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CAROLINA POWER AND LIGHT COMPANY l .- BRUNSWICK STD31 ELECTRIC PIANT UNITS 1 AND 2 I MSIGN RDORT 50, 7
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CONIAlle(Drf MSIGN REPORT
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December 31, 1970 i i e e F' 8709280444 870921 PDR FOIA MENZG7-111 P,DR - - - - - - - - - g
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1 BSEP-1 & 3 TABLE OF CONTTNTS Seetton Tit 1e Page 1.D INTRODUCTION 1-1 2.0 CENERAL DESCRIPTION OF BRUNSWICK PRESSURE 2-1 SUPPRESSION CONTAINMENT 3.0 DESIGN AND ANALYSIS OF CONTAINMENT 3-1 3.1 Program Abstract l 3.2 3-1 Method of Analysis 3.3 Program Technical Description 3-2 3.3.1 References 3-3 3.4 3-9 l Description of Program Generated Loads .i 3.4.1 Body Forces 3-l' ) 3.4.2 3-11 Pressure Surface Tractions 3-11 l 3.4.3 Thermal Loads 3.5 3-11 Description of Program output 3-12 1 4.0 DRYWEl.L a 4-1 4.1 Physical Description of Dryvell 4.2 4-1 Idealization of the Structure 4-3 4.3 Design and Analysis 4.4 4-6 Procedure for Determining Drywell Seisamic Loads 4-8 s 4.5 Material Constants
, 4.6 Concrete Cracking Analysis 4-11 j 4.7 Drywell Analysis Sequence 4-13 i
, 4.8 Pedestal Hoop and Radial Reinforcing 4-14 4-15
;j 4.9 4.10 Drywell Wall Meridional and Noop iteinforcing 4-17 4.11 Investigation of Liner Yield and Buckling 4-18
- Investigation of Concrete and Reinforcing 4 19
! 4.12 Stresses at Full Yield Condition of Liner Concrete Radial Shear 4.13 Analysis of Drywell Head 4-20 4.14 Load and Stress Plots 4-21 I i 4.14.1 Typical Force Resultant Calculation 4-21 i 4.14.2 4-23 4.14.3 Force Resultant Plots - Discussion and Summary 4-24
{ Stress Plots - Diseussion and Summary i 4-27 5.0 SUPPRESSION CHAMBER 5-1 5.1 Physical Description of Suppression Chamber 5.2 5-1 Iden11: asion of the Suppression Chamber 5.3 Design and Annlysis 5-2 5.4 Material Properties 5-1 5.5 Seismic Loading 5-5 5.6 5-5 Hoop snd Meridional Reinforcing 5.7 Liner Yielding and Buckling 5-6 5.M 5-R Suppression Chamber Load and Stress Plots 5-8 5.8.1 Torce Resultant Plots 5.8.2 Stress Plots 5-10 5 - 1.8 i 6.0
SUMMARY
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l BSEP-1 & 2 [ LIST OF FICURFS Figure No. Title 1 Dryvell Conditions following l_ ass-c.f-Coolant-Accident 2 Section Through Primary Containment Structure 3 Axisynenetric Modelling Elements 4 Finite Element Idealization of Axisyssetric Elements 5 Clobal and Stress coordinate Systeams j 6 Drywell Reinforcing and Details 7 Drywell Axisyuusetric Finite Element Model 8 Enlarged Finite Element Model of Drywell Head and Dome 9 Drywell Computer Cecnetry Plot 10 Head and Dome Computer Geometry Plot 11 Drywell Transformed Material Coordinate System l l 12 Drywell Thermal Cradients and Associated Pressures 13 Concrete Cracking Planes and Associated Coordinate Directions 14 Assumed Initial Drywell Crack Pattern l 1 15 First Drywell Cracking Iteration (with Liner) 16 second Drywell Cracking Iteration (with Liner) 17 First Drywell Cracking Iteration (without Liner) 18 Second Drywell Cracking Iteration (without Liner) 19 Tinal Drywell Crack Pattern 20 Drywell Deformation Plot (Radially Unrestrained Pedestal Base) 21 Drywell De forma t ion Plot (Radially Restrained Pedestal Base) 22 Vertical Displacement Pat tern at E1se of Pedestal 23 Drywell Local Yield Zone Stress Tabulation
, = ._ = u i _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ _ . . .
88EP.1 & 8 LIST OF FIGURES (Continued) Figure No. Tith 24 Drywell Shear Reinforcing Detail
.25 ,
Iecation of Drywell and Suppression Chamber Sections Used in Calculating and Plotting Force and Moment Resultants 26 Typical Cross-Section for Calculating Force and Homent Assultants 27 Drywell Load Plot 1.0D (Dead Load) 28 Drywell Load Plot 1.0P (Pressure) 29 Drywell Load Plot 1.07 1.00 (Temperature) 30 Drywell Load Plot 1.0T1.25 (Temperature) 31 Drywell Load Plot 1.0T1 .50 (Temperstm) 32 Drywe11 Imad Plot 1.0E (Operating Basis Earthquake) 33 Dryw ' ad Plot 1.0E' (Design Basis Earthquake) 34 Drywell Load Plot . 9D + 1.00P + 1.07g,00 + 1.00E' ;. 35 Drywell Load Plot . 9D + 1. 25P + 1. 07 1. 25 + 1. 25 E 36 Drywell lead Plot . 9D + 1. 50P + 1. 071. 50 37 Drywell Stress Plot .9D + 1.00P + 1.073 ,00 + 1.00E' 38 Drywells Stress Plot . 9D + 1.25P + 1.0T 1. 25 + 1. 25E 39 Drywe l l St res s Pl o t . 9 D + 1. 50P + 1. 0T1 . 50
'40 Drywell Stress Plot .9D + 1.15P (Pressure Test) 41 Suppression Chamber Reinforcing and Details 42 Suppression Chamber Finite Element Model 43 Suppression Chamber Computer Geometry Plot 44 Suppression Chamber Transformed Material Coordinate System j
( l (< t _m_.
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BSEP.1 & 2 LIST OF FIGURES (Continued 4 Figure No. Title 45 Suppression Chamber Thermal Gradients and A-iatef Pressures 46 Suppression Chamber Deformation Plot (Radially Unrestrained Base) 47 Suppression Chamber Deformation Plot (Radially Restamined Base) 48 Suppressica Chamber Imad Plot 1.0D (Dead Emed) 49 Suppressica Chamber Load Plot 1.0P. (Pressuure e 50 Suppressica Chamber Load Plot 1.071 .00 (Tempesraturet 51 Suppression Gamber lead Plot 1.071.25 (Temymaraturen 52 Suppressica Chamber Imad Plot 1.071.50 (Tempesraturet 53 Suppression Chamber Load Plot .9D + 1.00P + LOT 1.03 + 1.00E' 54 Suppression Chamber Load Plot .9D + 1.25P + LOT 1.25 + 1.25E 55
, Suppression Chamber Imad Plot .9D + 1.50P + L OT1 .50 56 Suppression Chamber Stress Plot .9D + 1.00r + 1.071 ,3 + 1.00E' 57 Suppression Chamber Stress Plot .9D + 1.25F + I.0Tg ,3 + 1.25E
-58 Suppression Chamber Stress Plot .9D + 1.50r + 1.0Tg,g (andially Unrestrained Base, Liner Included) 59 Suppression Chamber Stress Plot .9D + 1.50F + 1.0Tg,g (andially Restrained Base, Liner Included) 60 Suppression Chamber Stress Plot . 9 D + 1. 50P + 1.0T 3,33 (aadially Unrestrained Base Liner Not included)
}
, 61 Suppression Chamber Stress Plot .9D + 1.15P (Pressure !<st) e d~' '
'd ) _ - - , ,, . . _ _ . .ds.w-- w-- __ .~ Ek
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BSEP-1 & 8
1.0 INTRODUCTION
1 l l This report includes the detailed physical description, design, and
, ; analysis of the primary containment structure. The report is divided into i
l two major sections:
- 1. Containment Drywell '
! 2. Suppression Cham' Each section includes
;etailed physical description, design crite- i i
e ria, analysis procedures, discussion, and graphical presentation of analyti-cal results. ne design approach involves an iterative series of computer runs that strive to achieve mexiomas governing stress states in the structures for the established seximum loading criteria. Although the objective of the j design and analysis is to arrive at maximum stress states, the stresses and ' l associated cross sectional force and moment resultants are isesstigated. <,)
]
Force and moment equilibriuse checks at all points in the structure present an excellent checking procedure and an added insurance that the associated I , stresses are correct. ne design philosophy for this type of approach includes bounding the results over the maximum possible range of design assumptions. 31s eliminates the possibility of overlooking a situation which can govern the design. l Extreme conservatism is introduced into the design of the drywell anti suppression chamber in several areas. I 11 e _ ym , -
B8EP.1 & 2 An initial series of runs concentrates solely on establishing a maximum conservative concrete cracking situation in the dryve11. nis is 9 achieved by subjecting the drywell to the loading criteria associated with the loss-of-coolant accident condition. If tensile stress occurs in the concrete, it is assumed to be cracked. De cracking which occurs in the plane norme! to the tensile stress is simulated by setting the elastic modulus associated with that stress direction to zero. This approach is repeated in a series of computer runs until the struture reaches a state of equilibrium in which all uncracked concrete is in compression. %is final crack pattern is used as the basis for obtaining maximum reinforcing stresses. Since maximum con-crete cracking does not result in conservative radial shears, an independent series of ruca concentrates on this area. For the same combined loadig l 1; cases all concrete is assumed to be uncracked. Mis, in coe61 nation with a fully restrained pedestal base, results in a maximum radial stress condition in both the drywell well sad drywell pedestal, nese maximaan radial shears are used as the design shears. I We degree of fixity at the bottom of the drywe!! pedestal and the bottom of the suppression chamber is uncertain. This uncertainty is elimi-nated by bounding the maximum stress state at all points in the structure.
%is is achieved by performing two independent runs for each loading condi-tion. The maximum stress state resulting from these limiting assumptions is established 6a the governing design case:
- 1. Complete fixity at the bottom surfaces
- 2. Unrestrained radial growth of the bottom surfaces I'
I i
, 1-2 l .
,g ,i m a ef-- - _ 4M% U MWdO11 b " "N
-l BSEP.! & 2 Conservatism occurs in the combined thermal loading cases. The steel liners in both the drywell and the suppression chamber are initi-ally assumed to have stif fness properties and load carrying capabilities.
To a!!ow for the sutremely conservative case in which the entire liner would be in the compressive yield state, a series of runus are performed for the same thermal loading cases, assuming that the liner contributes no stiff-ness to the overall structure and has no load carrying capabilities. The theras! gradient through the reinforced concrete remaina esachanged. This insures that the reinforced concrete portion of the structure can maintain its structural integrity without any assistance from the Itser. Under these assumptions, stresses are limited to (W Fy) to assure overall structural in-tegrity. In the suppression chamber locally there exists an overstreas of 51 above (6 Fy) but those stresses remain less than the minimas specified yield stress. This is extremely conservative for the following re.asons:
- 1. T'e suppression chamber liner is in complete rens ion under acc ident loading as shown in the runs in which its stiffness is included.
- 2. The load equations, on which the design of the suppression chamber is. based, initially contain pressure components which are unrealist ically conservative. 11:e design pressure of 62.0 psi, tron which the factored pressures of 77.50 psi ,
l
.m d " l.t' P,51 f or t he 1.25P and 1. 50P load a qua t ions, respec t-l l
ivelv. are th rived, are identical to those used in t he drywe l l . 1 v l-3 i;
v BSEP 1&2 1 ' The suppression chasser can never see an incident pressure of 62.0 psi. An incident pressure of 28.0 psi is associated l t with the accident condition in the suppression chamber. ' Also, in the drywell load equations, use of the drywe11 peak temper-ature, rather than the liner tasperature ceincident with the peak design pres- !
, sure, results in a large marsim of conserystism, as shown in rigure 1.
1he pressure test condition of .9D + 1.15P is included as a test case i in the analysis of the drywell med suppression chamber. This case did not . control the design.
"he seismic toed contributions to the load equations are considered.
The seismic leeds are not determined as part of the analysis discussed in this report. Seismic forces and moments are furnished in an independent dynamic analysis of the containssent structure contained in RSEP DR-4, "Seissaic Analysis of the Primsry and Secondary Containments". As discussed in the following text, these forces and mts are distributed over the drywell and suppression chamber and superimposed on the results obtained from the axisyun-metric analysis. 5 I l-4 ;
BSEP-1 & 2
- 2. 0 CENERAL DESCRIPTION OF BRUNSWICK PRESSURE SUPPRES$10N CONTAINMENT
, The primary containment structure consists of the drywell and the 1
pressure suppression chamber. Both are reinforced concrete pressure vessels t l
, with leak-tight steel liners on all inside surfaces of their reinforced i , concrete shells. The drywell and suppression chan6er.are founded on a reinforced ' concrete est supported on a layer of dense sand and a limestone i
foundation. The primary containment structure and the reactor building share l a common foundation. i The drywell is cogosed of a series of vertical right cylinders and truncated cones with inside diameters varying approximately betweem
- I 36'-0" and 65'-0"; forming a configuration similar to the conventional e
- steel containment " light bulb" shape. The overall height from the top of the foundation est to the drywell head flange connection is approximately 111'-0". The reactor vessel and its shield structure are supported on a cylindrical concrete pedestat in the center of the drywell and will be '
restrained laterally by the drywell at the top of the shield structure. The pressure suppression chamber consists of a. steel lined rein-forced concrete shell that is composed of 16 inter-connected cylindrical sections. The liner has a circular cross-section with an internal diameter-
! of 29'-0", and a major diameter of approximately 109'-0".
The suppression chamber encircles the bottom of the drywell, and is physically independent of it under all axisymetric loading conditions l including dead load, pressure, and temperature loads. The structures are i l keyed to behave as a single combined unit during an anti-symmetric (seismic ) <l l 2-1 l Y t
BSEP-1 & 2
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1 loadins occurrence. . A sectional view of the primary containment is shoun in Figure 2. The physical material properties for the various components of the containment structure are:
- 1. Liner Steel ASTM-A516 CR.60 to A-300 Fy = 28 kai 9 7 = 333*F
- 2. Drywell Head Steel ASTN-A516 CR.70 to A-300 Fy = 28 kai 9 7 = 333*F
, 3. Reinforcing Steel
#14 and #18 bars ASTN-A615 - modified to CR.50 Fy = 50 kai
#6 to #11 bars A31N-A615 CR.60 Fy a 60 kai .
d5 and anstler bars , ASTM-A615 CR.40 Fy = 40 ka!
- 4. Concrete Minimum 28 day cgressire stress = 3000 pai .
Capacity reduction factors used are those stated in the PSAR:
- 1. W = .9 for Flexure and Tension
- 2. W = .85 for diagonal tension, bond, ar.chorage I
i l 2-2
...._ _ _ ~ - - -
_7_-__--_ . BSEP 1 & 2
- 3. 0 DESICal AND ANALYSIS OF CONTAI!9ENT The structural design and analytis of the pressure-suppression containment is based on a series of computer runs, using the finite element direct stiffness method for the analysis of axlayumetric solids of revolu-tion. The program used in the analysis was developed by Dr. E. L. Wilson of the 1kiiversity of California, Berkeley. The particular changes and additions to the general progs a, as required for this analysis, were per-formed directly by Dr. Wilson.
3.1 Program Abstract I - i The finite element method is applied to the determination of dis. placements and stresses within axisymmetric solids with orthotropic, temper. ature-dependent material properties. The continuous structure is replaced by a system of ring elements with TRIAIICt1LAR, Qt!ADR11ATERAL, or TitIJNCATED CONE (LINEAR) coastant thickness cross sections, as shown in Figure 3. Since the elements are of arbitrary shape and may'have different material properties, the procedure may be applied to structures composed of many different materi-als and of complex geometry. I The advantages of this approach are many. The method is completely 8 general with respect to geometric and material properties. Complex bodies composed of many different materials are easily represented. Displacement or force boundary conditions may be specified at any node. Arbitrary, thermal, wechanical, and accelerational loads are possible. Increased accuracy may be obtained by increasing the number of elements in the model.
.t,
- - -. ..- - . - {
BgEP.1 & 2
- 3. 2 Method of Analysis
%e continuous structure is replaced' by a system of axisymmetric elements or rings which are interconnected along circumferential joints called nodal circles. en whm n in Figure 4 Using energy principles and the direct stif fness approach, equailibr.
Lum equations for the complete structure in terms of the unknown nodat dis-placements are forusd. The program is limited to axisymmetric loading, in which the axial and radial displacements at the nodal circles are the un-knowns. Linear displacement functions are used to describe the displacement field. %is assures displacement compatibility between inter-element M artes since lines which are initially straight remain straight in their dis-placed position. Its curvature is allowed in the displacement equations. Therefore, when rotations may be of 'stanificance, the number of elements
- s. must be increased, since a curve any be approximated by a series of straight lines and accuracy increases as the the number of chords on the curve is inc reased, ne program can consider material orthotropy in the radial, tangential, and oxial coordinate directions, including angular transformation of the materi-o! properties to any desired local coordinate system in the planc described by the radial and vertical axes.
Body forces (gravity), surface tractions (pressure), temperature in-duced stresses, and externally applied forces, and any factored combination of these loads, can be considered.
)
3-2
.. i
U 88EF-1 & 2
- 3. 3 Program Technical Description The finite element method for the analysis of complex axiaymmetric solida and the general equations which govern the equilibrium of the system, defining the various terms involved, are derived on the following pages:
Equilibrium of couplete System The potential energy for an elastic solid is given b7 r V=U= wgFg dV -JwgFgdA
, 1 Area (g) i I
l I where U.is the total strain energy for the solid, and is gives by the lategral P 1 81 U* f f eg deg & Viol ,, O (2) In Eqs. (1) and (2), vg is the displacement field, Fg are the body forces, Fg are the surface tractions and e'g & eg are associated stresses and straire. For a finite element system which is composed of M elements, the potential energy can be written as a stannation over all the elements. F M m a a m V=[mal I/" - wg Fg dV- [ wi F1 dA (3) u Vol Area , The surface integral exists only 1' the mth e l emen t is oc. the boundary of the body and subjected to surface tractions Pt . In order to express the potential energy in terms of unknwn nodal circle displacements, it is necessary to assume a solution for the dispiscenent field within each ciceent. 3-3 u n - _.- - - _ - - A I l
, BS EP- 1 & 2 This assumed displacement field should satisfy compatibility between elements of the system. Based on this approximation, an expression for tho displacements within the element, (w) ,
in terms of a v ctor of all the nodal point dir, placements,[u] , is obtained, f (W)m = (d), [u] ~ (4 ) or. In transposed form, , [w),7, {,J T[g]gT (5) Also, an expression for element strains f*), is obtained, H. - N.N m J oes and, in transposed form, H!- [#fM! m For an elastic material, the stresses at any point within the element are expressed in terms of the corresponding strains and thernal effects by the elastic stress-strain ritlationship, [#), = ([C),H,). [rj. W The st rain energy,1:q. (2) can be written in matrix form as: l l b4 mm m -m .-
.d
BSEP.1 6 2 '
,. 1 Um - a d," dV = f fVol[ ),[C] [,], dV (9)
Vol u O i i ,
- Volf H
- H . *v i
The total potential energy, Eq. (3), can be written in matrix for, as: Y= [u] T[K) [u] - [u]T[Q] (10) , l ifhere [K] is the stiffness .strix for the complete finite
. e.t s,st :
M i l, [K]-se!I [u), cli > The individual ele.ent stiffness is: T i f*]. ~ N. M. N. dv (22) l 4 And the load vector [Q] is defined as: ( M M f0) * [ fL)m +m=1[ m=1 R], (13 ) g I The body force matrix is: T T fl m
- Vol m m
+
- m a dV (14 )
J l 2 3-5 l9 + _- _-g r,?wwr, , . _ -~_ .T~
- T-
f BSEP-1 & 2 l l And' the load vector due to surface forces is: T- [R), = f [d) , [P) , dA . (15 ) Area ,
, I-
'Ihe potential energy is made stationary by requiring that .
dv = 0 g i = 1 ..... , N (16) i* where I is the total number of unknown nodal point dis. placements. After consideration of the rules for matriz differentiation with respect to a vector, the application ^ of Eq. (10) yields the following relationship for the equilibrium of a finite element system: [K) [u] = [Q) (17) 1 Boundary Conditions I Eq. (17) represents the relationship between all nodal point forces and all nodal point displacements. Mixed boundary conditions are considered by rewriting Eq. (17) in the following partitioned form: a
' Q, K K ab u, ' l
. . . . a ----4-- . - -
.. b. ,I b
k6 l __ *J _3 Ar (A , dg a4aM)
7-
. BSEP- 1 & 2 where:
Q, = the specified nodal point forces I
; 'Q
, b.
= the unknown nodal point forces
'u, =
the unknown nodal point displacements u b
= the specified nodal point- displacements he first part of Eq. (18) can be written as a separate
; equation,
'Q, =
'K,, 'u, + 'K IIII
, , . . . . . . ab. ;
,"b.
i e and then expressed in the following reduced form Q "
'E ,,' 'u, (20) where the modified load vector is given by Q " - K Q,
, "b, (21)
. . . . ab.
{ l After Eq. (20) is solved for the nodal point displace-nents, the strains within any element in the systen are evalu. ated by the direct application of Eq. (16). De corresponding l l' stresses are calculated from the stress-strain relationship, Eq. (8). he elemental modeling stiffness astrices [k] refer to the library of three geometric shapes available in the program: 3rJ l 1 o 1 1 ___m_.m __ . . _ _ _ .. - - - - - - - -
/ -
\ e BSEP-1 & 2
- 1. TRIANCLE 2.
1 QL'ADRI LATERAL ap ,
- 3. TRUNCATED CONE (Membrane Element)
Properties and limitations of each element type: TRIANGLE Linear displacement function Plane strain assumption Constant temperature change. within element in all 3 orthogonal directions. QUADRILATERAL Internally stab-divides quadrilateral into four triangular elements with fifth common sode at the centroid. Above mentioned triangle assuaqptions apply to quadrilateral elements. TRUNCATED C0KE Tuo nodes Linear displacement function Plane stress assumption Constant thickness No joint rotation or bending stiffness No transverse stiffness Constant temperature rise in the axial and hoop directions only.
; The material coordinate system for triangular and quadrilateral elements is sesumed coincident with the global, R-Z systeu, but any be trans-formed on an individual element basis, or all elements of one specific material may be transformed.
i s t 3- 8 '
. )
*1 J
.f. 44 wash .
. . efM s 5db-e N $* - "M i
BSEP.1 & 2
- 3. 3.1 References
- 1. Clough, R. W. , "We Finite Element Method in Plane Stress Analysis",
Proceedings. 2nd ASCE Conf. on Electronic Computation, Pittsburgh,
& (September 1960).
l
- 2. Turner, M. J. , R. W. Clough, H. C. Martin, and L. J. Topp, " Stiff. l ness and Deflection Analysis of Complex Struccures," Journal of the Aeronautical Sciences, Vol. 23, he. 9, pp. 805 823 (September 1956).
1
- 3. I Melosh, R. J. , "A 5tiffness Matrix for the Analysis of Din Plates in tending," Journ't of the Aeronautical Sciences, Vol. 28, pp.34. l 42 (1961),
- 4. - Clough, R. W. and Y. Rashid, " Finite Elemen. Analysis of Axisyuw metric Solide," Journal of the Engineering Mechanics Division ASct, pp. 71 85 (February 1965).
l
- 5. )
Rashid, Y. R. , " Analysis of Axisyuumetric Composite Structures by the Finite Element Method," Muclear Ensinaering and Design, Vol. 3, pp. 163 182 (1966).
- 6. Wilson, E. l , " Structural Analysis of Axisynsnetric Solids," AIAA Journal, Vol. 3, pp. 2269 2274 (1965). ;
j 1 7 Crafton, P. E. and D. R. Strome. " Analysis of Axisyuumetric Shells by the Direct Stiffness Method," AIAA Journal, 9o1, 1 No. 10, pp. 2342 2347 (October 1963). 8 Percy, J. H. , T. H. N. Pian, S. Klein, and D. R. Marvaratna,
" Application of the Matrix Displace,nent Method to Linear Elastic Analysis of Shells of Revolution," A1AA Journal, Vol. 3, No. II, pp. 2138 2145 (November 1965).
9 Melosh, R. J. , " Structural Analysis of Solids " Joisrnal of the Structural Division ASCE. pp. 205-224 (August 1963). 10 Argyris, J. H. , " Continua and Discontinue," P,roceedings of the Cont ., on Matrix Methods in Struc tural Mechanics, Dayton, Ohio, Oc t ooer 1965 AFTDI-TR.66.no , pp. 11- 189 (November 1966),
- 11. Herrmann. l.. R. , "Elastle *iorsional Analys is of irregular Shapes."
Jmrn.11 of the Engineering Mechantes Di v i s i on A SCT , Vol. 6, pp. 11 IV (December 1965) .
- 12. 21enkin ir. , O. C. and Y. V. Cheung, "Fini t e Elemen t s in the Solution of Field Problems," The Engineer. Vol. 220 (196)).
._ L9
. O l
BSEP- 1 4, 2 13. i Wilson, E. L. and R. E. Nickell, " Application of the Finite Element Method to Heat Conduction Analysis," Nuclear Engineering and Design, Vol. 4, pp. 276-286 (1966). 14 Wilson, E. L. , " Finite Element Analysis of hso-Dimensional Strue. tures," Doctor of Engineering Dissertation, Uniwrsity of California, Berkeley, California, June 1963.
- 15. . Sokolnikoff, I. s. , Mathematical Theory of Elasticity, McGraw-Hill Book Company, Inc. , pp. 385 386 (1956).
16. Testabull, H. W. and A. C. Aitken, An Introduction to the Theory of Conical Matrices, Dover Publications, Inc. , New York, pp. 173 1 74 (1961). I I 3 10
,___ . _m.__------ - - -
88EF-1 & 2 , l 3.4 Description of Prorram Genersted Loads l 3,4.1 Body Forces i Based on the furnished mass density (5) for each material type and an arbitrary Z acceleration, gravity loads are generated for a one l iadian segment of the structure and distributed uniformly at the nodat i points describing an element. 3.4.2 Pressure Surface Tractions A different uniform pressure can be specified as acting on any arbi-trary boundary surface of as element. Pressures are automatically lumped at the modal points based on the contributory area of the boundary surfaces on uhich the' pressure is applied. 3.4.3 ,Vigruelt.osde Thermal loading is initially established by specifying actual temper-atures at the nodal points. True temperatures are calculated as the difference betueen the spect-fled nodal temperatures and the reference temperature. The at used in calculating the initial thernst strains ( e 4 T ) is the average of the o T's at the nodal points ' describing an element. QUA DR i t>. TERA 1. (4 NODES) AT - inWe j + Tnode 2 + Tnode 3 + TnWe t. - Tref erence 4.0 I 3-11
--- 4
BSEP-1 & 2 TRIANCIE (3 NDDES) Tnode 1 + Tnode 2 + Taode 3 of =
- Treference 30 TRUNCATED CONE (2 NDDES)
Toode 1 + Teode 2 AT =
-Troference 2.0 Uniform thermal expansion in the R, 9, and 2 directiona is assumed using the calculated average temperature.
The restrained temperature induced model forces are calculated and applied to the structure .in the following masser:
.t ..of
?* 1
't =[D] eof l Pt at [DhoYl A l .
- re
't = Thermal strain e = Coefficient of theriaal expaaston in the A, e, Z directions A *1 = Element temperature as defined above
't =
Thermal stress
=
Pt Nodal forces due to tesaperature [D] = ' Elastic constants f(E,e)
=
A Cross-sectional area in R, e, Z directions E = Modulus of elasticity e = Poisson's Ratio 3.5 Description of Pronram Output The co ordinate systems to which all program input and output are ( referenced are the global (R, 9, 2) system and the transformed (N, e, S) j 3-12 ' h
=_ ,
"U
BSEP-1 & 2 sys t ess. Displaeaments'in the global (R, e, Z) and Transformed (N, e, S) Coordinate Systems are printed for each nodal point in the model shown on Figure 5. Stresses are calculated and printed for each elemment in the A, e, Z 1 Clobal Coordinate System and the N, e, S Transformed Coordinate System, to l Witch the material properties are referenced. i Stresses arc calculated at the geometric centroid of each element. Shear stresses in both the R-Z I (Clobal) and N-S (Transformed) planes are printed for each element. Maximum and minimuss principal stress, and the angle defining the I plane on which these principal stresses act, are calculated and printed, based am Mohr's Circle. The printout of exact element stresses in ecocrete and reinforcing elements perm $ts direct interpretation and comparison of results with no inter-mediate or post-run calculations. Force resultants are obtained by inte-i stating the particular elemental stresses and associated areas about the neutral axis at any desired cross-sectional location in the snodel. 'f i f Actual and exact element stresses are calcula ted and priated since
+
the true nnaterial properties, thicknesses, and dimensions of the physical structure are used in the idealized model. l 3-13 i _.g_ _ _ 3- . - _ . _%. . M
SSEP-1 6 2 4.0 DRWELL 4.1 - Physical Description of Drywep The drywell consists of two vert h.n right cylinders joined by two ?- truncated conical sections with a solid cylindrical base pedestal. We top of the drywell is closed by a continuous steel done bolced to the top of the drywell. The pedestal, a 17'-0" thick solid cylindrical disk, contains top { and bottom reinforcing over the total plam-form area. The top.reinforcias grid consists of radially aued circimferentially spaced bars. The circumferential bars are continuous elesed circular hoops. The radial bars are terminated on the outside face of the pedestal by a 900 hook extending into the depth of the pedestal. Near the pedestal center line, where concrete tensile stresses are minimism or in crapression, the radial bars are terminated and lap-spliced with an orthogonal grid of reinforcing bars. Figure 6 shows the drywell reinforcing and associated details. ne bottom pedestal reinforcing contains two orthogonal layers of reinforcing bars uniformly spaced over the full area of the pedestal. The bot tom reinforcing grid is terutnated in the concrete compression tone near the outside perimeter of the ast. In .iddit son to the top and bot tom reinforcing . hand of closed hoop b.ars, evenly spiced through the f ull depth of the pedcar al . runs along the outside face. l 4-1 5 . _-,-mm__-w
3 g7-1 & 2 I l
' Pedestal shear reinforcing is provided through the depth in the l
} ;. l form of bent Z bars inclined at an angle of 450 as shown on Figure 6. 3
- 1 t
i 1he drywell wall reinforcing consists of circumferential closed I hoops on each face along its full height, and continuous meridional rein. t l forcing on each face. The wwer cylindrical section of the drywell contains two large circular openings approximately 10'-0" in diameter. These openir.gs are directly opposite each other (1800 apart). gight 7'.0" diameter vent openings are c'o ncentrically spaced at 450 in the lower contr.a1 section of the drywell. 1 The closed hoop reinforcing bars that would normally occur within i these openings are evenly grouped or banded above and below the openings. l The continuous main meridional reinforcing, which would be interrupted by l the presence of the openings, is grouped or bent around the openings to maintain the continuity of the reinforcing. l Supplemental U bars are placed around the upper and tower perisseters of the openings in the areas lef t void of the meridional bars. Additional closed rings are provided around the openings to accommodate stress concen- l tration effects. Where required, supplemental shear reinforcing and fill hars are provided. Typical reinforcing details around large openings will be presented in a fortheimiing BSEP Design Report. Shear reinforcing in provided in the drywell wall in the form of inclined and horizonts! hooked bars through the depth of the wall, as shown in Figure 6. A diagonal grid of reinforcing is provided at the outside face
't 4-2 !} '
f
)
J.WL _ , w - _kK =- -
. Y- SM - A W N '
BSEP-1 & 2 of 'the drywell wall to resist tangential (in plane) seismic shears. The splicing of bars is by the Cadweld process. Minimum stagger for Cadweld splices is shown on Figure 6. Where bars are no lunger required to carry the design loads, they are developed using a length equal to two tismes the embedment length required by ACI 318-63 f f stratsht, or 1.33 times the embedment length required by ACI 318-63 if booked. N steel done covering the top of the reinforced concrete drywell and the liner extension are furnished with contineous machined flanges. The dase is securely fastened to the liner with 84 synssetrically spaced pre-tensioned bolts. A continuous plate along the top of the drywell is provided with cadweld sleeves which are spliced to the meridional reinforcing in the upper cylinder portion of the drywell. 168 vertics' L u l!.n.tr plates around the drywell head support the liner extension.
, 4.2 Idealization of the structure i
The initial phase of the analysis invo1* *s the modeling or idealiza-
, tion of the structure into a camposite mesh of interconnected ring elements I
using quadrilaterals, triangles, and truncated cone elemments. The compoalte i i i model exactly duplicates the physical dimensions of the containment. Idealised i reinforcing elements within the structure are positioned to exactly duplicate their physical location in the structure. A cross-section of the axinynsectric finite e1<went model used in the analysis is shown on Figure 7. An enlarged model of the done and head area
*
- n hown on Fi gu r e 8. A table of equivalent plate thicknesses associsted with all idealized line r and reinforcing elements in also included on Figure 6.
I-4-3 .m n ha;z Em n. d W % % - -rwa ~ m W
58EP-1 & 2 e The coordinates of all nodal points are given in the diagranenstic
> sketc5 adjoining the actual models. Encircled numbers identify the element 1
numbering. The remaining numbers identify the nodal points. 1 t The liner and reinforcing (membrane elements' are shown as heavy lines. i,' The data describing the idealized model was numerically and visually checked. A ctanputer plot of the complete model (Figure 9), and an enlarged e
! plot of the dame and head area (Figure 10) insure that the element and node
! point data 'used in the analysis is correct.
' 1 Elements whose material axes are not coincident with the globst g I l and Z axes are rotated to correspond to their local material coordinate
+
(41rections in the structure. This coordinate transformation permits the l use of true reinforcing thicknesus, and elastic constants in areas of the structure where axial directions deviate from the global Z direction. Retr > forcing and concrete material properties in the cone areas of the structure are rotated to the N-S coordinate system, coincident with the axis of the structure as show t on Figure 11. Each quadrilateral, triangle, and membrane element can be given an independent set of material properties and an independent local coordinate system. This' allows complete flexibility in changing individual elemental cracking assumptions in the concrete and permits arbitrary change of re-inforcing without altering the modeling scheme. The liner is idealized as ! a continuoun isotropic steel plate on the inside surface of the model. Each ! .er of hoop and meridional reinforcing is ideslized as an 4 independent uncoupled orthotropic steel plate with a material stiffness (E) f- 44 a S L____ _ _ _ _ _ _ - - -- m
f-BSEP-1 & 2 only in its axial direction. Because each layer of reinforetag is inde-pendent, exact thicknesses representing the actual area of reinforcing per foot can be used. The equivalent thickness of each of these plates is adjusted according to the actual size and spacing of' the bars at that loca-tion and is calculated using the following relationship: By t,
=
A,/12 where:
= equivalent plate thickness t,
2 t. A, = reinforcing area in in /f of width An exception occurs in the bottom pedestal reinforcing ukich is physically placed as a rectangular met. These ort.Sf ---11y placed bars are transformed into radial and hoop components in the cylindrical coordinate system to maintain a consistent global coordinate systesa. The concrete portice of the structure is idealized as triangular and quadrilateral elements. Five elements through the cross-section describe the cylindrical and conical portions of the well. Five elements with linear displacement functions give a good representation of the bendias which occurs at the discontinuity areas. To account for the large openings and the b.anding of the reinforcing, the structure was deliberately modeled to determine their inflorece on the remaining areas of the structure. The concrete elements withis the openings retain the samme concrete properties as the drywell wall . The heep reinforcing in each particular zone of penetrations in the drywell well is handed above and below the openings. The banded reinforcing is idealized by increasing i < f l-4- 5 l
"4 'wds ' %&
- ,,9 . * *; *Q f .
sSYM&rifr) l
F 58EF-1 & 2 the equivalent thicknesses of the hoop reinforcing elements above and below these penetration zones. The idealized meridional reinforcing remains un-
; changed since it is physically bent around the openings and is not inter-i j rupted.
ft the steel dome, is idealized as a continuous isotropic steel plate with two layers of quadrilaterals representing its actual thickness. The gasket flanges and drywell cover plate are also modeled as ' t quadrilateral steel elements. the plate stiffeners are orthotropic steel triangles, and quadri-laterals. All plate stiffeners are given sero boop stiffness because they are not continuous rings. A single membrane element with no hoop stiffness represents the holddown bolts. Its equivalent thickness is deternained by equating the total cross-sectional area of the 84 bolts to the cross. i sectional area of the idealized ring element: ; Ab t = where: tb = thickness of bolt element 2#Eb used in model A b
= total cross-section area of the 84 bolts Ab
= radius of the bolt circle 4.3 Design and Analysis The total design and analysis of the containment structure involves 39 computer runs. The design philosophy involves an iterative approach to maximize reinforcing stresses for all combined load equations over a com-plete range of boundary conditions and concrete cracking. Based on this I
, i 4.s l . i
\
e 6 .h ,
^
h
. , . . eeke ? ,, T.\* 4* h
BSEP-1 & 2 approach the 39 computer runs are reclassified into series of runs, each series concentrating on either maximizing reinforcing or concrete stresses in a specific portion of the structure or a complete stress investigation in the total structure for all given loading cases. The loading criteria used as the basis for all runs involve three governing combined load equations containing factored pressure, thermal, dead load, and seisesic components.
- 1. Ug =
0.9 D + 1.50P + 1.0 T1 .5
- 2. U2 =
0.9 D + 1.25F + 1.0 T 1.25+ 1.25 E )
- 3. U3 =
0.9 D + 1.00P + 1.0 71 .00+ I'# E' f
* =
0.9 D + 1.15F (Pressure test condition) 7, Where: I U .= ultimate required load capacity of the structure ! D = dead load of structure including fixed equipment i
; P = loss of coolant accident pressure T1 .0 T1 .25 ,= normal gradients associated with each of the factored loss of coolant accident pressures T1 .5 , (1.0P,1.25P, and 1.5P)
E = Earthquake loading (ope. cing basis earthquake) E' = Et.rthquake loading (design basis earthquake) See detailed description of seismic loading below
- Included as test case; does not govern design. ,
I, l. 4-7 f I 4 ..
~
y , 4 M g b k *A ' . ?
BSEF-1 di 2 The pressures and temperatures associated with the above load equations are shown on Figure 12. 4,4 , Procedure for Determining Drywell Seismic Loads The combined load equations require consideration of earthquake I load components E and E' where: E a operating basis earthquake E' = design basis earthquake I A dynamic analysis of the containment structure for the operating and design basis earthquakes produces modal frequencies, deflecticas, shears and moments at the representative mode points of the dynaanic model. The response spectrum approach was used to obtain the maximuss 1 equivalent static deflections, shears, and moments for the seismic occur-l 1 rene.s. l Def na 0"" 4 Shear,, ,a< Sn. , x PA( x where: a _ Momen t =_ _
= .
De f,, Sh ea r, Moment na .= nodat deflections, shears and manents for node n and mode a 1 (na' I nn' *nm, a modal deflections, shears and moments frcn the dynsmic analysis l l pag = participation. fac tor for mode a j Sam
=T a spectral acceleration f rom response chart ( see PSAR, appendts A l. ) f or percent of s ritical damping (a), divided by the l circular frequency (w) squared.
48 l 1
I BSEP.! & 2 l ( l l 1 The individu.tl modal results are combined as fo!!ows: M 2 o n s [(Defnm) m=1 V, a f (Shear,,) where: as= 1 2 5 ,,, g * *ns) l o,V,M
=
n n n W root-mean. square combinations of modal deflections, shears and moments for the nth node of the structure for all significant modes a = 1. . . . . M. The significant ! I
' modes fall within a closely spaced range. I
& sbove procedure is repeated for each node in the dynamic model.
The groes nodat memet.ts and shears, refereoced to the centerline of ; the containment structure, are distributed over its actual cross-sectiam { based on the following relationships: The maximum tangential design shear at each node is obtained as follows: N Slin i
- g i
t. I i' 4.9 l
.. J,
-U '
/
38EP-16 2 Where: Hn
= maximum tangential design shear at node n.
Vn = root-sert-square seismic shear at node n.
'n =
centerline radius of containment at node n. , the maximum vertical uplift force at each mode due to the seismic moments is obtained as follows: Ze (k/in) . % 8 "" In Where: In = maximum vertical uplift force at mode a. M. = root-meen-square seismic assent at mode n, r = n { radius of centstaneet at mode a. to .= cross-sectional well thickness at mode n.
=
In Ntn 'm Vertical earthquake contributions associated with each of the two horisontal earthquakes are considered. (E, and E4). The vertical ground acceleration is established as 2/3 of the horisontal ground scesteration
'shown in the response charts.
i' The maximum vertical design forces due to the vertical ground acceleration are obtained from
= Sm Zn l
1 4-10 ) 4g . 4 l .
a 58EP.1 6 2 Where: M = maximas vertical design force at node n associated with the vertical earthquake components.
=
mass of structure at mode n. 5
, 8,,,, =
vertical acceleration for mode m. The above equations refer only to horisosts! and vertical components typical of the cylindrical areas of the drywell. The shears and moments in the conical areas of the drywell are transformed to ognivalest forces atoms the axis of the come and shears normal to the surface of the come. The seismic loadiss ets are a superposition of the horisontal o and vertical earthquakes.
;t ij E =
E, + Ey I E' = I'q + E'y E and E' are the total seismic contributions to the coms'ined l load equations. EH and E'g are the maximum horisontal earthquake components. Ey and E'y are the maximum vertical earthquake components. 4.5 Material Constants The stress-strain relationships in all computer runs are based on f e the following material constants where: En , E, , a nd ( refer to the elastic moduli in the N,0.5 material coordinate system (see Figure 11). y = Poisson's ratio C = Shear modulus in pel ,, i l 4 11
\
. .--- - ,~.e
.- a '(
38EP-1 & 2 The following material constants were used: [ Concrete: (uneracked ) Concrete: (cracked) En , and/or E,, and/or E, = 3.0 x 106 pet E,, and/or E,, and/or E, = 0.0 psi V'= .15 V = 0. 0 C = 1.3 x 10 psi C = 1.3 x 100 psi l Uur
' 0 E,, and 5 , = 27.4 x 10 poi E,, = 0.0 (not applicable,s, = 0.0)
,,, = .30 C = 10.5 x 100 pet i!
l Reinforcing (Noop) toinforcina (Heridional) a 0 l I, = 29.0 x 106 ,,g Is = 29.0 x 10 psi l!
= = =
E, = E, 0. 0 E, E, 0. 0 { 1 = 0. 0 W = 0. 0 C = 0. 0 C = 0. 0 Head Bolts E, = 29.7 x 100 pai En " E = 0. 0 9 V = 0.0 C = 0, 0
?
I t a f t a.n . O G
- -- ~ _ _ . - -
A
58EF.1 & 2 4.6 Concrete Cracking Analysis psi Based on the original estimate of required reinforcing, a series of f six computer runs determined the final crack pattern in the drywell well and base pedestal. The procedure for determining the crack pattern involves a check of the stresses is each idealised comerete element after each sec. cessive run. If taka concrete stresses in the global R, 9, Z system or the N, 9, 3 transformed coordinate system are in tension the elastic properties in that directica (E and assosisted if) are set equal to 0.0. As shown on Figure 13, this orthotropic concrete element has no stiffness camponent on I the plane normel to this corresponding coordinate direction. Figure 14 shows the assumed' initial crack pattern prior to the first run. The entire pedestal la ascused uncracked in all directions. The drywell tell is assumed to be fully creaked in the meridional and hoop directions. Successive runs indicate that the coecrete in the drywell well remains in radial cog resstaa. The iterative approach to determine the extent of concrete cracking involves two liner stiffuesa assumptions for the factored load equattom. U =
. 9 D + 1. 50 P + 1. 0 T1 . 5
- 1) Liner has stiffness and load carrying capability.
- 2) Liner has aero stiffness and no load carrying capacity.
The above-mentioned liner assumptions are discussed in detail in See t ion 4. 9 below. Figures 15, and 16 show the progressive crack pstterns of the runs including the liner stiffness. I 4 13 :
~ _m
388F.1 & 2 Figures 17, and 18 show the progressive cr$ck patterns of the runs not including the liner stiffness. f Figure 19 shows the final iteration and crack pattern by the super. position of the cracked sections associated with the two above mentioned liner staffness assumptions. This final iteration results in complete pedestal cracking ta the 31ebel R & 9 directions, except for a vertical ~ band along the outside face dich remains uncracked in the radial direction. The total hoop and radial cracking in the pedestal in combination with total hoop and seridional cracking ta the drywell well produced a maxi. aus stresa condition in the corresponding radial, hoop and seridions! rein. forcing. The dryuell reinforcing sises and spacing were then re-estina'ted baseo on the her stresses obcained from the first series of rune. This reinforcing and the finst crack pattern shown on Figure 19 are used la all successive runs. 4.7 Drvve11 Analysis Seemenes The objective of tis drywell analysis la to maximise stresses for the required load egostions; consequently, the runs are subdivided into 5. independent series of analyses. Each series is specifically concerned with maximizing the stresses in a particular area of the drywell containment; they are:
- 1. Pedestal hoop and radial reinforcing
- 7. . Drywell well meridional and hoop reinforcing !
f
- 3. Liner yielding I 4 14 j
. l
> m_ . - www
, 883P.1 & 2 4 Pedestal concrete shaar
- 5. ' Dryuell concrete shear Each s'pecific area of investigation is analyzed for the load combi-nations listed in Section 4.3, "Destga and Analysis".
4.8 Podestal Epop and Radial Reinforcias Se pedestal reinforcing consists of top radial bars and closed hoops, a bettman orthogonal est over its full area and closed hoop bars along the vertical entside face. The stresses la all the compements of reinforcing are not maximised with a enigme set of boundary conditions. M:gr1y, several runs are made to obtain the maximum stresses associated with each set of reinforcias. The concrete in the pedestal is assumed fully cracked 'in the radial and hoop directions. (Se<< Concrete Cracking Analysis, Section 4.6 above. ) This assures that oudy the reinforcing will resist all radial and hoop
- forces.
Free amiform expansion with no radial restraints aloog the base of the pedestal gins maximum hoop strains la che outside face hoop rain forcing, conaepntly maximizing these hoop stresses. The unrestrained radial growth also produces maximum stresses in all bottom reinforcing and top radial reinforcing closest to the physical centerline of the containment structure. Figures 20 and 21 show the original and the deformed structure graphically plotted for the radially unrestrained and radially restrained 4-15
\
d
/
BSEF- 1 & 2 pedestal base boundary assumptions. Concrete cracking assumptions in these l 1 plots are as previously discussed. The combined load equation, U = 9D + 1. 50P + 1. 107 .5 is the loading used in both plots and is con. sistent with that used to maximize the pedastal hoop and radial reinforcing. We deformations are increased by a factor of 50 for visual clarity. i 1bo exceptions to maximising stresses in the top radial reinforcing I occur: the stresses in the top radial reinforcing near the discontinuity I at the intersection of the drywell well and the pedestal are not suszimized when the base of the pedestal is allowed to move freely in the radial direc-tion. The well and pedestal tend to displace radially as a unit. These { stresses are maximised in the second set of rama in editch the base of the pedestal is restrained from ar.y radial growth. The radially restrained I pedestal creates the maximum bending stress situation near the us11 and I pedestal intersection. The relative difference in radial displacement between the restrained bottom and free top caused by the discontinuity soment produces the maxisans strain and stress condition in the top radial reinforcing near the intersection. An investigation of the area near the base of the pedestal shows vertical tensile (+Z) stresses in the (uneracked) concrete. The specific boundary assumptions related to these tensile stresses are restrained radial and ve r t ica l g rowt h. The third set of runs in this series involved successively releasing the vertical restraints where tensile stresses in the concrete occurred. This fl
!l 4 16 1 N- _C F
./-
Bagpol & 2 set of runs produced an "inve'rted saucer" displacement pattern with only the outside section of the base bearing in compression as shown on Figure 22. 4 This condition maximizes the stresses in the top radial reinforcing
' between the centerline and wall intersection.
4.9 Drywell Wall Meridional and Hoop Reinforcing The drywell well reinforcing consists of continuous meridional and hoop reinforcing rm both faces of the well. The us11 is completely crarW in the hoop and meridional directions consequently producing a maximum stress condition in these components of reinforcing. ; i The well-pedestal discontinuity and the various restraint conditions " associated with maximising stresses in the pedestal reinforcing have little f effect on the stresses in the upper portion of the drywell well. A comparison of the drywell well displacements and reinforcing stresses show that the dis-t continuity moments and stresses dampen quickly. The initial cracking analysis resulted in a meridional compression sone near the outside base of the well due to the discontinuity moments occurring in this area. Successive cracking of the concrete elements at the base of the well produced an increase in the meridional reinforcing stresses to the point where complete cracking produced maximum bar stresses. The stresses in the meridional reinforcing along the inside face of the drywell wall near the pedestal intersection are maximised when the base is fully restrained. This again occurs as a result of the maximum bendies stress situation inherent in the fully restrained base . 1 4-17 W= ; - 3
BSEP-153 l l
, 4.10 Investigation of Liner Yield and Boekling Liner stresses are checked for possible liner buckling in all runs f
previously discussed. Based on the ASME pressure vessel code, the allowable yield stress at 333 F (liner temperature for T1.5 1 ad condition) is 28,000 psi.
}-
The yield stress condition is based on the distortion energy - I theory for the two-dimensional stress esse: 0
! ,1 2 _ e l ,2 +<2 2- , , ,,2 i
where 8 a1 = Liner hoop atress e2 = Liner meridional stress l
- yp = 28,000 psi Based on the above yield criteria cee zone of local compressive yielding occurs in all runs involving the combined load equations. The combined effects of the fully restrained pedestal base, and the relatively high degree of fixity at the intersection of the pedestal and well prevent free thermal growth and consequently produce this local yielding condition.
The equilibrium state of the liner and reinforced wall due to the 1.5P + Tg,$ load condition results in compressive liner hoop stresses dich exceed the yield stress in a 9'-0" section of the liner shown on Figure 7. The liner elements in the area where yielding occurred were modified and given zero strength properties (E, = E, = \/j a 0.0) to ~ideallre this yield condition. This is consistent with the annumption that the liner will contribute no struc tural stiffness in the yield state. A run was made (I) Elements of Strc9gth of Materinta by Timoshenko and Young 4th edi t ion, January,1962.
. 4-18 I
~
s .
1 i BSEP-1 & 2-with the same boundary conditions and loading case. A check of the resulting liner and reinforcing stresses in the vicinity of the' liner yield area indi-cated that only slight stress changes occurred due to the redistribution of the externally applied loads. '!he redistributtreo causes no additional yielding in the areas above and below the local yield zone. This con-clusively indicates that the local yieldi: 3 which occurs -remains local and no total structural failure will occur. The liner and reinforcing stresses above and.below the area where ths loss of liner occurred are tabulated on Figure 23, comparing t'he pre- and post-yielding conditions, i l 1 4.11 Investigattom of Concrete and Reinforcing Stresses at Full 7 f Yield Condition of the 1.iner , It is difficult to predict an accurate pattern of local yielding. ! b It might be revealed by the analysis or it may occur in areas not shown by ll ' the analysis. If the latter case is true, the actual behavior of the struc- ,' ture differs from that analytically deteratined, thus leading to larger or san 11er internal forces and/or noments at these points. Consequently, it is shown that the reinforced concrete portion of the structure can provide suf ficient neebrane and flexural resistance for all factored load combi. nations that include thermal effects, with the liner assumed to have no strength properties. A series of runs for the three governing load equations listed on page 4-7 using the conservative assumption that the liner contributes no load resinting strength to the structure were made. An investigation of the rein-forcing stresses from these runs shows that the reinforced .ontrete portion of , the structure alone is adequate to resist the established loading criteris. '
,ir f
o ,I l 4-19 i i d tM - ,W .
'A
_ __ a
I I BSEP-1 6 3
- t. 12 Corn re te Fadial She.ar The individual neries of analyses thus far discussed concentrated solely on maxisiiring the reinforcing stresses in specific areas of the s t ruc ture. The same cont rr t e cracking anstamptions were typical for each case which governed the reinforcing stresses.
To assistre cemcrete shear stress resultants and determine the ) location and position of shear reinforcing, a series of rauns was made aantaming the cancrete to be ecsupletely cracked in the plane morsial to J the hoop (M direction only. Consplete fixity at the b.ase of the pedestal was annumed since resistance to radial growth would put tbc pedestal con-crete into a maxisua shear staic. Radial shear forces are obtained by integrat inn the normal shear attrines through representative cross s ec tions . The stresses are acting in a pl.ine perpendicular to the axis of the structure. The allowable concrete f.hiar ntress in based on the proposed revialem of ACI 11 A-6 3. Code, "hullifing Code Rceguireecnt s for Reinforced Conc rete". { in the pedi at.11, where hoop concrete tensile s t re s se s a re nasa l t , she.ar rein f orc ing is stred to resist.that porti m of the shear that exc eed = the .illow.ihic conirrte atrons. In the drywc!! wall, where the allowable conc ret e s t reo l a. en erded the following two criteria are used to sire shcar rrini or: ing
- 1. At the t op .in41 bot t ern of the drvvell w.111 whear reinf ort ing is l s i . rit t o i .o r r) the t o f .i t s hr ei r iorce. These two seetiona resiat larre bending misnent s whish result in diagon.il t em s t on c rac k p.i t t e rt' < .
l 4-20 l
'^
._ _;; *i .
-_-.-_-__m___._______m.__
MSEPol 6 2
- 2. In the remaining sections the shear reinforetng is sized for the
, net shear force,1.c. credit , as permitted by the ACI Code, is taken f or the concrete. The- cract pat terns are horizontal, indl.
cating that diagonal tension is not a significant tactor. Based on the above criteria it i s de t ermined t ha t she.ar reinforcing is required in the dryvell wa!! and in the outer areas of the pedestal as shown on Figure 6. The shear reinforcing is mechanically anchored by hooking the ends of f the st irrups around the main meridional bars. I la the peh stal and at rhe top and bottom of the drywell well the st( m g9 y 4 sp*ma parallel to the anzimum principal stress direction (perpendicular to the crack plane). I in the remsinder of the drywe!!' the stirrups are placed horizontally, if a crack should develop between the stirrups, the clasping action of the meridional rebar is used to develop shear friction across the section. The required meridional reber is sized considering the axial force plus the total shear f orce. Figure 24 shows the two stirrup pat terns used. 4.13 Analysis of tifywell Head The analysis of the steel dnme and .sssociated at i f feners it.volves a
, series of runs to check the adequacy of pre-tensioned hold.down bolts and the et!cct of the loail eqtut tons in the head and the drywell.
The pre-loid in t he bol t s insured that the gasket flange rem 11ns in compression.
'. 1 1. l.oad and St rem s Plot s the graphic presentation of the drywell design and analysis results consists et inree resul tant and actual st ress plot s.
l h 4 21 ' m m_
/
RSEFo1 '& 3 In both cases results are plotted for all coordinate directions in-
, depetiden t l y (!;, S, & e) . Individual load comiponents for gravity, pressure, i
therwaal, seismic loads, and comibined load equations are presented. The load
, and stress plot diagrauss consist of an envelope of curies for each load con-dition.
The centroid of the effective material is used as the reference plane for all force resultant plots. In the wall area, positive force camponents are plotted to the right of tise centroidal aria ased negative forces to the lef t. positive soonnents are plotted to the lef t and sesative sensets to the right of the centroidal axis. Positive and nesstive stresses are plotted respectively to the right 'and lef t of the siner or reinforcing centerline. The sigs convention for sones's and ' shears is based on the fo!!owing assumpt ians. All clockwise moments are considered positive, and counter-clockwise moments are considered negative. Positive shear forces are those whose vertical face coup'ies tend to rotate an element or section in a counter-clockwise direction and negative shear forces are those ednese vertical face couples tend to rotate an elssent or sec tica in a clockwise direc tion. All plotted values in the pedestal are referenced to its centerline. The procedure used in obtaining force resultants involves integrating the elemental stresses through the full cross section of the structure, about the material centroid of the cross-section being ccmsidered. t.oad intensities are calculated and plut ted at 14 representative loc a t ion s in the dryvell wall, 6 l oc a t i sm s in the drywell ped.ntal. and 3 locationn in the drywell head. The dimens t mal locations of these points are shown on Figure 25. 1 i 4*22 l l ser* " n-
BSEP-1 & 3 4.14.1 Tvoleal Fogt J.esultant Calculation Figure 26 tb.ws a cross-section and associated stress components used in calculating force resultants and smosasnts. The centroid of the effective material at any cross-section is calculated as: n I Ei tgsg
= where:
C i=1 a T Etti i=1 Eg (i = 1...n) = elaatic moduli of liner, reinforcing, or cosse. rete. ti (i = 1...s) = equivelset plate thicknesses of liner and reinforcing, or actual thickness of concrete layer. Zg (i = 1...a) = distance from reference plane to centerline of element. When concrete layers are assumed cracked in a specific direction, the centroid depends only on the properties of the liner and reinforcing eves in that specific directies. The total axial force at any particular cross-section is calculated as: n
=
Fgg [ et tg i=1 where: F = axial force resultant at sectice 1-1
,g (i = 1,n) = stress cautponents in liner, reinforcing, and concrete (if uncracked)
The total sectics moment is calculated as: n M gg = [ et tg x i i=1 where: Mgg = total moment at section 1-1 referenced to the centroid of the effective material ( 4-23 .
BSEP-1 & 3 i X a perpendicular distance from centroid to associated layer of the cross-nec t ion . f St: cases are plotted directly from the program calculated results. Strest, commponents are plotted individually for the liner and each layer of meridional and hoop reinforcing. The individual stress and force resultant plots are ccryrised of values obtained fram seve ral runs. These values are plotted frasi the parti-cular run which maximized each reinforcing camponent or concrete stress component. The sections of the report pertaining to reinforcing and comeretc shear discuss in detail the restraint and cracking assumptions editch maximize 4. the stresses in the various cosponents of re'aforcing. 4.14.2 Force Ilesultant Plots - Discussion and Summary st. An explanation and related discussion for each force resultant (load) plot is included in this section, r.ach plot contains the following force and scenent ccuiponents:
- 1) Meridional forces
- 2) Hoop forces
- 3) Meridional bending moment
- 4) Hoop bending moment
- 5) Radial shear forces (O Di splacement plot.
The dine arca of each force resultant plot crmtains two independent sets of resulta. The solid line. Plots are the forte resultant s of the spct ific li6nt i ond.c lon asnos tat ed wi t h the plot . The dashed line plota are a super-posit ion of the .insa iated load condition and the pre-load forcen in the h ad b.Its. t.-24 & -c _
.m.s-+f
BSEP 1 & 3 Figure 27 (1.0 Dead ' oad Dr.ly)* 1,oading consists of the gravity vector in the 2 direction only. The base of the pedestal was assumed to be fully rescrained in the R 6 Z directions. Concrete was assumed cracked as discussed in Section 4.6. Although concrete cracking does not occur during the " dead load only" case, maximum reinforcing stresses are obtained. The inclusion of dead toed in the combined load equations requires consistent cracking assusytions. Fleure 28 (1.0 Pressure only) Leading consists of 62 psi internal pressure only. Cracking assump-tions are identical to the final crack pattern sa shown on Figure 17. The force and manent componenta are a superposition of 2 runa: Rtm #6 = Base of pedestal fully restrained in R & Z directions RtTN 825 - sese of pedestal free la radial direction only Maximum meridional and hoop forces, and sneximum meridional and hoop bending are obtained in the drywell wall with the base of the pedestal fully restrained. Maximum forces and masents in the pedestal are obtained uhen the base is allowed to displace in the radial direction. Meridional acments and radial sheurn are plotted for RttN #18 in which the base of the pedestal was fully restrained, and the entire drywell well and pedestal cracked only in the hoop direction. This case produces maximum meridional moments at the wall-pedestal intersectism, but does not result in a maximum reinforcing stress condition. The nurginal notations on this and subsequent load and str.u plots refer to comput er run numbers and their associated load conditions. 4 25 I
\ &k k s%;2%%dd.";"ll,12WCWW JLA A usEC4&? " $'NY # 5 '
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* ' .l f BSEP-1 & 2 i
, {
, Timures 29. 30, 31 (Tg ,o, 71 .25, 71 .50 '"I F )
This series of plots for the three temperature loading conditions consist of six runs (2 for each temperature condition): i
, )
RUN 9 (T ) 10 (T ) final concrete crack pattern as shown on Figure 19 l 7 (7 .5) 1 base fully rescrained EUN 22 (71 .5' 23 (7 1 5) same cracking assumptions as above 24 (71.0) base free to displace radially The meridional forces, hoop forces, meridional bending, and hoop bending in the drywell wall are maximited for all three thermal load con-ditions with the base of the pedestal fully restrained. Radial and hoop forces in the pedestal are maximiaed when the base of pedestal is allowed to displace radially. The meridional and hoop moments in the pedestal contain maximum values from both the fixed and free base 5 condition. i ' Ftrure 32 and 33 (E & E') The seismic forces for the two required earthquakes are calculated as described in Section 4.4 4 Seismic mcuments arc distributed through the cross-section as axial l forces. The plots for the sciamic loads contain axial forces and tangent i.it i shears for the horirontal carthquaken, and axial foreca for the vertical earthquake components. t (. I 4-26 hW N !
58EP-1 & 2 '. Figure 34 (,9D + 1. 0P + Tg , o + 1. 0E') Figure 35 (,9D + 1. 25 P
- T1 . 25 + 1.257)
Figure 36 (. 9D + l. 50P + T 1. 50) The three ccambined load plots are a superposition of runs involving e i two independent sets of boundary conditions.- Drywell well forces and moments are smaxianized with the base of the pedestal fully restrained. The pedestal f orces and moments are maxinaired with its base free to displace radially. All forces and moments are maximised with the following crackius assusytions: 1 wall and pedestal fully cracked in the plane' normal to the hoop directica 2, well cracked in plane normel to the meridional direction,
- 3. pedestal cracked in the plane normal to the radial direc-tion, uncracked in plane normal to the 2 direction The above cracking assumptions are identical to the final cracking pattern shown on Figure 19 W radial shears are maximised and plotted for the cases in which the well and pedest.at are fully tracked in the hoop direction. Tangential shears occur'only from earthquake Ioading and are plotted for the load equations in which they occur.
4.14.3 St ress Plot s . Discussion .ind Smry The inalysis procedure involves direct calculation and printout of element al st ressen. Ttw objective of the analysis is to concent rate solely I
- 4 27
.scr.i 6 2 I
I on maximizing these stresses 6a the combined factored load equations. Con. f 4, sequently, stress plots are presented for the pressure test condition and the three factored load equations. ' The stresses in all cases are mexinuam stresses with associated boundary conditions and cracking assumptions as 2 discussed in the sections of t'he report dealing with pedestal and ws11 6 reinforcing, liner stresses, and concrete shear. Included in this seciton is an explanation and related discussion for each stress plot, 1he following individual ' components are plotted for each combined ,!and equatium:
- 1) Liner meridions! stress .l
- 2) Liner hoop stress :i
- 3) Inside seridional reinforcing stress
- 4) Inside hoop reinforcing stress !l
- 5) Outside meridional reinforcing stress
- 6) Outside hoop reinforcing stress j Seismic tangential reinforcing stresses are also plotted for the
{ j two er.uations in which E and E' occur. l The three plots included are:
- l. Finure 37 9 D + 1. 00P + T1 00 + 1. 00 E '
- 2. Finure 3M . 9D + 1. 25P + Tg,25 + 1. 25 E ;
- 3. Finure 39
. 9D + 1. 50P + T l. 50 E.ich load equit ion with its associated pressure and temperat ure con.
dition cont iins auximum stresses of run cases with two independent sets of bound.ary concations. 4-28 s " -
' . _ _=A ___ _Y_
l BSEP.-1 & 3 I The concrete cracking assumptions are typical for a!! stress plots: Drywell wall completely cracked in planes normal to meridional and boop directions. Pedestal is completely cracked la planes normal to the radial and hoop directions; uncracked la the Z directica. I . The maxisman well reinforcing and liner stresses occur when the
. base la fully restrained. All well reinforcing and liner stresses are plotted for the run cases using the restrained base assimiption. j l
Pedestal radial and hoop reinforcing stresses are asximised when the l l , I base is allowed to displace radially, and are pit ited for run cases with this boundary assumption, I Seismic tangential reinforcing stresses are calculated as follows: l The diagonal steel in one direction is designed to resist the full horizontal component of the seismic tangentist shear forces. The liner and reinforcing stresses associated with the pressure test l case (.9D + 1.15P) are plotted on Figure 40.
~
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8817 1 & 2
- 5. 0 SUPPRESSION CHAMBER 5.1 Physical Description of suppression Chamber l The suppression chamber (SC) is a hollow reinforced concrete shell of rectangular cross-section er. circling the lower portion of the drywell containment structure. The concrete encloses a continuous 16-sideo steel l
liner of circular cross-section. The major centerline diameter of the SC is 109'-0". N cross-section diameter of the circular liner is 29'-0". The SC is structurally independent of the drywell (DW) containment. The W and 3C are both supported os the same est foundation. A paper joint is provided between the bottom of the SC and the est foundation to alloer radial sapansion of the SC. Jertical keys are provihd along the outside perimeters of the drywell pedestal to allow independent, unrestrained ex-pension of the SC when suLjected to symmetric loading conditions. Dader anti-synestric loads the keys will force the DW and SC to respond as a single unit. The SC is reinforced with a single layer of continuous closed hoop reinforetas aquelly spaced around the perimeter of the liner. Meridional reinforcing in the form of closed rings is provided. The meridional reinforcing is placed radially to the centerline of the containment. Diagonal seismic reinforcing is located along the vertical faces of the SC. Figure 41 shows the typical suppression chamber reinforcing. The SC contains eight symmetrically located vent openings corres-ponding to the vent. openings in the drywell. The main hoop reinforcing which would normally occur in the openings is banded above and below the openings, the meridional bars which would normally occur at these locations are grouped on either side. The vents are locally reinforced by closed rings around I 5-1
- 5
, 88EP-1 & 2 the perimeter of each opening. Fill steel is also provided where necessary.
5.2 Idealization of the Suppression Chamber ne idealized model used in the analysis is representative of a ; 1 section of the suppression chamber, ne mesh of interconnected elements ! are assumed to be true tings. A cross section of the axisymmetric finite element model used in the analysis is shown on Figure 41. j J A table of equivalent plate thicknesses associated with all idealised liner and reinforcing elements is also included on Figure 42. ne data das. cribing the idealized model was numerically and visually checked. De com.- puter plot of the model shown on Figure 43 insures that the geometric j
. data used in the analysis is correct. .
Se liner, which is a series of 16 interconnected cylinders, is idealised as a pure circular torus of isotropic steel and a constant thickness of 3/8". He hoop and meridional reinforcing is idealized as independent layers of truncated come elements of toroidal shape. D ese layers of rein-forcing are given stiffness properties only in their respective hoop and meridional directions, ne concrete portion of the SC is modelled as inter. connected triangle and quadrilateral ring elements. De coordinates of the exte nal surfaces of the SC model exactly duplicate the physical dimersions ' of a section of the structure taken normal to one of its 16 sides. Material coordinate systems of the truncated cone elements that describe the liner and reinforcing are assumed to be coincident with the meridional axis of cach liner or reinf orcing element.
$~2 m.
F f BSEP- 1 & 2 ne mater ial axes of the concrete elements are transformed f roci che globst (R, e, 7.i coordinate system to an (N, e, S) system consistent with a radial generator from the center of the suppression chamber as shown on Figure 44. The concrete elements are assumed to be ful:y cracked .n the hoop (e) and in the meridional (S) directions. The equivalent plate thickness of the idealized meridional rein-forcing rings varies along its perimeter. 21s is caused by the physical variation between the minimum inside radius and maximum outside radius of the SC while maintaining a radial pattern. Se equivalent thickness of ti.e hoop reinforcing elements above and below the vent openings is increased to compensate for the banding, The con-crete elements within the bond of vent openings are assumed to contribute no stif fness to the suppresstor chamber. De banded meridional rings are continued through the vent area. 5.3 Design and Analysis The complete design and analysis of the suppression chamber includes to independent computer runs. The design philosophy for the suppression chamber involves an iterative approach to maximize reinforcing stresses for
~
all combined load equations, nis approach includes 2 separate runs for each load component or load equation; the first allows the base of the SC to treely displace in the radial direction; the second allows no radial dis-
; plac enw n t at the base, in both cases the base is restrained in the global 7.
1 { i l 4 5-3 l Q
y
?
4 flSEP-1 6 2 direttion. This aporoach bounds the solution for each load condition due to che unca risinties in the degree of restraint at the construction joint between the suppression chamber and the foundation mat. The loading criteria used as the basis for all runs involved three governing combined load equations containing factored pressure, therisa? , dead load and seismic component ~.
- 1. - l'g =
. 9 D + 1.50P + 1.07 1.5
- 2. =
U2 9 D + 1.25P + 1.071 .25 + 1.25E ,
- 3. U =
9 . 9 D + 1. 00 P + 1. 0T1 . 0 + 1. 00 E'
- Tp =
.9 D + 1.15P (Pressure test only) uhcre:
U = ultimate required load espacity of the structure D = dead load of structure including fixed equipment P = loss of coolant accident pressure 1.50 T 1. 2 5 p = Wrmal gradients ass ciated with each of the T 3 factored loss of coolant accident pressures (1.0P,1.25P, and 1.SP) E = earthquake loading (operating basis earthquake) E' = earthquake 1.>ading (design basis earthquake) See detailed description of sa Ismic loading M 'ov. The pressures and tetrperatures associated with the above load equat ions are shown on Figure 45. 1 i l
- Inc luded onl y 4s test case, and does not govern design. l e
5-4 9
BSEP-1 & 2 Extreme conservatism occurs in these load equations. The pressures associated with each of the 6hree thermal loading conditions are based on the factored loss-of-coolant accident pressure of 62.0 psi. This pressure is initially factored before it entere tne ultimate load equations. The suppression chamber during a loss-of-coolant accident condition I.as a peak pressure of 28,0 pai. The pressure components in the load equations are con-servative since the SC never sees the 62.0 psi pressure load used in the a na lysis, 5,4 Material Properties The elastic properties E, NI, and G used for the liner, reinforcing, and concrete in the SC are consistent with the elastic material constants used in the analysis of the drywell presented in Section 4.5.
- 5. 5 Seismic Loading The combined load equations require consideration of earthquake load cwponents E and E' where:
E = operating basis earthquake (.083 horizontal ground acceleration E' = design basis earthquake (.16g horizontal ground acceleration) The dynamic analysis of the containment structure produced modal accelerations at each node of the dynamic model. The average root-mean. square accelerations at the elevation of the suppression chamber are used to determine the tangential shear: Vn " Mn A where: Vn
=
horizontal seismic shear at nodc 1 Mn = mass of suppression chamber at nods . 1
, A =
overage scismic acceleration at centerline elevation I of suppression chamber. 5-5 { {
l l BSEP-1 & 8 ) The maximum tangential design shear uniformly distributed over the height is: Yn i V (k/in) = q , whera: Rn
= centerline radius of the suppression chamber j
5.6 aoop and Meridional Reinforcing The main reinforcing in the suppression chamber consists of large I closed hoops around the perimeter of the liner and small closed rings ! J parallelling the liner in the meridional direction. l I i To create a maximum stress condition in the suppression chamber ' reinforcing a set of runs for each load component and each factored load equation were made based on the following assumptions: i (
- 1. All concrete is cracked in the elemental meridional and l hoop coordinate directions, t
I l The concrete retains shear stiffness and s:if fness in the l
; direction normal to the liner, l d
- 2. Two runs are needed for each load case with respect to the j I
bottom surface boundary conditions: j A. unrestrained radial displacement condition at base R. fully restrained base condition i
- 3. A set of runs are required to show that the reinforced concrete portion of the structure can independently provide sufficient mem-brane and flexural resistance f u all factored load equations that 5-6 e 'g
*& i
f , AgEF- 1 & 8 i l 3 ' include thermal effrets, asstasing the liner has no stif fness or load carrying capacity. This asstanption effectively presents a conservative approach with respect to'the dif. l t j ficulty incurred in accurately predicting local compression yielding or buckling. ' The reinforcing stresses in each successive run indicate that a maxisman hoop and nord.dional reinforcing stress condition occurs when the base is allowed free radial displacement, and the liner has no stif fness or l load carrying capacity. 1 The maximiss hoop bar stress with the liner having no stiffness (47 Irsi) and the assimum meridional bar stress (43 kai) occur near the l- outside center of the suppression chamber, These maximum stresses are below
~
1 the yield stress (Fy,= 50 ksi). i The near yield anximum stresses in the hoop and meridional rein-forcing occur as a consequence of the extremely conservative asstaption j that the liner has no stiffness or load carrying capacity. The comparable j
?
maximum bar stresees for a run with the same load conditions, in which the i stiffness of the liner is included, are 26 kai in the hoop and meridional bars, versus 43 and 47 kai in the same reinforcing using the zero liner assumption. 1 All liner stresses in this rur are in tension in both hoop and eneridi- { onal direc t ions. This defialtely shows that no compressive yielding or buck-ling occurs in the liner, and verifies the conservatism resulting from the l zero liner assumption. 5- 7 1
,o .
l 4-
'BSEP.1 5 3 I
\;
- 5. 7 Liner Yielding an'd Buckling Due to the uncertainty in the degree of fixity at the construction joint between the bottos surface of the suppression chamber and the founda.
tion ma t , a duplicate series of runs for the same loading conditions was performed assuming the base to be completely fixed. We restrained radial displacement produces a maximum compressive stress condition in the liner at the bottom of the SC ( 47 kai in meridional direction, and -42 kai in hoop direction). D e inhibited thermal strain is the dominant factor that produces this local yielding of the liner. The fixed base condition that produces the local yielding is again an extremely conservative assumption. the frictional resistance to radial movement at the base would be overcome before the maxisman stress condition is reached. The uncertainty in the degree of fixity at the bottom surface of the suppression chamber is eliminated by using the maximum stresses obtained from both the restrained and unrestrained run cases as design stresses, in sussnary, the unrestrained base produces maximum hoop and meridional reinforcing stresses, and the restrained base produces the maximum stress condition in the liner. Figures 46 and 47 show graphically the original and the deformed structure plotted for the radially unrestrained and radially restrained boundary assumptions. The deformations are increased by a f actor of 50 for visual clarity,
- 5. 8 ,S_upyression u chamber load and Stress Plots The graphic presentation of the suppression chamber design and 5-8 f
. . . . - - ~ - ~ ~ ~ - ~ - -
I '
- 4 i BSEP-1 & a i i
1 analysis results consists cf load (force resultant) and actual stress plots. A In a;l cases results are plotted for all coordinate directions in- ) f dependently (N, S, and e). Individual load components for gravity, pres. 1
; sure, thermal and ceismic loads, and combined load equations are presented.
As in the drywell the plots consist of an envelope of curves plotted
, for each load condition.
Since all concrete is cracked in the meridional (S) and hoop (e)
' directions, the centroid of the reinforcing steel and the liner is used as the reference plane for all force resultants. Positive hoop, meridional, and shear forces are plotted external to the reference line, and negativw hoop, meridional, and shear forces internal to the reference line. All moments are plotted on the tension side of the centroidal axis.
The procedure used in calculating moment resultants involves inte. grating the stresses about the centroid through a section normal to the liner. The method is identical to that used in calculats.ig the drywell noments. Force resultants are calculated and plotted at nine representative locations around the perimeter of the SC. The dimensional location of these points is shown sn Figure 25. Stress diagrams are plotted directly from the program calculated result s, and are plotted independently for the liner and each layer of rein-forcing. i l 4 I 59 { a _ __, ._ - - - - - - . --- 2
BSEP.! & 2 5.8.1 rorce Resultant Plots An explanation and related discussion for each force resultant plot t is included in'this Section. The assumptions, typical of all plots, unless otherwise noted are:
- 1. Concrete fully cracked in meridional (S) and hoop (e) ,
directions i
- 2. Base free to displace radially l
t
- 3. Liner stiffness included '
I 1he forces and moments plotted for each load case include: '
- 1. Meridional forces a i;
- 2. Hoop forces i
- 3. Meridional bending
'k' j
4 Hoor bending 1
- 5. Radial shear forces 1
- 6. Displacement plot Figure 48 (1.0 Dead Load Dr.ly)
Includes only gravity loads in Z direction. Concrete cracking J I does not occur under dead load. Consistent cracking assumptions must be used in the combined load equations of which dead load is a part. Fi gu re !.9 (1. 0 l' On l y) In this load case and plot the suppression chamber is subject to an internal pressure of 62.0 psi only. The combination of the liner, rein-forcing, and concrete stif fness in the normal direction approxinkitely q 5-10
, _ __ ______ _ d
BSEP-1 & 3 idealize a classic torus. The 1.0 P load case presents the basis for com-paring the suppression chamber results with the classic solutfor. The following four items reflect the specific areas and physical :ondicions in the suppression chamber that cause a h' variation from results expected in a classical torus solution: 1.
, We variation in the spbeing of the closed risig meridional' +
bars increases from a minimum spacing on .the inside of the SC to a maximus spacing on the outside, this causes a vari-ation in equivalent plate thickness of the elements ideal. isini these meridional bars. I j 2. ne banding of the hoop reinforcing in the vicinity of the vent openings results in an increased equivalent thickness f of the hoop reinforcing elements on either side of the openings.
- 3. We restrained vertical displacements along the base of 1
the SC prevent free uniform growth. 4 ne concrete stiffness in the direction normal to the liner varies with the normal thickness through the i 1 concrete. l
,j ii ne variation of the hoop and meridional forces about the I horizontal centerline occur as a consequence of the vertically rest rais ed I;ase. The hoop fotces would be identical around the i) perime'..r of a circular cross-section if unrestrained radial
)
growth existed. I 1 l 5-11 _u - .- - - . . - - -- ._ U
BSEP-1 & 2 , figures 50, 51, 52
-(71.0, 71.25, and T1.50 resPectively)
The forces and moments for the three thermal load conditions are plotted separately. They do not occur as unique loading conditions; the plots serve only to give an indicaticm of the effects of the themel loads in the combined load equations. Flaures 53, 54. 55
.9D + 1.0P + 1.071.0 + 1.gE'
. 9D + 1.25P + 1.071.25 + 1.25 E
.9D + 1.50P + 1. Ort.5 The forces and moments for the combined factored load equations include the superposition of all factored components and the tangential seismic shears in the two equations con-taining E and E' .
4 5.8.2 Stress Plots Stress plots are presented for the combined factored load equations > l only on Figures 56, .
- 0. 9D + 1.00P + T1 .00 + 1.00 E' 57, 0.9D + 1.25P + T f!
1.25 + 1.25 E
$8, 0. 9D + 1.50P + T g f Since the objective of the analysis is to maximize stresses, these j
plots reflect the maximust values on which the design is based. 1 in all cases the third cambined load equation resulted in maximum stresses
.90 + 1.50P + 1.0T1 .50; therefore, the stresses associated with this load condition govern the design. i !
! i The maxiinum design stresses involve a comparison of three different I 0
runs for thin load equat lon: 5-12 i
^
' ' ~~ ' ~ ~~
BSEP-1 & 8
- 1. Figure 58 (unrestrained base, lines included)
- 2. Figure 59 (restrained base. -liner included)
- 3. Figure 60, (unrestrained f base withouc liner)
As described in the section cm hoop and meridional reinforcing, the
" unrestrained base with zero liner" assuspelon governed the hoop and seridi-onal reinforcing stresses. De extreme conservariam in the zero liner case is evident when comparing the stresses in the hoop'and meridioma! reinforcing in the first and third runs, figures 58 and 60. The satire timer is shown to be in tension in both the hoop and seeridional directions. N refore, eliminating the liner to s!!ou for compressive yielding and buckling is ;
conservative. He second run, Figure 59, with the base of the 3C fully restrained I and. the liner included, shows the maximum liner stress condition in which a . local overstress exceeding Fy occurs near the bottom of the ansppression ) c hambe r. His occurs as a consequence of the fully restrained base. ~ ne overstress creates no significant problem in the SC because:
- 1) De physical joint detail at the base of the SC will not allow complete base fixity.
l i
- 2) Be design pressures and associated temperatures are unrealistically c onserva t ive.
E linrr and reinforcing strenses assoc 'ted with the pressure test s a r c . .o n
- 1 151') are plotted on Figurc 61.
I 5 13 L a
/
BSEP-1 & 2 6.0
SUMMARY
4 The primary objective of, the foregoing analysis was in all cases the maximization of the v.arious component stresses in the reinforcing, liner, .and c onc ret e. This objective was achieved by a series of runs using an iterative approach to bound the solution of the specific area being considered. The intent of this iterative approach was to cover the full range of possibilities
. med assumptions concerning the structural integrity of the ec.itainment.
ne anstytical tool used in the analysis was efficient and accurate. The espability' to create an idealized model th t was secusetrically and mathematically comparable to the physical structure, developed a high degree of confidence in the analysia and its results. He program require-- ment that the structure and toeding be axisyur -ric did not interfere with the scope of this analysis. Sufficient etress margins were included in the axisyunscric analysia to compensate for the effects of jet forces and local stress concentrations. ! , 6-1 f
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i I END i l 1 DATE FILRIED
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o _- g, m p- c. - w 0 Carolina Power & Light Company Raleigh, North Carolina 27602 N N f November 9, 1970 Q DOCKETED USAEC
'.Ar "J
NOV161970 g ~5
* - REGULATORY gp Malt SECTION ,,,,
DOCKET CLER g/ s i! Dr. Peter A. :lorris Division of Reacter 1.icen4in;
': . 9. Atomic Energy Di/trien
/ II'lI [ y
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Washington, D.C. 235 5
' s
-Regulatory File Cy. !
RE: DnCKET ::0[ 50- .ind 50-325 v .f
Dear Dr. 'lorris:
]., '
WI
' The Compan:, herewith transaits forty (-0) copies of a
<locument containing the followin;; Design Reports:
- 1. Design Report No. J. (BSEP DR-4) titled " Seismic Analvsis of the Primary and Secondary Containments." !
- 2. Design Report 'M. 3 (BSEP DR-5) titled " Final Report, Tcst i
Fili Pro; ram Recults and Baci< fill Reco..9datians for Clasr I 3,chfil', Brunr.eick Steam E:cctric Plant, sou t'ipor t . North Carolina." ,
'. Desier Report No. 6 (BSEP DR-6) titled " Final Report, Test Fill Program Rcsults and Backfill Reco mindations for i
- Class II Backfill, Brunswick Steam Electric Plant, Southport , l North Carolina."
Design Report No. 4 is submitted in partial fulfillment ol' our c ontmi tmen t in ree,ard to ACRS Concern No. 3.0 as stited in P.iAR Suppl. ment -p . .:.u No. 6 The report addresses itself specifically: ti Ite- I where a comi t ted t o provide a "summar:. o f cen t a i nmi nt ana' ,i- i "c ! nd i n,; ca'ed o.u! p l o t .:" ind to 1 ten 2 wher6 m co- ittud to pro..' " a r ' h <pi .n . 4 <n mi c l
. :di ng condi t ion < on critica ! centrol, and i n t rumen t ,: "
't h i - r. u.rt
, '.o partially fulfills Tschnical I-sue 8.3.1, "3tructu . :n inn," a t.ited in the "S.i f c t y C"a ! ...n t en B: the Di ision .f '
4c : ,r 1. i c .e n . ;, U . :) .
.\ tomic 1:nergy Co .ni s s ion , in thi htter of Care;!na Pwi r o Light Company, , l
' lirun c. i ck S t eam F. !i c t r : c Plant -
L* n i t s 1 and 2, Doc' ..t - ' . . . 50-12!. and 30-325." Thi report 1. r e - p.>n i ' i to pa r a,;r aph - on, ar.d tv. of Technical l I s. >.u c 8.3.1. which stats. l 1
"We ., f11 re'.ic. the a p p l i c .2n t ' s d,n_tmic loading '
criteria and tht atress and deformation charact.ristics employed ,in di<ign o!' the critical in s t rumer.t a t ion ind ,,. . control equip ent. (Di vi s ion of Reac tor 1.icens in;; (DRL) ' question 5.1.11). The Spplicant agreed to submit the 8 information prior to .' lay 1970. kV -
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x Dr. Peter A. Morris i 2-November 9, 1970 "We will reviev the details of implementation of the dynamic analyses reflecting the interaction between the. ground and: the structures to assure that symmetrical and unsymmetrical lateral forces, which may arise from carth-quakes and which may act on those parts of the structures
,o. located below ground,.will be suitably provided for in the design.
(DRL questions 5.2.1(b).. The applicant agreed to furnish this information for our review prior to construction." is necessary to revisc our structural backfill material it to requiremen permit usage of backfill' material which has a percentage of fines passing a No.;200 sieve in.the range of 15 to 40 percent. has been designated as Class II backfill to differentiate This backfill materi.il from Class backfill which was originally proposed in the PSAR. Investigations atI the
.insite-have disclosed that there is not sufficient structural fill available reasonable proximity =to the plant with the physical pfoperties designated as Class I fill. It ,
should be emphasized that:the Class'I and Class II designations systems and that in no way refer to seismic classifications of structures and in the field. the use of these designations has been adopted for simplicity We informed your staff that extensive test fill compaction programs have been conducted at the si'te of the Brunswick Steam Electric Plant to establish construction control procedures and optimal compaction methods for Class I and Class II backfiil materials. The test fill programs have shown vided that'placement specific Class II procedures hackfill can arebe substituted for Class I backfills followed. pro-we submit the reports describing the test programs and recommendationsYour for staf f req
; placement and compaction of Class I and Class II materials for their revie w.
The' reports designated Design Rcport No. and Design Report No. 5 pertaining to Class I backfill, response t'o their request.6 pertaining to Class II backfill are submitted in Brunswick project, Dr. Robert V. Whitman, Soils Consultant for the curs with the findings expressed therein,has reviewed draf ts of Design Report September 30, 1970, attached to Design Report as noted No. S. in his letter dated As noted above, this 6 letter transmits Design Reports Nos. 4 ., 5, ,nd Ii We have adopted this numberleg system to facilitate tion transmitted to you in response reference to informa-Permit review and subsequent requests by your staff. to commitments made during the Con:.t ruct ion > mitted and 3: to you the following reports now designated as Design Report Nos.W 1, 2, have previou
- 1. 1 i Design Report No. I titled " Proposed In-service Inspection I Program and 2," for the Brunswick Steam Electric Plant, Units !
t ransmitted to you by our let ter of March 9, 1970. ~
- 2. Design Report No.
2 titled " Proposed Containment Design i Pressure 2," for the Brunswick Steam Electric Plant, Units 1 and also transmitted to you by our letter of March 9, 1970. I . e" - _ _ . b
Ti l* % r-
. ., ;. . . / ,
- i. Dr. Fat;r A. Morris- ' November-9, 1970
.3. Design Report No. 3 titled " Additional Information -
Pre'ssure Suppression Concept - Test-Data Report" transmitted to you by our letter of May 19, 1970. We hope that the reports transmitted by this letter will enable your staff to conclude their review of the items outlined above. Yours very truly, 7 '% J. A. Jo. s Senior Vice President Engineering & Operating JAJ/lef-6 1
, 4 o Attachments
, AFFIDAVIT OF J. A. JONES I,'J. A. Jones, being duly sworn, depose and state that I am Senior Vice President and Group Executive fo? Engineering and Operating, Carolina Power & Light Company, and am fully cognizant of the contents of.the attached document containing the reports titled, " Seismic Analysis of the Primary and Secondary Containments," " Final Report, Test Fill Program Results and Backfill. Recommendations for Class I Backfill, Bruns-wick Steam Electric Plant, Southport, North Carolina," and " Final Repodt, Test Fill Program Results and Backfill Recommendations for Class II Backfill,*' Brunswick Steam Electric Plant, Southport, North Carolina" and that the contents of these reports are true and correct to the best of my knowledge.'
- (I74. W J. U ' Jones Subscribed and sworn to before me this 9th day of November, 1970, at Raleigh, North Carolina
.ih')(adec) / ). 0 % -
CNotary Public My commission expires: O 44j h /N 1 m'_ _ _ _ . _ . _ _ _ _ _ _ _ _ _ _ . .
Commonwea: Edison Quad-Cities Nuclear Power Station Post Office Box 216 Cordova. Illinois 61242 @ latory ,@# C" Telephone 309/654-2241
/ D l\ I a # ,*/ F M, ), -
& coextTto usAtc BBS-73-32 '
6- 'V/' r ' 5y- i d: MAR 1 1973 * [5
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*: .A d REGULATORY MAIL SECTION l1 February 26, 1973 4 " " ' # , 4Ndh
- Mi' v
lo A s a co n jco/ l Mr. Angelo Giambusso l Deputy Director of Reactor Projects Directorate of Licensing U. S. Atomic Energy Commission i Washington, D. C. 20545 Re ference: Quad-Cities Nuclear Power Station, Unit 1 Docket No. 50-254, License DPR-29, Appendix A Section 1.0. A.2 and 3 2. A.
Dear Mr. Giambusso:
The purpose of this letter is to inform you of the details concerning the setpoint drift of dPIS-1-2353 The setpoint was found to be at 151 inches of water increasing during a routine calibration on February 15. 1973 This is contrary to Section 3 2. A. of the Technical Specifications ; which requires the switch to trip at 1300% (150 inches of i water). The switch is a Barton Model 280. This switch is used to measure a dP across an elbow tap in the steam line of the HPCI s:; 3 tem and will isolate the system on a high flow. The one inch drift of the switch is considered insignificant with respect to the isolation function. Drift problems have been encountered previously with this instrument and a setpoint change from 150 inches to 147 l inches of water was made in January. Apparently this is still an inadequate margin to prevent drif t beyond the LCO ; during the interval between calibrations. A further decrease in setpoint is being reviewed to increase this margin. I Very truly yours, COMMONWEALTH EDISON COMPANY ? QUAD , CITIES NUCLEAR POWER STATION 1 4/L , e on Superintendent -d /2 28/o 1403 j 9 BBS/zm I _ _ _ _ _ _ _ _ _ _ _ _ _ _ TJ
c Commonwealth son Quad-Cities Nuclear Power Stction _pow fatory File @ Post Office Box 216 Cordova, Illinois 61242 Telephone 309/654-2241 u;, . . W D , g
! A 1
1 ~ DOCYETED UWC f ((gSj
=
% j973 r m IE FEB211973 * {3 7!
BBS-73-22 I *
' mouuTORY blAIL SECTION DOCET CERK 6 /
February 16, 1973 co Mr. A. Giambusso
- Deputy Director for Reactor Projects Directorate of Licensing U.S. Atomic Energy Commission Washington, D. C. 20545
Reference:
Quad Cities Nuclear Power Station Unit 1 Docket No. 50-254, License DPR-29 Appendix A, Section 1.0. A.2, 3 1, and 6.6.B.
Dear Mr. Giambusso:
The purpose of this. letter is to inform you of the details concerning the failure of a limit switch on a Unit 1 Main , Steam Isolation Valve. The function of the limit switch is to provide a valve closure input signal to the Reactor . Protection System. This abnormal occurrence was reported to you by telegram on February 3,1973 4 DESCRIPTION OF INCIDENT On February 3,1973 with the Unit 1 reactor operating at 35% power, a functional test was conducted on the Main Steam Isol.ation Valve -(MSIV) instrument channeis. This monthly surveillance test is conducted by removing a fuse in one ' instrument channel and then closing an MSIV in a second steam line to the 90% open position to verify that an RPS channel trip is obtained. When MSIV 1C was exercised, an RPS trip was not received. All other valves properly de-energized their respective RPS relays. INVESTIGATION Each MSIV has limit switches which de-energize two RPS trip ! auxiliary relays. The relays are then combined in logic such that isolation of two steam lines is allowed before a full reactor scram would be initiated due to valve closure. \ M. wwesey 1211 0 ;
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- Mr. A. Giambusso February 16, 1973 The limit switches are normally closed and open.when the valve is <90% open. Investigation revealed .that one of the two RPS relays monitoring the C steam line was not dropping out when the 10 valve was exercised. The fuse was removed for this relay, 590-102D, to place the channel in a tripped condition.
Prior to a drywell entry to inspect the val've on February 3, it was exercised again and this time the limit switch operated s atis f actorily . The valve was then operated repeatedly with the fuse re-installed and the 590-102D relay dropped out properly on each occasion. . The surveillance test was also repeated and the RPS trip occurred as expected. The switch continued to operate properly during the twice weekly valve exercise surveillance until the unit was shut-down the same week to repair a condenser tube leak. On February B,1973 during the outage the 1C MSIV limit switches were removed by station maintenance personnel and disassembled. The switches operated freely by hand and' appeared to be in good condition. What appeared to be dried lubricant was found under sone of the contact arms, however, it is not believed that this could have caused the failure. The switches were all cleaned, reassembled and installed. The I' valve was then tested satisfactorily several times. CONCLUSIONS The safety significance of this failure is that a closure ! l of the 10 valve combined with isolation of the A steam line and B or D lines would not have produced a scram. Isolating either combination of these three lines however, would have resulted in a high flux scram or a,high flow isolation. Any isolation would have produced a scram from the remaining operable valve limit switches. The only feasible event which could cause an isolation of only three lines would be an air system f ailure or a coincident failure of both MSIV's in one line when an isolation occurred. The air system failure would result in a very slow closure of the va3ves and would affect all four steam lines. 7he probability of three lines isolating before the valve in the fourth j j reaches 90% open is considered to be extremely remote. The 4 coincident failure of two valves in the same line is also j extremely unlikely since all valves are exercised twice a week. I o
l i . j c, Commonweal. idison Que.d-Cities Nuclear Power Station Post 0" ice Box 216 . Cordova, Illinois 61242 Telephone 309/654 2241 2,0 Th ' d BBS-73-23 2 f % y
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(o cp Regulatory File Cy. Mr. A. Giambusso Deputy Director for Reactor Projects Directorate of Licensing U. S. Atomic Energy Commission Washington, D.C. 20545
SUBJECT:
Quad-Cities Nuclear Power Station License DPR-29, Docket No. 50-254, Section 6.6.C l
Dear Mr. Giambusso:
The purpose of this letter is to inform you of a condition relating to the operation of the station which took place on January 16, 1973 On that date Unit 1 was being brought up in power with recirculation flow for 100 per cent test-ing purposes. .With the unit at a core thermal power level of 2300 Wt and generator output of 770 We, a noise and relatively high frequency vibration were heard in the Unit 1 reactor building in the vicinity of the pressure suppression chamber. These vibrations persisted between recirculation pump speeds from 84 per cent to 87 per cent. An effort was made to pinpoint the vibration, and it was determined that the vibration was greatest at the point where the "B" LPCI pipe penetrates the drywell. On January 17, following a One-Recirculation Pump trip test, the unit was brought up in power to duplicate the plant conditions of the previous day. A drywell entry was made and excessive vib-rations were noticed on the "B" recirculation loop piping. The "A" recirculation loop piping and the "B" recirculation pump motor did not appear to have any abnormal vibration. Af ter the Two-Recirculation Pump trip test on January 18, reactor power was decreased and the unit was placed in the hot-standby ccndition. A thorough inspection of the Unit 1 Drywell was conducted by Quad-Cities, General Electric, and Sargent & Lundy personnel. No evidence of any damage or
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Mr. A. Giambusso February 16, 1973 abnormal conditions was detected on the recirculation or RHRS piping. The . recirculation pumps, pump motors, and piping supports were also examined and no defective con-ditions .were observed. A vibration detector with a recorder was installed on LPCI valve 1-1001-29B over the pressure suppression chamber for the purpose of monitoring vibration during power increases..
.On January 20, the unit was placed in the RUN mode and load was increased at the rate of 10 MWe/hr. On January 21, when the power level and pump speed conditions were again duplicated, no abnormal vibrations or noise occurred.
Recirculation pump speed tests were run between 80 and 90 per cent, and the vibration recorder on valve 1001-29B showed no abnormalities. Reactor pressure was also varied with no effect. . Following a load decrease to 70 per cent, vibrations were again monitored in the 80 -90% pump speed range on January 22. No excessive vibrations were observed. Further vibration monitoring has been discontinued for the present time. The Commonwealth Edison Company Mechanical and Structural Engineering Department has been consulted on this problem and has been requested to design and install a permanent monitoring system. This installation will in-clude a permanent vibration detector with indication in the control room, and associated annunciator. Any information obtained in ,the future on this subject will 'be reported to you. Very truly yours, COMMONWEALTH EDISON COMPANY QUAD-CITIES NUCLEAR POWER STATION
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$.tjREFERENCE1 QUAD CITY NUCLEAR STATION UNIT 1 DOCK IT 50-254, DPR-29 APPENDIX %.SECTIONI.O.A.2, 3.2.B.DURING A ROUTINE CALIBRATION QM$pFj.THE:ECCS/
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i 1 4 ' November 27, 1972 Dockat No. 50-254 Dr. Chester P. Siess Chairmn, Advisory Comittee on Beactor Safeguards U. S. Atomic Ehergy Com1csicri Washington, D. C. 20545
Dear Dr. Siess:
sixtema (16) copies of the following are transmitted for the Comittee's informatient i TWK from Commonwealth Edison Compaar dated November 25, 1972, reporting that two main steam line flow evitches drif ted beyond the Quad-cities Unit 1 Technf cal Specifi-cation limits. Sincerely, Diptribu tion W ocket File , Branch Reading
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October 16, 1972 Docket No. 50-254
' Dr.'Chester P. Siess Chairman, Advisory Conunittee on Reactor Safeguards U. S. Atomic Energy Coasnission Washington, D.- C. 20545
Dear Dr. Siess:
Sixteen (16) copies of the following are transmitted f information: or the Conunittee's
- Letter from Cossmonwealth Edison Company dated October 3 1972, reporting a condition in which three of the Quad-Cities Unit i suppression chamber to drywell vacuum breakers were found to' have an incorrect position indication. I
_ Dis tribution SIDC8f*17s d ocket File Branch Reading
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Donald J. kovholt Assistant Director
, for Operating Reactors Directorate of Licensing
Enclosures:
'As stated '
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October 4, 1972 Docket No. 50-254 Dr. Chester P. Diess Chainnan, Advisory Comnittee
.on Beactor Safeguartis U. S. Atomic thergy Coninission Washington, D. C. 20545
Dear Dr. Siess:
#8 Sixteen (16)
Conmittee's information:
'1%'I from Comanonwealth Edison Company dated September 29, 1972, reporting an incident regarding the 12th suppression chamber to drywell vacuum breakers.
. Distribution
*Tocket File Branch Reading Sincemly,
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/ k Donald J. olt Assistant Dimetor for Operating Beactors Dimotorate of Licensing As stated l .h omer , L:0R , ,, , , , , , ,
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[ REGION til 799 ROOSEVELT ROAD yggjgpgogg GLEN ELLYN. ILLINOIS 6o137 0 12) 858-2660 September 25, 1972 J. G. Keppler, Chief, Reactor Testin'g and Operations Branch Directorate of* Regulatory Operations, Headquarters COMMONWEALTH EDISON COMPANY LICENSE NO. DPR-29 QUAD-CITIES UNIT 1 DRYWELL TO TORUS VACUUM RELIEF VALVE - FAILURE TO CLOSE The attached inquiry report is forwarded for information. The licensee intends to notify L within ten days in accordance with Section 6.6.B.3 of the Technical Specifications, <, This item hhs been added to the OII list and information related to the testing, modifications, and corrective action will be obtained during the next routine, unannounced inspection. Drywell to torus vacuum breaker (vacuum relief valve) malfunction during surveillance testing at Quad-Cities Unit 2 is reported in Inquiry Report No. 050-265/72-1S. Jh srkwW D. M. Hunnicutt, Chief Reactor Testing and Startup Branch
Attachment:
RO Inquiry Rpt No. 050-254/7?-13Q ' (2 cys) ( I cc: H. D. Thornburg, RO J. B. Henderson, R0 R. B. Minogue, RS (3) R. S. Boyd, L (2) R. C. DeYoung, L (2) D. J. Skovholt, L (3) H. R. Denton, L (2) P. A. Morris, RO R. H. Engelken, RO RO Files DR Central Files
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REGULATORY OPERATIONS, REGION III~ A. RO Inspection Report No. Transmittal Date ,a Distribution: Distribution: R0 Chief, RT&OB or RO Chief, RCB RO Chief, M&FFB RO:HQ (5) RO:HQ (4) j DR Central Files L:D/D for Fuel & Materials Regulatory Standards (3) DR Central Files Licensing (13)- B. RO Inquiry Report No. Transmittal Date : Distribution: Distribution: R0 Chief, KT&OB or. R0 Chief, RCB (2) .R0 Chief, M&FFB RO:HQ (5) RO:HQ DR Central Files
- DR Central Files Regulatory Standards (3)
Licensing (13)
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C. Incident Notification From: Commonwealth Edison Company - Docket No. 50-254 (Licensee & Docket No. (or License No.) Transmittal Date. : December 1. 1972 Distribution: Distribution , R0 Chief, RT&OB or RO Chief, RCS R0 Chief, M&FFB RO:HQ (4) RO:HQ (4) Licensing (4) L:D/D for Fuel & Materials IR Central Files DR Central Files t 9
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a U. S.~ ATOMIC ENERGY COMMISSION'
' DIRECTORATE OF REGULATORY OPERATIONS REGION III
.R0 Inquiry Report No. 050-254/72-13Q -
Subject:
. Commonwealth Edison Company
' License No. DPR-29 Quad-Cities Unit 1 Drywell To Torus Vacuum Relief Valve - Failure to Close
' Prepared By:
'sh?N}h M D. C. Boyd 9/If/71 f V(Date')
. A. Date and Manner AEC'was Informed:
- 1. TWX,Mr.F.A.-PalmertoRegionIII,rece,ihedon9-25-72.'
- 2. Telecon, Mr. F. Palmer (Station Superintendent) to Mr. D. Boyd .
on September 25, 1972.
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B. Description of Parti ular-Event or Circumstance: On Sep tember . 22, 1972,- the Unit I reactor was shutdown for the purpose of, modifying the drywell to torus vacuum relief valves. After cooldown and de-inerting, an entry was made to determine the undisturbed'(as found) position of each of the twelve vacuum relief valves. - Nine were found to be fully closed, as required. However, three of the vacuum relief valves were found to be partially open (two were v1/16" open; one was v3/4" .open). C. Action By the Licensee: The licensee is in the procese of' modifying all twelve of these
-vacuum relief valves in the same manner as the Quad-Cities Unit 2 valves were modified. This modification includes repositioning of shaft key such that it does not bind on the shaf t packing; repositioning of the counterweight arm and counterweight to assure that the valves have sufficient closing force; and readjustment of the position-limit switches'to assure that the switches accurately indicate the fully open and fully closed positions.
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GLEN ELJ.YN ILL (s.C/ 3 7 REFERENCE QUAD-CITIES NEC ULAT POWER STATION UNIT 1 LICENSE DPR-29 APPENDIX A SECTION 1 0. A.2, 3 2.A AND 6 6. A.3 WHILE CONDLCTING ROUTINE SURVEI'LLAbCE CALIBRATION ON THE UNIT 1 MAIN STEAM LINE FLOW SWITC HES, TWO SWITC HES WERE FOUND TO HAVE DRIFTED BEYOND THE TEC H SPT.C L1 HIT. ThE SET POINT EQUIVALENT TO 120 PERCENT FLOW IS 100 PSID DPIS 1-2 G1-2C TRIP AT 104 PSID DPIS 1-2G1-2,D' TRIP AT 102 PS1D. M rs SWI AC nta nnVr., bEEN " REC ALIBRATED - N J KALIVI ANAKIS u.= m NOV 0 01972
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ATOMIC ENERGY COMMISSION g y DIRECTORATE OF REGULATORY OPERATIONS t REGION lli
. e e d. 799 ROOSEVELT ROAD m,a GLEN ELLYN. ILLINOIS Sol 37 (alg) ese-asse September 1 , 1972 J. G. Keppler, Chief, Reactor Testi* gn and Operations Branch Directorate oS* Regulatory Operations, Headquarters COMMONWEALTH EDISON COMPANY, LICENSE NO. DPR-29 QUAD-CITIES UNIT 1 RCIC SYSTEM - - LIMITORQUE VALVE FAILURES The attached inquiry report is forwarded for information.
During the monthly surveillance testing of the RCIC system, three limitorque valves failed to operate properly. One limitorque valve failed to operate in the RCIC system on Unit 2 (
Reference:
Inquiry Report No. 050-265/72 11, dated July 28, 1972) during a test on July 27, 1972. A third failure of limitorque valves is discussed in RO Inquiry Report Nos. 050-282/72-03 and 050-306/72-02 dated June 15,197?. The licensee intends to report this occurrence to L within ten days in accordance with Section 6.6.B.3 of the Technical ~ Specifications. This item has been added to the OII list and information related to this occurrence will be obtained during the next routine, unannounced inspection. hM knm5csf D. M. Hunnicutt, Chief Reactor Testing and Startup Branch
Attachment:
RO Inquiry Rpt No. 050-254/72-12 (2 cys) cc: H. D. Thornburg, RO J. B. Henderson, RO [9 D R. B. Minogue, RS (3) ,,
- f [l R. S. Boyd, L (2)
R. C. DeYoung, L (2) Tai U l [% c'd r 4 D. J. Skovholt, L (3) --> H. R. Denton, L (2) SEP141972
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' s i~t P. A. Morris, RO 9 . g!:fIJ[3 R. H. Engelken, RO -'l F ,i . d)
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l b f-l,., U. S. A'IOMIC ENERGY COMMISSION DIRECTORATE OF. REGULATORY OPERATIONS REGION III , RO Inquiry Report No. 050-254/,72-12
Subject:
Commonwealth Edispn Company
- l License No. DPR-29 j.
Quad-Cities Unit 1 RCIC System - - Limitorque Valve Failures Prepared By: D '#' ' (Date) i A. Date and Manner AEC was Informed:
- 1. Telecon, Mr. B. B. Stephenson, Assistant Plant Superintendent, to Mr. D. C. Boyd, on September 3, 1972.
- 2. TWX, Mr. B. Stephenson to Region III, received on September 5, 1972.
B. Description of Particular Event or Circumstance: - The licensee reported that on September 3,1972, during the monthly surveillance testing of the RCIC System, three limitorque valves failed to operate properly. Valves 1301-26 (torus supply to pump - normally closed) and 1301-48 (inboard valve on main coolant injection line to the RPV - normally open) failed to close during the testing l cycles. Valve 1301-49 (outboard valve on the main coolant injection line to the RPV - normally closed) failed to open. C. Action by Licenseei The licensee declared the RCIC system inoperable and immediately performed the required functional testing of the HPCI (Technical Specification 3.5.E.2) to permit continued operation of the reactor. 4 1 , . i o e
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'The licensee reduced power level from 50 percent, at which the unit was operating, to 25 percent to permit personnel access to these valves which are located in the steam tunnel.
The licensee's investigations nevealed the following:
- a. 4301I 49 (outboard valve on the main coolant injection line to the RPV - normally closed) - - - the cause of failure of this valve to open was determined to be due to a faulty torque.
switch. This torque switch reset spring (Limitorque Model SMB-00) had failed to return, thus rendering the valve inoperable. A new torque switch was installed and the valve was functionally tested four times to verify its operability.
- b. 1301-48 (inboard valve on the main coolant injection line to the RPV -- normally open) - - - The cause of failure of this valve was determined to be that the 250 volt D.C. breaker cpunter spring was out of adjustment such that the bresker would not pull in by itself. This counter spring was adjusted, and the system was functionally tested four times to verify -
. its operability. '
- c. ' 1301-26 (torus supply line to pump suction -- normally closed)
- - - The cause of failure of this valve was determined to
, be that the 250 volt D.C. breaker counter spring was out of adjustment such that the breaker would not pull in by itself.
This counter spring was adjusted and the breaker was functionally tested four times to verify its operability. Following the above repairs and testing, the entire RCIC surveillance test was repeated., All portions of the test were performed successfully. The licensee plans to inform the Directorate of Licensing of this occurrence by letter within ten days as required by Technical Specifications Paragraph 6.6.B.3. t *
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l JUL 2 61972 Docket Mos. 50-254 i - and 50-265 Commeemaalth Edison Company ATTN 4 Mr. Dyron Lee, Jr. Assistaat to the President Post Office Box 767 Chicago, rihis 60690 Gentlemen; I y
-uring recent meetings with representatives of your Company, required changes were developedin the and Quad-Cities discussed.Units 1 and 2 Technical Specifications (TS)
A number of these changes will be made prior to our issuance of a licasse forscontimmeus operation at 90% of rated your letterpower dated with the7,1972.}}