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Evaluation of Facility Containment to Determine Limiting Internal Uniform Pressure Capacity. Prepared for American Electric Power Co
	ML17326A883
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Site:	Cook
Issue date:	03/16/1981
From:	; STRUCTURAL MECHANICS ASSOCIATES
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ML17326A884	List: ML17326A883; ML17331A707; ML17331A708;
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	SMA-80C129-1, NUDOCS 8104290520
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{{#Wiki_filter:SMA. 8OC129-1 SECTION 1 EVALUATION OF D.C. COOK CONTAINMEiNT TO OETERMINE LIMITING INT""RNiAL UNIFORM PRESSURE CAPACITY Prepared for: american Electric Power Co. 2 Broadway New York, N.Y. 10004 16 March 1981 Prepared by: Structural Mechanics Associates 3645 'i(arrensville Center Rd. Cleveland, Ohio 44122 (216) 991-8841

TABLE OF COilTEl<TS 1.0 Introduction

1.1 Purpose and Scope

of Report 1.2 Evaluation Criteria 1.3 Containment Description and Design Basis 1.4 Material Properties 2.0 Identification of Limiting Failure i~iodes Associated with Uniform Static Internal Pressure 3.0 Potential Failure Mode Analysis 3.1 Shear Failure in Base Hat 3.2 blembrane Hoop Tension Failure of Concrete Cylinder 3.3 Pressure Capacity of Equipment Hatch Closure 3.3.1 Splice Plate Homent Capacity 3.3.1.1 Hand Calculation 3.3.1.2 Finite Element Analysis 3.3.2 Equipment Hatch Cover Plate Pressure Capacity 3.4 Pressure Capacity of Personnel Hatch 3.4.1 End Plate Closure 3.4.2 Door 4.0 Summary and Conclusions

1. 0 INTRODUCTION PURPOSE AND SCOPE OF REPORT-The object of this report is to determine a best estimate of the limiting uniform equivalent static internal pressure capacity of the containment structures for the O.C. Cook Nuclear Generating Units No. 1 and 2. The evaluation reported is limited to the reinforced concrete base mat, the reinforced concrete right circular cylinder and hemispherical dome as well as major containment penetrations including the equipment and personnel hatches. This report completes Phase I of a three phase effort which will include as Phase II an upper and lower bound estimate of the internal uniform equivalent static pressure capacity of the "as built" containment structures and Phase III which will consider potential time dependent localized non-uniform pressure load effects.

1.2 EVALUATION CRITERIA The evaluation to determining the limiting best estimate uniform static pressure capacity of the containment structures is based on a linear elastic analysis of critical por tions of the structure up to stress levels limited by "as built" mean value samples of the yield in steel and ultimate strength of the concrete. It should be understood that the structural and leak tight integrity of a steel lined concrete containment shell and slab structure should be maintained well beyond actual yield of the steel reinforcement. This is due primarily to the relative high ductility of the steel liner, (ie. 20-23 percent uniform ultimate strain at rupture ) compared to .the 40 ksi steel reinforcement (ie. 8-11 percent uniform ultimate strain at rupture) and strain hardening in the reinforcement. Hence the liner in general ~ould be able to accomodate relatively large nonlinear deformation of the concrete structure before significant leakage would occur. However, since deformations beyond the yield range are difficult to predict and localiz d deformations in the structure can significantly exceed those calculated globally, the limiting internal pressure has been determined conservatively considering only assumed elastic response up to the initial yield of the materials used. It is the author's opinion based on the observed results of model tests that significant leakage (> 1.0 p rcent of containment volume) of the containment would not occur until oressures exceeded the limiting oressures calculated in this study by at least 20 percent. Assuming a composite coefficient of variation of 10 percent and a log normal distribution of material properties the probability of significant leakage at the pressure level defined in this study would be approximately 0.01. The probability of significant leakage and upper and lower bounds on pressure will be evaluated in more detail in Phase II.

CONTAIililENT OESCRIPTIOH Atl0 DESIGN BASIS The reactor containment structure is a reinforced concrete vertical cylinder with a flat base and a hemispherical dome as shown in Ficure 1: A ductile welded steel liner with a thickness of 1/4-inch on the containment base and. 3/8-inch'on the Cylinder is stud attached to the inside face of the concrete shell to insure a high degree of leak-tightness. The design objective of the containment structure is to contain all radioactive material which might be released from the reactor coolant system following a postulated loss of coolant accident. The structure serves as both a biological shield and a pressure container. The structure consists of side walls measuring 113-feet from the liner on the base to the springline of the dome, and has an inside diameter of 115-feet. The side wall thickness of the cylinder at tne base is 4.5 ft. tapering to 3 .5 ft., seven feet above the base mat and continuing at 3.5 ft. to the springline. The reinforced concrete thickness of the dome varies uniformly from 3.5 ft. at the springline to 2.5 ft at the apex of the dome. The inside radius of the dome is equal to the inside radius of the cylinder. The flat concrete base mat is 10-ft. thick with an outside diameter, of 140'-0" and with the bottom liner plate 1/4" thick located on top of this mat. The botto~ liner plate is covered with a 2-ft. structural slab of concrete which serves to carry internal equipment loads and forms the floor of the containment. The base mat is supported directly by relatively stiff soil. The basic structural elements considered in the design of the containment structure is the base slab, side walls and dome acting as one structure under all loading conditions. The liner is anchored to the concrete shell walls by means of stud anchors so that it forms an integral part of the entire composite structure under all membrane loadings. The reinforcing in the structure has an elastic response to all primary design loads. The base mat is 10'-0" thick and 140'-0" in diameter. The reinforcement in the top of the base slab consists primarily of one layer of 818S bars at 12" c/c in the hoop and 2 layers of 58S bars at 9" c/c in 'the radial directions. The bottom reinforcement consists of 2 layers of 818S bars at 12" c/c in the hoop and 3 layers of alternate 818S and 811 bars at 9" c/c in the radial directions. The base slab was poured in two five foot lifts which are tied together in order to transmit horizontal shear induced by bending moments by shear keys and vertical All bars at 6'-0" c/c spacing.

The membrane hoop (horizontal) reinforcement in the cyclinder walls is generally in two rows, one on each face consisting of N18S at 18" c/c circumferentially extending to 20'bove the base mat reduced to 9" c/c spacing between 20'nd 57'bove the base and then increased to 12" c/c spacing between 57'nd above the base mat. 113'springline) The membrane meridional, (vertical) reinforcement in the containment shell consists primarily of two layers one on each face of 818S bars on 18" c/c. In the region of discontinuity base mat the amount of vertical reinforcement is doubled to 4 at'he layers of ~18 bars at 1S" c/c and at the cylinder dome intersection one vertical staggerd row of 411 bars at 18" c/c is added to the existing membrane vertical reinforcement to provide discontinuity bending moment resistance. In addition to the vertical and horizontal membrane and bending reinforcing steel, in plane diagonal reinforcement has been provided to carry seismically induced membrane shear. The 45 degree diagonal bars consist of 811 bars spaced 3'-0" on the horizontal c/c placed in two rows in each face and in each direction. The diagonal reinforcement is embedded in and extends from the base mat to 4-3" above the springline into the dome for alternate bars and 7'-0" for the rest of the diagonals. The dome reinforcement consists of 818 bars at 18" c/c in each face in each direction. The containment structure encloses an ice condenser containment system which is designed to limit pressurization o the containment under design basis accident conditions to 12.0 psi. Other significant design load parameters are the equivalent safe shutdown earthquake loading of 0.2g zero period ground acceleration and a design basis tornado of 360 mph. wind and and 3.0 psi differential pressure. Hot process pipe penetrating the containment are anchored in the containment shell with the anchors designed to resist the postulated rupture of the process line without loss of containment leak tight integrity. A load factor of 1.5 (additional safety factor) is used with the internal pressure component of design load while a load factor of 1.0 is used with both the SSE and Tornado loading. !)ATERIAL PROPERTIES The particular specified minimum materials properties used in the construction of the containments are summarized as follows: (a) concrete fc = 3,500 psi at 28 day (ACI 308-63, 301,66, and 214-65) (b) reinforcing rod - fy = 40,000 psi (ASTN A 15)

(c) liner plate =- fy = 32,000 psi; fu = 60,000 psi (ASTtl SA 442-Gr.60) (d) equipment hatch fy = 38,000 psi; fu.= 70,000 psi (ASTH SA 516-Gr.70) (e) personnel hatch fy = 38,000 psi; fu = 70,000 psi (ASTt< SA 516-Gr.60) (f) hatch bolts - fy = 105,000 psi; fu = 125,000 psi (ASTH SA 193-6r.87) In Table 1 can be found a summary of the "as built" strengths as well as a measure of the dispersior. associated with the materials used in the containment construction, based on a limited sample of existing test record data. As part of Phase II of this evaluation a more detailed evaluation of "as built" material property data will be developed. 2.0 IDENTIFICATION OF LItiITING FAILURE NODES ASSOCIATED MITH UNIFORi~l TATIC N RNAL P E U In selecting the potential limiting failure modes associated with equivalent static uniform internal over pressurization of a PPR reinforced concrete ice containment a number of existing analyses have been reviewed'. These include the following references: Harstead, G.A. "D.C. Cook Nuclear Power Plant, American Electric Power, Estimate of Ultimate Pressure Capacity of Containment Structure", Harstead Engineering Associates, Report prepared for the NRC Staff, September, 1980. (See Attachment A) (2) Von Riesemann, M.A. et.al. "Structural Response of Indian Point 2 and 3 Containment Oui ldings" Summary of Draft Report. results presented to !IRC Staff, Technology-Exchange t~eeting 5, 17 June 1980. (3) United Engineers and Constructors "Evaluation of Capability of Indian Point Containment Vessels Units 2 and 3" presented to NRC Staff, Technology Exchange ileeting 5, 17 June 1980. (4) American Electric Power Service Corp., "D.C. Containment Design Calculations, AEP, 1969. (5) S. 3arnes et.al. Indian Point !!uclear Generatin Unit No. 2 Containment Desi n Repor~, '!estinghouse Nuc ear Energy ystems, United Engineers and Constructors, !larch, 1969. (6) Shulman, J. "Analysis of TVA Sequoyah Containment Shell to Determine Response of a Critical Panel to Uniform Internal Pressure", Offshore Power Systems, September, 1980.

Based on this review the following areas have been identified as potentially limiting the containment capacity to carry uniform internal pressure load. (1) Bending shear in the reinforced concrete containment base mat adjacent to reinforced concrete cylinder walls. (2) flembrane tension in hoop direction in the reinforced concrete cylinder adjacent to the base mat (assuming no rotational or shear restraint by the cylinder). (3) Bending moment in equipment hatch end plate. (4) Bending moment in personnel hatch end plate. 3.0 POTENTIAL FAILURE blODE ANALYSIS 3.1 SHEAR FAILURE IN BASE YAT The program used to determine net shear and tensile forces in the base slab is "GENSHL" which was developed by the Franklin Institute Research Laboratoryof Philadelphia. The program consists of a multi-layered static shell formulation where each shell layer may have different stiffness oroperties and can consider elastic foundation support conditions. This is the same program that was used in the original design and analysis of the base slab for design basis loadings. A uniform soil reaction distribution is used for dead load plus internal uniform pressure case. Results of the analysis are summarized as follows:

1. Specified minimum design strength of concrete at 28 days = 3,500 psi

2. t<ean Sample Value at 28 days = 4,950 psi

3. flinimum Sample Value at 28 days = 4,156 psi Foundation Slab:

T = 120 inches d = 114 inches From computer output as shown in Tables 2 and 3 at sections indicated in. Figure 2:

Evaluation for lowest measured concrete strength value: Nxz Qxs Comp+ k/in k/in Run Case 12.0 psi, internal pressure 1.898 - 2.948 Soil Par. 1 Assume 49.5 psi, internal pressure 7.829 -12. 160 Soil Par. 1 Oead Load 1.300 0.193 Soil Par. 5 OL + 49.5 psi pressure 9.129 -12.353 v = i~ = 12.353 x 1000 = 108.36 psi where: Nxi = membrane tension in base slab Q ia = maximum vertical shear in base slab v = maximum shear stress in base slab From AStlE Section III Oivision 2 and ACI-359-80 Code for Concrete Reactor Yessels and Containments CC 3421.4.1 r I Using lowest measured mean value of concrete strength: r, vc 2r0 p fc (1 + I Or002 Nu/Ag j j rc = 2 0 ~4166 ! 1 + [0 002 -9.129 x 1000 xt 1 flt vc = 2(64.46)[1 - 0.152] = 103.59 psi Note: Internal Pressure Capacity wherever noted as "Psi" means "Psig"-

                                                                                              ~gggQgllXCLXt%
                                                                                              ~RX2~ M

Evaluate for mean of measured concrete strength values Qx.a Comp. k/in k/in Run Assume 53.8 psi, internal pressure 8.509 -13.217 Soil Par. 1 Oead Load 1.300 0.193 Soil Par. 5 OL + 53.8 psi pressure 9.809 -13.410 v = 13.410 (10 vc = 2 74950 )1 + I 0.002 (- 9.809 x 1000~, 1 x 114

balll 1 v "- 117.63 psi vc = 2 (70 ~ 356) C1 0.1635j vc = 117.71 psi In like manner it can be shoNn for a specified minimum concrete strength fc = 3500 psi that the internal pressure capacity is 46.4 psi. In this evaluation no credit is taken for the vertical /Ill bar at 6'/c in the base mat nor is any credit taken for shear capacity of the fill slab above the base mat. In table 4 can be found the limiting pressure capacity adjusted for the assumption of minimum specified and minimum sampled material properties as defined in Table 1. In Reference 1 the Hars ead report Pg. 5-1-1 identified a failure mode based on the assumed pull out of the vertical membrane steel in the cylinder wall from the base mat as having a containment internal pressure capacity of 46 psi. Tne pull out failure mode capacity of 46 psi internal pressure capacity of the containment was determined ivithout consideration of the radial shear (diagonal tension) capacity of the concrete vihich is permitted by the ACI-359 code even in presence of membran tension. To ignore the shear capacity of the concrete is not in accordance riith normal design nor analysis procedures. Hence the failure pressure in the concrete containment of 53.8 psi as defined by the calculations performed in this section is limiting.

3.2 MEHBRA</E HOOP T HSIO'l FAILURE OF COtlCRETE CYLII>OER i~lembrane load due to containment pressurization in the horizontal (hoop) direction P =p R where: P = membrane load in lbs/in of wall p = uniform internal pressure R = mean radius of wall (57.5 x 12 = 690 inches) Hembrane load capacity of reinforced concrete cylinder at its base neglecting discontinuity moment transfer: Available Reinforcement

1) 2 Layers 818 bar hoop reinforcement at 18" c/c =

8 in = 5.33 in2/ft of wall ft

2) 3/8" Liner plate = 3/8" x 12 = 4.50 in2/ft of wall

3) 2 Layers 811 bar diagonal reinforcement at 36" c/c considering only those bars acting in tension 2 x 1.56 x ~2 = 1.47 in2/ft of wall 333 ft From Table 1 of this report the mean value of the reinforcement yield = 49.8 ksi and liner plate = 48.3 ksi P = (5.33 in2 x 49.8 ksi) + (4.50 in2 x 48.3 ksi)

              +  (1.47 in 2     x  49. 8 ks i )
              = 265.4 k + 217.4 k + 73.2 k
              = 556.0    kips/ft     =  46.33    kips/in From Eq. 1 p =   46,330    lbs/in    =  67.1 psi 690   >n In Table 4 can be found th limiting pressure capacity adjusted for the    assumption of minimum specified and minimum sampled material properties           as defined in Table=       l.

3.3 PRESSURE CAPACITY OF THE EQUIPMENT HATCH CLOSURE The equipment hatch closures used on the D.C Cook Containments have been identified (Ref.l) as potentially limiting the capacity of the containment to carry internal pressure loads. The reasons for this limitation are identified as follows:

1. The end closure is in the form of a flat plate hence pressure induced loading must be carried by bending rather than membrane shell action.

2. A bolted splice is used in a region of high bending moment which may limit the capacity of the hatch cover to carry pressure load.

3. The far spaced bolt pattern and the relatively low rotational .

stiffness of the equipment hatch barrel result in little rotational stiffness or fixed end moment capacity of the equipment hatch cover-barrel attachment. This requires that the hatch be analyzed essentially as pin connected (allowed to rotate) rather than fixed (moment resistant) at its supports thereby significantly increasing center span moments in the hatch cover. Because of the presence of the unsymmetric splice and the unsymmetric insertion of the personnel hatch into the equipment hatch cover as shown in Figure 3 the evaluation of the equipment hatch uniform pressure capacity cannot be performed with a high degree of accuracy without recourse to a finite element formulation. Two such analyses were performed, one of the cover plate splice and the other of the equipment hatch cover plate including the effect of the splice and the inserted personnelhatch to determine. their maximum internal pressure carrying capacities.

                                                            ~r 3.3.1 Splice Plate    t1oment Ca  acity 3.3.1.1    Hand   Calculation   considering 1" full penetration weld   detail as shown in Figure 4(1~

Before proceeding to a review of the finite element analysis of the splice plate shown in Figure 4, a hand calculation was performed in order to have a basis of comparison with the more detailed finite element calculation (1) flote the Harstead report neglected the weld geometry in its calculation of stresses.

Given: Splice as shown in Figure 4 check section at top of weld (a) 95 1" A-193 Gr 87 bolts on a 224" length of splice = 2.38" spacing between bolts on tension side of splice Limiting capacity of splice at top of weld is assumed at mean yield in outermost fiber of 2 inch splice plate on tension bolt

            , side of splice M2  PL = sZ =     (53.2 ksi)     1   (2.38)(4)  = 84.41   k-in/2.38 in. of splice Limited tensile capacity of splice plate Tx =  t'I~ x = 1o5 in+

T ='M/1.5 = S4.41/1.5 = 56.27 kips/bolt T' jd = (56.27) x (2.5 + 4.0 + 1.875) = 471.26 k-in/2.38 in. of splice Mgoint Moment capacity /in of splice 471.26/2.38 = 198.01 k-in/in Moment capacity of 4" plate without splice M4<<PL = sZ = 53.2 ksi(1)(1) 16 = 141.87 k-in/in < 198.01

                                                              ~s
                                                                       .".4" plate governs design Capacity of Splice         =  198.01    ]39,6    of 4" plate
                                     ~87 Check  section at base of weld

Limiting capacity of splice at base of weld is assumed at mean yield in outermost. fiber of 2 inch splice plate plus 1" weld. (Minimum Specified F of the Held material = 60.0 Ksi) M2<< PL + 1<~ weld = s Z = (53.2 Ksi) 1 (2.38) 9 = 189.92 K in/2.38 in. of splice Limited tensile capacity of section Tx = M; x = 2.5 in. T = M/2.5 = 189.92/2.5 = 75.97 Kips/bolt Since 75.97 > 56.27 Kips top of weld limits design Check capacity of bolt From Table 1 Mean Yield of 1" bolt = 121.3 Ksi Tensile area 1" bolt = 0.605 sq. in. Pyie]d = 121.3 x 0.606 = 73.51 Kips/bolt > 56.27:. OK ~

3.3.1.2 Finite Element Analysis In Figure 5 is the finite element model of the equipment hatch splice joint showing plate elements. Using the computer program AtlSYS for an applied moment to the 4 inch hatch cover plate equal to a reference containment internal pressure of 40 psi, .the. maximum outer most fiber stress in the 2 inch splice plate is 27.82 ksi in element 76. The maximum outer most fiber stress in the four inch plate is determined as 46.27 ksi in element 145. It appears therefore that the 4 inch rather than 2 inch plate at the joint controls design. This is due primarily to the weld which significantly increases the effective thickness of the splice plate at its connection to the four inch plate. 3.3.2 E ui ment Hatch Cover Plate Pressure Capacity In reference 1 Harstead determined the equipment hatch capacity of 53.0 psi uniform pressure loading based on simple support boundary conditions of the cover as a uniform 4" thick circular plate having a diameter of 19'-10". 8ecause of the effect of the unsymmetric splice and personnel hatch insert a finite element analysis of the plate is performed. A finit el ment modeling of the plate which included the splice is shown in Figure 6. The personnel hatch because of its rigid equiva'lent 12" thick support ring connection to the equipment hatch is assumed to transmit only reaction loads due to pressure to the equipment hatch. The splice is modeled as an equivalent 12" x 4" beam parallel, to the splice and equal to the stiffness of the four inch plate across the splice. Using the computer program A"(SYS the maximum stress intensity in the cover plate is d termined in element 95 as shown in Figure 7 adjacent to the splice. The resultant limiting internal pressure load at element 95 is 45.1 psi for an "as built" mean yield stress of 53.2 ksi in the plate. From Sections 3.3.1.1 and 3.3.1.2 of this report it is determined that the splice plate has a greater moment capacity than the four inch plate. The limiting internal pressure capacity of the Equipment Hatch Cover Plate is therefore limited by the capacity of the four inch plate at 45.1 psi. In Table 4 can be found the limiting pressure capacity adjusted for the assumption of minimum specifi d and minimum sampled material properties. 3'.4'RESSURE CAPACITY OF PERSOiklEL HATCH 3.4.1 End Plate Closure

Hecause of the unsymmetric stiffening of the personnel hatch cover plate as snown in Figure 8, a finite element analysis of the plate is performed to determine its internal pressure retaining capacity. As in the case of the equipment hatch the loading from the personnel hatch door is transmitted to the personnel hatch closure plate as a reaction line load at the ooint of support. Also the plate is conservatively assumed simply supported rather than fixed end supported at its connection to the personnel hatch barrel because of the relative low rotational stiffness of the barrel. In Figure 9 is found the finite element model of the hatch showing all elements. The plat and stiffner system is analyzed using the computer program A ISYS. The maximum outermost fiber stress is determined in the door stiffner at element 87 as 79.3 ksi for a ref rence 70 psi internal pressure load. The pressure capacity p of the personnel hatch closure is determined: p = 70 x 53;2 = 47.0 psi 79.3 3.4.2 Door The personnel hatch door is shown in Figure 8. It acts essentially as a one way spaning simoly supported stiffened plate. The total span of the 1/2" tliick plate is 42". The plate is stiffened by 3" x 1-1/4" solid plate stiffners on approximately 15 inch centers. Assuming a composite T section with the effective outstanding flange leg of tee equal to 8 times the flange thickness, the moment of inertia of the T section is 6.93 in4 and distances to the outermost fibers of the section are 1.03 and 2.47 inches respectively. tlaximum applied bending moment:

    )~i = 1 b p  1 2    1 (15) (p) (422)     3307.5 F

Homent Capacity of Stiffen Ooor Section: H = sZ = (5 ,200) I = (53,200) 6.93 = 149,261 c 22iT Limiting internal pressure p = 149,261 = 45.1 psi 3307.5

In Table 4 can be found the limiting pressure capacity adjusted for the assumption of minimum specified and minimum sampled material properties. , 4.0 SUt"'u~'IARY Af'lD CONCLUS I OH From the summary results of the analysis presented in Table 4 it can be seen that the current limiting internal pressure capacity of the D.C. Cook Containments are the equipnent hatch closure plate and the equipment hatch door at 45.1 psi based on the use of mean "as built" material properties. It should also be pointed out that even if specified minimum material properties had been used as was the case reported in Ref. 1 by Harstead the minimum capacity of the D.C. Cook

     'Containment is 32.3 psi based on the more detailed analyses reported herein rather than the 23.5 psi reported in Ref. 1 which were based on more approximate hand calculations.

It should also be emphasized that the analytical assumption used in the more rigorous analyses reported in this study of the equipment and personnel hatches whose limiting failure modes were in bending still considered only elastic behavior and section properties. It has long been established in the behavior of plate elements during test and as the'asis for the 1.5 increase in allowable bending versus menbrane stress limits of the ASt<E Boiler and Pressure Vessel Code that pl.ate and shell bending elements behave essentially elastic (small deformations) until the plastic section modulus is reached. Since the plastic section modulus for rectangular shapes associated with the hatch plates is 1.5 times the elastic section modulus there is significant additional safety margin in the hatch analysis which is not applicable to the membrane or shear type failure modes identified in the containment concrete shell and base mat. To quantify the effect of the plastic section modulus of the equipment hatch on the internal pressure capacity of the containment a non-linear elastic-plastic finite element analysis of the hatch cover plate using the computer Program Af"SYS was performed for the assumed fy = 50.3 Ksi material property. Evaluation at 70 psi internal pressure or 1.64 times the elastic capacity of the cover plate indicated .that the maximum deflection of the plate is still linear and the maximum plastic strain was 1.8 times the elastic strain at yield. Therefore it is our conclusion that the D.C. Cook Containments as presently designed and constructed constitute a balanced design. That is, the true pressure retaining capacity of the hatches when the 1.5 factor discussed previously is applied is approximately the same as that of the concrete limiting portion of the containment, approximately 54.5 psi. On this basis we do not recommend any modification of the existing D.C. Cook'ontainment hatches.

0 0

Figure 1 O.C. Cook Containment Oimensions and General Arrangment Figure 2 Shear Failure Planes and General Arrangment of Reinforcement in the 8 base mat Figure 3 General Arrangment of the Equipment Hatch Closure Plate Figure 4 General Arrangment of the Equipment Hatch Closure Plate Splice Figure 5 Finite Element 51odel of Equipment Hatch Closure Plate Splice Figure 6 Finite Element tlodel of the Equipment Hatch Closure Plate 'Figure 7 Oetailed Finite Element i~lode of the Equipment Hatch Closure Plate Figure 8 General Arrangement of the P rsonnel Hatch Closure , Plate Figure 9 Finite Element i'lodel of the Personnel Hatch Closure Plate Table 2 Computer Calculated Resultants Forces in the Containment Base Slab Oue .to Oead l!eight Table 3 Computer Calculated Resultant Forces in the Containment Base Slab Oue to a Reference 12.0 psi Internal Pressure

TABLE 1 SU &iARY OF tiItfIMUN SPECIFIED AND AS BUILT MATERIAL PROPERTIES

1. LINER PLATE SA442 ~

GRADE 60 YIELD ULTIMATE ksi SAMPLE SIZE = 6 SPECIFIED MINIMUt 1 32.0 -60.0 S = 2.27 MEAN SAMPLE VALUES 48.3 64.7 Cov. = 0.047 i~lINIf'lUM SAMPLE VALUES 45.8 62.4

2. EgUIPMEiNT HATCH SA516 GRADE 70 SAMPLE SIZE = 5 SPECIFIED MINIMUM 38.0 70.0 S = 2.74 i'lEAfl SAtlPLE VALUES 53.2 81.2 Cov. = 0.051 MI fIMUM SAiiPLE VALUES 50.3 80.2

3. BOLTING SA193 GRADE 87 SAtiPLE SIZE 2 ea.

1/2" x 2-1/2" SP EC IFI lI ED i N Ii~lUM 105.0 125.0 i~lEAN SAMPLE VALUES 119.0 137.0 1" x 5-1/2" (SPLICE) SPECIFIED MIiVIMUM 105.0 125.0 i~iEAth SAMiPLE VALUES 121.3 141.0 l-l/4" x 10" (COVER) IFIED tlINIi~iUi~l

                                                              'PEC 105.0      125.0 flEAN SAMPLE VALUES                 120.1      140.3

e. REINFORCING ROD A15 GRADE 40 18S SPEC I FIED tlI N IMUM 40.0 70.0 SAMPLE SIZE 9 t1EAN SAMPLE VALUES 49.8 81.8 S = 3.34 MINIMUM SAt'lPLE VALUES 44.3 75.5 Cov. = 0.067

5. 28 DAY STRENGTH UflIT 1 and 2 CONCRETE SPECIFIED t'lINIflUil '.5 SAMiPLE SIZE 29 flEAN SAt'1PLE VALUES 4.956 S = 0.508 YINItlUtl SAMPLE VALUE 4.112 Cov. = 0.103

WV64VJI140 ~ 409 ~ i999)]l 0 lll 974 ~ ~ ii0 vi Table 2 Computer Calculated Resultant es in the SOLUTION FUNCTIONS IH SYSTEtt REFEREHCE FRAttf Containment Base Slab Due to De Weight 1 0.207258K 04 0.0 0.179783E 04 0.834138E 05 -0.120054E"DR 0.0 -0.719304E-DR 0 9 163836. E 0 3 R 0.220710K 04 0.0 0.1816692E 04 0.548537K 05 0.148695C-OR 0.0 0.70479OE-OR 0.16218iiE 03 3 0.2340638K 04 0.0 0. 183450 E 04 0.243167E 05 0.414161K-02 0.0 -0 '85N9E-02 0.159343E-03 0.249031E 04 0.0 0.185155K 04 0.8N924E 04 0.675165E-OR 0.0 -0.64346.7E-02 0.156762E 03 5 0.263875E 04 0.0 f

0. 186 704 O4 -O.42e949E 05 0.9304051E DZ 0,0 0 '36066K-DR 0. 152896 E-03 6 0.279153E 04 0.0 0. 188343 E 04 -0.796689E 05 0 ~ 117869E-0 1 0. 0 -0.603897K-OR 0. 148194" E-03 7 0. 2940CEi5 04 0.0 0.189335E 04 -0.11C619E 06 0.141846K-ol 0.0 0.5646ii9E 02 0.142618E-03 8 0.310?2'?E 04 0.0 0.191265K 04 -0.159789E 06 0 ~ 164i827E-Dl 0' -0.56200034'.E-02 0. 135108f -03 9 0.32737"E 04 0.0 0.19263>>E 04 -0.20322lf 06 0.186456E-01 0.0 -0.4757042E-DR 0.126616f-03 10 0. 344161E 04 0.0 0. 193942K 04 -0.24S953K 06 0.2071452-01 0.0 0.421537K-DR 0.120092E-03 11 0.361263E 04 0.0 0. 195193K 04 -0.29)069K 06 0.226201E-ol 0.0 -0.361094K-02 0.110472K-D3 ACTUAL STRESS RESULTAIITS-SIIELL REFERENCE FRAtlf-BODY 7

                                                     %AT CEIITROIOc STATIOtl CEIITROIOS                       Hll                 tt12                H22            013             Q23           Hll               H12                 tt22 tlO;    tlfRID ~ HOOP               LB/IH                LB/IH               LB/IN          LB/IH           LB/IH       IH-LB/IH           Itl-LB/IH            IH-LB/IN 62.662     62.BII 0. 179783 E             04 0.0                  0.216753K 04-0.207258K 04     0.0         -0.192634E 05 0.0                        0,189855E 06 R   62.86R     62.811 0. 1S1662E              04 0.0                  0.216602E Q4-0. 220710K 04    0.0         -0.488963f 05 0.0                        0 ~ 177062E 06 3   62.65      6 .Sll 0. 183450 K             04 0.0                  0. 216593E 04-0.234638E 04     0.0        -0.80454ioiE 05 090                      0.163588E 06 62.85R     62.811 0. 185155E              04 0.0                  0.216720E 04-0.N9031E 04      0.0         -0.113994E 06 0.0                        0,149402E Ob 5   62.6562    62.811 0.166784E               04 0.0                  0. R16971E 04-0.626'3875E 04 0.0          -0.149570E 06 0.0                        0.134474E 06 6   62.R62     62.Cll      0. 160>>343K 04        0.0                  0.217339E 04-0.279153E 04 0.0             -0 ~ 187235E 06 0,0'                     0.11877ef 06 7   62.862     6".Sll      0.189335K          04 0.0                  0. 217818E 04-0.2948ii E 04 0.0           -0.227037E 06 ~ 0                        0. 102285K Ob 8   62066R     62.611 0.191 4"f               04 0.0                  0.218402E 04-0.310929K 04 0.0             -0.269024K 06 0.0                        0 849739E   05 9   62.862     62.811 0.192634if              04 0.0                  0.219084K 04-0.327378E 04 0.0             -0. 313238E     06   0.0                 0.668169E 05 10   60.06 "61 011 0;19 W."E                   04"0;0                  0:039060r04=0".344161E"04 D.E             -0.359717E      06   0.0                 0.47793if   05 ll   62.862 62.811 0.195193E                   04 0.0         -

0.220726E 04-0.361263E 04 0.0 -0.408547E 0 06 0.0 0.278602K 05 RESULTANT STRESSES-PSI BOGY S STATION LAYER STRESS Sll STRESS Sll STRESS S12 STRESS S12 STRESS S22 STRESS S22 HO. HO. INSIDE OUTSIDE INSIDE OUTSIDE INSIDE OUTSIOE I 1 0.14413K 01 0.13126E 01 0.0 0.0 0.36946K 02 0.33971E 02 2 -0.46754E 02 -0.46667E OR 0.0 0.0 0.99142E-05 0. 98951E-05 3 -0.1609iE-05 -O.16O44E-O5 0.0 0.0 0.28696K 03 0.28607E 03 4 -0.465i7E 02 -0.46276E 02 0.0 0.0 0.9C645K-05 0 93099E 05 5 0.12971E 01 0 ~ 1113CE ol 0.0 0.0 0 ~ 3>>4313K 02 0.29374E 02 6 <<0.40609E 02 0 403r59E 02 0.0 0.0 0.85744K 05 0.85194K-05 7 "0. 13917E-05 -0. 13820K-05 0.0 0.0 0.24?obf D3 0.24529E 03

3 Computer Calculated. Resultant Forces i Containment Base Slab Due to a Referen .0 psi COOK PLAHT SOIL PARAtlETER STUDY tIO. 1 12-30-80 LOAOItlG 3 DEAD IIEIGIIT Internal Pressure SOLUTIOII FUIICTIOIIS IH SYSTEti REFfREtlcE FRAtiE 1 0.362819E 04 0.0 0.12<i021E 04 0 ~ 197442E 06 "0.181325E 01 0.0 -0.936208f-02 0. 193628E-03 2 0.300629E 04 0.0 0.125169E 04 0.1477Z6E 06 -0.161004E Ol 0.0 -0.9 >0639E-02 0 195016E 03 3 O.R40632E 04 0.0 0.126290E 04 0. 108624E 06 -0.180663E Ol 0.0 -0. 939061E-02 0 195358E 03 4 0 18 6rSZE 04 Q.o 0 12732rDE 04 0.600810E 05 -0.180361E 01 0.0 -0.93292QE-OZ 0.1'94910E-03 5 0 ~ 126535E 04 0.0 0.1"6295E 04 0. 609194E 05 0.1600<>OE 01 0.0 -0.923599E-02 0. 193908E-03 6 0.721415C 03 0.0- 0.129201E 04 0 roQ246E 05 -0.179?RZE 01 0.0 -O. 912352 E-02 0.1925?GE-03 7 0.19$ 453E 03 0.0 0.1300<>5E 0E-02 0.191125E-03 8 -0.319680E 03 0.0 0.13OGReE O4 0 560519E 05 -0.179091E 01 0.0 -0.866667E-02 0. 18975<IE-03 9 -0.618996E 03 0.0 0.1315<>?E 04 0.705932E 05 -0.178760E 01 0.0 -0.878656E-02 0. 166653E-03 10 -0.13054ZE 04 0.0 0 ~ 132209f 04 O.926286C 05 -0.1764i69E 01 0.0 "0.871393E-OR 0. 180003E" 03 ll -0.176029E 04 0.0 0 ~ 13261ZE 04 , 0.121891E 06 -0.178159E 01 0.0 -0. 667<i6<iE-02 0. 187975E-03 ACTUAL STRESS RESULTAHTS-SIIELL REFEREtICE FRAtlE-BODY 7

                                              /AT CEIITROIO<

STAT IOII CEHIROIDS till O I '> ll22 R13 O23 till ti12 . ti22 HO. ttCRIO. IIOOP LG/IH LG/IH LB/IH LB/IH LG/IH IH-LG/IH IH-LB/IH IH-LG/IH 1 62.662 62.811 0 ~ 124021E 04 0.0 0.163855E 0<>-0.362619E 04 0.0 0.126612E 06 0.0 0.250910E 06 2 62.662 62.611 0.125169E 04 0.0 0.16391<>E 04-0.3006 9E 04 0.0 0.7622GOE 05 0.0 0. 236913E 06 3 62.862 62.611 0.12629OE 04 0.0 0.16320?E 04-0.2<i0632E 04 0.0 0.3669?GC 05 0.0 0.223955E 06 62.662 62.611 0.1273 5E 0-0.12653rSE 04 0.0 -0.123520E 05 0.0 0.202044E 06 6 62.662 6"..811 0. 129201E 0if 04-0.'/21<>15E 03 0.0 "0.2296<i2E 05 0.0 0.193441E 06. 7 62.662 GZ.I>11 0.13004. E 03F 03 0.0 -0.2<>9$ 14E 05 0.0 0.186555F 8 62.652 CR.GII 0.13062of 0<i 0.0 0.16095?E 04 0.3196QCE 03 0.0 -0.166653E 05 0.0 0.161563E 06 9 62.662 62.611 0.1 515<i?C 04 0.0 0.159956E 04 0.81699GE 03 0.0 -0.453569E 04 0.0 0.17652ZE 06 10 62.CGZ 62.811 0.132209E 0:GZOE 04 0.130542E 04 0.0 0 1712"1E 05 0 0 0.177466E 06 ll 62.662 62.611 0.13"612E 0SIDE OUTSIDE II>SIDE OUTSIDE 1 0.2<i416E 01 0.22035E Ol 0.0 0.0 0.60108E 02 0.5419?E 02 2 -0.73659E 02 -0.73<i81E 02 0.0 0.0 0.15811E-04 0 ~ 15773E-04 3 -0.25338E-05 0 2r5241E 05 0.0 0.0 0.457<ilE 03 0.4556<if 03

-0.?319QE 02 -0.72691E 02 0.0 0.0 0.1571ZE"04 0 15603E 04 5 0.2)749E 01 0.18357E Ol 0.0 0.0 0.53466E OR 0.45064E OR 6 -0.61216E 02 -0.60709E 02 0.0 0.0 0. 131<ICE-04 0.13037E-04 7 -0.20934E-05 -0.20738E-05 0.0 0.0 0.37609E 03 0 '7<>56E 03
TABLE 4 SUf1NARY OF LIHITING INTERNAL UNIFORt1 PRESSURE CAPACITY OF D.C. COOK CONTAINMENT INTERNAL PRESSURE CAPACITY (ELASTIC ANALYSIS) (See Subsection 4.0 for Plastic Analysis) CRITICAL FAILURE SPECIFIED MINI tlUN LOlJEST ii/EASURED MEAN SAMPLE 110DE PROPERTIES SNlPLE PROPERTY PROPERTY
l. Bending Shearing fc = 3500 psl; fc = 59.16 fc = 4100 psi; fc = 64 03 fc = 4950 psi; fc = 70 36 Concrete Base t1at Limiting internal Limiting internal Limiting internal pressure = 45.8 psi pressure = 49.6 psi pressure = 54.5 psi

2. Membrane Hoop Tension fy = 40,000 psi f> = 44,300 psi fy = 49,800 psi in Concrete Cylinder Limiting internal Limiting internal Limiting internal pressure = 50.2 psi pressure = 61.2 psi pressure = 67.1 psi

3. Bending Capacity of f> = 38,000 psi f> = 50,300 psi fy = 53,200 psi Equipment Hatch Limiting internal Limiting internal Limiting-internal pressure = 32.3 psi pressure = 42.6 pressure 45.1

4. Bending Capacity of fy = 38.000 psi fy = 50,300 psi fy = 53,200 psi Personnel Hatch-(a) Closure Plate Limiting internal Limiting .internal Limiting internal pressure = 33.6 psi pressure = 44.4 psi pressure = 47.0 psi (b) Door Limiting internal Limiting internal Limiting internal pressure = 32.3 psi pressure = 42.6 psi pressure = 45.1 psi Note: Internal Pressure Capacity wherever noted as "psi" means "Psig"
CCNTAILlHEMT Ql Qll STEEL LILlE ll
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leo'-0" OD GFCTtohJ A-A GECTIQN4,L F.LEYATIQQ Figure 1 D.C. Cook Containment Dimensions and General Arrangment
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Figure 2 Shear Failure Planes and General Arrangement of Reinforcement in the Containment Base Mat
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1 11 72 S9 36 39 33 29 5 32 35 33 3iQ 12 16 821 28 31 37 ?48 56 3 14 720 2'? 14? 55 62 33 13 619 26 046 54 61 20 12 518 25 QiS 53 60 7 16 ea 6 11 417 24 844 52 59 70 17 9 1 5 10 316 23 743 51 58 6< 69 18 70.8 -5'?.8 -44.8 -31.9 -18 9 -5 ' 7.1 20.1 33. ~ 46.0 SQ 0 Ala LOCK PLAY PLAYS AtlALVSls CEO?1ETRVANSYS Figure 9 Finite Element Model of the Personnel Hatch Closure Plate
SECTION 2 Phase II of the D.C. Cook Internal Pressure Containment Anal sis-Probabilistic Anal sis In this effort the variability of the "as-built" material parameters on the best estimate capacity of the containment to carry static uniform internal pressure is being evaluated. Four potential limiting failure modes have been identified by deterministic analysis. 'wo of the modes involve potential failure by plate bending of the equip-ment and personnel hatch closure plates. The other two potentially limiting failure modes are by membrane tension failure of the main steel hoop reinforcement at the base of'he containment shell and shear (diagonal tension) failure of the concrete base met. The ACI-359 Code equation governing diagonal tension failure is based on test results hence it is also being evaluated in a probabilistic manner. Results of this statistical analysis will be probability density function of containment resistance defined for the two different contain-ment "as-built" material properties and in the case of shear in the base mat the statistical nature of the code defined failure equation. This evaluation should be completed by May 15, 1981. SECTION 3 Phase III of the D.C. Cook Internal Pressure Containment Anal sis-Localize D namic Loads In this evaluation dynamic analytical models of the contain-ment structure assuming localize dynamic pressure loading input are being prepared. The containment areas where the dynamic models are being de-veloped include the equipment and personnel hatch closure plates, the shell portion of the containment shell adjacent to the base mat and the bise mat adjacent to the cylinder shell juncture. The development of the dynamic models should be complete by May 30, 1981. Then using the internal pressure time history forcing functions, a dynamic analysis will be done to determine the forces and moments at the critical sections of the containment.
DONALD C. COOK NUCLEAR PLANT UNIT NOS. t AND 2 ATTACHMENT NO. 2 TO AEP:NRC:00500A SECOND QUARTERLY REPORT ON HYDROGEN MITIGATION AND CONTROL
0 2.0 Distributed I nition S stem 2.1 Introduction Indiana 5 Michigan Electric Company (ISMECo.) has decided to install a Distributed Ignition System (DIS) in Unit Nos. 1 and 2 of the Donald C. Cook Nuclear Plant. The DIS utilizes thermal resistance heating elements (glow plugs) located throughout the containment building. Operation of the DIS will be accomplished by means of manual control switches located in the main control room. 2.2 Distributed I nition S stem Desi n The DIS is a two-train system employing sixty eight (68) igniter assemblies located throughout the containment building. Each train of thirty four (34) igniter assemblies is further divided into two groups one group of sixteen (16) assemblies in the general lower volume area and a second group of eighteen (18) assemblies in the general upper volume area - including the ice condenser upper plenum volume. Each igniter assembly consists of a General Motors type 7G AC glow plug and a Dongan Electric control power transformer (model 52-20-435) mounted in a sealed box housing as shown in Figure 2. The igniter box is a water tight enclosure meeting NEMA-4 specifications. A copper plate is employed as a heat shield to minimize temperature rise inside the igniter box and a drip shield is utilized to minimize direct water impingement on the thermal element. The transformer is seismically mounted to the igniter box using unistrut. The entire igniter assembly is seismically mounted so as to prevent any possible interferences with safety-related equipment during/after a design basis seismic event.
The normal and emergency power sources for each train of igniters meets Electrical Class lE specifications and the electrical train separation criteria commensurate with a Class 1E system are maintained in the DIS design. The DIS will be a manual system controllable from the main control room. Two control switches per train will be located on auxiliary relay panels A7 and A8 in the main control room. The control switches are of the two-position type,
'off'nd 'on', and red and green indicating. lights are provided above each switch. Control room annunciation will be provided to indicate loss of power and failure to operate due to hypothetical control circuit equipment mal:functions.
2.3 JIIAIAb1 The igniter assembly is a 16" x 12" x 8" enclosure meeting NET-4 specifications. The igniter is protected from direct water impingement by a 1/8" steel plate (10" x 18" galvanized steel) drip shield welded to the top of the enclosure. The igniter is mounted to the enclosure through a 6" x 4" x 1/4" copper, plate to reduce the temperature rise. inside the enclosure during. periods of combustion. All electrical connections inside the igniter assembly; its associated condulet box, and the two splice boxes per train utilized in the DIS are protected with heat shrink tubing to enhance system performance in an adverse environment. In addition, all DIS cables inside containment are routed in conduit and hence are protected from the environment associ'ated with hydrogen combustion.
'3 Access to the interior of the igniter assembly is through a hinged cover plate secured with screws. A bead of silicone rubber will be placed around all bolt holes in the igniter assembly. Details of the igniter assembly and its condulet box are given in Figure Nos. and 2. 1
2. 4 I niter Assembl Locations Igniter assemblies are distributed throughout the containment to promote combustion of lean hydrogen/air/steam mixtures. The DIS will minimize the potential for hydrogen accumulation and preclude detonations in the unlikely event of a degraded core cooling event similar in nature to the TflI-2 accident involving substantive hydrogen generation. The containment air recirculation/hydrogen skimmer system, in conjunction with upper and lower volume containment sprays, provides sufficient mixing so as to prevent the stratification or pocketing of hydrogen in the various compartments of the containment building.
Approximate igniter assembly locations are listed in Table 2-1. A general view of. the containment structure is provided in Figure
~
3 and approximate igniter locations shown in Figure Nos. 4, 5 and 6. The locations given are
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for D. C. Cook Unit No. 2 and are typical for Unit No. 1.'inor'.variations'n ig-
~ ~ ~ ~ ~ ~ ~
niter locations may be required in. Unit iVo. 1'in consideration of physical inter-ferences with. existing equipment. A'schematic representation of the DIS electrical network inside containment is provided in Figure Nos. 7 and 8. One of the questions raised by members of the NRC staff during our meeting of March 18, 1981 dealt with the need, or lack thereof, to install igniter assemblies in the instrument, room. The results of our reviews. performed to date indicate that except for potential in-leakage there is no communicatio'n between the instrument room and either the general lower volume or the pipe tunnel (annulus region) with the exception of the flow path-through the hydrogen skimmer ductwork.
The above notwithstanding, it should be noted that any leakage into the instrument room would, in all probability, be significantly less than the hydrogen skimmer flow (100 CFt1 per train) out of the room, thus preventing I the accumulation of hydrogen to combustible levels. It should also be noted, that the effects of hydrogen combustion on 'required'quipment located in the instrument room, pressurizer pressure and pressurizer level transmitters, is, for all intents and purposes, bounded by the calculations contained in Attachment No. 4 of this submittal.
TABLE 2-1 Sheet 1 of 2 IGNITER ASSEMBLY LOCATIONS* TRAIN TRAIN
'B'om 'A'om No. artment/Area-El evati on No artment/Area-El evati on A-1 Ice Cond. Upper Plenum Ice Cond. Upper Plenum A-2 Ice Cond. Upper Plunum B-2 Ice Cond. Upper Plenum A-3 Ice Cond. Upper Plenum B-3 Ice Cond. Upper Plenum Ice Cond. Upper Plenum B-4 Ice Cond. Upper Plenum Ice Cond. Upper Plenum Ice Cond. Upper Plenum A-6 Ice Cond. Upper Plenum 708'09'09'09'09'10'09'86'86'86'86' B-6 Ice Cond. Upper Plenum A-7 Ice Cond. Upper Plenum B-7 Ice Cond. Upper Plenum A-8 Inside ¹1 SG Enclosure B-8 Inside ¹1 SG Enclosure A-9 Inside ¹2 SG Enclosure B-9 Inside ¹2 SG Enclosure A-10 Inside ¹3 SG Enclosure B-10 Inside ¹3 SG Enclosure A-11 Inside ¹4 SG Enclosure B-11 Inside ¹4 SG Enclosure A-12 Inside PZR Enclosure B-12 Inside PZR Enclosure 686'59' Outside ¹1 SG Enclosure B-13 Outside ¹1 SG Enclosure Outside ¹2 SG Enclosure B-14 Outside ¹2 SG Enclosure A-15 Outside ¹3 SG Enclosure B-15 Outside ¹3 SG Enclosure 709'09'09'09'09'09'09'86'86'86'85'82'62'59'59'59'59'42'37'36'36'37'45'30'29'23'34'18'60'60' A-16 Outside ¹4 SG Enclosure 662'62'62'62'47'48' B-16 Outside ¹4 SG Enclosure A-17 Outside PZR Enclosure B-17 Outside PZR Enclosure A-18 Primary Shield Wall B-18 Primary Shield Wall A-19 Primary. Shield Hall B-19 Primary Shield Wall A-20 Primary Shield Wall B-20 Primary Shield Hall A-21 Primary Shield Hall B-21 Primary Shield Wall A-22 Primary Shield Wall B-22 Primary Shield Wall A-23 Primary Shield Wall 648'48'41'48'31'29'34'18'60'60'-1 B-23 Primary Shield Wall A-24 East Fan/Accumulator Room B-24 East Fan/Accumulator Room A-25 East Fan/Accumulator Room B-25 East Fan/Accumulator Room A'-26 West Fan/Accumulator Room B-26 West Fan/Accumulator Room A-27 Hest Fan/Accumulator Room B-27 Hest Fan/Accumulator Room A-28 Vicinity of PRT B-28 Vicinity of PRT A-29 Upper Volume Dome Area B-29 Upper Volume Dome Area Upper Volume Dome Area B-30 Upper Volume Dome Area
Sheet 2 of 2 TRAIN TRAIN
'B'om 'A'om Ho. ar tment/Area-El evation No. artment/Area-Elevation A-31 Upper Volume Dome Area - Upper Volume Dome Area -
A-32 Volume Dome Area - 760'pper B-32 Volume Dome Area - 760'pper A-33 Volume Dome Area - 748'-31 748'pper B-33 Volume Dome Area - 748'pper A-34 Volume Dome Area - 748'pper B-34 Volume Dome Area - 748'pper 748'EY: SG - Steam Generator PZR - Pressurizer PRT - Pressurizer Relief Tank locations given are for Donald
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C. Cook Unit Ho. 2 Containment Plan Above Elevation 715'
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DONALD C. COOK NUCLEAR PLANT UNIT NOS. 1 AND 2 ATTACHMENT NO. 3 TO AEP:NRC:00500A SECOND QUARTERLY REPORT ON HYDROGEN MITIGATION AND CONTROL
3.0 Inade uate Core Coolin H dro en Control E ui ment 3.1 Introduction There are two primary concerns associated with an inadequate core cooling (ICC) event similar to the TMI-2 accident involving the release of substantive amounts of hydrogen and subsequent combustion uti li zing the Distributed Igni tion System (DIS). These concerns involve, (1) the abi lity to achieve and maintain the reactor coolant system in a safe shutdown condition and (2) maintenance of containment integrity through adequate hydrogen control. The equipment located inside reactor containment required to perform the above functions is identified in this section. The survivability of the equipment discussed herein during periods of hydrogen combustion is addressed in Attachment No. 4 of this submittal. The containment response to hydrogen combustion is contained in Offshore Power System (OPS) Report No. 36A05 previously transmitted to the Commission as Attachment No. 2 to our first quarterly report on hydrogen issues (AEP:NRC:00500 dated 12 January 1981). The analyses performed by OPS utilizing the CLASIX computer code clearly indicate that the peak pressure resulting from hydrogen combustion is well below the ultimate strength of the Cook Plant containments. 3.2 ~Ei EE Table 3-1 lists the active components inside containment required to function during and (or) after periods of hydrogen combustion. The location of these components and their susceptibility to hydrogen combustion effects are addressed bel ow.
(1) Steam Generator Narrow-Ran e Level Monitors
.Three safety-grade differential pressure transmitters (tL P) are employed on each steam generator to monitor narrow-range steam generator water level. The kP transmitters, manufactured by ITT 8arton, are fully qualified for post-accident use inside containment (LOCA/MSL8 qualification). These transmitters are located in the general lower volume, with two transmitters per steam generator mounted nearly eleven feet below the maximum containment flood level of 614'levation.
Clasix run JVAC4 (see Attachment No. 2 to our AEP:NRC:00500 submittal - OPS Report No. 36A05) represents the minimum time to for the S2D cases run to date .and hence represents the case
'ombustion for which the minimum containment water level would exist at the time of'nitial combustion. Figure No. 32 of the OPS report shows the initial combustion to occur in the lower compartment approximately 4,600 seconds into the S2D event sequence. Assuming that water is transferred to the containment from the refueling water storage tank (RWST) solely via two containment spray pumps, it is cle'ar that the minimum usable Rl<ST volume specified in the Plant Technical Specifi-cations (350,000 gallons) would have effectively been delivered to the containment pump long before the onset of combustion. In addition, the OPS report shows that approximately 22.4X of the initial ice inventory has been melted during the LOTIC portion of the analysis; up to a time of 3480 seconds. Assuming the initial ice inventory to be the Technical Specification minimum value of 2.37 million pounds;
~
it is thus shown that in excess of 530,000 pounds of ice has been
.melted prior to combustion. This ice melt is equivalent to approxi-
. mately .80,000 gallons of additional water in the containment. Combining the ice melt with the Rl<ST water yields a total containment "water inventory of 430,000 gallons, well-in excess of the water inventory which would result in submergence of two level transmitters per steam generator. Thus, it is clear that the steam generator narrow-
.range level monitoring function would not be susceptible to the effects .of a hydrogen combustion environment. '(2) Pressurizer Pressure and Pressurizer Level Monitors The pressure transmitters and the kP transmitters utilized for .the pressurizer (PZR) pressure and level monitoring functions, respectively are located in the instrument room. These transmitters, -manufactured by ITT Barton, are fully qualified for post-accident use -.inside containment (LOCA/MSLB qualification). As stated in Section '2.4 of Attachment No. 2 of this submittal, our reViews performed to indicate that there is no comnunication between the instrument 'date room and either lower compartment or the pipe tunnel (annulus region) -;other than the hydrogen skiomer ductwork. In addition, the CLASIX H =analyses do not predict combustion in the dead-ended volume, of which the instrument room is a part. Hence, the information available at -this time indicates that the PZR pressure and level transmitters would II not be exposed to a hydrogen combustion environment in the unlikely event of a degraded core cooling event involving the generation of substantive amounts of hydrogen.
I ~ e
(3( ~333 tll -3 3 T The RCS wide-range pressure transmitters are located in the lo.rer compartment nearly eleven feet below maximum containment floodup level. The transmitters, manufactured by ITT Barton, are fully qualified for post-accident use inside containment (LOCA/MSLB qualifi-cation). For reasons set forth in Item (1) above, these transmitters would be submerged prior to initiation'f combustion and hence would not be exposed to a hydrogen combustion environment in the unlikely event of a degraded core cooling event involving the generation of ,substantive amounts of hydrogen. (4) Core Exit Thermocou les The effects of a hydrogen combustion environment on the core -exit thermocouple cable is addressed in Attachment No. 4 to this submittal. (Ri ~RCS ( RT The hot leg and cold leg RTQs, located in the lower compartment, -are fully qualified for post-accident use (LOCA/MSLB qualification). 3 The cable associated with the RTDs is addressed in Attachment No. 4 to this submittal. tl =(6) Air Recirculation H dro en Skimmer Fans air recirculation/hydrogen skimmer fans are located in the ('he uppe~ cd.,partment and the Pan motors are fully qualified for post-accident use'(LOCA/MSLB qualification). (7). Distributed I nition S stem DIS Com onents The DIS components inside containment are the igniter assemblies; splice boxes and condulet boxes, and the ancillary cable. All DIS cable inside containment is routed in conduit and thus is protected
.from a hydrogen burn. All electrical connections inside the igniter .assembly, its associated condulet box, and the two splice boxes per train utilized in the DIS are protected with heat shrink tubing to .enhance system performance in an adverse environment. The igniter assembly itself is a sealed enclosure meeting NEMA-4 specifications. h
;.Table 3-1 ,;Donald C. Cook Nuclear Plarlt Unit Nos. 1 and 2 .Inade uate Core Coolin /H droqen Control E ui ment* -{1,) ',Narrow-range Steam Generator Level Monitors .(2) Pressurizer Level Monitors
{3) Pressurizer Pressure Monitors
-':(4) RCS Wide-Range Pressure Monitors -.:{5): Core Exit Thermocouples -{6) RCS Loops RTDs --{7) Air Recirculation/Hydrogen Skimmer Fans =-.(8) Distributed Ignition System Components
inside reactor containment
DONALD C. COOK NUCLEAR PLANT UNIT NOS. 1 AND 2 ATTACHMENT NO. 4 TO AEP:NRC:00500A SECOND QUARTERLY REPORT ON HYDROGEN MITIGATION AAD CONTROL
4.0 E ui ment Survivabilit This attachment to the quarterly report addresses the issue of the survivability of equipment exposed to a hydrogen combustion atmosphere inside containment. Heat-transfer models have been developed to determine the effects of hydrogen burns on critical components (see Table 3-1 in ). The models are presented in this attachment followed by a calculation made for a representative piece of equipment. Particular attention has been devoted to a number of individual pieces of equipment, each of which is discussed separately. 4.1 ~G1 A In order to characterize the environment to which a piece of critical
~
equipment is subjected during
~ ~
and subsequent to a hydrogen burn, two heat-transfer models have been developed. The first heat-transfer model is a time dependent heat-transfer analysis which calculates the lower compartment environ-ment as a result of a hydrogen burn. This model takes into account the presence of structural heat sinks and sprays in the lower compartment and assumes that during a hydrogen burn energy is removed by the ice condenser. The burn itself is modelled by an energy input rate to the compartment. At the onset of the combustion, the lower compartment is assumed to be isothermal; energy is then introduced into the compartment for a duration of 20 seconds, comparable to the time of a hydrogen burn in the containment. As a result of the burn, the temperature of the compartment atmosphere begins to rise rapidly; concurrently, heat is being transferred to the structural heat sinks and removed by the ice condenser and by )he lower compartment sprays. Heat transfer to the containment sinks is characterized by both convection and
radiation.
~ ~ Conservative assumptions have been made in the calculation with regard to parameters such as gas emissivity and configuration factors.
After 20 seconds, the atmosphere temperature is observed to decrease exponentially, whereas the containment wall temperature continues to rise over the next twenty seconds (see Figure 4-1) until the time when the atmosphere temperature falls below the wall temperature. The maximum atmos-phere temperature calculated does not exceed 500 F. Sensitivity studies of various parameters used in the analysis are presented in Figures 4-2 and 4-3. Figure 4-2 depicts the results obtained when the heat transfer coefficient,"h", from atmosphere to wall is varied; as "h" vanishes, the peak atmosphere temperature approaches the CLASIX results. It can also be noted that, in general, the peak temperature is fairly insensitive to small variations in the values of the heat transfer coefficient chosen. Perturbations in the spray flow rate also reveal small increases (n 15Ã) in the peak temperature, see Figure 4-3. These analyses clearly show that if containment structural heat sinks are considered, the containment environment is not expected to experience temperatures in excess of 500 F. The equipment included in the critical list of components (Table 3-1) is qualified for LOCA and MSLB events; which includes exposure to 340 F for a period in excess of one hour. Comparison between the MSLB conditions and the data presented in Figure 4-1 indicates that equipment, which is 'subjdcted P to a hydrogen burn of the magnitude predicted by CLASIX, will*experience environmental conditions no more severe than those of a MSLB event. The second heat-transfer model attempts to describe and define the environmental condi ti ons for equi pment which is located in the path traversed by the hydrogen flame. A Barton pressure transmitter has.been selected as a representative piece of equipment to be investigated.
Prior to hydrogen ignition, the transmitter casing and its internals are assured to be in thermal equilibrium with the containment environment. At the onset of a hydrogen burn, it is postulated that ignition occurs in -the vicinity of the transmitter and the casing is subjected to a very high hydrogen flame temperature (~2000 F) initially as the flame front moves away from the component. The temperature to which the transmitter surface is exposed will then decrease gradually and will eventually approach long-time results calculated by the previous heat-transfer model. This temperature profile will provide the outside boundary condition needed to evaluate the temperature rise on the inside surface of the transmitter. The one-dimension .time-dependent conduction heat transfer equation is evaluated assuming that the inside surface-is an adiabatic boundary. This model treats the trans- -.mitter casing as a one-dimensional slab. The time dependent temperature to be used on the outside surface is imposed as a convective boundary
'profile condition.
Two different temperature profiles, which reflect the environment temperature to which the transmitter is exposed, have been employed in this calculation. The first profile represents a hydrogen flame temperature of '2000 F for a duration of one second at the onset prior to a linear decay to
-1000 F in the next second; temperature continues to decrease to 300 F from two to six seconds and eventually approaches 150 F after 10 seconds (see Figure 4-4), curve A. This temperature profile is similar to the one used by TVA in its equipment survivability calculations. The other profile, see Figure 4-4, curve 8, decays exponentially from 2000 F to 150 F over a period of 18 seconds and is similar to the one used in the Duke analysis. A computer code was used to analyze the temperature rise in a 1/4" carbon steel casing given the aforementioned boundary conditions. The heat transfer coefficient assumed in the code includes both convective and radiative transport.
0 The temperature transients at the inside surface calculated from the two temperature profiles are depicted in Figure 4-5. Curve (A) of Figure 4-5, -which corresponds to the curve.A of Figure 4-4, showed that the initial
-temperature rise is very abrupt during the first few seconds; later on the inside surface reaches a maximum temperature of 171 F at 10 seconds prior to a gradual decrease. The temperature response depicted by curve (B) of
.Figure 4-5 indicates .that there is a more gradual rise over the initial 15 seconds and that the temperature reaches its maximum of 175 F at about 30 seconds before a slow decay begins. Based on this analysis, one can assume that for a single hydrogen burn, the inside casing temperature will rise no more than 30 F.. Additionally, if one assumes that there is a total of eight consecutive burns and that between each burn the inside casing surface temperature is held constant, the temperature profile will be a stepwise function similar to the one presented in Figure 4-6. Each temperature increase (30oF.) can be interpreted as the heatup of the casing resulting from one hydrogen burn. Between each burn, the temperature I at the inside casing is assumed to be constant which implies that no credit is given to the cooling of the component subsequent to any burn. In addition, the time interval between combustions is assumed to be substantially shorter
'than what is predicted by CLASIX; only 100 second intervals are used in this calculation. Based on the stepwise curve, a conservative linear heatup temperature profile at the inside surface of the casing is used, see Figure 4-6.
Utilizing this linear temperature response at the inside of the trans-
~
mitter casing, a heat transfer analysis has been performed to'evaluate the heatup rate of the air and the subcomponents inside the casing. Results
indicate that the heatup rate of the air inside is slightly below the temperature of the casing and that the heatup rate of the subcomponents is estimated to be approximately 50 F over seven burns, or,7 F per burn.
.It is important to bear in mind that conservative assumptions have been .used in obtaining the above results.
The .heat transfer analysis clearly indicates that for most equipment
-.which is environmentally. qualified for LOCA or HSLB events, elevated temperatures resulted from hydrogen burns of the magnitude and duration
.discussed do not appear to pose any threat to its abi lity to sur vive in a
=~2D,-type event. -4e2 Survivabilit of Particular Pieces of E ui ment This section of Attachment 4 discusses
~
the survivabi lity of particular -:pieces s of equipment
~
needed for the mitigation and control of a S2D-type sequence. These pieces of equipment require either particular evaluations or, else, the analysis presented in Section 4.1 does not apply to them.
.a) Cables '-The burning of hydrogen inside containment by use of a Distributed Ignition System (DIS) results in very short duration exposure fires and may -involve cables which are exposed in trays.
Inside the Cook containment buildings power and control cables are -either installed in conduits or in cable trays. Cables installed in conduits are not likely to burn as a result of exposure to short-duration exposure fires. These cables cannot propagate a -fire even if they burn since the flame resulting from the. combustion is -entirely confined to the conduit
~ ~
and cannot cause failure of cables in adjacent enclosures.
~ ~
-In the case of the control cables where the current carried by the conductors is small relative to the thermal rating of the conductors, .the cables are installed in trays with solid steel sides, bottoms and covers. .Hence, it is not likely for a hydrogen burn inside containment to ignite any control cables installed in trays . However, upon exiti ng a tr ay, either mid-span through a hole in the tray cover or at the end of the tray span, a portion of the cable becomes exposed for a very short length until the cables either enter a conduit whi ch faci litates entry into terminal devices .or until the cables are connected to the device or containment penetration
,(below flood level ).
.All control cables inside containment needed for inadequate core cooling mitigation equipment are qualified for flame resistance in accordance .:with either IPCEA Standard S-19-81 or IEEE-383. Hence, for the exposed portions
of the control cables and cables entirely contained in trays or conduits, it is extremely likely that the cables will survive hydrogen burns inside contain-s ment. Furthermore, the cable will be wet due to the actuation of containment sprays making the possibility of ignition from a short duration exposure to
,fire even more r'emote.
For the case of power cables, they are installed in conduits or in
-expanded metal trays without covers and are sized to accommodate the full .load current of connected equipment without exceeding their continuous rated.
temperature. 1Jhen installed in expanded metal cable trays,'he cables are laid typically one layer deep with spaces between adjacent cables and secured to the bottom of the tray to maintain this spacing. The power cables for ICC
'equipment may be exposed to hydrogen burning inside containment but they are
q ualified for flame resistance in accordance with IEEE-383 or S-19-81. Further, since the power cables are exposed (open trays) they will be wet due to the effect of containment sprays. Testing results have been reported by L. J . Klamerus of Sandia on IEEE-383 cables . Private communication with Nr . Klamerus revealed that the cables used in the experiment were X-link polyethylene cables. They were selected f'r the test because they were believed to be most susceptible to exposure fi.re fai lure. Reported results indicate that the time to electrical short for these cables ranges from five to nine minutes. Review of ICC equipment power cables at Cook confirms the fact that they are either insulated by g palon or a synthetic compound made by Kerite. Both types of materials are believed to exhibit superior fire resisting capability than those tested by Sandia Laboratory. Therefore, despite the fact that power cables at Cook might be exposed to a two to three minutes total duration of hydrogen burns experimental evidence support the contention that it is very likely that they will be able to survive hydrogen burns typical of those discussed for a S2D-type event. b) Air Recirculation Fans There are two air recirculation fans at Cook and both of them are located in the upper compartment. These two centri,fugal fans have a total capacity of 80,000 cfm and discharge the flow into the two fan/accumulator rooms. At the exit of each fan there is a backdrop damper which opens as a result of flow through the fan. The damper is gravity loaded and is expected to close if there is an"overpressure in the fan/accumulator room. The CLASIX results predict burns in the upper compartment with pressure differentials
1b unaccounted for in the design of the system. Fan integrity is being evaluated both from the point of view of casing damage and overspeeding - of the wheel and motor. c) Steam Inertin and Pol urethane. Insulation Burn In a S D-type event, hydrogen release begins approximately 3800 seconds after the onset of a small break. Results obtained from the March code for Sequoyah indicate that during the initial 700 seconds, the steam con'centration at the lower compar tment reaches a maximum of 78/ prior to decaying to 45/, see Figure 4-7. Subsequently, the steam concentration
-continues to decrease to approximately 25/ at onset of the hydrogen release.
Data reported by the U.S. Bureau of Nines indicate that little change
.to the lower flammability limit of hydrogen is noted when steam concentration in the mixture is kept below 308. Therefore, with a 254 steam concentration in the lower compartment, the effects of steam upon hydrogen combustion should be minimal.'oreover, lower compartment sprays at Cook would further serve to enhance condensation of steam and to promote rapi d temperature r eduction in the. lower compartment. Thus, it is expected that the steam concentration in the Cook lower compartment will be substantially lower than .what has been presented in Figure 4-7. Therefore, it is unlikely that Cook will experience steam .inerting in a S2D-like event except possibly during the initial 1000 seconds.
In addition, data presented by Lawrence Livermore Laboratory in their igniter test program clearly show that steam concentrations up to 40Ã do not inhibit the ignition of hydrogen by the glow plugs nor the abi lity of the igniters to function as designed. In spite of the fact that there would be a higher steam concentration in the lower compartment, evidence indi cates 1
.that the glow plug igniters ~
will perform their
~
intended functions as
~ -required. ~
It is conceivable that at the upper plenum of ice- condenser, a
higher hydrogen concentration may be present as a result of steam stripping by the ice condenser. It has also been postulated that combustion may first occur at that location and that it may even burn in a continuous manner.
However, it must be pointed out that the likelihood of the above scenario =diminishes if the assumption on steam inerting at the lower compartment is considered unrealistic. Given the complexity of this issue, the question of burning in the
.upper plenum of the ice condenser will continue to be investigated by AEP. 'Moreover, upcoming results from the modified version of CLASIX should be able to provide additional information on this subject. If hydrogen combustion is
-assumed to occur at the upper plenum for an extended period of time, it has been postulated that the integrity of the polyurethane insulation may be threatened by the presence of hot gases. This question is being addressed
-at AEP simultaneously with the upper plenum burn issue. The results of our evaluations will be transmitted to the NRC in the next quarterly report.
0
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< 1 ill I l
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References:
(1) Klamerus, L. J ., "Fire Protection Research," quarterly Progress Report, October - December 1977, NUREG/CR-0366. (2) Private Communication, L. J . Klamerus to K. K. Shiu, March 1981. (3) Hertzberg, M., "Flammability Limits and Pressure Development in H2-Air Mixtures," U.S. Bureau of Mines, PRC Report No. 4305, January 1981. (4) Lowry, W., "Preliminary Results of Thermal Igniter Experiments in H2-Air Steam Environments," Paper presented at the workshop on the impact of Hydrogen on Water Reactor Safety, Albuquerque, New Mexico, January 1981. (5) Sequoyah Nuclear Plant, Core Degradation Program, Yolume 2, Report on the Safety Evaluation of the IDIS, December 15, 1980.
0 0
DONALD C. COOK NUCLEAR PLANT UNIT NOS. 1 AND 2 ATTACNENT NO. 5 TO AEP:NRC:00500A SECOND QUARTERLY REPORT ON HYDROGEN MITIGATION AND CONTROL
5.0 Current Research Pro rams Several research programs have been undertaken by AEP to investigate hydrogen control related phenomena; some of these programs were discussed in the last quarterly report. In this section a number of the current research programs will be reviewed; program status, revised test plan and program schedule of each effort wi ll be discussed individually.
~RPRI P AEP., along with Duke and TYA,are co-sponsors of f'our EPRI research programs in which fundamental flame studies wi 11 be made; research and development on various igniter types will be pursued; mixing and distribution of hydrogen in prototypic containment environments will be investigated and additional glow plug testing will be performed.
a) Mhiteshell Nuclear Research Establishment This research facility is operated by 'Atomic Energy of Canada Limited. Two research programs will be pursued independently at this facility; namely, the hydrogen combustion phenomena study and the research and development of different igniter types. Both of these programs will be undertaken with the collaboration of Ontario Hydro as an additional financial contributor to the work. I The first experimental program is designed to investigate various hydrogen combustion phenomena and can'be divided into four parts. The first part of this experimental effort entails performing nineteen ignition tests on lean hydrogen mixtures. The hydrogen concentration to be examined in these tests will vary from 5.05 to 30Ã by volume. A spark ignition source which is
i n the order of 0.5 joule will be used to ignite the mixture. Details of the experimental set up and test vessel dimensions have been presented in the previous quarter ly submittal. Fast response pressure transducers, thermo-couples and ionization probes will be employed to monitor and record various important test parameters. Of the= nineteen tests planned the majority of them will be conducted with the ignition spark located near the bottom of the spherical test vessel. Two tests are planned in which the ignition source will be located at the center of the vessel and one test is planned with the ignition source near the top of the vessel. These three tests will be used to assess the effect of igniter location. These tests are anticipated to require approximately three weeks to complete. According to the latest estimate provided by HNRE, system shakedown is being performed on the test vessel and on the data acquisition system; it is expected that data collection will begin by arly Nay. Part II II of the hydrogen combustion program includes a total of eighteen tests which are intended to study spherical deflagrations of a hydrogen flame. The hydrogen concentrations that will be investigated range from 105 to 42K whereas the steam concentrations will vary from 0 to 30Ã. With the exception of two tes ts in which i gni ti on will be ini ti ated at the bottom of the test vessel all tests will be performed using center ignition. The'time required to complete these tests is approximately one month. II Subsequent to these tests the test vessel will be modified for the study of turbulent effects on hydrogen combustion. Two weeks have been scheduled in the program plan to accomplish these modifications.
The primary objective of the Part III tests is to investigate turbulent effects upon completeness of hydrogen burns, and upon pressure and temperature responses. Turbulence in these tests will be created by two different means: 1) two 16" diameter vaniable speed fans and, 2) gratings. The fans are rated at 1500 cfm each and consequently are capable of creating a very turbulent environment. The gratings are made of 1/4" perforated plate with 50% porosity and they are used to simulate obstacle-induced turbulence. Six tests will be devoted to examining lean hydrogen combustion under turbulent conditions; ignition will be initiated at the bottom of the vessel. Four additional tests will be conducted using 14% and 20/ hydrogen-air mixtures when the ignition source will be placed at the center of the test vessel. The time needed to complete these tests is expected to be about one month. Part IV of the hydrogen combustion program entails a total of six 7 tests. Prior to performing these tests, a week's time is needed to set up the vest rig which includes a sphere used in the previous tests. Ignition for these tests will be initiated at either the center of the sphere or at the end of the pipe f'r hydrogen mixtures of either 8% or 20%. In addition to collecting the temperature and pressure data, ionization probes will be used to record flame propagation from one compartment to,another. The final two tests using this test geometry include studying hydrogen combustion characteristics from a 8% or a 10% mixture to a 6% mixture. In these tests the pipe will be I filled with a 8/ or'10% mixture, while the sphere is filled with a 6/ mixture. Ignition will be initiated in the pipe section. The duration of these tests is anticipated to be about three weeks.
The second experimental program that will be carried through at the llhiteshell facility involves research and development effort on various igniter types. The objective of this work is to perform extensive benchmark tests in a six cubic foot spherical test vessel to identify igniter types and to demonstrate their combustion capability in a prototypic environment. The testing program will begin in May and last about four months. Based on test data obtained, a'selection of igniters will then be further tested in a larger scale test vessel (600 ft3 ) at Acurex. Presently, besides the GMAC 7G glow plugs, a few resistance-heating glow plugs developed by Tayco will also be examined.
,b) Acurex In the Acurex program, the test plan can also be divided into two parts; the first part is designed to examine the effectiveness and the performance of. glow plugs in igniting hydrogen under various prototypic contain-ment conditions . In these experiments, hydrogen flow rate, steam flow rate, water sprays parameters and ignitor locations will'e varied to provide parametric studies on the ability of glow plugs'o ignite hydrogen mixtures .
The effect of micro-fog on glow plug ignition and pressure transients will also be investigated. A number of the experiments will attempt to provide data to correlate fogging as a pressure suppressant with spray volume, spray drop size, and hydrogen concentrations. A strong ignition source, e.g., electric match, will be used in all the fogging-related tests. A second part of the test plan calls for testing a selected number of igniters developed at the Whiteshell Nuclear Research Establishment. These will be large scale confirmatory tests for ignition devices which have demonstrated a superior potential in igniting lean hydrogen mixtures and in
\
replacing the existing glow plug designs in the future. Their effectiveness
~ ~ ~ ~ ~ ~ ~
in ~ a
~
spray environment will be
~
evaluated at Acurex's 600 ft vessel. ~ Prior to carrying through the above described test 'plan, a series of shakedown tests will be performed to provide checks for consistency and accuracy of all instrumentation; specifically, results will be compared with those obtained at Mhiteshell and from the available literature. c) Hanford En ineerin Develo ment Laborator HEDL The objective of this effort is to experimentally investigate aspects of hydrogen mixing and distribution in a simulated ice condenser lower compartment geometry. Hydrogen release into the compartment will be modelled by two approaches. In the first approach, steam and hydrogen are introduced as a jet into the compartment simulating a pipe break; in the second approach, hydrogen and steam are added to the compartment as a diffuse source similar to pressurizer relief tank release. In order to extend the rang of hydrogen concentration beyond "4%%d, helium will be used as a simulation fluid in place of hydrogen. Confirmatory tests will be performed to demonstrate that helium can indeed be used to substitute hydrogen in these mixirig studies. The first test is scheduled to begin some time in mid June and the whole test program is expected to last approximately two months. In the meantime, similitude and scaling calculations are being done so as to properly model the necessary parameters that are vital to the investigation of mixing and distribution. Some of the non-dimensional groups that are being examined are: the Richardson number, the Reynolds number, and the Grashof number. d) Factor Mutual Research AEP, Duke, TVA and EPRI recently came to the conclusion that in order to better understand fogging as a means of hydrogen control and to eventually
0 render a decision on its applicability as a viable solution to hydrogen mitigation, they would contract with Factory Mutual Research to undertake a research program to investigate fogging. The objective of this program is to determine the effects of micro-fog upon the lower flammability limit (LFL) of hydrogen, to provide a relationship between dropsize and fogging density on LFL and to correlate the concentrations of lean hydrogen air mixtures with various fogging parameters. In order to ensure that the effects of fogging on LFL are properly reproduced, a strong ignition source has been proposed and is likely to be used to initiate ignition on all LFL tests. The range of droplet sizes that is of interest to the utilities varies from a few microns to hundreds of microns, whereas the fogging density varies from zero to a.few percent. Test parameters that will be measured
~
include temperature, pressure, dropsize distribution.and fog.density distribution.
~ ~ ~
schematic of. the experimental
~ ~ ~ A set up is shown in Figure 5-1.
A detail test plan is being prepared by Factory Mutual Research with aid from AEP and the other participants. The test vessel is scheduled to become available for test in approximately three weeks. Finally, it is also the intent of this effort to provide the necessary and pertinent information 'o assist in the selection of test parameters in the Acurex fogging tests. e) CLASIX In the AEP-NRC meeting on March 18, 1981, the staff expressed interest in reviewing a -number of additional CLASIX runs . The first concern centers around the unique lower containment spray capability at Cook and its possible effect upon other compartment responses during and subsequent to a hydrogen burn.
0 Reviews at AEP indicate that in the CLASIX sensitivity study submitted to the NRC, spray parameters such as spray flow rate, droplet size, heat transfer characteristi'cs to the drop and spray temperature were varied; minimal effects on the containment pressure and temperature responses were noted. Thus, the available information from CLASIX, points out that variations in spray parameters would not significantly affect containment temperature and pressure response. Another possible CLASIX run discussed in the above mentioned meeting involved initiating hydrogen combustion at 10% with 50% burn fraction. Experimental measurements on completeness of hy'drogen combustion reported in the literature show that in spite of the large scattering in data around 5% to 7%, an initial 10/ concentration consistently results in an almost 100% (1) burn. In addition, it has been shown that turbulence will further enhance completeness of combustion for lean hydrogen mixtures . Therefore, if the probability of incomplete combustion of 10% is indeed negligibly small, as it seems to be, its effects upon the containment need not be investigated. It was suggested by the staff that a case with ignition initiated at 10% and then propagating to a 8% hydrogen concentration region should be studied. Both types of combustion would assume a 100% burn fraction. Close examination of the various cases presented in the CLASIX sensitivity studies reveals that there is one case (JVD15) which uses the exact input parameters requested by the staff. One burn was observed in the upper compartment with an estimated maximum pressure of 57 psia (only one air recirculation fan was assumed to be operational in the run). This maximum pressure is very close to the Cook containment elastic limit. However, since heat sinks have not been included in these
~
sensitivity calculations, the results are likely .to be overly conservative. ~
Floivmeter Flowmeter Mixer Air pressure Regulator Flash Arrostor Water Air Solo n old Op o r at ed Valve
/
Solon old Oporatod Valve H>- Air Mix Supply Line
/I X Fog Nozzles / II / / Spark Gap ~ Eloctrodes loniza tl on Prob os I ih r'(I ii /il 4 Thermocoeploe, ~
For bRopSfhE: hl~sug~NQ7) peal~ Flamo Spood / I lAoasuromonts I I i 6" Dlcmeter x ~ Longth I I I Drain FIGURE 5-1 EXPERIMENTAL ARRANGEl1ENT OF FOGGING TESTS
References:
(1) Liu D. D. S., et al, "Some Results of WNRE Exper iments on, Hydrogen Combustion," Water Reactor Safety Workshop on the Impact of Hydrogen, Albuquerque, New Mexico, January 1981. (2) Hertzbert, M., "Flammability Limits'nd Pressure Development in H2-Air Mixtures," U. S. Bureau of Mines, PRC Report No. 4305, January 1981.
DONALD C. COOK NUCLEAR PLANT UNIT NOS. 1 AND 2 ATTACNENT NO. 5 TO AEP:NRC:00500A SECOND QUARTERLY REPORT ON HYDROGEN MITIGATION AND CONTROL
t~ 5.0 Current Research Pro rams Several research programs have been undertaken by AEP to investigate hydrogen control related phenomena; some of these programs were di.scussed in the last quarterly report. In this section a number of the current research programs will be reviewed; program status, revised test plan and program schedule of each effort will be discussed individually. 1.1 ~E AEP, along with Duke and TVA,are co-sponsors of four EPRI research programs in which fundamental flame studies will be made; research and development on various igniter types will be pursued; mixing and distribution of hydrogen in prototypic containment environments will be investigated and additional glow plug testing will be performed. a) Whiteshell Nuclear Research Establishment This research 1 facility is operated by Atomic Energy of Canada Limited. Two research programs will be pursued independently at this facility; namely, the 'hydrogen combustion phenomena study and the research and development of different igniter types. Both of these programs will be undertaken with the collaboration of Ontario Hydro as an additional financial contributor to the work. The first experimental program is designed to investigate various hydrogen combustion phenomena and can be divided into four parts. The first part of this experimental effort entails perfororing nineteen ignition tests on lean hydrogen mixtures. The hydrogen concentration to be examined in'these tests will vary from 5.0$ to 305 by volume. A spark ignition source which is
0 in the order of 0.5 joule will be used to ignite the mixture. Details of the experimental set up and test vessel dimensions have been presented in the previous quarterly submittal. Fast response pressure transducers, thermo-couples and ionization probes will be employed to monitor and record various important test parameters. Of the nineteen tests planned the majority of them will be conducted wi th the ignition spark located near the bottom of the spherical test vessel. Two tests are planned in which the ignition source will be located at the center of the vessel and one test is planned with the i gnition source near the top of the vessel. These three tests will be used to assess the effect of igniter location. These tests are anticipated to require approximately three weeks to complete. According to the latest estimate provided by WNRE, system shakedown is being performed on the test vessel and on the data acquisition system; it is expected that data collection will begin by early Nay. Part II of the hydrogen combustion program includes a total of eighteen tests which are intended to study spherical deflagrations of a hydrogen flame. The hydrogen concentrations that will be investigated range from 105 to 421 whereas the steam concentrations will vary from 0 to 30Ã. With the exception of two tests in which igni tion. will be initiated at the bottom of the test vessel all tests will be performed using center ignition. The time 'required to complete these tests is approximately one month. Subsequent to these tests the test vessel will be modified for the study of turbulent effects on hydrogen combustion. Two weeks have been scheduled jn the program plan to accomplish these modifications.
The primary objective of the Part III tests is to investigate turbulent effects upon completeness of hydrogen burns, and upon pressure and temperature responses. Turbulence in these tests will be created by two different means: 1) two 16" diameter variable speed fans and, 2) gratings. The fans are rated at 1500 cfm each and consequently are capable of creating a very turbulent environment. The gratings are made of 1/4" perforated plate with 50/ porosity and they are used to simulate obstacle-induced turbulence. Six tests will be devoted to examining lean hydrogen combustion under turbulent conditions; igni tion will be initiated at the bottom of the vessel. Four additional tests will be conducted using 14/ and 20/ hydrogen-air mixtures when the ignition source will be placed at the center of the test vessel. The time needed to complete these tests is expected to be about one month. Part IV of the hydrogen combustion program entails a total of six tests. Prior to performing these tests, a week's time is needed to set up the test rig which includes a sphere used in the previous tests. Ignition for these tests will be i nitiated at either the center of the sphere or at the end of the pipe for hydrogen mixtures of either 8/ or 205. In addition to collecting the temperature and pressure data, ionization probes wi ll be used to record flame propagation from one compartment to another. The final two tests using this test geometry include studying hydrogen combustion characteristi cs from a 8Ã or a 10Ã mixture to a 6X mixture. In these tests the pipe will be filled with a 8Ã or 105 mixture, while the sphere is filled with a 6X mixture. Ignition will be initiated in the pipe section. The duration of these tests is anticipated to be about three weeks'
~ ~ The second experimental program that wi 11 be carried .through at the Whiteshell facility involves research and development effort on various igniter types. The. objective of this work is to perform extensive benchmark tests in a six cubic foot spherical test vessel to identify igniter types and to demonstrate their combustion capability in a prototypi c environment. The testing program will begi n in May and last about four months. Based on test data obtained, a selection of igni ters will then be further tested in a larger scale test vessel (600 ft3 ) at Acurex. Presently, besides the GMAC 7G glow plugs, a few resistance-heating glow plugs developed by Tayco will also be examined. b) Acurex In the Acurex program, the test plan can also be divided into two parts; the first part is designed to examine the effectiveness and the performance of glow plugs in igniting hydrogen under various prototypic contain-ment conditions . In these experiments, hydrogen flow rate, steam flow rate, water sprays parameters and ignitor locations will be varied to provide parametric studies on the ability of glow plugs to i gnite hydr'ogen mixtures . The effect of mi cro-fog on glow plug ignition and pressure transients wi 11 also be investigated. A number, of the experiments will attempt to provide data to correlate fogging as a pressure suppressant wi.th spray volume, spray drop size, and hydrogen concentrations. A strong ignition source, e.g., electric match, will be used in all the fogging-related tests. A second part of the test plan calls for testing a selected number of igni ters developed at the Whitqshell Nuclear Research Establishment. These will be large scale confirmatory tests for ignition devices which have demonstrated a'uperior potential in igniti ng lean hydrogen mixtures and in
h replacing the existing glow plug designs in the future. Their effectiveness in a spray environment will be. evaluated at Acurex's 600 ft vessel. Prior to carrying through the'above described test plan, a series of shakedown tests will be performed to provide checks for consistency and accuracy of all i nstrumentation; specifically, results will be compared with those obtained at Whiteshell and from the available literature. c) Hanford En ineerin Develo ment Laborator HEDL The objective of this effort is to experimentally investigate aspects of hydrogen mixing and distribution in a simulated ice condenser lower compartment geometry. Hydrogen release into the compartment will be modelled by two approaches. In the first approach, steam and hydrogen are introduced as a jet into the compartment simulating a pipe break; in the second approach, hydrogen and steam are added to the compartment as a diffuse source simi lar to pressurizer relief tank release. In order to extend the range of hydrogen concentration beyond 45, helium will be used as a simulation fluid in place of hydrogen. Confirmatory tests will be performed to demonstrate that helium can indeed be used to substitute hydrogen in these mixing studies. The first test is scheduled to begin some time in mid June and the whole test program is expected to last approximately two months. In the meantime, similitude and scaling calculations are being done so as to properly model the necessary parameters that are vital to the investigation of mixing and distribution. Some of the non-dimensional groups that are being examined are: the Richardson number, the Reynolds number, and the Grashof number. d) Factor Mutual Research AEP, Duke, TVA and EPRI recently came to the conclusion that in order to better understand fogging as a means of hydrogen control and to eventually
render a decision on its applicability as a viable solution,to hydrogen mitigation, they would contract with Factory Mutual Research to undertake a research program to investigate fogging. The objective of this program is to determine the effects of micro-fog upon the lower flammability limit (LFL) of hydrogen, to provide a relationship between dropsize and fogging density on LFL and to correlate the concentrations of lean hydrogen air mixtures with various fogging parameters. In order to ensure that the effects of fogging on LFL are properly reproduced, a strong ignition source has been proposed and is likely to be used to initiate ignition on all LFL tests. The range of droplet sizes that is of interest to the utilities varies from a few microns to hundreds of microns, whereas the fogging density varies from zero to a few percent. Test parameters that will be measured include temperature, pressure, dropsize distribution and fog density distribution. A sch'ematic of the experimental set up is shown in Figure 5-1. A detail test plan is being prepared by Factory Mutual Research with aid from AEP and the other participants. The test vessel is scheduled to become available for test in approximately three weeks . Finally, it is also the intent of this effort to provide the necessary and pertinent information to assist in the se1ection of test parameters in the Acurex fogging tests. e) CLASIX In the AEP-HRC meeting on March 18, 1981, the staff expressed interest in reviewing a .number of additional CLASIX runs. The first. concern centers around the unique lower containment spray capability at Cook and its possible ffect upon other compartment responses during and subsequent to a hydrogen burn.
Reviews at AEP indicate that in the CLASIX sensiti vity study submitted to the NRC, spray parameters such as spray flow rate, droplet si ze, heat transfer characteristics to the drop and spray temperature were varied; minimal effects on, the containment pressure and temperature responses were noted. Thus, the available information from CLASIX, points out that variations in spray parameters would not significantly affect containment temperature and pressure response. Another possible CLASIX run discussed in the above mentioned meeting in'volved initiating hydrogen combustion at 10/ with 50% burn fraction. Experimental measurements on completeness of hydrogen combustion reported in the literature show that in spite of the Ilarge scattering in data around 5X to 7X, an initial 10Ã concentration consistently results in an almost 100'5 (1) burn. In addition, it has been shown that turbulence will further enhance completeness of combustion for lean hydrogen mixtures . Therefore, if the probability of incomplete combustion of 105 is indeed negligibly small, as it seems to be, its effects upon the containment need not be investigated. It was suggested by the staff that a case with ignition initiated at 10/ and then propagating to a 8/ hydrogen concentration region should be studied. Both types of combustion would assume a 100/ burn fraction. Close examination of the various cases presented in the CLASIX sensitivity studies reveals that there is one case (JVD15) which uses the exact input parameters requested by the staff. One burn was observed in the upper compartment with an estimated maximum pressure of 57 psia (only one air recirculation fan was assumed to be operational in the run). This maximum pressure is very close to the Cook containment elastic limit. However, since heat sinks have not been included in these sensitivity calculations, the results are likely to be overly conservative.
ll ~ y t
Floemeter Flowm et or Mixer Alr Pror."o uro Regulator Flash Arrostor V/ator Air Solenoid Operated Valve
/
Solenoid Oporatod Valve H Air Mix Supply Line 2
/gi i Fog Nozzles / II Spark Gap ~ Eioctrod os /i iia loni za tl on Pro b e s /(I ii /' ihermocouplos, ~
l For Flamo Spood bROPS!~ M EAsuk lNQ PEVlcC I Moasurom ants I I O" nromolor x~ Lnnnth I I I Drain FIGURE 5-1 EXPERIMENTAL ARRANGEMENT OF FOGGING TESTS
References:
(1) Liu D. D. S., et al, "Some Results of WNRE Experiments on
- Hydrogen Combustion," Water Reactor Safety Workshop on the Impact of Hydrogen, Albuquerque, New Mexico, January 1.981.
(2) Hertzbert, M., "Flammability Limits and Pressure Development in H2-Air Mixtures," U. S. Bureau of Mines, PRC Report No. 4305,. January 1981.
DONALD C. COOK NUCLEAR PLANT UNIT NOS. 1 AND 2 ATTACHMENT NO. 6 TO AEP:NRC:00500A SECOND QUARTERLY REPORT ON HYDROGEN MITIGATION AND CONTROL
I 6.0 Core Coolin Ca abilit Subse uent to H dro en Combustion 6.1 'ntroduction The write-up below addresses the existing components necessary to achieve and maintain a safe shutdown condition subsequent to a reactor trip and to maintain a safe shutdown condition and contain-ment integrity via adequate hydrogen control during and after a hypothetical degraded core cooling event. I 6.2 Safe Shutdown The three primary functions to be performed in order to achieve and maintain a safe shutdown condition subsequent to a reactor trip are: (1) circulation of reactor coolant (2) residual heat removal (3) control of RCS pressure
'he methods 6y which. each. o$ these functions can be'erformed,.-,
and the necessary equipment located inside containment, are discussed below. 6.2.1 - Circulation of Reactor Coolant Circulation of reactor coolant is provided by natural circulation with the reactor core serving as the heat source and the steam generators serving as the heat sink. Water is provided to the steam generators via the safety-grade Auxiliary Feedwater System (AFS) or, if offsite power is available and sufficient steam is available, via the normal feedwater system.
. The AFS can be aligned to take suction from the Essential Service Water Syst'm, which i itself,takes suction from Lake
Michigan, thus assuring a virtually limitless supply of cooling water for the steam generators. Steam release paths include turbine bypass (if offsite power is available) using the main condenser, the main steam safety valves, and the main steam power operated relief valves. Those portions of the reactor coolant system, main feed-water system, auxiliary feedwater system, and main steam system inside containment contain no active components required to operate to assure coolant circulation and operation of said systems wo'ul'.d not be .adversel'y -affected'y;"a hydrogen combustion environment. The equipment located inside containment needed to . assure adequate reactor coolant circulation is listed below. The susceptibility of this equipment to a hydrogen combustion environment and the effects of such an environment on equipment operation are addressed in Attachment Nos. 3 and 4 of this submittal, respectively.
l. Steam Generator Narrow-Range Level Monitors

2. Pressurizer Water Level Monitors

3. Pressurizer Pressure Monitors

4. Loop RTDs

5. Core Exit Thermocouples

6. RCS Wide Range Pressure Monitors
1 6.2.2 Residual Heat Removal Residual heat is removed via the steam generators utilizing the methods and equipment described in 6.2.1 above. For the same reasons set forth in 6.2.1, this function is not adversely affected by a hydrogen combustion environment. 6.2.3 RCS Pressure Control Subsequent to a reactor trip, RCS pressure is maintained utilizing the '.natural circulation'quipment described above, with the pressurizer (PZR) safety valves serving as high pressure protection. The PZR safety valves are self contained, spring loaded valves and would not be adversely affected by a hydrogen combustion environment. A second aspect of RCS pressure maintenance deals with isolation of the various branch lines attached .to the RCS. Each of these potential leakage paths, including the method of isolation, is discussed below. (1) Pressurizer Power 0 crated Relief Valves PORVs Each PORV is normally closed and i.s designed to fail closed upon loss of air or loss of power. In addition, a block valve is located upstream of each PORV to assure RCS isolation in the event that PORV leakage were to develop.
(2) Letdown Line Letdown isolation is provided by three parallel fail-closed air operated valves located inside contain-ment and a fail-'closed air operated valve outside containment. These valves will automatically close on a sa fety in j ecti on si gnal . (3) Excess Letdown/Seal Mater In 'ection Flow from the excess letdown heat exchanger is directed. to the reactor coolant pump seal water return line (connection inside containment) which is isolated by two motor operated valves in series, one inside reactor containment and one outside containment. These valves will automatically close on a safety injection signal. (4) Residual Heat Removal RHR Letdown'he RHR letdown line is isolated by two normally closed motor operated valves in series located inside reactor containment. Both valves are interlocked with RCS wide-range pressure to automatically close on increasing pressure above 600 psig and cannot be opened until RCS pressure has decreased below 426 psig. In addition, the valve control switches are administratively key locked closed'in the main control room during power operation.
(5) Reactor Vessel Head Vent The reactor vessel head vent system consists of two
-.redundant parallel paths, each path containing two normally closed, solenoid actuated valves in series for isolation. These valves are designed to fail closed upon loss of power.
6.3 H dro en Control E ui ment Operation of the containment air recirculation/hydrogen
.:skimmer (CAR/HYS) fans and the DIS in conjunction with the containment spray system (CTS) further assures the combustion of'lean hydrogen mixtures without posing a threat to the containment structure via overpressurization. The portion of the CTS inside containment .xontains no active components and hence CTS operation is not adversely
affected by a hydrogen combustion environment. The active components inside containment used for hydrogen control are the CAR/HYS fans
-,and the DIS. The electrical hydrogen recombiners would be used to 'remove residual hydrogen (less than 4 volume percent) from the .containment subsequent to DIS operation.
6. 4 ECCS Injection Subse uent to Combustion

An evaluation has been made to verify ECCS injection capability subsequent to hydrogen combustion inside containment. The results o'f this evaluation indicated that high-head safety injection (SI)
(charging pumps) flow path via the BIT and the intermediate/low head
'SI (SI and RHR pumps) flow path to the RCS cold legs will be unaffected
by hydrogen combustion. These flow paths contain motor operated valves inside containment. These valves receive a signal to open on a SI signal despite the fact they are normally in the open position, thus providing further assurance of ECCS injection capability. No mechanism has been identified whereby the environment associated with hydrogen combustion would result in closure of these valves. With the refueling water storage tank (RWST) available, twelve weight percent boric acid can be delivered to the RCS by aligning the suction of the charging pumps to the RWST and aligning the pump(s) discharge to the boron injection tank (BIT). A second flow path involves alignment of the charging pump suction to the discharge of the boric acid transfer pumps, which are themselves aligned to take suction from the boric acid tanks with the discharge of the charging pumps again aligned to the BIT. Neither of the above described flow paths utilize components (eg. valves) inside containment which are required to change position/function in a hydrogen burn environment. In the event that the contents of the RWST had already been injected coolant injection is achieved by aligning the charging pump(s) suction to the discharge of the residual heat removal (RHR) pump(s); with the RHR pump(s) taking suction from the containment recirculation sump. This third flow path does not utilize any active components inside containment which are susceptible to a hydrogen combustion environment.
The subject valves are fully qualified for post-accident
.use inside containment (LOCA/MSLB qualification). In addition, ;the analyses described in Attachment No. 4 to this report clearly show that the environmental conditions associated with hydrogen .combustion are less severe than the environment to.which they have
been qualified; thus assuring maintenance of the aforementioned
.flow paths. The normally closed motor operated valves in the intermediate/low head SI flow path have also been qualified for use -in a LOCA/NSLB environment and would be expected to remain in operation
subsequent to hydrogen combustion; thus providing another ECCS injection path.
DONALD C. COOK NUCLEAR PLANT UNIT NOS. 1 AND 2 ATTACHMENT NO. 7 TO AEP:NRC:00500A SECOND QUARTERLY REPORT ON HYDROGEN MITIGATION AND CONTROL
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Preliminar Safet Evaluation Indiana 5 Michigan Electric Co. ( IQ1ECo.) has decided to install a Distributed Ignition System (DIS) in the Donald C. Cook Nuclear Plant Unit Nos. 1 and 2. The DIS in conjunction with operation of existing safety-related equipment provides additional hydrogen control capability in the extremely unlikely event of a degraded core event similar in nature to the TMI-2 accident involving the generation of substantive amounts of hydrogen. The DIS, described in detail in Attachment No. 2 of this report, is designed to assure combustion of lean hydrogen/air/steam mixtures and hence will minimize the pressure and temperature transients associated with hydrogen combustion. Conservative analyses of the containment response have previously been submitted via our first quarterly report (AEP:NRC:00500). The results of these analyses indicate that deliberate ignition of lean hydrogen mixtures using the DIS will result in pressures below the ultimate strength of the Cook Plant containments. The effects of a hydrogen combustion environment on necessary equipment located inside containment has been evaluated and the results of this evaluation presented in Attachment No. 4 of this report. It is clear from our evaluation that the temperature effects oF deliberate hydrogen combustion are less severe than those to which most of the necessary equipment has been qualified (LOCA/MSLB qualification). It has also been shown that the ability to inject emergency core cooling water is not affected by hydrogen combustion.
The extensive plant modifications and enhanced operator training implemented subsequently to the TNI-2 accident have effectively reduced the already low probability of occurrence of events which could result in the generation of substantive amounts of hydrogen at the Cook Plant'. The DIS, in conjunction with existing plant equipment>will provide an additional level of mitigation capability for hypothetical events well beyond the design basis of the Cook Units, further enhancing the defense-in-depth .philosophy. Installation of the DIS provides further assurance that operation of the Cook Plant will in no way adversely effect the health and safety of the general public.
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ML17326A883: Difference between revisions

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ML17326A883: Difference between revisions

Latest revision as of 02:10, 24 February 2020

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1.1 Purpose and Scope

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