate the range of consequences stemming from hydrogen combustion, coupled with the knowledge that the Mark I and II containments have successfully operated for a number of years with a pre-inerted containment.

A separate provision of the rule, 46 FR 58484 (December 1, 1982), requires those utilities relying upon the various methods of purging to control hydrogen to provide a recombiner capability. This provision is applicable to all containments, including the Mark I and II designs, which rely on purge/repressurization systems to prevent the combustion of hydrogen. It does not apply to those plants which already have recombiner systems and a backup purge system.

The genesis of this requirement was again the experience following the TMI accident but from a different perspective. It was clear, though the TMI accident

-3- was a degraded core accident, that even if a less severe accident were to occur, a design basis accident for example, flexibility in handling hydrogen control is a desirable option. It became obvious that even if purging meant release of rela- tively small amounts of radioactivity to the environs, under certain circumstances it is preferable to control combustible gases with recombiners. Thus it was not the fact TMI was a degraded core accident but the realization that design basis accidents may dictate the need for options other than purging.

Operating licenses have recently been issued for the first units of the LaSalle and Susquehanna plants, both with Mark II containments. Both of these units will be operating with inerted containments and with recombiners. The first Mark I plant to be licensed since the TMI accident will be Fermi, Unit 2, which will be licensed on the basis of an inerted containment and a recombiner system.

The Mark I Owners Group has filed a report to demonstrate that they need not rely on their purge systems for hydrogen control under certain prescribed conditions. The Commission's decision on this is pending.

B. Intermediate Size Containments

The interim requirements related to those ice condenser and Mark III plants for which a construction permit was issued prior to March 28, 1979, were published as a proposed rule on December 23, 1981 (46 FR 62281). The rule would require that these plants be provided with a hydrogen control system capable of handling an amount of hydrogen equivalent to that generated from a 75% fuel cladding-water reactor without loss of containment structural integrity.

-4- It has been staff's position that hydrogen control systems should serve to maintain containment integrity by preventing gross leakage from overpressurization and that hydrogen combustion, deliberate or otherwise, should not cause failure of equipment necessary to achieve and maintain a safe shutdown condition. We believe that the hydrogen control systems chosen for such a task need not be designed to the exacting standards required for conventional safety grade systems. However, while there may be no need for the designs to meet the rigorous redundancy or seismic design criteria, the known conditions of degraded core accidents must not assuredly lead to failures of the mitigation systems.

1) Ice Condenser Containment

The deliberate ignition concept has been the subject of review since TVA initially proposed such a system in mid-1980. At this point all ice condenser utility owners (TVA, Duke, AEP) and Offshore Power Systems have chosen, after investi- gating various alternative approaches, deliberate ignition as the solution to the hydrogen control issue for degraded core accidents. To date, the staff has approved the igniter systems installed in the Sequoyah and McGuire plants and, on an interim basis, those installed at the D. C. Cook nuclear plant.

The approach taken by the ice condenser utilities to establish the acceptability of the distributed ignition system was to select an accident sequence based on its significance and characteristics from the standpoint of hydrogen threat, and then to vary key aspects of the containment analysis parametrically. A small- break LOCA with failure of safety-injection, the S2D event, was chosen as the base case. Calculations of the containment pressure and temperature response to the base case scenario were performed by utilities using the CLASIX computer code and the steam and hydrogen releases calculated from the MARCH code. Staff confirm-

-5- atory analyses were performed using a hydrogen burn version of the COMPARE code developed by Los Alamos National Laboratory (LANL).

In addition to the base case, sensitivity studies were performed to assess the effects of partial operation of the containment air return fans and sprays, heat removal by ice, and hydrogen release rates. The effects of such postulated phenomena as fogging reducing the burn completeness in the upper plenum, and steam inerting the lower compartment were also analyzed.

In all cases analyzed, the peak containment pressures are well below the ultimate containment pressure capacity and are typically below the containment design pressure.

As part of its hydrogen control evaluation, the staff identified a number of technical concerns that it will be continuing to investigate as confir- matory items. The confirmatory items are:

1. flame acceleration and local detonations; 2. containment code work; 3. equipment survivability for a spectrum of accidents; and 4. combustion phenomena, including large scale effects and igniter operability in a spray environment.

The subject of flame acceleration and local detonations in confined regions of containment is currently under investigation as part of the NRC hydrogen research and technical assistance programs with Sandia. As part of this effort, tests are planned by Sandia and their contractor McGill to address such topics as the effects of steam addition and scaling on the requisite

-6- hydrogen concentration for flame acceleration. A substantial portion of this work will be conducted in the FLAME facility and the heated detonation tube at Sandia. In addition, analyses of the likelihood and consequences of local detonations will' be performed. The flame acceleration and detonation work is considered confirmatory in nature because (1) mixing of the containment atmosphere in conjunction with igniter operation at low hydrogen concentrations will preclude the formation of detonable mixtures and (2) recent analyses performed by Sandia using the CSQ code and a refined structural analysis indicate that the ice condenser containment can withstand the postulated detona- tion of a 20 volume percent hydrogen mixture in the upper plenum. The Sandia investigation should be completed by late 1983.

Confirmation of the CLASIX containment code has in large part been provided via comparison with results from the LANL-developed hydrogen burn version of the COMPARE code. Additional confirmation of the CLASIX code has been provided by means of a generic hydrogen combustion code, HECTR, currently under development at Sandia as part of the NRC hydrogen research program. The HECTR (Hydrogen Event Contai~ment Transient Response) code is comparable to CLASIX and COMPARE in terms of analytical capabilities for ice condenser plants, but will in future versions provide for more detailed treatment of mixing, transport, and flame propagation than the present codes permit~ The staff will continue to assess adequacy of the CLASIX code as part of its technical assistance program with the Los Alamos National Laboratory and its hydrogen research program with

-7- Sandia. This containment code work is considered to be confirmatory in light of the staff's findings regarding the adequacy of the CLASIX models and the reasonable agreement obtained between CLASIX, COMPARE, and HECTR.

We will also continue to investigate equipment survivability for a spectrum of degraded core accidents. This investigation will be carried out as part of the NRC Hydrogen Burn Survival Program already in place at Sandia. The favorable results of the survivability analyses for the more temperature-sensitive pieces of equipment provide the bases for classifying this item as confirmatory.

The staff will monitor the results of other ongoing NRC and EPRI hydrogen research programs to confirm (1) the margins provided by deliberate igni- tion systems (2) the absence of significant flame acceleration at large scale, and (3) the reliable operation of thermal igniters in a spray environment. Research programs to address these concerns will be per- formed at the Nevada Test Site and at Sandia. These programs are considered confirmatory because similar test programs have been com- pleted at small scale with acceptable results.

2) Mark III Containments The use of deliberate ignition systems in Mark III containments began when the Mississippi Power and Light Company (MP&L) selected a glow plug igniter system for the Grand Gulf containment in April 1981. The staff completed an interim evaluation of the system in July 1982, and found the proposed Hydrogen Ignition System to be acceptable. Similar to the approach used in licensing ice condenser plants, the staff has

-8- required MP&L continue research on the efficacy of deliberate ignition for the Grand Gulf plant. The owners of subsequent Mark III plants, the Clinton and Perry nuclear plants, have also stated their intention to use ignition systems for control of excessive hydrogen generation. In both practice and principle, Mark III igniter systems are quite similar to those employed in ice condenser plants. In as much as the Grand Gulf Nuclear Station is the lead Mark III plant, this discussion will address Mark IIIls generically by describing the issues as they were addressed for the Grand Gulf plant.

Based on Commission guidance, the NRC staff required in December 1980, that MP&L provide additional hydrogen control features to mitigate degraded core accidents consistent with general requirements set forth in the proposed interim hydrogen rule. MP&L, as part of its hydrogen control program, evaluated several hydrogen .control techniques. The various systems considered include pre-inerting, large capacity recombiners, purge & filtered vent systems, post-accident inerting, Halon injection systems, oxygen depletion and thermal igniters. The criteria for evaluation included effectiveness, operational consequences, reliability, testa- bility, availability of design and equipment, as well as cost and schedule impact. Against these criteria thermal igniters were chosen for the control of degraded core hydrogen releases.

Concurrent with the Grand Gulf licensing activities, the BWR Mark III Hydrogen Control Owner's Group (HCOG) was also formed to pool resources in response to staff requests for demonstration of hydrogen control system adequacy. The owners group indicated that igniters were the generic choice for degraded core hydrogen control.

-9- Similar to the approach taken on the ice condensers, the NRC has indicated its intent to review Mark III hydrogen control in a two-phase process, with an interim and final evaluation. The staff completed an interim evaluation of the Grand Gulf igniter system and issued a SER with favorable findings in July 1982. As part of our interim evaluation of Grand Gulf, license conditions will be imposed requiring certain additional research and development. The research effort will include the following items: (1) improved sensitivity studies, (2) investigation of combustion in the drywell, (3) investigation of combustion above the suppression pool, and; (4) tests of equipment survivability. The HCOG research program has been developed to address these technical issues.

In recent meetings with the HCOG, the NRC has been apprised that small scale tests conducted by the HCOG to investigate hydrogen burning indicate that sustained diffusion burning of hydrogen, i.e., standing flames, just above the suppression pool may occur. This situation may create a locally severe environment, which in turn could jeopardize nearby essential equipment. It is the utility contention that severe conditions are produced as a result of very conservative hydrogen releases which have served as the bases for evaluating igniter systems. In that regard, past licensing reviews conducted on ice condenser plants have considered the adequacy of igniter systems for hydrogen release rates of up to approximately 200 lbs. per minute. The BWR Owners have argued that mechanistic analysis of degraded core accidents would yield less hydrogen both in rate and total amounts. The HCOG has presented results of preliminary analysis using a recently developed BWR core heatup code. For unmitigated accidents, the calculated metal water reaction was estimated at 7 to 25% of the active cladding on a normalized basis, for recoverable accidents. Peak sustainable hydrogen releases rates are estimated to be less than 20 lbs. per minute. It is important to note that the phenomenon

-10- of a standing flame on the surface of the suppression pool has only been seen in tests with hydrogen release rates higher than approximately 25 lbs. per minute. Therefore, industry positions are developed around the principle that realistic hydrogen release rates would not create the phenomenon that yield potentially severe conditions.

The NRC has considered these matters and has concluded based on preliminary evidence that several actions should be taken. Foremost among these is the requirement for additional larger scale tests to investigate hydrogen combustion above a suppression pool. It is anticipated that larger scale tests more repre- sentative of Mark III plants may provide a more accurate assessment of hydrogen burn behavior. It is the staff's judgement that essential equipment located in the region of potentially severe conditions be considered for relocation, or have additional thermal protection provided. As an example, radiation shields may be installed to decrease the heat flux seen by certain equipment. Alternative and supplementary hydrogen control features may also deserve additional attention. Partial pre-inerting, which involves the reduction of initial oxygen inventory but at a life supporting level, may also reduce the thermal loads on equipment by limiting the amount of hydrogen which may be burned.

C. Large Dry Containments

The proposed interim hydrogen control requirements related to those plants for which a construction permit was issued prior to March 28, 1979, were published on December 23, 1981 (46 FR 62281). This rule would require owners of the affected plants to perform and submit by two years after the effective date of the rule or the date of issuance of a license authorizing operation above 5 percent of full power, whichever is later:

-11- (i) analyses to assure that during degraded core accidents, containment structural integrity will be maintained; and (ii) equipment survivability analyses to assure continued containment integrity and safe shutdown cap- ability. These postulated degraded core accidents will be assumed to produce hydrogen releases to the containment resulting from the reaction of up to and including 75 percent of fuel cladding surrounding the active fuel region with water. To date, and pending issuance of an effective rule on the matter, the staff has issued a number of operating licenses for reactors with large dry containments without requiring any hydrogen control measures or analyses for postulated degraded core accidents. Moreover, the final version of the interim rule is likely to have the above deta~led proposed requirements deleted, pending completion of the staff's longer term work dealing with severe accidents.

D. NTCP & ML Applications

A separate set of requirements was issued as a final rule on January 15, 1982, (a7 FR 2286) to address hydrogen control requirements for pending construction permit and manufacturing licensee applications. Some of these requirements go beyond those needed for dealing with DCA's. They were imposed with the anticipation that future efforts will require them because the impact of their imposition prior to construction was minimal. In this regard the principal requirements are:

(1) provisions of a dedicated penetration for possible use with a filtered- vent system;

(2) containment structural integrity is assured for internal pressures of at least 45 psig;

-12- (3) analyses of alternative hydrogen control systems are to be performed assuming 100 percent fuel clad-water reaction and the resultant containment uniform hydrogen concentration must be less than 10 percent by volume or the mixture rendered non-flammable;

(4) essential systems must be shown to survive the environments associated with the accidents considered.

To complete its reviews for the CP-stage, the staff has found acceptable applicant commitments to abide by the above requirements and to submit the results of analyses by two years following the issuance of the con- struction permit or manufacturing license.

-13- PAPER FOR CSNI SPECIALIST MEETING ON WATER REACTOR

CONTAINMENT SAFETY - TORONTO, 17-21 OCTOBER 1983

ENTITLED: THE SIGNIFICANCE OF CERTAIN PHYSICAL EFFECTS IN CONTAINMENT TECHNOLOGY INCLUDING THOSE ASSOCIATED WITH HYDROGEN

W H L PORTER

ABSTRACT

This pape~ examines some of the physical effects which may be significant in evaluating the pe~formance of a containment. Such as the possibility that when fluid is first discharged into a containment dep~essu~isation may occu~ which would delay the ope~aticn of pressu~e sensors installed to detect such an eventually. Also the significance of diffusion of hyd~ogen gas between compa~tments is examined in relation to possible othe~ mixing phenomena and an evaluation made of heat absorption of the containment walls.

The significance of the equilibrium dissociation:

H2 + 1/2 02 = H20V is evaluated together with the extent to which the solubility of hyd~ogen in wate~ may be expected to influence the hyd~ogen concentration in the containment.

This paper the~efo~e gathe~s togethe~ some selected physical phenomena and examines to what extent they are relevant in dete~mining the pe~fo~mance of the containment and in particula~ the behaviour of hyd~ogen within it.

1 INTRODUCTION

This pape~ conside~s some of the physical effects which a~e ~elevant to containment technology in a manner which permits an app~eciation to be made of thei~ significance. [t is hoped that the mate~ial p~esented will assist in the gene~al app~aisal of specific ~esults de~ived from computer codes and allow a more developed unde~standing of the importance of certain physical effects in computer modelling. [n ce~tain situations some effects a~e found to be quite negligible, but this also is important in an overall appreciation of containment technology. Not all significant effects are presented in this pape~ but it is hoped that readers will find those mentioned relevant and of inte~est to their work.

EFFECT OF HEAT TRANSFER COEFF[CIENT AT THE SURFACE OF THE CONTAINMENT WAU..S

Figure la shows for a specific case the peak pressure when a steam wat.er mixture is injected into a containment vessel for a period of 180 seconds as a function of the heat transfer coefficient between the completely mixed containment environment and the steel wall. For very low heat transfer coefficients no heat is transferred to the steel and a peak pressure of 4.1 bars is achieved when the inflow ceases. However quite modest increases in the surface heat transfer coefficient produce a significant reduction in the maximum pressure until the stage is reached, at high surface heat transfer coefficients, where the heat flow is effectively controlled by the underlying thermal resistance of the steel; the pressure under these latter conditions flattens out at a saturation value of about 1.5 bars.

Figure lb indicates that for a steel wall with high thermal conductivity, if the same total inflow is introduced to the containment over a shorter period of time, the final pressure is practically the same provided the product of the surface heat t~ansfer coefficient and the period o£ the release is kept constant and the other parameters are unaltered. This shows that the conductivity of the underlying steel is sufficiently high for the temperature rise at its surface to be insignificant compa~ed to the temperature rise of the containment contents. However the figure does indicate some departure f~om this generalisation for heat t~ansfer coefficient well up to the upper end of those to be expected in practice (> 100 kW m-2 K-l) indicating that the conductivity of the steel wall is beginning to influence the result.

The effect of discharge rate is much more apparent with concrete .the lower conductivity of which makes the overall generalisation for the steel containment untenable; the low conductivity of the underlying concrete now significantly affects the peak pressure even with low heat transfer coefficients.

2 The above generalisations become inapplicable when the volume or area or the absorbing concrete or steel is so small that the bulk material produces an inadequate heat sink. This situation can readily be checked by comparing the overall ther.mal capacity of the walls with the heat release. Quite modest increases in the heat transfer coefficient still reduce the peak pressure but the saturation pressure at high heat transfer coefficients is now affected not only by the resistance of the underlying wall material to heat transfer but also by its lack of heat capacity, which makes it relatively ineffectual as a heat sink.

THE DISSOCIATION OF WATER VAPOUR AT HIGH TEMPERATURE

The generation of hydrogen in water reactors, particularly within the core where it is produced by the oxidation of the zircaloy clad with water vapour:-

Zr + 2H20 = Zr02 + 2H2

and also by radiolysis gives rise to the possibility of a pressure rise if it burns in contact with the oxygen in the atmosphere. However this process is necessarily limited by thermodynamic requirements for equilibrium wjthin the reaction:

H2 + 1/2 02 = H20 (vapour)

This reaction produces an enthalpy change of 57,798 cals/gm mole H2 at 18oC, 1 atmosphere, or 241.5 KJ mol-l.

Where the reaction takes place in a constant volume, then the change of internal energy:

~E18oC = ~H18oC + ~n RT

where ~E = 58087 cals/gm mole H2

= 243 KJ mol-l

~n is the moles of additional gas produced in the reaction = - 0.5

R is the gas constant = 1987 cals/gm mole OK

= 8.315 J mol-l K-l

and T is the absolute temperature in OK

The following values are taken for the specific heat at constant volume:

3 Cv ca1s/gm.mo1e J.mo1-1

02 and N2 4.733 19.81

H2 5.013 20.97

H2O 6.653 27.84

In addition the following table gives published values for Kp and a the degree of dissociation of H20 at equilibrium (1).

a3 p ] 1/2 [ (1-a)2(2+a) where P is the total pressure of the reactants - bars a is the degree of dissociation of H20 to H2 and 02 Kp is the dissociation constant -

lOOa - Ine Kp

l325 15.844 0.00325

l393 14.716 0.0069

l433 14.113 0.Ol03

l474 l3.644 0.Ol4l

1531 12.753 0.0255

156l 12.325 0.0340

1783 9.806 0.182

1968 8.326 0.518

2257 6.393 1.770

2505 4.96l 4.5

2684 4.466 6.2

2731 4.033 8.2

2834 3.750 9.8

3092 3.320 13.0

4 These results are plotted on Figures 2a and 2b; and a value for 6H can be derived from the relationship

6H R

giving 6H = 58400 cals/gm mole H2 or 244.5 KJ mol-l which is in reasonable agreement with the published va~ue(2).

It should be noted in Figure lb, that the a is dependent on P which is the sum of the partial pressures of the reactants.

p : PH + Po + PH 0 (vapour) 222

This is not the same as the total pressure which will have contributions from other gases such as the nitrogen in the air. However for lower temperatures, giving rise to low values of a, the dependence on a of pressure is not marked.

To illustrate what might well happen in practice a calculation is now performed in which a stoichometric mixture of hydrogen and air at an initial temperature of l80C (2910K) is burnt.

If a fraction F of the hydrogen is burnt, the relative abundance is given by

H2 l-F

02 0.5 (l-F)

H20 F

N2 79 N2 1.881 (Air: molar ratio is --) 02 21

Using the values of specific heat and internal energy change already presented the specific heat of the products of combustion per mol of initial H2 is:-

Cv = 16.282 - 0.7265 F

Hence Cv (TOR - 291) = F 6E

5 But at TOK Kp is known from Figure 1:-

where K P

(I-F) P where ~ PH 3.381-0.SF 2

~ ~ F.P Po 0.5PH PH a 3.381-0.5F 2 . 2 2

1.881P P ~ N 3.381-0.SF 2

P ~ [3.381-0.SF] [TOK] P 3.381 . 291 0

and where P is the initial pressure. o

A trial and error solution of these equations gives for

Po = 1 atmosphere F -= 0.83 P = 10 atmospheres

and T = 33000K

It is apparent from this calculation that at the higher temperatures which can readily be achieved with the rapid combustion of near stoichometric mixtures of hydrogen, the thermodynamic considerations concerning the equilibrium condition indicate that 100% combustion is impossible although the oxidation of the hydrogen will be encouraged to approach completion as the temperature is lowered and, to less effect, if the pressure is increased. The equilibrium constant for the dissociation of water vapour is reproduced in Figure 2a, which allows the effect to be evaluated.

DIFPUSTION MIXING OF HYDROGEN

The major dispersal of hydrogen within a containment will be by thermal convection, turbulent mixing and forced convection by circulators and sprays. Molecular diffusion may be expected to be of consequence only in extremely still conditions when the other effects are minimal.

Gilliland's equation (3) gives an estimate of the mass diffusion coefficient for one gas through another:-

6 0.43 T3/2 1 D = +.L p (Vl/3 + Vl/3)2 Ml M2 1 2 V where 0 . Diffusivity - mm2 s-l

T ':; Temperature - OK

P ':; Pressure bars

MI and M2 .= molecular weights of the two gases

VI and V2 -= molecular volumes of the two gases at their boiling points

Effective values for V are as follows (3) :-

V (air) 29.9 x 103 mm3 per gm atom

V (H2) .= 3.7 x 103 mm3 per gm atom

V (steam) ':; 18.8 x 103 mm3 per gm atom

There is also a tabulated value (3) available for the diffusion of hydrogen in air at 2980K (0 .= 41.0 mm2 s-l). I

I For a range of water-air mixtures the Diffusion coefficient for hydrog~n can be taken as approximating to the value:

D ':;O.01TI.5 mm2 s-l

If a simple case of considered of a cubical compartment with sides of 10 m (volume 103 m3) possessing a passage of 10 m2 area passing through concrete wall I m thick. Then the depletion of hydrogen out of the compartment is given by:

dc DA ':; dt c VL where A is the area of the passage m2

L is the length of the passage m

V is the volume of the compartment m3

dt is the time interval - sec

and dc is the concentration difference for H2 along the passage (arbitrary units)

7 In the case unde~ conside~ation if the tempe~atu~e is 3910K, and if there is no buildup of hydrogen in the receiving compartment, the maximum diffusion rate is 8% of the compartment hydrogen content per day. This rough calculation supports the contention that except in conditions of quite exceptional stillness molecular diffusion of the hydrogen is of very secondary importance and can in general be ignored.

THE SOLUBILITY OF HYDROGEN IN WATER

The solubility of H2 in pu~e wate~ at 400C is given by (4):-

S : 1.33 PH2 gm H2/tonne of wate~

where PH . is the pa~tial p~essu~e of hyd~ogen in ba~s. It is self evi~'ent that even if la~ge quantities of wate~ with a la~ge su~face a~ea fo~ hyd~ogen abso~ption a~e available, such as might be envisaged if sp~ay whe~e used, solubility effects will not be of significance. This is especially so as the hyd~ogen will only be ~emoved by the sp~ay if it is effectively deso~bed befo~e the wate~ is ~ecycled into the sp~ay system. This argument is now . ..~einfo~ced with an example:-

At 400C the pa~tial p~ess~~e of hyd~ogen is given by:-

p -= 0.013C bar

whe~e C is the hyd~ogen concent~ation in the containment atmosphe~e in gm m-3

If the sp~ay flow ~ate is Q'Tonnes/day and the containment volume is V m3:

dc Then V dt = 1.33 P Q = 0.0173 CQ

whe~e dt is the time inte~val in days.

Taking the volume of the containment (V) as being 105 m3 and the spray flow (Q) as being (1500 Tonnes/hour) or 36000 Tonnes/day gives a fractional reduction in the hydrogen concentration of 0.6%/day.

Again this illust~ates that hyd~ogen solubility in the spray is unlikely to be of any ~eal significance as a ~emoval p~ocess unless chemical additions fo~ 'getting' the hyd~ogen a~e int~oduced.

8 THE EFFECT OF INJECTING HIGH TEMPERATURE WATER INTO A .CONTAINMENT WITH A HOT ATMOSPHERE

This effect, which has already been reported (5), is of considerable interest. The design of water cooled reactors sometimes requires that pipework be protected by concrete which forms an enclosure or hot-box containing hot air. If high pressure saturated water is injected into such a hot-box then it is possible to obtain a significant reduction in pressure as indicated in Figure 3. During the initial stages when the pressure is declining, the air, in establishing equilibrium with the flashed steam, is unable to supply enough latent heat of evaporation to the water without a substantial reduction in temperature and an overall loss in pressure.

This mechanism is .however only apparent if there is a degree of mixing; if at the opposite extreme to complete mixing the expanding steam displaces the air without mixing then the hot-box will experience a rise in pressure consistent with the gas laws for adiabatic compression .

.CONCLUSIONS

The main conclusions to be drawn from this paper for a containment in a loss of cooling accident are:-

1 During the blowdown period particularly near the breach where high surface heat transfer coefficients are encountered, materials such as steel will have a far greater effect in reducing the pressure rise in the containment than comparable areas of concrete. In the longer term when the surface heat transfer coefficients are lower concrete becomes important as a massive heat sink.

2 Significant dissociation of water into hydrogen and oxygen can occur at high temperatures ~ 10 mlo at 30000K and ~ 1 mlo at 25000K. In degraded core accidents the steam could dissociate into hydrogen and oxygen when in contact with molten fuel and this would give a continuous method of producing hydrogen which on entering a relatively coel region would only recombine on ignition.

3 Diffusion mixing of hydrogen through stagnant air or steam is so slow that it is most unlikely to be a significant method of mixing other than in conditions of extreme stagnation which might arise with a temperature inversion.

4 The possibility that saturated water at high pressure may give rise to a pressure depression when injected into hot .air is of great importance if it is intended to discover leaks with a pressure monitor. The leak could give rise to a decrease in pressure rather than an increase.

9 5 Although spray is capable of reducing local hydrogen concentrations by encouraging mixing, it will not take significant quantities of hydrogen into solution and so reduce the overall hydrogen inventory unless it is dosed with a hydrogen 'getter'.

REFERENCES

1 DORSEY, N. E. Properties of Ordinary Water Substance. Reinhold 1940. Monogram Series 81, Table 6, Page 29.

2 Large Handbook of Chemistry. 1946, Page 1545.

3 Chemical Engineers Handbook. McGraw-Hill Publishing Co Ltd. 1950, Page 538.

4 A Comprehenisve Treatise of Inorganic Chemistry. Mellor, Vol 1, Page 302, Longmans 1957.

5 PORTER, W. H. L. Some Basic Thermohydraulic Calculation Methds for the Analysis of Pressure Transients in a Multi- Compartment Total Containment Enclosing a Breach Water Reactor Circuit . •.AEEW - M 1400, May 1976.

Safety and Engineering Science Division AEE Winfrith

July 1983

10 5 70 VOLUMETRIC SPECifiC HEAT VOL UM E 6500 m ) [230,0000 J) Of STEEL WAll 3800 KJ m-3 t<-1 ORIGINAL TEMPERATURE 27°( [BOOf) SATURATED MASS INFLOW 47 Kg/see [103 Ib/see) fOR 160 SECS 60 4 ENTHALPY INflOW 127000 KJ/see [120,000 Btu/see) 2 VI AREA OF WAll 2500 m [26,910 ft2) d L .- [) TUI CKNESS OF WAll 025 m [ 0.82 tt) .0 ~ CONDUCTIVITY OF WAll 005 Kw m-1 K -1 (STEEL) 50 a.

w 3 0: w :::> 40 0: l/l :::> Vl Vl w Vl 0: w Q.. 0: 2 30 Q.. L: :::> L: ~ :::> - 20 L: x -

o 00 05 10 15 20 25 HEAT TRANSfER COEFFICIENT - Kw m-2K-1

FIG.1a. EFFECT OF WALL-SURFACE HEAT TRANSfER COEFFICIENT ON PEAK PRES SUR E CONDITIONS AS IN FIG 10 EXCEPT:- 1 CONDUCT IVITY Of CONCRETE WAL L :; 0.0014 k W m- 1<"1 t VOLUMETRIC SPECifiC HEAT OF CONCRETE WAll:; 2000 KJ ffi3 K-

5 10

______f:; 100 ....60 d

VI L. 4r d n - - f = 10 Jso : n:: I f = 1 ::J w VI n:: 3 V) ::J w V) . ~STEEl a: V) +0 n. w n:: MULTIPLY: MASS AND_. n. ENTHALPY INFLOW BY FACTOR f :£ :J :£ 2 :THE HEAT TRANSFER ~ ::J - ~ COEFFICIENT BY FACTOR t f = 100 x - f :;

o o 0.01 0.1 1.0 10.0 xt HEAT TRANSFER COEFFICIENT- Kw ni2 K-1

FIG. lb. EFfECT OF WALL-SURFACE HEAT TRANSFER COEFFICIENT ON PEAK PRESSURE REACTION: 2 Z Cl P 0 7 ~~t ~' HZ t t 02~ H2 vap Kp;: rHzOJ ;: !1""aI2(2tU) r lSi

6 K ;: DISSOCIA liON CONSTANT a. ;:DEGREE Of DISSOCIATION OF H2O P ;: SUN OF PRESSURES Of REACTANTS - bars 5 dloe Kp - 6H SLOPE ;: - R 10 -I d (t) 4 a. GIVES 6H ;: 56400 cals/gm q. ~ ~ mal H2 REACTED (lJ .-0 c:: OR 6H ;: 245KJmal-1 - 01 3~ H2 REACTED I

5 J 2

1 TEMPEHATURE oK 5000 4500 4000 3500 3000 2500 2000 1500 o 2 3 5 6 7 1 T0l( lC 10"

fiG. 20. DISSOCIATION Of WATER VAPOUR TO HYDROGEN AND OXYGEN 3000

NOTE: THESE RESUnS ARE FOR 1 BAR PRESSURE OF REACTANTS (p=11 (SEETEXTl WHER€ P = PHltP02 "'PH 0= 1 bar 2 THE TOTAL PRESSURE OF THE SYSTEM INClUDES 2500 L THE PRESSURES OF OTHER GASES SUCH AS NITROGEN

::.::: Q'

I

UJ 0:: :J ~ 2000 0:: u.l n. 1: u.l I-

1500

1300 10-5 10-4 10-3 10-2 10-1 1.0 a- fRACTIONAL DISSOCIATION OF H20 VAPouR

FIG.2b DISSOCIATION OF WATER VAPOUR TO HYDROGEN AND OXYGEN. 1.25 18

11 Kg/1000mj 100 200 300 400 500 10 o 10 20 30 40 50 WT. Of WATER PER 1000 ftl AIR - Ib

FIG.3. DEPRESSURtSATION EfFECT IN A HOT BOX PREDICTIQN OF THE TIMING OF

CONTAINMENT PRESSURE DEPENDENT SIGNALS

by

M.S. Quraishi Containment Branch Safety Analysis Department

Paper to be presented at OECD/CSNI Specialist Meeting on Water Reactor Containment Safety Sheraton Centre Hotel, Toronto, Canada 17th - 21st October 1983

Sponsored by

OECD Nuclear Energy Agency

Committee on the Safety of Nuclear Installation

Hosted by

Atomic Energy of Canada Limited - 1 -

PREDICTION OF THE TIMING OF CONTAINMENT PRESSURE DEPENDENT SIGNALS

1.0 INTRODUCTION

Many reactor designs use elevated containment pressure and temperature as potential signals for automatic recognition of abnormal s~eam/hot water discharge in the operating reactor containment. Initiation of these signals within containment is dependent on whether or not containment pressure and/or temperature reaches pre-determined set points. This paper will discuss the basis for prediction of the time at which these set points are attained.

From a safety design point of view a knowledge of whether or not the signal is initiated and if so, how long after the break the signal can be expected to be initiated is important. Pressure and temperature within containment are direct functions of the energy content of the containment atmosphere, which in turn depends on the balance between heat sources and sinks within containment. The determination of signal times necessarily involves a study of the heat source/sink balance within the containment system.

The major emphasis in containment analysis has generally been the overprediction of containment peak pressures to demonstrate containment integrity and prediction of release. OECD and CSNI standard problems remain geared toward peak pressure estimate verification. Much less emphasis is placed on the analysis of containment response (timing) for set point purposes.

From a conservative point of view, the aim of the analysis defined here is to overestimate the time at which a given pressure or temperature set point is attained. This is achieved by underpredicting the rate of increase of the containment atmosphere energy content and as a consequence the rate of increase of containment pressure. Notably this also results in an underprediction of the peak pressure. Many assumptions deemed conservative from the point of view of estimating peak pressure, must be re-evaluated for this analysis. - 2 -

The major assumptions involve the treatment of breakflow energy addition and the heat removal/addition capability of sinks and sources.

In evaluating heat sinks and sources within containment the heat removal capability of each heat sink is maximized and the heat addition capability of each heat source is minimized.

2.0 DETAILS OF AN ANALYTIC APPROACH

2.1 Spatial Discretization

Areas of some reactor containments are designed to be accessible under certain operating conditions. To reduce the radiation in these areas they are separated from each other and from the rest of containment by doors and partitions.

The application of containment analysis codes includes, implicitly, assumptions with respect to mixing of steam discharged into the containment. These assumptions originate from the nodalization of the containment volume. Perfect mixing is generally assumed in each node but different nodes can have different air-steam composition and different temperatures.

For a conservative prediction of signal initiation times an appropriate spatial discretization of the containment geometry is required. This spatial discrttization of the containment physical layout into a node-link structure should consider the location and status of the internal partitions and the doors and the location of the sensors. If there are doors or break open partitions separating the breakroom from another room not containing the signal sensors then the door must be open and the break open partitions should open at their minimum specified design opening pressure. If the door or break open partitions separate containment signal sensors from the breakroom then the door must be closed and break open partitions should start opening at their maximum specified design opening pressure. The panels composing a partition should be allowed to break sequentially until the pressure differential across the partition falls below the maximum specified design opening pressure.

Air-steam segregation is possible close to the break location. The influence of this can be assessed by modelling the break node as a small sub-node of the break room volume. - 3 - 2.2 Energy Sources and Sinks

For a typical CANDU (CANada ~uterium ~ranium) reactor containment atmosphere (Figure 1) energy sources include the break, motors, instrument air, and hot parts of the reactor. Energy sinks include local air coolers, vacuum building and spray systems. The systems which can act as energy sources or energy sinks, depending on containment environmental conditions, are metallic' and concrete surfaces, the outside atmosphere, and liquid pools.

The energy and mass transferred from the hot break discharge to the containment atmosphere is taken as the lowest possible for a given physical size of the break. This is achieved by using lower bound estimates for enthalpy and break flow rate. The break flow rate is calculated on the low side, for a given physical size of break, by considering high resistance and low pressure drop across the break locations. The break is considered to be located at a point where fluid enthalpy is lowest. The energy addition to the atmosphere from the break can be minimized by presuming the break flow does not achieve thermal and vapour pressure equilibrium with containment atmosphere. In other words, the break fluid is assumed to separate into a steam-water mixture which is at the liquid boiling point at containment pressure.

Other energy sources included in the analysis are those which are either related to major energy sinks or required to bring the reactor to a safe cold shutdown stage. For a typical CANDU reactor, they include electric motors on local air coolers and instrument air. The amount of instrument air released to containment, at any time, is taken as the minimum req~ired for safe operation.

The most important heat sinks present inside contair~ent, for this analysis, are the local air coolers. Normally, for peak pressure analysis, cooler heat removal capability is taken on the low side. In this analysis their maximum capability is assumed. The maximum number of coolers which can possibly be operating before an accident are considered. If any cooler can automatically start on a post-accident signal it is included in the analysis on receiving the first available signal. If the cooling capacity of any cooler is designed to increase on post LOCA signal, by increased coolant or air flow rate, it is increased once the signal becomes available. The cooling capability of local air coolers is also dependent on cooling water temperature. In the Canadian climate - 4 -

the cooling water temperature can potentially vary from 20°C in summer to laC _.in winter. For conservative analysis the temperature of the service water is ,- taken as its minimum value.

Local air coolers present inside the break room are capable of removing additional energy if they are enveloped in a stearnrich atmosphere. An over-estimate of their energy removal rate can be achieved by presuming they are enveloped in stearndischarged in the break node.

If the spray system is activated then it is considered operating with its maximum possible thermal utilization efficiency. The spray system water is at its lowest expected operational temperature.

If a vacuum building is activated then it is considered available at its lowest pressure and temperature.

In analyzing energy transfer to or from surfaces, the surface area and thickness used in the analysis should represent all the available surfaces, steel and concrete. wnen heat transfer is from surfaces to the containment atmosphere, the lowest possible heat transfer coefficient, representing a dry surface covered by a dry air layer, is used. When heat transfer is from the containment atmosphere to the wall, the heat transfer coefficient is the maximum possible for a wet surface covered by a boundary layer whose air-stearn composition is the same as that of the containment atmosphere.

Energy transfer from the liquid pool to the containment atmosphere can be conservatively assumed negligible.

The flow resistance of all connections between the containment and the outside atmosphere is considered to be the lowest possible.

For many reactors the initial pressure inside containment is slightly subatmospheric. In those cases, maximum subatmospheric operation is assumed to define the initial containment pressure. This gives the maximum pressure difference between initial pressure and the pressure set point. - 5 -

In some reactors, due to continuing operation of ventilation systems before containment isolates, the outside atmosphere may act as a heat and mass sink or source. As a heat and mass source (i.e., ventilation inflow) it is considered totally dry at its lowest possible temperature so the incoming air will add minimum possible energy to containment.

3.0 AN APPLICATION

We have studied a number of factors which can effect the containment behaviour including the effect of nodalization. Some of those results are presented here. The behaviour of a real containment is expected to be similar, if not the same, depending on the geometric details and other design details.

3.1 Spatial Discretization

Figure 2a shows the containment building of a hypothetical reactor. This type of model lumping the volume of individual rooms into a single big volume and assuming the pressure, temperature and composition are uniform throughout containment, is called single-node representation of containment.

To. study the influence of spatial variation of the pressure, temperature and composition, the containment may be divided into many nodes. This division, called multi-node representation, can logically be based on volume segregation due to natural barriers inside containment. For the hypothetical reactor under consideration,a three-node representation of the containment is appropriate (Figure 2b). The reactor building is divided into three areas; the Boiler Room, the Vault and the Other Rooms. Loss of coolant can occur either in the boiler room or in a vault. The vault volume is 7% of total containment volume and 11% of the boiler room, ~he boiler room volume is 63% of the total containment volume. - 6 -

To study the implications of imperfect mixing within the vault and within the boiler room and possible segregation of steam close to the break location, these nodes can be further subdivided. Figure 2c shows such a representation of our hypothetical reactor where the vault and the boiler room are subdivided into two equal volumes.

3.2 Energy Sources and Sinks

a) Initial Pressure and Temperature.

The normal inside operating conditions for the hypothetical reactor considered are a dry atmosphere at -0.25 kPa differential pressure and variable temperatures. This is an accessible containment with a ventilation system continuously operating. At design conditions the inside temperature is 36°C maximum when cooling water is at 23°C (summer conditions) and it is at 21°C when cooling water is at 4°C (winter conditions). The ventilation system operates continuously throughout the duration of the analysis.

b) Ultimate Heat Sink Temperature

The effect of winter and summer operating conditions, due to differences in cooler water temperature and initial containment temperatures, on containment pressure transient for a constant break discharge at 30 kg/s is shown in Figure 3. This analysis is based on single node representation of containment (Figure 2a). The pressure drop during the first 20 seconds is due to evaporative cooling of the dry containment atmosphere. The pressure shows a peak at a point where the heat removing capability of the heat sinks is balanced by heat addition capability of the heat sources. The maximum peak pressure reached for winter conditions is lower than that for

summer conditions. - 7 - c) Initial Relative Humidity

The effect of variation in initial relative humidity is shown in Figure 4. Containment is represented by a three-node model (Figure 2b) and contains a 50% greater condensation area than normally present. A 30 kg/s break in the boiler room is simulated during winter operating conditions. In one case the initial containment atmosphere is considered dry resulting in an initial pressure drop due to evaporative cooling. When the containment atmosphere is considered initially saturated (100% relative humidity) there is no pressure drop due to evaporative cooling. d) Pressure vs. Steam Discharge

Figure 5 shows the variation in containment pressure transient as a function of constant break discharge rates. A dry containment is represented by a three-node model (Figure 2b). The break is assumed to occur in a vault during summer operation. The initial drop in pressure at time 0 from 101 kPa to almost 100.6 kPa is due to evaporative cooling in the vault. The next two lowest values are due to evaporative cooling in the boiler room and other rooms. The peak pressures are predicted to occur at about 270 s and the subsequent pressure decline is attributed to increasing moisture content (and subsequently enhanced cooler capability) in the containment atmosphere.

If peak pressures are plotted as a function of break flow rate (or break cross section area), the trend can be represented by a straight line as shown in Figure 6. Figure 6 also shows the variation of peak pressures due to nodalization differences (single node, three node), variations in initial conditions (summar operation, winter operation) and location of the break (boiler room, vault). The peak pressures are lower during winter than in summer. Similar peak pressures occur at a later time and at high break flow rate when the break node is smaller, because the smaller break node leads to a high steam fraction, which enhances cooler capability. - 8 -

e) Imperfect Mixing in Break Node

Further subdividing the nodes from three (Figure 2b) to five (Figure 2c)

has little effect for breaks in the vault. This is shown in Figure 7

for winter operating conditions. The insignificant effect of

dividing the vault into two volumes is a result of the already very small volume of vault. During the early stage of the transients the vault

air was purged by stearn and consideration of imperfect mixing (via

further nodalization) in the vault had little effect. The peak

pressures in the boiler room did show a decrease in value and an increase in time of peak pressure as a result of further subdivision

of that node.

f) Passive Heat Sink

The case studies also include a 50\ increase in condensation area

by adding additional steel structures not normally present in our hypothetical reactor. A comparison (Figure 8) shows that the effect of the

additional condensation surface area. For the time span of the analysL-

the effect is similar to the additional cooler capability but in the long

run their effect is expected to decrease gradually as their temperature

increases and they no longer absorb energy from the atmosphere.

g) Source And Sinks Energy Balance

Figure 9 shows the containment energy balance at various times for

one case analyzed. The share of the total energy in containment is

constantly increasing for coolers and decreasing for containment

atmosphere and condensate. - 9 -

4.0 SUMMARY AND CONCLUSIONS

The above mentioned theoretical considerations result in an analysis which slightly overestimates the energy removal capabilities of major energy sinks and slightly underestimates the energy added by major energy sources. This, in turn, results in keeping the steam and energy content of the containment atmosphere at a minimum. The temperature in containment follows the trend set by the variation of the energy content of the containment atmosphere. The pressure variation inside containment is directly proportional to the product of absolute temperature and number of moles of air-steam present. Thus, this approach to analysis generates an upperbound estimate of signal initiation times based on temperature or pressure set points.

Quantitative results for a limited number of cases show the effect of the variation of these factors on the response of a hypothetical reactor. For these analyses the effect of blowout panels, which might be installed between rooms, is not analyzed. Analysis or the summer and the winter operating conditions shows little effect on the containment response due to the lower winter operating temperatures.

Pressure rises to a maximum value for the small discharges considered but is brought down by the coolers due to increasing cooler capability which is due to increased moisture content of the air as a result of continued operation of the ventilation system.

Pre-saturation of the containment atmosphere resulted in slightly lower peak pressures.

The analytical detail considered which affected the results most was the nodalization. Decreasing the break node volume results in less energy transfer from the break and much improved cooler capability in the break node. Both are a consequence of high temperatures and moisture content in that small node. - 10 -

FIGURE 1 HEAT SOURCES AND SINKS FOR THE CONTAINMENT ATMOSPHERE 8/2 ....----.... /' ..... , (B) , \ I I BOILER I I ROOM I 8/2 I 33,000 m3 I I REACTOR I I I I BUILDING I I 52,000 mJ ..... (V) I ..... VAULT V I V 3,500 " I " 2 I 2 m3

(0) OTHER ROOMS 15,500 m3 o

./ %

(a) (b) (c)

FIGURE 2: CONTAINMENT NODALIZATION REPRESENTATION - 12

105 ,,-....---...., __ ~D •••• O ... " ...... " ... -- ~ 104 _.-../ , ...... ~, ~--~-- ~(j41 .-I ~~,o • I a..',.., 'T :. a., ~- 103 : I ..aa.7/~, , :/ ,~~ ' •. / "a~ :.. / ...• 102 : I ...... e. as ! I no ; I ~ ~ 101 w ; I a: :::l : I (I) -(1)100 : I w a: i I no : I : I 99 ...!,I ,.

98 V ' '. :, ,' SUMMER WINTER ,' INITIAL TEMP 36 2'1' 97 "V COOLING WATER TEMP 23 ,4

I 96 o 100 200 300 400 500 TIME, S

SINGLE NODE CONTAINMENT REPRESENTATION. iNITIAL PRESSURE.~O.25 kPag, RELATIVE HUMIDITY 0% CONSTANT BREAK DISCHARGE 30 kg/so

FIGURE 3 EFFECT OF SUMMER AND WINTER OPERATION ON

CONTAINMENT PRESSURE TRANSIENT - 13 -

104

103

w .~a: '(/) ,(/) Ia:'w ~Q. 101

)'.,,' -. 99 o 100 200 300 400 500 TIME,s

THREE NODE CONTAINMENT REPRESENTATION. INITIAL PRESSURE -0.25 kPag, TEMPERATURE 21°C. CONSTANT BREAK DISCHARGE IN BOILER ROOM 30 kg/so 50% ADDITIONAL CONDENSATION, SURFACE AREA

FIGURE 4 EFFECT OF INITIAL RELATIVE HUMIDITY ON CONTAIm'IENT PRESsURE-TRANSIENT-: - 14 -

,..-- !,

105

104 CONSTANT BREAK DISCHARGE RATE 50 kg/s

as 103 ~ ~ W a: ;:, en ...... \", en ...... ~ &&.I •••• 40 kg Is . a: ..- . ~ 102 .••.... -....~.~.... •• •• •• •, .. :.-- 101 • • ;. • " .~. • , '. •• :; :" r •• I •• ,. •• '\ • -•• 1 100 o 100 200 300 400 500 TIME, S ._-~--I ..-,. THREE NODE CONTAINMENT REPRESEt'ST.ATION. INITIAL PRESSURE -0.25 kPag, TEMPERATURE 36°C, INITIALLY DRY ATMOSPHERE, BREAK IN VAULT

"i't- .•.

FIGU~ 5 EFFECT OF VARIOUS BREAK DISCHARGE RATES ON CONTAINMENT

PRESSURE TRANSIENT - 15 -

108

CONTAINMENT INITIALLY DRY AT -0.25 kPag

107

• 106 ~ ~ W a:: ~ (f) (f) w g: 105 lll:. ct. W' ~.

, 104

-!

, ,-_.~-- .

102 20 "'! •• 3tt, 40 ,', ~, 60 70

<. . CONSTANT BR~K D~$Ct-fARGE. kg/s

CONTAINMENT SYMBOL BREAK INITIAL PEAK PRESSURE REPRESENTATION LOCATION CONDITIONS REACHED AT

• SINGLE NODE SUMMER 220 5 0 .. WINTER 160 s THREE NODE BOILER SUMMER 250 s • .. .. , ~.; 0 " BOILER' WINTER 150 s" ... .. VAULT SUMMER 2705 A .. VAULT WINTER 2105

FIGURE 6 EFFECT OF CONTAINMENT NODALIZATION, BREAK LOCATION, DISCHARGE RATE AND INITIAL CONDITIONS ON PEAK PRESSURES - 16 -

105 ,'''- ..~".- '~"""-'-""~'---' - .... -_._.

I I, Cll. a. ~ 104

102 :__ ....2Jt._~~...~~....._.. 30...... ;._ ..__ 40.... __ ._..~'_. .so...... _._~ .. __ 60__ ._.l._..l0-_ ... 80

'..L -'. ;:~; CONSTANT'BREAK DISCHA'RGE, kg/sjr, c. >..:.. ,?"Pt<."'\3iC :1'-1;:1: ";.~f.l"':..'~;'"

-"-CONTAINM"ENT:""1JREAK'" '-INITIAI---'-- ""PU1CP'ffisURE ,J)!h~B9:L .~EPRE~EN~~!~.N:r~, LOCATION. ~O~!q.f.rIONS ~~~ao AT -_ ...~ .-.- --THREENOOE-80llER"---' W1NTEfr._ ..... ---130'5' ~ FIVE NODE " BOILER ~':;'W1NTER .:. 260 s o THREE':""OoE' ~...' VAULT .' WINTER 205 s

_i~';.;(' ~':FIVE NODE' ""'." VAULT 1l:'W1NTER 225 s

....- . I "". ~ '

FIGURE 7 .EFFEC.TOFCON'XAINMENT NODALIZATION AS A FUNCTION .. , ,,' - . - ... OF BREAK NODE VOLUME - 17 -

_._-_ ...... \ 'r1"" .... - ._-._--. i.::j . 105 o

i ; t ',t: Cll / __ ~J.'-";~ a. \ ~ 104 f W- \. e: ;:) UJ UJ w e: Y a. -. .' . ~ i w< 103 If' ~;... a.

'.(" 102

20 30 .2 '1 t,: ~ S ~'.~ "i;"':: ~.: ;\1; ..3.:. ~~. . ..-t t":' -: ':: 60 70 80 . CONSTANT BREAK DISCHARGE, kg/s

....~+A@~(~d~~=CO(~1~1~~'~~.tiijII'I~~~Y~I~~X'Ar.~-O.2~ kPag, _____ ... ~ .'_ '.... _._~'-INT;RJ:..9NDlTJqNS._.__ _.. __

::l;.;::~,;53;;':;{ }:;;;,;SIt. .J, .... , :".. n :..J.... ,.: •• ~::: •••• ~\ •• i.: ,-

- '.Jr;CSYMBO[ ;_.>;C.,iJ~~~ ~'t.). ,"'::-.! CdNDENSATION SURFACE)\R'~ ___ . _ =..t.OCATION.. .. -..--- ._. . _ --- ..- ..-.

? $&1 • : i/~" ..\. ...' ..>C{ .::.. r i' r ~ :.::" BR.~LER . 100% Cq~~RET.E", 0 . - . VAULT 100% CONJ:RE;TE :- .. :,'1 \ q"" • • .... '. ~- . , , ~,:., .A ~. ~~,.~qILER I;.:"'~: 100% C()~CRerE + 50% S~EL • ... .. -'~ VAUU .._ .... _-_..--~~CONCRETIi-+ -59-%-~ER:-'

FIGURE 8 EFFECT OF ADDITIONAL CONDENS~1'ioN'STiR.FACEAREA? - 18 -

103

~• de 102 W ~ ::> B (I) (I) UJ I: ~

101

THREE NODE CONTAINMENT REPRESENTATION' 50kg/s CONSTANT BREAK DISCHARGE IN VAULT DRY CONTAINMENT AT -0.25 kPag DURING SUMMER

100 0 200 . 400 600 . 800 1000

TIME, S

ENERGY BALANCE (°10 OF TOTAL)

TIME, SECON OS 1.2 (A) 20 (B) 81 (C) 243(0) 913 (E)

STEAM FROM BREAK 97.7 97.7 97.7 97.7 97.7'

ADDITIONAL HEAT SOURCES 2.3 2.3 2.3 2.3 2.3

TOTAL ENERGY SOURCES 100.0 100.0 100.0 100.0 100.0

ENERGY TO WALLS -0.1 0.4 2.2 3.8 4.8

ENERGY TO OUTSIDE -0.9 2.9 0.0 5.2 8.3

ENERGY TO COOLERS 3.0 7.3 15.9 26.5. 43.4

ENERGY TO CONTAINMENT 98.0 89.4 81.9 64.5 43.5 ATMOSPHERE AND CONDENSATE

TOTAL ENERGY SINKS 100.0 100.0 100.0 100.0 100.0

FIGURE 9 CONTAINMENT ENERGY BALANCE AT VARIOUS POST LOSS OF

COOLANT ACCIDENT TIMES