ML20079G353
| ML20079G353 | |
| Person / Time | |
|---|---|
| Site: | Clinch River |
| Issue date: | 06/01/1982 |
| From: | Longenecker J ENERGY, DEPT. OF |
| To: | Check P Office of Nuclear Reactor Regulation |
| References | |
| HQ:S:82:038, HQ:S:82:38, NUDOCS 8206080410 | |
| Download: ML20079G353 (98) | |
Text
e, Department of Energy i
3, Washington, D.C. 20545 Docket No. 50-537 HQ:S:82:038 JUN 011982 y
Mr. Paul S. Check, Director CRBR Program Office 1
Office of Ruclear Reactor Regulation U.S. Nuclear Regulatory Commission Washington, D.C.
20555
Dear Mr. Check:
RESPONSES TO REQUEST FOR ADDITIONAL INFORMATION - STRUCTURAL ENGINEERING
Reference:
Letter, P. S. Check to J. R. Longenecker, "CRBRP Request for Additional Information," dated February 26, 1982 This letter fonnally responds to your request for additional information /
contained in the reference letter.
Enclosed are responses to Questions CS 220.4, 7, 8, 9, 11,,12, 13, 14, 16, 17,18,19, 23, 26, 28, 29, 31, 32, 33, 36, 37, 38, 40, 41, 42,'43, 44, and
- 45. The responses to the remaining questions will be provided by June 14.
These responses will also be incorporated into the PSAR Amendment 69; scheduled for submittal later in June.
Sincerely, m.
Cr
- Le JohnR.Longeneckr, Manager Licensing & Envitonmental Coordination Office of Nuclear Energy Enclosures cc: Service List Standard Distribution Licensing Distribution N
O 8206080410 820601 PDR ADOCK 05000
Yag) - D (CZ-U154)LU,2LJ - 935
}
J v
Duestion CS220.4 f
Seismic design is presented in Section 3.7 supplemented:by Appendix 3.7-A.
However, a review of the section and its 'ippendix will-reveal that there is quite some repetition in the Appendix of the materials presented in the body i,
of the section.
- t.
/<,
The-Appendix also mentions loads and load combinations which are delineated in Section 3.8.
The presentation of the materials in this manner not ony consumes effort in preparation by you and in review by the staff unnecessarily, but also may lead to contradiction and confusion.
It is therefore proposed that Appendix 3.7-A be revised to eliminate materials which are not contained in Section 3.7 and 3.8.
/
f.
Resoonse:
Appendix 3.7-A is a self-contained appendix to the body of the PSAR Section 3.7.
As an appendix, it is, by its very Intent, not required to eliminate Information presented in the body of the PSAR but to supplen,ent it with more details. This necessitates a certain amount of duplication and repetition.
We believe that the additional information will be helpful in the NRC review.
To assist the NRC staf f in their review, the applicant,nas prepared a cross ref erence (ref. Table CS220.4-1) of inf ormation for your use.
~
l
/
/
QCS220.4-1 Amend. 68 May 1962
Pago - 0 (62-0154)LU,22J - ase Table CS220.4-1 SUBJECT PSAR SECTION(S) 1.
SEISMIC DESIGN PARAMETERS 2.5, " GEOLOGY AND SEISMOLOGY" 3.7-A PARAGRAPHS 3 AND 4, " SITE DESCRIPTION" AND "0PERATING BASIS FOR SAFC SHUTDOWN EARTHQUAKES".
1.
DESIGN GROUND MOTION a) Design Response Spectra 2.5.2.10 and 2.5.2.11 3.7.1.1 and 3.7.1.2 b) Des!gn Time History 3.7-A.4.3 3.7.2.1.1, 3.7.2.1.2 and 3.7.2.3 2.
CRITICAL DAMPING VALUES 3.7.1.3 3.
SUPPORTING MEDIA FOR 3.7.1.5, 3.7.1.6 CATEGORY I STRUCTURES l
l i
l QCS220.4-2 Amend. 68 May 1982
Pcge - 7 (62-0184)L8,22J - #36 a
1 l
SUBJECT PSAR SECTION(S) i ll.
SEISMIC SYSTEM ANALYSIS 1.
SEISMIC ANALYSIS METHODS a) Dynamic Analysis Methods 3.7.2, 3.7-A.6 and 3.7-A-A b) Equivalent Static Load Method 3.7.2.1.2 and 3.7.A.2 2.
NATURAL FREOUENCIES AND 3.7.2.2 RESPONSE LOADS 3.
PROCEDURES USED FOR 3.7-A-A ANALYTICAL MODELING 4.
S0ll-STRUCTURE INTERACTION 3.7-A-C, 3.7.1.6 5.
DEVELOPMENT OF FLOOR RESPONSE 3.7.2.6 SPECTRA 6.
THREE COMPONENTS OF EARTHOUAKE 3.7.2.1.1 and 3.7-A-B MOTION 7.
COMBINATION OF MODEL RESPONSES 3.7-A- A.1.3 3.7.3.7 8.
INTERACTION OF NON-CATEGORY I 3.7.3.13 and 3.7-A.6 STRUCTURES WITH CATEGORY I STRUCTURES l
9.
EFFECTS OF PARAMETER VARI ATIONS 3.7.2.1.1 and 3.7,2.8 i
ON FLOOR RESPONSE SPECTRA 10.
USE OF EOUlVALENT STATIC 3.7.3.9 and 3.7-A-A.2 FACTORS l
11.
METHODS USED TO ACCOUNT FOR 3.7.3.11 TORSIONAL EFFECTS
- 12. COMPARISON OF RESPONSES 3.7.2.11 l
l 13.
ANALYSIS PROCEDURE F0F DAMPING 3.7.2.14 l
l 14.
DETERMINATION OF CATEGORY I 3.7.2.13 STRUCTURE OVERTURNING MOMENTS l
l l
I QCS220.4-3 l
Amend. 68 Ch71962
P ge - 8 ( 82-0184)[8,22] - #36 SUBJECT FSAR SECTION(S) 111.
SEISMIC SUBSYSTEM ANALYSIS 1.
_ SEISMIC ANALYSIS METHOD 3.7.2.1 2.
DETERMINATION OF NUMBER 3.7.3.1 EARTHOUAKE CYCLES 3.
PROCEDURES USED FOR 3.7-A-A ANALYTICAL MODELING 4.
BASIS FOR SELECTION OF 3.7.3.2 FREOUENCIES 5.
ANALYSIS PROCEDURE FOR 3.7.2.14 DAMPING 6.
THREE COMPONENTS OF 3.7.2.1.1 and 3.7-A-B EARTHOUAKE MOTION 7.
COMBINATION OF MODEL 3.7.3.4 RESPONSES 8.
INTERACTION OF OTHER 3.7.3.13 SYSTEMS WITH CATEGORY I SYSTEMS 9.
MULTIPLE-SUPPORTED EOUIPMENT 3.7.2.14 and 3.7.2.7 AND COMPONENTS WITH DISTINCT INPUTS 10.
USE OF EOUlVALFNT STATIC 3.7.2.1.2 and 3.7-A.2 FACTORS
- 11. TORSIONAL EFFECTS OF 3.7.3.11 ECCENTRIC MASSES a2, CnitGORY l BURIED 3.7-A-C. 3 PlPING. CONDUlT AND TUNNELS
- 13. METHODS FOR SElSMIC 3.7.2.12 ANALYSIS OF CATEGORY l DEMO.
QCS220.4-4 m
Page - 9 (82-0184)[8,22] - #36 SUBJECT PSAR SECT 10N(S)
IV.
SEISMIC INSTRUENTATION 1,
COMPARISON WITH REG.
3.7.4.1 GUIDE 1.12 2.
LOCATION AND DESCRIPTION 3.7.4.2 OF INSTRUMENTATION 3.
CONTROL ROOM OPERATOR 3.7.4.3 NOTIFICATION a
4.
f0MPARISON OF MEASURE AND 3.7.4.4 PREDICTED RE'SPONSES 5.
INSERVICE SURVElLLANCE INSERVICE INSPECTION PROGRAM FOR SEISMIC INSTRUfENTArl0N WILL BE PROVIDED IN SECTION 16.4 0F THE PSAR IN A FUTURE AENDMENT.
QCS220.4-5
. Wo%k2BL.
rigg - so scz-unco /LO,44J - FJo Question CS220.7 in Table 3.7-2A damping values are related to the shear strain values.
Indicate how such relations are obtained.
Resoonse:
The dynamic properties for compacted granular fill used in seismic design were established based on laboratory and in situ testing on compacted granular fill used at another nuclear plant site.
Detalis were provided in response to Question CS324.7 and were subsequently incorporated into Section 2.5.4.5.1.5 4
of the PSAR.
PSAR Table 3.7-2A has been revised f or cons'stency with values in Section 2.5.4.5.1.5.
QCS220.7-1 Amend. 68 Ch? D OM
rage zu Lo,itaras TABLE 3.7-2A DAMPING RATIOS FOR FOUNDATION MATERIALS (internal Damping)
CLASS l FILL Shear Strain Damping
'{- I n./ I n.
5
~6 5 x 10 5.6
-5 1 x 10 5.6
-5 2 x 10 5.6
-5 5 x 10 5.6
-4 1 x 10 5.6
-4 2 x 10 5.7
-4 5 x 10 6.7 l
-3 1 x 10 8.0
-3 4 x 10 14.8
-2 1 x 10 17.2 BDCE SSE: 2% Damping OBE:
1% Damping 1
I i
l I
3.7-24a
__pw.vn _f.R
Paga - 15 (82-0184)[8,22] - #36 Question CS220.8 (3.7.1.6)
It is stated that the input motions shall be applied at the surf ace level (finished grade) on an assumed rock outcrop and shall consists of the rock motion used in the analysis of the Nuclear Island and that no credit shall be given for soli cover on overburden in the deconvolution. Clarify this statement and provide a full discussion on how the analysis will be done.
Resoonse:
The application of the input motion at the top of an assumed rock outcrop is the result of NRC Instructions (Ref. QCS220.8-1) with regard to a FLUSH analysis of the Nuclear Island contemplated during 1976.
For consistency in the input, this method was used in the analysis of the seismic Category 111 structures (Turbine Generator and Radweste Buildings,)
adjacent to the Nuclear Island. The computer program FLUSH was used in these analyses.
The deconvolution consists of determining the input motion to be used at the base of the FLUSH mathematical model (of the soll-structure system) for an input motion defined at the surface. This was doen by using a profile representative of the " free-field" In which the rock was extended to the finished grade, disregarding the overburden or backfill. The FLUSH models for the soil-structure interaction, however, represented the actual profile including soll and rock.
Strain dependent properties of the soll were used.
.'..._:,___. r; ;;rf er..:d for a range of soil-rock properties and the results enveloped.
Since the FLUSH anelysis is two-dimensional, separate models for the North-South and East-West directions are required.
For the Diesel Generator Building, which is supported on soll, the soil-structure interaction will be performed with FLUSH.
In the deconvolution, the input motion will be applied at the finished grade using i
the actual profile including soll and rock. The response spectra at the "f ree-field" foundation level will envelope the Design Response Spectra.
rann sections a.i.e.o and 3.7.2.1.1 have been upgraded to include the above information.
Pc'e-care:
QCS220.8-1 Letter of July 2,1976 from Themis P. Spies of USNRC to Lochlin W. Caffey of CRBRP.
l QCS220.8-1 Amend. 68 w
rvra
Paga 3 L8,22J739 3.7.1.6 Soll-Structure Interaction 3.7.1.6.1 Structures Sueoorted on soll The seismic analysis of the Diesel Generator Building (DGB) which is a seismic Category I structure supported on soII, will be conducted using the finite element computer program FLUSH. The analysis accounts for the strain dependent properties of the soll. Since FLUSH is a two-dimensional computer program, analyses for the North-South and East-West directions will be required.
The input motions will be applied at the surf ace (finished grade) of the
" free-field" and will consist of the motions used in the analysis of the Nuclear Island. The strain dependance of the soll properties will be accounted in the analysis.
Analyses will be performed for a range of soll properties and the results will be enveloped. The envelope response spectra of the foundation level, at the
" free-field", will not be below the Design Response Spectra.
As described in Section 3.7.2.1.1.3 the motions calculated at the foundation of the structure will be used in a three-dimensional analysis of the structure. The peaks of the In-structure spectra envelopes for the range of soll properties will be widened by i 10%.
For Category I buried structures, duct banks, pipes and electrical manholes, the nethod described in Section 3.7.3.12 will be used.
Two Category 111 structures adjacent to the Nuclear Island (Turbine Generator and Radwaste Buildings) are supported on soll and are seismically analyzed for the SSE.
Following the NRC Instructions of Ref.13, the FLUSH analysis was performed assuming the input motion applied at the top of a rock outcrop at l
the finished grade elevation.
Further description of the analysis at these structures is given in Section 3.7.2.1.1.5.
3.7.1.6.2 Structures Sueoorted on Rock The major Category I structures are founded on competent rock with an average ch::r wave velocity of 4000 ft/sec.
3.7.1.6.2.1 NucIsar Island
)
The Nuclear Island consists of the fol:owing interconnected buildings with a l
common foundation mat: Reactor Containment (RCB), Confinement, Reactor Service (RSB), Steam Generator (SGB), Control (CB) and Electrical Eqiupment 1
(EEB).
The rock structure interaction of the Nuclear Island will be analyzed as follows:
i 3.7-3 Amend. 68 May 1982
Pcge 4 Le,22Ja39 The characteristics and properties of the rock are describec !n Section 2.5.4.4.
The material underlying the foundation consists of layers of slitstone and limestone dipping at an angle of approximately 30 with the horliontal, from West to East. The Nuclear Island Is located directly upon the siltstone. The geological profile shows an upper zone of weathered rock above the continuous (sound) rock. The foundation level (Elevation 715.0 ft) is about 20 to 35 f t.
below the top of the continuous rock and 100 ft. below finished grade (elevation 815.0 f t).
The excavation for the foundation consists predominantly of vertical or near vertical cuts, except at the West side where the excavation profile consists of a relatively small vertical cut at the lower elevations with the rock and soll above slope cut on a 2:1 slope along the bedding plane. The space between the side of the excavation and the plant will be backfilled with lean concrete from base level to the top of the vertical cut. Compacted Class A fill will extend from the top of the fill concrete to grade level.
(Figures 3.7-19, 20)
The rock-structure Interaction has been represented in the analytical models for the seismic analysis by equivalent massless foundation springs and dashpots.
Due to the inclined configuration of the rock strata, half space theory is not directly applicable to this site. To calculate the foundation springs a static finite element approach was used. This type of approach has been suggested by Whitman (Ref. 8) and has been used to calculate foundation springs for embedded structures (Ref. 9,10).
Since the Reactor Containment, Containment, Steam Generator, Electrical Equ!Pr.cni, Centrol and Reactor Service buildings are supported on a common foundation met only one set of rock-springs and dashpots will charactrize the rock-structure Interaction of these structures. The set consists of springs and dampers for translation in the North-Soutle, East-West and Vertical directions and for rotation about axes in the same three directions. The translational spring constants were calculated as the total force required to impose a unit displacement on the foundation mat and vertical boundary of the embedded structures a unit rotation.
In imposing the unit displacements or rotations the foundation mat and vertical boundarles of the embedded structure were assumed to be rigid.
The computer program STARDYNE was used for the finite element calculations.
Tnree oasic mocets were used:
- 1) Model A (North-South Direction)
This is a plane strain model based on a section through a North-South plane along the center of the Containment (Figure 3.7-21) and was used to calculate the horizontal translational and rocking springs for the North-South direction and the vertical translational spring.
3.7-3a Amend. 68 May 1982
Pege > Lts,22Jm The materlei properties are given in Table 3.7-5.
The side and bottom boundaries were located suf ficiently away from the structu o minimize their ef fects; the nodes on the side boundaries, and the nod.
+'a bottom boundary were restrained against both vertical and horizon ment.
Becau
'egular shape of the foundation mat, it was necessary to defir
.m ent rectangular foundation. The length (L) and the width (W) cf tt er foundation are such that the area and moment of inertia abou5
.oldal EastWest axis are the same as those of the actual four
.,n; the following equations were used:
LW (1)
A
=
3 p
(2)
I
=
area of actual foundation mat Where:
A
=
moment of inertia of acutal foundation mat about centroidal i
=
EastWest axis From equations (1) and (2), "L" and "W" were calculated.
For the translational springs (vertical or horizontal) uniform translational vertical or horizontal displacements were applied at the nodes that represent the foundation-and the required forces for a unit displacement were calculated. For the rocking spring a unit rotation was applied and the moment of the required forces was calculated.
In imposing the unit displacements or rotations the foundation mat and vertical boundaries of the embedded structure were assumed to be rigid.
Since the plane strain model assumes a unit width, the calculated values were multiplied by the equivalent width of the foundation "W" to obtain the total spring forces for the two dimensional model: k, kg and k.
H y
To check the adequacy of the model against available solutions for an elastic halfspace and to account for the influence of three dimensional ef fects since the calculated spring stif f nesses were based on a two dimensional analysis the following procedure was foflowed:
a) The model described above was adapted to a halfspace solution by giving to the soII (rock) elements above the foundation level (Elevation 715 ft) a shear modulus G=0 (which is equivalent to the elimination of the embedment) and he material below that level a constant shear modulus G=55, 385 k/ft2 (which is the average for the siltstone). The spring (k, g, k )
constants per unit width of foundation were calculated.
H y
b) Using halfspace theory, for the same "L", "G",
and "V" values as in the modelof(a),thevaluesofthespringconstantperunitwidth(k'Y,.k' k' ) were calculated for dif ferent aspect ratios of the foundation The y
equations used, based on Table 10-14 of Reference 11, were the following:
3.7-3b Amend. 68 May 1982
rege o Lo,4ZJF W 6b (3)
Vertical spring constant:
k'
=
v w (i y)
Eb b
(4)
Horizontal spring constant:
k'
=
0b (5)
Rocking spring constant:
k'
=
W (T-V)
Where G = shear modulus of soll v = Poisson's ratio of soll The parameters B, B, B6 are a function of the ratio and are given in Figure 10-16 of Itef eEence 11.
c) The ratios betweer the values calculated in (b) and those of (a) were calculated (k'/k).
Table 3.7-6 shows the results of (a) and (b) and (c) for the horizontal trerst ational spring and Figure 3.7-22 shows a plot of k'H/kH es a function of
+ha eenoc+ ratio of the foundation, (W/L).
It can be seen that as the ratio W/L increases, the ratio k'" condition of the approaches Since a large W/L ratio approaches the plane strain unity.
finite element calculation and for large W/L ratios the two methods give similar result, a good correlation has been proven.
I From the curve of Figure 3.7-22 for the aspect of the foudnation )( = 0.67 the
= 1.33.
This value gives the correction f actor for three ratiok'NNeffectstobeappliedtothepreviouscalculatedstiffnessk*
dimensio H
I Similar calculations were performed for the rotational and the vertical spiings.
- 2) Model "B" (East-West direction)
This is a plane strain model similar to "A" but based on a section through an East-West plane (Figure 2.7-23) and was used to calculate the horizontal translational and rocking springs for the East-West direction and the vertical translational spring. The procedure was similar to that described in (1) above.
3.7-3c Amend. 68 W
Paga 7 L8,22Jf39 The translational vertical spring constant was averaged with the one obtained with model "A" and this gave the final design value.
- 3) Model "C" (Three dimensional model for calculating the torsional spring)
A three dimensional model was required to calculate the torsional spring.
An equivalent axisymetrical model was used with ons y one - half quadrant and appropriate boundary conditions (Figure 2.7-24).
i 3.7-3cc Amend. 68 May 1982 l
P;ge 7 L8,22Jf39 The nodes on the periphery of the model were restrained against displacwement in their three degrees of freedom. The nodes on the radial planes that limit the model were made free to displace in the tangential direction and restrained in the other two directions. The embedded structure was represented by a circular rigid cylinder with a radius such that the polar moment of inertia of a horizontal section was equal to the moment of inertia of the actual met about its centroid. To find the torsional spring, a unit rotation was applied about the centroidal vertical axis and the moment of the forces required was calculated.
To find the damping coef ficients associated with the springs, for each type of spring the shear modulus of an equivalent elastic half-space that gives the same spring stiffness was calculated.
With this G value, and based on the dimensions of an equivalent base, the damping coefficients for an elastic half-space were calculated and these values were used in the analysis.
Three sets of sp.-Ing stiffness and damping coefficients were calculated which correspond to the upper bound, average and lower bound of the rock (soll) properties. The seismic design is based on the envelope of the results obtained with the three sets of constants.
A verification of the spring stiffness based on half-space theory is shown in the response to NRC Question 130.53.
Since Seismic Category I structures (except for the Diesel Generator Building) are supported on competent rock with an average design shear wave velocity of 4000 fos, a fixed base approach would be Jutified, consistent with established accep.aave u itar io.
However, equivalent lumped springs and dashpots calculated using a finite elanent static analysis are used to account for soll-structure Interaction ef f ects, it is noted that for CRBRP structures the spring constants and associated damping coef ficients are calculated for average of values assigned to dynamic rock properties (125% from the average value). The envelope of responses obtained for this range will be considered i
in design.
l l
t t
3.7-4 Amend. 68 May 1982
Pag 3 0 LU,22JF39 i
The spring constants were calculated as the total force or moment required to f
Impose a unit displacement or rotation, respectively, on the foundation mat i
and vertical boundary of the embedded structures. Thus, not only the rock underlying the structures but also the soft (relative to rock) compacted fill matrefal around them were included in the analysis. To verify that the influence of side springs is negligible, side springs were calculated (using results of the static finite element analysis) and attached at elevations 752.7 f t.,
765.0 f t., 779.0 f t., and 816. f t. of the prel iminary mathematical models.
Table 3.7-3 and 3.7-4 compare natural frequencies, model shapes and associated participation f actors for the governing modes of the mathematical models with and without side springs; it can be seen that the ef fect of the side springs is not significant.
3.7.1.6.2.2 Other Seismic Categorv I Structures The Emergency Cooling Tower (ECT) is the other Seismic Category I structure supported on rock.
Since on the Nuclear Island the foudnation spring stiffnesses calculated by elastic half-space theory were in good.sgreement with those calculated by the static finite element analysis, the damped parameters (spring and dampers) for the ECT rock-structure Interaction will be calculated by the equations f or an elastic half-space.
3.7-4a Amend. 68 May/ 1982
Pcg3 9 LB,22JF39 3.7.2.1 Seismic Analvsis Method 3.7.2.1.1 Categorv i Structures A complete analysis will be performed on each of the Seismic Category I structures to predict its behavior during an earthquake. Tr SSE and OBE will l
be considered; each of two orthogonal horizontal directions and the vertical direction will be treated separately and the results combined. The input motions are described in Section 3.7.1.
3.7.2.1.1.1 Nuclear Island A lumped-mass formulation, with direct integration of the coupled equations of motion will be used.
The Buildings of the Nuclear Island: Reactor Containment (RCB), Confinement, Reactor Service (RSB), Steam Generator (SGB), Electrical Equipment (EEB) and l
Control Building (CB) have a common foundation mat with the bottom at Elevation 715'-0"; finished grade is at El. 815'-0".
Except for the RW Interior structure, the Category I structures are tied together up to the roof level. The RC8 Interior structure is tied to the adjacent structures at the mat level and at the operating floor level (Elev.
816'-0").
The structures and foundation materials will be represented in terms of lumped-masses amd massless springs and dashpots.
The inertial preperties are characterized by the masses and mass moments of inertia which will be lumped at points selected to assure proper representation of the dynamic behavior of the structures.
l The mass points will be, in general, at the elevation of the floors and will l
be located at the center of mass of the contributing elements.
The stif f ness properties are characterized by the areas, shear areas and moments of inertia (for bonding aad torsion) of the members and by the moduli of ata=+!cib/ and Poisson's ratios. The flexible members (beam elements)
Lol.oe.. I;eer lesels will be assumed located at the shear center of their sections and their ends are defined by the elevations of the mass points. The ends of beam elements are connected to the mass points by horizontal rigid elements.
The soll (rock) structure interaction will be represented by equivalent springs and dashpots. The stiffnesses of the foundation springs will be calculated as described in Section 3.7.1.6.
The damping values to be used for the structures in terms of percent of critical damping are given in Table 3.7-2; the combined damping ratios for th.
structures (steel containment and concrete buildings) will be calculated baseo on the equation:
b _~.
{El g-(_,') f#[
K
[#T7(c{@J 3.7-5 Amend. 68 C23tER
Nge 11 Ls,22JF59 Where:
i (K) = assembled stiffness matrix for fixed base structure
[j=equivalentmodaldampingratioofthejthmodeforthefixedbase structure (E)=modifiedstiffnessmatrixforthefixedbasestructureconstructedfrom element matrices formed by the product of the damping ratio for the element and its stiffness matrix.
f = Jth normalized modal vector for the flxed base structure These damping ratios together with the damping coefficients associated with the foundation dampers will be used in the formulation of the damping matrix for the soil-structure system.
The damping coefficients for the foundation dampers (translational, rocking and torsional) will be calculated based on the equations for geometrical damping in an elastcl half-space using equivalent half-space dynamic properties derived from the spring stiffness. The equations of Appendix 3.7-A, Sections C.3.1 and C.3.2 will be used.
[
Three basic mathematical models will be used, one for each directional I
component of the earthquake.
Figures 3.7-16, 3.7-16A and 3.7-16B respectively l
show sketches of the mathematical models for the North-South, East-West and Vertical directions.
Figure 3.7-16c shows a plan of the Nuclear Island and the system of coordinates. Table 3.7-7 gives the coordinates of the mass points (nodes).
The mathematical model consists of four main parts:
- 1) The RSB; 2) The Confinement; 3) The RSB; 4) The EES, CB and SGB.
l The four parts of the model are supported by the foundation met which is assumed to be rigid. This assumption is justified because the mat acts as a diaphragm and is stiffened by the vertical walls of the buildings.
The buildings above the mat are Interconnected by flexible ties which include ern== cou,nl Ing between interconnected nodes.
The stif f ness of the fle'xible ties that Interconnect the nodal points of the four main parts of the structure will be calculated by finite element analysis with the computer program Mll/STARDYNE.
Three dynamic degrees of freedom per node will be allowed on the mathematicle models for the two horizontal components of the earthquake: translational and rotational along the direction of the earthwuake and rotational (torsional) about a vertical axis.
Each of the models for the horizontal components (North-South and East-West) has three foundation springs: translational and rocking along the direction of the motion and torsional about the vertical axis through the mat centroid and the corresponding dashpots (dampers).
3.7-6 l
Amend. 68 l
May 1982
vage il Lo,LLJuay The model for the vertical direction will allow one dynamiv degree of f reedom per node and has only one foundation spring and dashpot (vertical); in this model the dome of the steel containment has been idealized using equivalent springs which account for the " breathing" (shell type of vibration) of the dome during a vertical vibration.
To account for " breathing," a stif f ness matrix of the dome with cross-coupling terms was derived f rom an axiasymmetrical shell model of the dome, using the KALNINS computer program the equivalent springs represent the terms on the stiffness matrix.
The Reactor Vessel has been coupled to the models.
Two computer programs: HETHA for horizontal motions and VETHA for vertical motions, will be used to calculate the structural responses.
Using a formulation for the equations of motion similar to that proposed by TSAl (Ref erence 5), the programs solve the coupled equations of motion by direct integration to obtain acceleration time-histories at each one of the mass-points for the dynamic degress of freedom assumed in the model.
With the acceleration time-histories, floor response spectra will be calculated.
Spectral values will be computed for the set of frequencies given in Table 3.7-1.
In addition, spectral values will be calculated at the natural frequencies of the structures.
P.c:;;..:: Op::t. c will be computed for critical equipment dampings of 2%, 35, 4%, and 7% for the SSE and 15, 2%, and 4% for the OBE.
In addition, to account for the ef fect of possible variations of the structural material properties and danping, and for the relative accuracy of the dynamic calculations, the computed floor response spectra will be smoothed and peaks will be widened within a 105 band.
The responses will be calculated for the upper and lower bounds of the range of foundation mater!al properties; the design response spectra will be the envelope of the corresponding widened spectra for the upper and lower bounds.
Ine responses will be calculated f or nodal points which correspond to centers I
of mass. To find the response at points away from the nodal points, additional linear accelerations caused by rotational and torsional accelerations will be added.
The ef fects of the three earthquake directions will be combined by the rule of the square root of the sum of the squares.
3.7-6 a Amend. 68 l
May 1982
Pag) 13 L8,22Jf39 The time-history of the forces acting on the structures (shears and moments) were calculated using the computer program STARDYNE using mathematicia models similar to those of HETHA and VETHA. The soll spacings and dampers were eliminated and teh acceleration time-histories calculated by HEATHA and YETHA at the foundation mat were used as imput. The spectra at dif ferent locations were calculated to check against those calculated by HEATHA and VETHA.
Peak values of the forces are Identified and envelopes of maximum forces were constructed f or the design of the structures. The envelo;as will be based on the results of the anlayses for the upper and lower bound of the range of the foundation material properties.
3.7-6aa Amend. 68 May 1982
ragO 10 LO,24Jm j
l 3.7.2.1.1.2 Emergenev CoofIng Tower i
Another Category I structure independent of the Nuclear Island and founded on rock is the Emergency Cooling Tower (ECT). A description of this structure is given in Section 3.8.4.1.5 of the PSAR.
The conditions of the foundation material under the ECT are similar to those under the Nuclear Island, i.e., inclined layers of silstone and limestone, with siltstone directly under the foundation mat.
A lumped mass analysis of this structure will be performed.
Since the analysis of the Nuclear Island showed a good correlation between the spring constants for the rock / structure Interaction calculated by the static finite element method and elastic half-space theory (with the properties of siltstone), the springs and dampers for ECT will be calculated by elastic half-space theory In a similar manner. Analysis will be performed for the Upper Bound and Lower Bound of the rock properties and also for a " fixed" base and the results for the three cases will be enveloped. The embedment springs will be connected at appropriate nodes of the structural model. The three input motions, (North-South, East-West and vertical) will be applied simultaneously on the three-dimensional, lumped mass model with six degrees of freedom at each node. The mass points will be located at the center of mass of the corresponding sections (in general, at floor locations). The beam elements, between mass points, will represent the axlal, bending, shear and torsional stif f ness of the structure and will be located at the corresponding shear centers. Time-history modal super-posifl' n analysis will be used with o
c&pmite rnodal dampings calculated by the equation:
f@
b id[
E
[4f'(4 [tG where:
l (K) = assembled stiffness matrix for structure
'.r.t modal damping ratio of the jth mode
(,D = modified stif f ness matrix constructed f rom element matrices formed by the product of the damping ratio for the elanent and Its stiffness matrix.
f = Jth normalized model vector The fluid in the ECT will be treated in accordance with Housner's theory (Ref.
12).
3.7-6b Amend. 68 May 1982
Page 15 LU,22JF39 The analysis will result in occeleration time histories at the dif ferent nodal points, forces in the structural members and floor acceleration response spectra.
(One for each of the six degress of f reedom). Analysis will be done f
for both OBE and SSE conditions.
3.7.2.1.1.3 Diesel Generator Buildings i
The Diesel Generator Building (DGB) is the only major Category I structure founded on soll. This structure is described in Section 3.8.4.1.4 of the PS AR.
The soil-structure Interaction will be treated as described in Section 3.7.1.6 using the computer program FLUSH. A three-dimensional lumped mass model of the structure will be generated using the computer program STARDYNE.
Condensed mass and stif f ness matrices consistent with the two-dimensional formulation of FLUSH, will be calculated f rom this model and used as superelements in FLUSH. Two separate mathematical models for the North-South and East-West directions will be used. The FLUSH analysis will provide seismic responses at dif ferent points in the structure.
A range of soll properties will be used in the analysis and these responses will be enveloped.
The enveloped response spectra at the foundation level will also envelope the design response spectra f or the site.
Since FLUSH is two-dimensional, to do a more detailed analysis of the structure, the acceleration time-histories calculated by FLUSH at the DGB foundation will be used for a three-dimensional analysis with six degrees of freedom per mode using the STARDYNE model. The acceleration in the STARDYNE model.
The responses from this calculation will be verified against those produced by FLUSH. Design response spectra and forces on the structure will be produced f rom this STARDYNE analysis.
Analysis will be done for both OBE and SSE conditions.
3.7.2.1.1.4 Miscellaneous categorv i structures Other Category I structures supported on soll are as follows:
Diesel Fuel Storage Tanks (described in Section 3.8.4.1.6), Electrical Manholes (described in Section 3.8.4.1.7)
Emergency Plant Service Water (EPSW) Pipes (described in Section 9.9.4) and Class IE Duct Banks (described in Section 3.10).
The seismic design of these buried structures will be in accordance with the method described in Section 3.7.3.12.
3.7.2.1.1.5 Categorv III Structures Two Category ill structures, the Turbine Generator (TGB) and Radwaste (RWB) buildings, because of their proximity to the Category I structures, were designed to withstand the effects of the SSE.
The TGB and RWB are supported on soII, and the soil-structure interaction approach is described in Section 3.7.1.6.
3.7-6c Amend. 68 May 1982
H:gs 16 Ls,22JF39 Three dimensional models of the buildings were constructed using the computer program STARDYNE.
Condensed mass and stiffness matrices consistent with the two-dimensional formulation of FLUSH were calculated f rom the models and used as supplements in FLUSH. FLUSH analyses were performed with models for the North-South and East-West directions, for upper and lower bound of soll properties. Response spectra at the foundations were produced and enveloped.
With enveloped response spectra applied at the foundation and using the three-dimensional STARDYNE models, the forces on the members of the structure were calculated by the response spectrum modal analysis method.
The response spectra used in the three-dimensional analysis were the envelopes of the spectra at the foundation level from 1) FLUSH analysis of the buildings
- 2) Seismic analysis of the Nuclear Island.
The latter was done to account for the ef fects of the motions of the massive Nuclear Island on the adjacent buildings.
The spectral envelopes were above the Design Response Spectra for the site.
Accelerations at the buildings calculated by the spectrum analysis were compared with those of the FLUSH analysis as a verification of the analysis.
3.7-6d Amend. 68 May 1982
P;ge - 16 (82-0184)[8,22] - #36 Duestion CS770.9 (3.7.1.6)
On page 3.7-3a, you mentioned the backfill of lean concrete and the use of compact Class A fill for the space between the side of excavation and the plant structure.
Discuss the merit of such a fill and how it is considered in your analysis.
Resoonse:
Lean concrete backf ill is being used to fill the gap between the vertical excavation and the structures. The lean concrete is used to fill narrow gaps between the rock and structure where the construction of compacted fill will be di f f icul t.
For larger areas compacted Class A fill is used.
Class A fill is placed around the str,uctures in areas where a sloped excavation is used, i.e. the west side of the Nuclear Island and on all sides above the weathered rock planes.
In the calculation of the lumped springs, the fill concrete was assumed to have the properties of the surrounding rock.
The mathematical models accounted f or the properties of the compacted backfill in the appropriate areas of the prof il e.
i l
l l
l QCS220.9-1 l
Amend. 68
P2ge - 1 (82-0184) [8,22] #38 Ouestion CS220.11(a) (3.7.1.6)
How dif f erent are the vertical translation soll-basemat Interaction spring constants calculated f ron the N.S. direction and the E. W direction soll foundation Interaction models? What physical ef fects are implied by this dif f erence and how are these ef fects accounted f or elsewhere?
Resoonse For the upper bound of f oundation material properties, the vertical spring constant calculated f ran the east-west model was 1.7% higher than that f rom the north-souin modei.
For the iower bound, the north-south modeI gave a value 5.5% higher than the east-west model. The physical of f acts of these dif f erencos are small and are due to the f act that the Nc6-th-South and East-West prof il es of the site are dif f erent.
l l
l l
l l
l QCS220.11(a)-1 Amend. 68 May 1982
i bge - 2 (82-0184) [8,22] #38 Ouestion CS220.11(b) f3.7.1.6)
On top of page 3.7-4, the word " rotation" is probably missing f ran the last sentence of the f irst paragraph.
Was the unit of the rotation applied to an elenent that was " rigid" by comparison, or by a rigid link method? or perhaps another method?
Resoonse The word " rotation" should be Inserted between " unit" and "was" on line 8 of the f irst paragraph on page 3.7-4.
The PSAR, page 3.7-4 has been revised to correct this error of emission.
The rotation was applied as tangential displacements at the boundary nodes of the structure, which would correspond to a net rotation applied at the centeriIne of the equivaient mass.
QCS220.11(b)-1 Amend. 68
Page / Lo,22JF39 The nodes on the periphery of the model were restrained against displacwement in their three degrees of freedom. The nodes on the radial planes that Ilmit the model were made free to displace in the tangential direction and restrained in the other two directions.
The embedded structure was represented by a circular rigid cylinder with a radius such that the polar moment of inertia of a horizontal section was equal to the moment of inertla of the actual mat about its centroid. To find the torsional spring, a unit rotation was applied about the centroidal vertical axis and the moment of the forces required was calculated.
To find the damping coef ficients associated with the springs, for each type of spring the shear modulus of an equivalent elastic half-space that gives the same spring stiffness was calculated.
With this G value, and based on the dimensions of an equivalent base, the damping coef ficients for an elastic half-space were calculated and these values were used in the analysis.
Three sets of spring stiffness and damping coefficients were calculated which correspond to the upper bound, average and lower bound of the rock (soll) properties. The seismic design is based on the envelope of the results obtained with the three sets of constants.
A verification of the spring stif f ness based on half-space theory is shown in the response to NRC Question 130.53.
' Since Selsmic Category I structures (except for, the Diesel Generator Building) are supported on competent rock with an average design shear wave velocity of 4000 fps, a fixed base approach would be jutified, consistent with established acceptance criteria.
However, equivalent lumped springs and dashpots calculated using a finite element static analysis are used to account for soll-structure Interaction effects.
It is noted that for CRBRP structures the spring constants and associated damping coefficients are calculated for average of values assigned to dynamic rock properties (125% from the average value). The envelope of responses obtained for this range will be considered in design.
r 3.7-4 Amend. 68
~
^
,66 IOD s
il
-~
u N
I?s
/
hs s s u[--<<N
(
7 s s s, l:
N
\\d\\/\\/\\
/N sJ
/
e s
N 5
m j
e s /.
g s
E i
s 1
J
\\
e N
e
'o E
e m
~
l
\\
s x
s 7
\\N S
l
. \\
9 e
~
~
8 i:
\\
\\
.x l
0220.11(c)-2
PIge - 4 (82-0184) [8,22] #38 fuestion CS770.11(d) (3.7.1.6)
How are the boundaries of the 2-D plane strain model detennined? Provide and justify the criteria used f or such determination.
Resoonse The boundaries of the 2D plane strain models were placed suf ficiently f ar f rom the structures to eliminate boundary ef fects. The vertical boundary (on ei ther si de) is at a distance, from the centerline, of more tnan 5 times the hal f width of the f o';ndation.
The depth is roughly 1 1/2 times the width of the f oundation.
The adequacy of the boundorles in the mathematical models (length and depth) was proved by the good co'rrelation of results f rom the f inite elenent and hal f-space calculations as described in Section 3.7.1.6 of the PSAR s(Page 3.7-3.c).
Since the half-space equations are based on boundaries being at an inf lpite distance, increasing the distances to the boundaries in the finite element model would have no f urther impact on the results (with the current boundaries).
QCS220.11(d)-1 Amend. 68 E%ETR
Pct.a - 5 (82-0184) [8,22] #38 Ouestion Cs??0.12 (3.7.2.1)
You used the terms gecrnetrical damping and critical danping. How ever, In Section C.1.2 of Appendix 3.7A, terms such as Interaction damping, radiation damping and internal damping are used.
Eyl ai n i n detall the di f f erences i n these terms.
If 'wo dif f erent terms mean the same thing, use one term consistently in order to avoid conf usion.
Resoonse DAYDING DEFINITIONS FOR THE PSAR in the equations of motion for the seismic analysis of Nuclear Power Plants, the energy losses are expressed by the damping f orces.
The energy losses include material damping and hysterstic losses within the structure or in the surrounding soll materials; f rictional losses at structural mnnections, equipment mounts or at soll-structure interf aces; and the radiation of energy away from the structure foundation and into the surrounding soll (radiation dampi ng ). For computational expediency the various energy losses are represented in dynamic analysis as equivalent viscous (relative velocity dependent) damping f actors.
The f ollowing def Ine the dif f erent terms related to damping used in CRBRP PSAR.
Damolng Coef f icient The damping f orce in the dynamic equation of a one degree of f reedom systera is g!cc ty:
-cy (1)
Where:
Mathematical constant - Damolng Coef ficient c
=
y Relative Velocity
=
l Critical Damolng IT i s Tne amount of camping that would completely eliminate vibration in a one degree of f reedom system, that is, the smallest amount of damping that will make the sy= tem return to the original position.
" crit - 2 N 2 Mw (2)
=
Where:
K Stif f ness of System
=
M
, Mass Frequency in Radians Per Second L)
1 1
QCS220.12-1 i
Amend. 68 May _1982_
..m Page - 6 (82-0184) [8,22] #38 DamoIng Ratio. DemoIng Factor This is the ratio between the damping coef ficient of a one degree of f reedom system and the critical damping Cc lf E-tu in experimental determinations of damping, the damping ratio is normally measured.
in a one degree of f reedom system if the damping ratio is known, the damping coef ficient may be calculated by equation (3).
Percent of CrItIcaf DamoIno This is the damping ratio expressed in percent:
C 0% = 100 (4)
Cecil Table 3.7-2 in the PSAR gives damping values in tenns of percent of critical dam pi ng.
+
DamoIng Matrix in a multidegree of f reedce system the damping coef ficients may be expressed in matrix f orm, by the damping matrix c
C=
t.1 clj cln l
c,y Cnl cj cnn n
l The coefficients c of the damping matrix represent the velocity dependent damping f orces devkloped at coordinate i by a unit velocity imposed at point J*
Moda1 DamoIno Modal Damping is the damping coef ficient or the percent of critical damping l
associated with a particular mode of the modal being analyzed.
In modal analysis each mode may have a characteristic or generalized damping which may be expressed as a generalized damping coef ficient er generalized damping ratio (or f actor). The generalized modal danping coef ficient is given by:
"hTfr C
(5) cr Where:
I cr Damping coef fIcient f or mode r
=
% = Mode shape vector f or mode r C
Denping Matrix
=
QCS220.12-2 Amend. 68 m
Page - 7 (82-0184) [8,22] #38 The generalized modal damping f actor (or ratio) is given by:
T dr C dr 4
P E Wr Mr (6)
Where:
j8r Generalized damping f actor (or ratio) for mode r.
=
Gr Model frequency for mode r in radians per second
=
j M
Generalized mass f or mode r
=
r To meet the orthogonality conditions that pennit the decouplirg of the modal equations the damping matrix must meet certain requirements.
However, in modal analysis the damping matrix was not used and damping ratios were defined f or the dif ferent modes based on the characteristics of the system.
Comoosite Modai Damolng - Combined Modal Damolng When a structure is composed of elenents with dif ferent damping ratios the modal daaping ratio was calculated by a weighted average based on the equation:
- r. ~,
B Ib $ N I (7)
N.# [ K [vi}
i Where:
BJ = Compost te Modal Damping f or mode j
[$j{ = Normalized modal vector for the Jih mode K
Assembled stif f ness matrix
=
5 ruusi e=u siiiiness matrix constructed from elements matrices f ormed by the product of the damping ratio f or the elenent and its stif f ness matrix.
Damping ratios f or elements were based on appropriate Table 3.7-2 values.
In the Nuclear Island model, the damping ratios f or the modes of the flxed based were calculated by this method.
The damping matrix used in the soll structure interaction of the Nuclear Island is a combination of these fixed base modal damping ratios and the damping coef ficients associated with the f oundation dampers. The damping matrices are constructed by the cxxnputer programs HETHA-VETHA f of IcmIng a f crmulatton similar to that proposed in Ref erence (5) of the PSAR.
Attachment A describes how this reference was use d.
QCS220.12-3 Amend. 68 l
May 1982
l
..o, Page - 8 (82-0184) [8,22] #38 Pronortional Damoing - Raleigh Damoing This is the system damping which is composed of material and interf ace shear damping.
In the seismic analysis of systems by time-history with direct integration of the coupled equations of motion, the operations are simpi tfied by using proportional or Raleigh damping.
By this procedure the damping matrix is f ormulated by a linear combination of the stif fness and mass matrices:
[ lKl (A l M C
+
=
(8)
Where:
C Damping Matrix
=
lMl = Mass Matrix lKl = Stif f ness Matrix 6 / 8 = Coefficients g &/ are selected f cr the range of f requencies of Interest with the appropriate damping ratio s crit cl.a.f are calculated with tha f ollowing equations M
80
+
s
=
crit 2 u) 2 Where:
Damping ratio - (Based on appropriate Table 3.7-2 values) s
=
crit R=
Frequency (Rads /sec)
_l'__
S !q Material Damoing internal or material damping is the hysteretic damping of the material.
For a soll foundation it is the damping characteristic of the soll mate-lal.
Although the Internal damping is not the result of a viscous behavior, it is expressed in terms of an equivalent viscous damping ratio.
Geometric Damoing - Radiation Damoing Goanetric or Radiation Denping represents the loss of energy that occurs through transmission of elastic wave energy from the footing to Infinity In a f ooting oscillating on an elastic half-space.
This damping may be expressed as a damp!.1g ratio or damping coef ficient.
l l
QCS220.12-4 Amend. 68 w
Page - 9 (82-0184) [8,22] #38
_ Interaction Damning The total foundation danping is the summation of the soll internal and radiation dampings.
In the CRBRP Nuclear Island analysis the rock internal danping was conservatively disregarded and the Interaction danping was made equal to the radiation damping.
This damping was expressed as a damping coef ficient.
Summarv in system modal analysis, modal dampings in terms of percent of critical were used.
The modal dampings are based on the values listed in Table 3.7-2.
When elements with dif ferent dampings were present in a system, the composite modal danpings were calculated and used.
In systems analyzed by direct Integration time-history analysis, proportional (Raleigh) danping was used.
in the analysis of the Nuclear Island (by direct integration) a denping matrix was constructed based on the composite modal dampings of the fixed base structure end the damping coef ficients (Interaction damping) of the f oundation dashpots.
i I
t l
l l
QCS220.12-5 Amend. 68 m
,QtJESTION 220.12 (3.7.2.1)
'At'tachment A,
)
1.~
DESCRIPTION OF HOW CRBRP HAS USED THE PAPER BY TSAI ON DAMPING The seismic analysis of CRBRP is performed with the computer programs HETHA (for horizontal motions) and VETHA (for vertical motions).
These programs, developed by Burns and Roe, solve the coupled equations of motion by direct integration allowing for a rigorous solution of the soil-stnacture interaction problem. The fannulation of the equations of action is similar to that used by Tsai in Reference Q220.12-1.
This approach considers the displacements of the nodes of the structure subjected to an earthquake as composed of tw parts: The displacements relative to the base which are expressed in terms of the modes of vibration of the " fixed" base structure and the displacements due to the base motion.
The equations of motion are expressed in terms of the mode shapes, modal frequencies and modal d".:npings of the " fixed" base structure and of the stiffness and damping coafficents of the lumped springs and dashpots that represent the effects of the soil on the structure.
As a result of this formulation the' number of coupled equations to be integrated is reduced from N+P to M+P, where N is the number of degrees of freedom of the " fixed" base structure, P the number of degrees of freedom of the base and M the number of modes of the " fixed" base structure selected to represent the structure.
For a large structures M is much less than N.
ilhile Tsai's formulation includes only two degrees of freedom (horizontal translation and rocking) the HETHA fannulation considers three degrees of freedom per node (horizontal translation, rocking and torsion).
VETHA considers one degree of freedom per node (vertical translation).
In Ref.1. Tsai proposest the use of the exact formulation to calculate equivalent modal campings for a modal ane. lysis of the structures by matching the responses of the exact analysis with those of a modal analysis.
In the CRBRP analysis, the coupled equations of motion, similar to equation (15) of Ref.1 are integrated directly using' the acceleration time history of the ground motion as input.
FQUATIONS USED IN HETHA The following equations were used in the computer program HETHA:
Let X be the " fixed" base displacement of the structure l
e i {r[
(1) 4 {e'3t lihcre:
d
= Horizontal displacement in the direction of the input motion, relative to the base.
M= Rotation about a horizontai axis perpendicuiar to the direction of the input motion, relative to the base.
8 = Rotation about a vertical axis relative to the base (torsion)
The equations of motion for a fixed base structure are:
0220.12-7
(M]
5 ]+ [C] (k) + [K] X} * - d (b1 [M] [a}
(2)
Where [M]ly[c] and [K]
are the mass, damping and stiffness matrices respective of the fixed base structure u(t) = Translational ground acceleration (free field)
I[I}
(A) = q [o})
jo[
~
~
tm;..,
8%IgIMa Where:
ing= mass of nodal point i Ij. = mass moment of inertia at nadal point i about a horizontal
, centruidal axis perpendicular to the direction of the motion Jg = mass moment of inertia at nodal point [ about a vertical centroidal axis (torsional)
.)
Assum;as U.L use fixed base structure has classical normal modes, the normal mode matrix C$2 satisfies the following conditions:
C4f CM3 Cd3 CI]
~
.(3)
[43'CKJ Cd3 Cd))
=
ttAj uj 3 b) 7 Cd3 CC3 Cd3
=
j (5)
Wh re uly are the modal frequencies andp; the modal fraction of critical damping for the fixed base structure For tr.e same structure with flexible foundation, the displacements of the structure relative to the ground motion are given by:
I 'h
- S* W h; + eel (6) l l
?g e 9'; + p (7) l
- +G (8) el 8 i
1>i>u Where:
S; Horizontal displacement of, point i in the structure in the direction l
=
of the input motion. relative to the ground motion.
,i l
0220.12 8 G
,r-
- - - - - - - - - - - - - - ^ -
^ ^ ^
- y Rotation at point I in the structure about a horizontal axis I = perpendicular to the direction of the input motion relative to the ground motion.
- 6. * = Rotation at point i in the structure about a vertical I
axis, relative to the ground motion.
4 = Horizontal translation of the base mat.
tp = Rotation of the base mat about a horizontal axis perpendicular to the input motion.
The axis passes through the center of rigidity, of the base mt.
6 Rotation of the base mat about a vertical axis. The axis
=
passes through the center of rigidity of the base mat.
h = Height of point j in the structure relative to the base mat.
Horizontai distance from point i in the structure to the vertical C.
=
8 axis of rotation of the base mat, measured perpendicular to the direction of the input motion.
Define 0W
%' =4(Y) ?
(9)
({eb
. With the equations of force equilibrium at the base written to satisfy
. the boundary condition, the equations of motion of soil structure interaction are (M) ( {9} +0.{A} H c cs {icj + c n3 { s} = [oj m (%+Q c.9s+ x.9 - -$m 1 7+cj.y. -j, m, 8; $,i (ig UJ.G.)-( I;7, g
5 3
(10)
Jb
' C 8 'K e=
mj t; (Ig 4G.).fJ;dg e
a 11 ist Where:
m = Mass of Mat b
I -Mass Moment of Inertia of Mat about the centroidal horizontal b
axis perpendicular to the direction of the input motion i
J6 = Mass Moment of Inertia of Mat about a centroidal vertical axis T = Mass moment of inertia at point i about a centroidal horizontal i
axis perpendicular to the direction of the input motion Jg = Mass moment of inertia at point i axis about a centroidal vertical Q220.12-9 o
a
~.
Equations (10) can be rearranged to give:
'Q]
I{'9} '
{
r er;- - -- -
- gw 1s
' Y1 4,
i, 3
_Lcl__]__-Jc_J LA.L____-IclLaht J
_ -Jc.Lael diti i
I
-Mi'tc) ': (c. lal'(c)(ap f ai'Ic) 981 (41'tc)fael 4:
)+>+
-lihnci!teeni t u.r n c u o n r s n c n. e3 I
- {Aaj'(c)
{ Aej'(c) {Al
{Aaj'(c][ Ahl (c..[Aaf(c){Aaj) is )
1 4
~
.mummene em
.musma, I
'fal'("I j r.ve{^i'Mib 1A5'tx] pal gl'tx] [6a5 g
I
/
- +
1 4 C"1 lIAhi'(x1fal 0(--tal'[x1@hD
{Ah!'(x) (Ae}
1
-leaf (x} i fael'ful(Al
[aal'[x] phj
(,u.. partnj p.3 e.)
b I --
,g (10
% o.
o o
I; J
l where:
If 1 I'l = J{eli fh;j1 f[ad )
J{i,; l 7
go;
[4h] -
l t
I,l*ll j,1cl,
- itj, Q220.12-10
~
Defina tha following transformation:
3 II (Q)
J %
j'/fEb
/
i i +i i
Vfr,-
j l, 9 I
8/f5.
I.T s
M+5s V",a(A){r}
(13)
Where H is number of structural modes and let
[r'! *(e[ (M) {a-}
{ r,1.(e r (u) [ A'}
{ r
- (#I (MI(A }
t (14) j -((4f (M){Ah}
I Ig
< r.. - $f CM)[Ae}
}{c)h
'[o j 'l Where:
[A}
{l} h
{A*j=<((ci'
{o}-)
,, t j,
Equation (11) ply equation (11) by (Af and substitute equation (13).
Premulti becomes:
{,E y + C 2.3 {h } + C E3 {f"} * -{, E} U.
i (15) j te-y w;
-t 63 [r.1/ rr,
-Eej3..{r.}/ m
. 1
-c a3 ( rJ /(5.
e a;2 3
3 s
a
-[rrta;) ' c c.+7,, a; r 'ym,
=
7,, a; r r;/tm r, 3,6; r; et;/4 j
j 3 o
ime
[.6 = ;{r }'t6 3 l u
a u
3,6;rr,3/en,t5 (ceg,ejrip/1, jc 6;ria e;/U's cu>
3 3
3 r
mb w
w w
l Jr,} [s;]
I =J j r,;/%2
. z a;r,; r /n,a, (c.9 nr1;y4 r
a Jat J'l J '8
~
~
Q2,20.12-11
--.1..
~
-, ~. -
' Matrix [Y.lis identical to C{3 when bj W[, c3 = K,, and C9=Ky
, and Ce = K e (17) tri
[ } * %a.)
(r} in eq.(13) is ann +1 by 1 column vector. Considering only the first N modes of the fixed base structure. (r[ becomes a M+5by I coitann vector.
Equation (15) is solved by Wilson's (2) step by step integration.
Let {r }}k} and{E } be known as vectors at time t, the acceleration t
g vector [g,y} is solved by the equation:
CH3 {F..y } + te.T (/.g.7 }+ t w J (e.,. } = [4 **r }
cis) s s
Where velocity and acceleration are replaced by:
(b 7} 'h ({P +7}-{rg])-i {f } -k (F }
(19) t t
t t
- rl $,(fr
- rl-Cetl)-p (*si -tt43 cao; i
s (Esv} = D d +1 (Ot + At } ~
f t
(E,) * - {' } U-b t
T ' s4
- A t (21) l Velocity and acceleration vectors at t+T are obtained by equations (19)and(20).
The acceleration, welccity and displacement att+4.+., are given by:
i (F + Atl * (l-f I {F } +j (*r:g +.7 }
t e
(h + 4.t3 ' C7 } + ^* C C F.e} + {F + AtD t
t t
g czz) l (f +At} * (rtl
- Ai ( t)*
s((~'s 4 st! +4 (E } )
P t
s The displacement and accelerations (c) are combined to obtain nodal 1
responses according to equation (13).
0220.12-12 1
Page - 11 (82-0164) [8,22] #38 i
Ouestion CS220.13 (3.7.2.1) 1 l
At top of page 3.7-6a, it is Indicated that in the model the dome of the steel containment has been idealized using equivalent springs which account f or the
" breathing" of the dome during a vertical vibration.
Explain clearly how through such a lump-mass model " breathing" of the dome can be taken into consideration. Def Ine " breath ing".
Resoonse:
" Breathing" ref ers to shell type of vibration of the dome.
To account for this ef fect, a stif f ness matrix of the dome with cross-coupling tenns was derived from an axisymmetrical shell model of the dome, using the KALNINS computer progran.
The equivalent springs represent the terms in the stif f ness matr1 x.
i l
l QCS220.13-1
Pego - 12 (82-0184) [8,22] #38 Ouestion CS 220.14 (3.7.2.1)
Two computer programs: HETHA and VETHA are mentioned.
Indicate if these two computer programs are validated in accordance with the procedure described in SRP Section 3.8.1 (P.3.8.1-10).
Resoonse The computer programs HETHA and VETHA have been validated consistent with the requirements of SRP Section 3.8.1 (Page 3.8.1-10), item II. A description of the programs is in PSAR Appendix A pages A169 to A182.
1 I
l l
l l
l
\\
l l
l l
QCS220.14-1 1
Amend. 68 f3rLJ1G92
bge - 14 (82-0184) [8,22] #38 Ouestion CS220.16 (3.7.2.6)
On page 3.7-9b, it is stated that two methods of combining the seven spectra, one by square root of sum of squares and the n+her by absolute sum.
Indicate the conditions under which each of the methout a ll i be used.
Resoonse l
First method:
1 Each of the seven spectra are applled Independently to find the final of facts t
(stresses, deflections, etc) and the resultant stresses are combined as shown by equations (8), (9), and (10) on page 3.7-A-B3 of the PSAR.
These equations l
add the total of fects of the north-south, east-west and vertical earthquakes by the square root of the sum of the squares.
Second method:
Instead of using each of the seven spectra Individually, the seven spectra are combined into three which correspond respectively to the north-south, east-west, and vertical directions, using equations (11), (12) and (13) on page 3.7-A-B3 of the PSAR. The final ef f ects obtained with these combined spectra are added by absolute values.
The selection of one or other method is optional to the designer.
l l
QCS220.16-1 cn%339
Page - 15 (82-0184) [8,22] #38
.puestion CS770.17 (3.7.2.13)
The discussion on overturning of seismic Category I structure appears to be f or a structure on a single f oundation mat not on a combined f oundation.
Indicate what your consideration will be f or the overturning of structures on a combined f oundation mat.
Express with the help of a figure the location of Hi and the distance of ht in the equation on page 3.7-10a and the basis of ffielr determination.
Re ~ nse:
Figure 3.7-19 in the PSAR shows the single f oundation mat of the Nuclear Island structures.
In the seismic mathematical model, the rock-structure Interaction was represented by lumped springs.
From the seismic analysis, the maximum foundation spring f orces, (translational and rotational) were cal cul ated.
These forces were applied in verifying the stability of the structure.
Thus, in the equation on page 3.7-10a in the PSAR:
Not1 Mrt + h1 Hj
=
Mr is the maximum reaction in the rotational foundat!on spring in the j
north-south direction.
H is the maximum reaction of the north-south j
translational spring, and h is the distave f rom the f orce H1 to the bottom 1
of thc f cendstlcn mat.
hj was conservatively made equal to the foundation mat th!-yaetr.
ui is the total overturning movment due to the north-south r
g earthqua ke.
Figure QCS220.17-1 shows a sketch of the f orces.
QCS220.17-1
M H
1 gl
\\
d' th 1
Overturning moment due to North-South Ft
=
rl earthquake (maximum reaction of rotational spring from seismic analysis) y Force on the horizontal translational spring H
=
under the North-South earthquake (Maximum reaction of translational spring from seismic analysis) h Vertical distance from the horizontal
=
y translational spring to the bottom of the foundation mat.
FIGURE Q220.17-1 Q220.17-2
i Page - 16 (82-0184) [8,22] #38
.0uestIon CS220.18 (3.7.2.14) in this section you discussed the analysis procedure for damping.
Indicate If there is any dif f erence between what you described in this Section and that in Section 3.7.2.1.1 on pages 3.7-5 and 3.7-6.
Explain in mathematical forms the f ol lowing: composite damping, modal damping, proportional damping, and their relationships to the critical damping values as specified in Table 3.7.2, i
together with the conditions under which they are used.
Resoonse:
See the Response to Question QCS220.12.
i i
I l
l i
l l
i i
l l
QcS220.18-1
~
Pago - / LGZ-UlG41 Le,tzJ sac
)
Ouestion CS 220.19 (3.7.3 1D1 in this section it is stated that the response spectra produced will be widened by 110% by frequency to account for uncertainties in the structural model and imput.
However, in Section 3.7.1.6 on page 3.7-3, it is stated that the response spectra will be widened by 1,15% in frequency.
Indicate which percentage of widening is actually used and if its use is in conformance with SRP Section 3.7.2 criteria.
Resoonse The design response spectra is the envelope of response spectra produced for the upper and lower bound properties of the rock (soll) properties.
Before enveloping, the spectral peaks were widened by 1,105. The 115% Indicated in Section 3.7.1.6, page 3.7-3 has been revised to 110%. The 110% is consistent with SRP which requires.a minimum of 10%.
In addition, CRBRP envelopes the results of upper and lower bound of rock (soII) properties.
i l
l l
QCS 220.19-1 Amend. 68 May 1982
P:ge 3 LU,22Jr39 3.7.1.6 Soll-Structure Interaction 3.7.1.6.1 Structures Sucoorted on Soll The seismic analysis of the Diesel Generator Building (DGB) which is a seismic Category I structure supported on soll, will be conducted using the finite element computer program FLUSH. The analysis accountr for the strain dependent properties of the soII.
Since FLUSH is a 1 so-dimensional computer program, analyses for the North-South and East-West directions will be required.
The input motions will be applied at the surf ace (finished grade) of the "f ree-field" and will consist of the motions used in the analysis of the Nuclear Island. The strain dependance of the soll properties will be accounted in the analysis.
Analyses will be performed for a range of soll properties and the results will be enveloped. The envelope response spectra of the foundation level, at the "f ree-field", will not be below the Design Response Spectra. As described in Section 3.7.2.1.1.3 the motions calculated at the foundation of the structure will be used in a three-dimensional analysis of the structure. The peaks of the in-structure spectra envelopes for the range of soll properties will be widened by i 10%.
For Category I buried structures, duct banks, pipes and electrical manholes, the rethod described in Section 3.7.3.12 will be used.
Two Category lli structures edjacent to the Nuclear Island (Turbine Generator and Radweste Buildings) are supported on soll and are seismically analyzed for the SSE.
Following the NRC Instructions of Ref.13, the FLUSH analysis was performed assuming the input motion applied at the top of a rock outcrop at the finished grade elevation.
l Further description of the analysis at these structures is given in Section 3.7.2.1.1.5.
3.7.1.6.2 Structures Suonorted on Rock Ti.
...ajv. Cai=wur y I structures are founded on competent rock with an average ch::r wave velocity of 4000 ft/sec.
3.7.1.6.2.1 Nuclear Island The Nuclear Island consists of the following Interconnected buildings with a common foundation mat: Reactor Containment (RW), Confinement, Reactor Service (RSB), Steam Generator (SGB), Control (CB) and Electrical Eqlupment (EEB).
The rock structure interaction of the Nuclear Island will be analyzed as follows:
3.7-3 Amend. 68 May 1982
~
P2ge 30 (82-0184) [8,22] #38
_Ouestleti CS220.23 (3.8.2.1)
It is stated to the ef fect that the steel shell in the lower portion of the contairvnent structure is sandwiched between two concentric concrete walls and neither of the two concrete walls are considered to be part of the contaltunent steel.
However, the outside concrete wall is Indicated to be designed to prevent the buck!Ing of the steel shelI.
Indicate what are the design criteria for the two concrete walls, especially the outside concrete wall, and which ACI Code will be used in your design.
Resoonse:
The walls are designed to resist all normal, seimic and accident Iced; in accordance with the load combinations specified in PSAR Section 3.8.3.3.
The design of the walls is done in accordance with ACI 349.
t l
l l
l I
QCS220.23-1 Amend. 68
Page - 33 (82-0184) [8,22] #38 Ouestfon CS??O.26 (2.8.2.2.2)
It is stated that potential corrosion of the portion of steel containment anbedded in concrete as a result of concrete cracking is precluded due to f act that there is a minimum of 22 inches of concrete embedment and the cracking under the worst of cases is minimal.
Indicate what size of cracks is expected as a result of the containment structural integrity test.
Note that these cracks will terminate at the steel containment shell.
Resoonse:
The sizo of possible cracks in the 22 inch concrete wall around the containment vessel for a test pressure of 11.5 ps! has been calculated.
it has been assumed conservatively that the internal pressure acts directly on the containment shell and that the load is shared by the steel shell and the 22 inch concrete wall in accordance with their relative stif f nesses without assistance f rom the interior 36 inch concrete wall. The calculations are based on the hoop stresses developed in the concrete by the 11.5 psig test pressure.
The analysis method of Ref erence QCS220.26-1 (Section 10.4.3) was used.
The calculated crack width is 0.009 inch.
In accordance with ACI Canmittee 224 (Ref erence QCS220.26-2) the following are acceptable crack widths in concrete:
In dry air:
0.016 inch Humidi ty, moi st ai r, sol l :
0.012 inch The calculated crack width is below the applicable limit which is 0.016 inch.
It should be noted, that the stresses in the steel shell and reinf orcing steel are below yield.
Theref ore, when the vessel is depressurized and the tensile stresses relleved, the cracks will close.
Based on the above discussion there is no concerl f or corrosion in the steel shell.
References:
QC47?n ?A 1 Park and Pauley, "Rei nf orced Concrete Structures", John W il ey, 1975.
QCS220.26-2 ACI Canmittee 224, " Control of Cracking in Concrete Structures", Journal ACl, Vol. 69, No.12, December 1972 pp. 717-753.
QCS220.26-1 crratkfE)
Page - 35 (82-0184) [8,22] #38 l
Ouestion CS 220.28 l
The applicant should substantiate the statement tnat the contairvnent will not be subjected to non-axisymmetric temperature distributions above the operating f l ocr.
Resoonse For the postulated containment vessel Design Basis Accident (DBA) the contalment vessel would be heated by convection f rom the hot gasses f rcrn Cell 102A entering the spaces above the operating floor, and by conduction throu0h the concrete. These gasses would enter this above operating floor space through potential paths that are widely distributed. One potential path is Cell 105U to Cells 109 and 113; another is Cell 105H to Cells 109 and 113; and another is CeII 105H fo CeII 110 on the opposite side.
(See Figures 1.2-14 through 1.2-16 of the PSAR). All of these paths are very unlikely and contain fire doors and other inpediments to the fire passage.
As shown in PSAR Section 6.2 the peak gas temperature above the operating floor is only 1400F which occurs only af ter a long time and at a very slow rate.
This is only a 700F rise above the normal operating temperature (and normal shelI temperature); thus, the magnitude of the temperature variation could only be a portion of the 700 which would create negligable stress in the shell, even if the heating source was localized.
It should be noted also that the design basis accident sodium fire analysis, which is used to postulate containment temperatures, assumes that hot gasses crc t. e.::.!tt:d directly to the space above the operating floor from the fire in Cel l 102A. This assumption maximizes the rate and total amount of heat transf erred into the area above the operating floor without taking credit f or the heat losses into the concrete and equipment along the actual gas flow pa th s.
This is a conservative assumption which cannot actually occur and theref ore adds even more conservatism to the design.
QCS 220.28-1 Amend. 68 J D L1191 R
Page - 36 (82-0184) [8,22] #38 Ouestion CS220.29 (3.8.2.4)
Indicate if there are any mechanical connections between the confinement structure and the containment above the operating floor that could transmit mechanical Iceds. Can the relative displacement between the confinement structure and containment become large enough to allow contact between the contairrnent and components attached to the conf inement structure (such as the partition supports)?
i Resoonse:
There are no mechanical connections between the confinement structure and the contairvnent above the operating floor that could transmit mechanical Iceds.
Further, the gaps between the containment vessel and components attached to the Confinement (such as partition supports) will be established to prevent contact under any Iceds, including Thermal Margin Beyond the Design Basis (T!EDB) accidents.
I l
l l
l I
1 i
l QCS220.29-1 Amend. 68
Psge - 14 (B2-0184) LB,22J 736 Ouestion CS 220.31 (3.8.'.1.3 )
On page 3.8-10, it is stated that the Interior surf aces of the cells are lined with carbon steel plates with the lower portion of the plate designed to contain hot sodium spills.
Indicate the dif ference in the design of the lower and upper portion of the cell lines.
Response
i The design of the cell liner system does not distinguish between wetted (lower) and unwetted (upper) zones of the I!ned cell.
As identified in PSAR Section 3.8-B, paragraph 3.1.1.9, the cell liners are desggned to withstand large sodium spills with Na spill temperatures up to 1015 F consistent with their application.
PSAR paragraph 3.8.3.1.3 has been revised to be consistent with the design of the cell liner system.
QCS 220.31-1 Amend. 68 May 1982
Page 38 [8,22]#39 3.8.3.1.2 Head Access Area (HAA)
The head access area is located below the operating floor level and above the reactor cavity.
The access area is of a square shape 44 feet long on each side and 14 feet high above the reactor head.
The head access area is a reinf orced concrete structure.
Steel franing will be provided in this area to support the EVTM operailons.
3.8.3.1.3 Primary Heat Transoort System (PHTS) Cell Each PHTS cell is a step type rectangular reinf orced concrete structure.
At its widest section, the cell is approximately 48 feet wide by 72 feet long.
The cell is approximately 58 feet deep in the area housing the primary sodium pump and Intermediate heat exchanger.
The Intericr surf aces of the cells are provided with carbon steel cell liners designed to contain sodium spills.
The cells are designed to withstand accident pressure and temperature conditions as noted in Table 3.8-2.
3.8.3.1.4 Reactor Overflow Vessel and Primarv Sodlum Storage Vessel Ce!l This cell is a step type rectangular reinf orced concrete structure.
The cell is approximately 26 feet wide by 69 feet long with 62 feet height at its deepest section. The interior surf ace of the cell is lined with carbon steel plate similar to the PHTS cells.
The cell is designed to withstand accidert pressure and temperature conditions noted in Table 3.8-2.
3.8.3.1.5 other CgLls These cells are reinf orced concrete structures with various sizes. The cells required to maintain a nitrogen atmosphere during operations will be lined and designed to the requirements noted in Table 3.8-2.
Cell liners are described i n Secti on 3. A.8.
3.8.3.1.6 Fill Slab l
i A structure fill slab of suitable thickness will be provided over the bottom l
containment liner plate.
l 3.9.3.2 Arrlicable Codes. Standards and Soecifications 3.8.3.2.1 Design Codes Applicable provisions both mandatory and recommended of the f ollowing codes 1
will be used in the design of the internal structures:
i t
3.8-10 l
Amend. 68
Page - to (92-u1541 Lu,22J Fan
_ Question CS 220.32 (3.8.3.2.1)
It is stated that concrete Internal structures will be designed in accordance with ACI 318-77. Since ACI 349, " Code Requirements for Nuclear Safety Related Concrete Structures" is specifically for the design of such structures and has been endorsed by NRC in Regulatory guide 1.142, use of this code is required.
Et1Dania Section 3.8.3.2.1 of the PSAR will be updated to state the design of the Internal structures complies with the requirements of ACI 349, " Code Requirements for Nuclear Related Concrete Structures," as endorsed in Regul atory Guide 1.142.
l l
l l
QCS 220.32-1 Amend. 68 May 1982
rcge - so sc4-uio## Lo,44J 0>3 Ouestion CS 220.33 (3.8.3.4)
The general structural analysis procedure using the strip method is not totally clear.
If possible, the method should be referenced to the ACI-349(76) code.
If not, more detail is needed on how the Interaction of surrounding cells will be handled when analyzing Individual cells. Can significant additive moments be Introduced from adjacent cells at a common Juncture? Will the method described take into account the high tensile loads developed on the diagonals of two-way slabs near the corners?
Response
The general structural analysis procedure using the strip method is based on ChapteF 8 of ACl-349(76) Code.
In general, strips are chosen such that the Interaction of surrounding cells are considered and therefore all effects are included.
The high tensile loads develop <a on the diagonals of two-way slabs near the corners apply only to exterior slabs f ramed by spandrel beams and does not apply to slabs Integrally framed into exterior walls and hence is not applicable to the Reactor Containment Building.
Section 13.57 of ACI 349(76) provides minimum reinforcement requirements for this condition. The reinforcement steel provided in the slabs exceeds, by far, the minimum requirement for slabs on spandrel beams.
4 l
QCS 220.33-1 Amend. 68 May 1982
Page - 52 (82-0184) [6,22] #38 1
l l
l Ouestion CS770.36 f a) ( Accendix 3.8-B)
)
l 1
i Corrosion ef f ect is included by reducing the plate thickness by 1/16 inch l
(3.1.1.5).
Is it possible that liner corrosion could introduce local flows that would not reduce overall stif f ness but introduce significant stress j
concentration points?
Resoonse:
For propagation of a local flow see the response to Question 220.40(b).
The corrosion allowance as described in PSAR Appendix 3.8-B, paragraph l
j 3.1.1.5, is based upon the guidelines established in ASME B&PV Code, Section
(
Vill, Division 1 for ownponents exposed to water or steam.
The cel l l iner system under normal operating conditions is exposed to an inert ahnosphere on l
one side, and is coated with inorganic zinc paint on both sides thus inhibiting corrosion.
On the back f ace of the liner plate, the liner is in l
contact with the 1/4 inch air gap at liner wall / celling locations and 1/8 inch l
ai r gap bel ow the l i ner f l oor pl ate.
As noted in the response to Question 130.95(1) the potential enount of moisture within the air gap is limited.
i Accordingly the consideration of the 1/16 inch corrossion allowance is conservative.
The possibility of a localized zone of behind the liner corrosion cannot be precl uded.
However, due to the conservative assumption of a corrosion allowance based upon ASPE B&PV Code, Section Vill Division 1 requirements f or steam and water vessels, coupled with the use of inorganic zine paint on all surf aces of the liner plate to prevent corrosion, the anticipated corrosion flew would have a thickness less than the corrosion allowance.
1 i
i e
s l
QCS220.36(a )-1
Pags - 52 (82-0184) [8,22] #38 Ouestion CS 220.36 (b) ( Anpendix 3.8-B)
Handling the corrosion by reducing plate thickness gives a lower stif fness.
However, could the stif f ness reduction erroneously give thermal stresses that are too low? Generally, the more flexible the structure, the lower the thermal stresses.
Resoonse in a fully restrained flat plate, the strains due to a uniform temperature are Independent of the plate thickness.
Because of local buckling, the liner strains dif fer from those of a f ully restrained flat plate.
An analysis with a mathematical model shown in Figure 3A.8-3 in the PSAR was performed using a reduced plate thickness (5/16 inch). The resultant strains are not significantly dif ferent f rom those shown in Figure 3A.8-3 for a 3/8 inch plate.
l l
QCS 220.36(b)-1 Amend. 68
[htE5P2
Pago - 53 (82-0184) [8,22] #38 Ouestion CS 220.36 (c)
(Accendix 3.8-B)
Near very stiff areas on the liner boundary, for instance close to the pipe penetrations, the neighboring structure will have to exhibit considerable "give" or the liner attachment to anchors could be over stressed (Ref erence Figure 3A.8-8 where a steel anchor is apparently within 3 in. of the penetration collar). This will occur because the liner plant cannot buckle for short, unsupported spans.
At what location in the structure will this situation be the worst and what are the shearing forces and displacements at the studs?
Resoonse Figure 3A.8-8 incorrectly depicted a stud anchor within 3 In. of the pipe penetration collar.
Figure 3A.8-8 has been revised to reflect the configuration used in the analysis and specified for the CRBRP cell liner anchorage.
The liner stud anchor nearest to the penetration collar and/or embedment plate is located a minimum of 10 Inches f rom the Interf ace of the liner and the penetration /embedment.
The worst location f or studs and plate is in the vicinity of "hard" spots such as embedments or penetrations. This has been covered in the reply to questions 220.36(d) and 220.40(a) which discuss the anchor and liner shear forces and displacements.
i QCS 220.36(c)-1 Amend. 68 l
l May 1982
3 ' LINER 8
Y a
I d
'eK
[-
4 o
o 7-"-
V SCHEDULE 80 4
3 f
i G v 14 I
b COLLAR J-W O
P
/
O1 J
9 y-
~
0 g
l 3
l 1
l 0
g
'd A
MODEL BOUN DARY a
s D
ys I e
[
__ y n
E MOD E L,' B O U N D ARY D (( STUD 5
'o >
}/}F I
e is-z s-
- ' 5 " 'A
.7+
1.21 8 '
l t
lel 3 4.5"
=
Sectional Elevation Of Wall Liner Penetration Figure 3A.8-8 3A.8-17
Page - 55 (82-0184) [8,22] #38 Ouestion CS 220.36 (d) ( Accend f x 3.8-B)
At the top of page 3A.8-6 an analysis is described in which the panel corners at the stud anchors are assumed to be rigidly supported. This requires no unbalanced l ateral f orces on the anchors.
If all panel sections buckle in the same direction this assumption is good. However, especially in the case of flat liner plates, the most probable buckling pattern may involve a shape where adjacent panels alternately buckle in and out.
Has this case been analyzed and are the resulting shear loads in the panel at anchor attachment poin+s acceptabl e?
Resoonse Analyses in which liner plates buckled in and out were perf ormed.
Fi gures 220.36(d)-1 and 220.36(d)-2 show two of the cases, which are in the vicinity of a penetration and an embedment respectively. The liner criteria is def ined in terms of allowable equivalent von Mises strains and the calculated values (shown in Figures Q220.36(d)-1 through 4) are within the allowable limits.
The principal shear stresses in the plate and in the stud were calculated and are well below one half of the ultimate strength of the material.
To prevent ilner tear by the stud, the ratio of stud disneter over plate thickness (d/t) was limited conservatively to a value of 2.0.
This is based on Ref erence QCS 220.36(d)-1 which reports the results of testing of plate stud systems.
Ref erence QCS 220.36(d)-1 reports that plate f ailure occurs at a plate thickness to stud disneter ratio (d/t) of 2.7 or greater.
Reference Q220.36( d )-1 G. G. Goble, " Shear Strength of Thin Flange Camposite Specimens," AISC Engineering Journal, April 1968 l
QCS 220.36(d)-1 Amend. 68 w
ULTIMATE STRENGTil AND STRESS AT TEMPERATURE Principal Shear Stress in Stud (Ksi) 9
=
i 3
g Principal Shear Stress in Liner (Ksi); Maximum at a Stud Location
=
,p 7.5" 15" 10.5" 3
f LINE OF SYMMETRY 1.43"
~~~
f
,s' m
_____...---['
'3 LINER DISPLACEMENT (IN) 1/4" gap g
//4 8
SECTION A-A i
I EMBEDDED PLATE o
18.lksi I
q 14.4 ksi
=
o q
=
w S
s o
I 0y STUD STRAIN Y
MAX. LINER STRAIN STUD STRAIN g
.046 (M+B)
.017(M+B
.024 (M+B)
.0174(M+B)
.028 (M+B)
.017 (M)
.006 (M)
.010 (M)
.007 (M)
.0.12 (M) 7w A A
^A a =17.0 q =18.1 p
p
\\
LINE OF SYMMETRY
/
M = Membrane Strain M + B = Membrane plus Bending Strain MAXIMUM EQUIVALENT STRAINS (IN/IN)
WALL LINER AT EMBEDDED PLATES, BUCKLING PATTERN FIGURE O220. 36 (d)-l
PANEL BUCKLES OUTWARD (UMSHADED)
//
PANEL BUCKLES INWARD (TOWARDS WALL) AND 3,, L I N E12 8
//
TOUCHES CONCRETE P L A Ti:_ -'~\\
C Y
\\
J
.I J
'a y f j/
PANEL BUCKLES INWARD j/
/,, j BUT DOES NOT 7
/,
TOUCH CONCRETE l
//
l 4.
UZ - OUT OF PLANE h
DISPLACEMENT UZ =1. 4 8
(33.6) o SCHEDULE 80 k
'/ /, '/
//
f V
st.s av a 0
v//
/,/,' /
o y
+9 G,,t l 4
/,',/
'/
/ -
d) y
's
/,,
C O L L A R.
V
\\,%
_ /
I'I/ /
' / $
r
'/
s
/ / /
?,
/ / / /
,I A
g
,Mg o E L,BO UN D Y
/
/ / / /'
as i
yj /
/,
UZ=1.19"
~lI, s.
/
/
/
/
/ /
m n
- y t
F:
g,;
li,p sruos e is" %
,1..
, 3..
73
, ['
q.
l l 216 '
ic
,w
=
WALL LINER AT PENETRATION; BUCKLING PATTER,N FIGURE Q220. 36 (d)-2 0220. 36 (d)-3
LINER STRAIN (IN/IN)
MEM E + BENDING
~
MEMBRANE STUD STRAIN (IN/IN)
MEMBRANE + BENDING
,I.006]
%" uned 3]
PLATE -s C
Y
\\
a a
d
- tn
.N I'
.023
.024 4
f'/
n
,.015
.01 Y
\\
- '/
.017 4
.018 o
i l
~
~
SCHEDULE So g
.0088 SLEEVE O
~
~
'm
.013 1
+
o i
C, = 1,4 I
b COLL AR
/
J-
/
l
.017
.012 1.LI I
.026
.017 O
P O _1_
o g,
5 1L I
d A
MODEL BOUN DARY D
s
'a
.017
_.0214/
.025 /,,,
.0075-
.02 "014 I,+,
.0162
. 024 e
o
~_y m
'o F
D[pSTUDS I
, s. '.
e is" *
,1 G*
~.007 ~ _\\
6.011 1.218 '
.014_
l i
18' 34.5"
=
WALL LINER AT PENETRATIONS; GENERALIZED STRTINS FIGURE O220. 36 (d) G3 0220. 36 (d)-4
ULTIMATE STRENGTH AND STRESS AT TEMPERATURE Principal Shear in Stud (K si) q
=
s Principal Shear in Liner (Ksi ) ; Maximum at a Stud Location q
=
p o
38 LINE12 PLATE 7 C
Y
\\
il
,1 d
g q, = 18.6 3
~
q = 27.3 c~
P l
l.
s
/,,;
se a*-
P
/
q =18.8 4 P---
- o g,
s A
SCHEDULE 60 V
q = 18.4 Q(f P
]
=
8 3-G.,v14 I
s g
C O L L A R.
(
\\
9s " 10*9 9 =15.5 O
s OJ
(
o Y
a 3,
S q = 14.7 Id_
p
,g P
0 l
A f
MODEL BOUN DARY l
'q =19.0 s
i.
\\
q = 15.3
- l +.
1 q =12.5 P
qs = 15.9 p
e o
_'\\o
/ /F E
D,[jdSTUDS MODE L, B O U N D AR.y s
e,,
@ IS" %
if 35' 7g a
i 1.21 6" !
- 8 345'
=
WALL LINER AT PENETRATIONS; SHEAR STRESSES IN LINER AND STUDS FIGURE Q220. 36 (d)-4 A.
0220. 36 (d)-5
Page - 56 (82-0184) [8,22] #38 Ouestion CS 220.36 f e) ( Accendix 3.8-B) in all analyses presented, unif orm temperature distributions over the liners are assumed. Are these areas where this assumption is not valid?
If so, are the stresses generated acceptable? Unequal thermal expansion on either side of a stud could generate considerable lateral force on the stud.
Resoonse The liner system has been analyzed assuming a unif orm temperature di stri buti on.
Localized hot spots resulting from sodium jets or localized sodium spills and the boundary between the wetted and non-wetted areas of the walls will be subjected to non-unif orm temperatures. These events could generate lateral forces on the anchors, when the hot plate tends to expand.
However, the liner plate in the colder areas provides restraint to thermal growth and will limit the possible lateral displacements of the anchors.
The response to Question CS 220.36(d) describes conditions in the vicinity of "hard" spots (penetration or embedment) where the thermal expansion of a thicker plate imposes lateral displacanents to the liner studs and shows that the studs and liner are capable to withstand those conditions.
A hot spot presents a similar situailon, but the colder liner plate around the hot spot will provide more restraint than the hot plate around the "hard" spot.
Therefore the stress / strain conditions of the stud and liner will be less severe.
QCS 220.36(e)-1 Amend. 68 May 1982 ____ _ _ _ ___ ___
I i
Page - 57 (82-0184) [8,22] #38 Ouestion CS 220.36 ( f) ( Accendix 3.8-B)
When the liner buckles and bears against the insulation concrete, considerable tensile loeds are generated in the studs.
If the studs don't give, the Iiner could f all in a shearing mode at the stud connection.
Has this possibil Ity been evaluated? If so, what are the results?
Resoonse The response to this question is included in the response to Question CS 220.40(a).
i l
l QCS 220.36( f)-1 Anend. 68 N WT*2
Page - 58 (82-0184) [8,22] #38 Ouestion CS 220.37 (a) ( CRBRP-3. Vo l. 2)
Table 3-14 gives results f or a submerged liner without creep. The applicant should quantify how much actual strains are reduced when creep is taken into account.
Does the applicant really mean creep, or is stress relaxation a more appropriate term? Under conditions of creep, will the ultimate strain capability of the liner material change significantly?
Resoonse The term " creep" was used in Table 3-14 to ref er to the " time ef fects" on the load def ormation characteristics of the material.
It is agreed that the term
" stress relaxation" would have been more appropriate f or a liner subjected to temperature ef f ects.
An analysis that included'the " time ef fect" on the load def ormation characteristics was done using a model of a restrained panel anchored to the concrete with a stud.
This analysis was done f or qualitative rather than quantitative purposes and Indicated a general relief of stresses and strains.
The analysis was ca rled only up to 16000F and a total time of 12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br />, however, due to the qualitative nature of the analysis certain simplifications were made in assessing the time ef fects.
Thus, instead of applying the time ef fect continuously it was lumped at certain selected levels of temperature.
The results Indicated that above T = 12000F there was suf ficient relaxation to prevent f urther plastic def ormations.
The time ef fects on the load def ormation characteristics result in relaxation of stresses and strains will have no significant ef fect on the ultimate strain capabil ity of the material.
i l
i l
l l
l l
QCS 220.37(a)-1 Amend. 68 M
Page - 59 (82-0184) [8,22] #38 Ouestion CS 220.37 (b) (CRBRP-3. Vol. 2) in Section C.3.4.4, the applicant proposes to use " von Mises ef fective strain" f or the liner f ailure criterion.
Keeping in mind that the strain is not necessarily linearly related to stress beyond yield, a rigorous def inition of what " von Mises strain" means beyond the yield point is needed.
e Ultimate strength Is of ten used to prediet f alIure when the f acture mode Is known to be simple cohesive f ailure.
For general ductile f racture, especially when shear fracture is a strong possibility, maximum shear stress is a pref erred f ailure criterion.
The liner can be expected to develop considerable shear stresses, especially near the anchor studs.
Considering the above comments, what is the justification for using " von Mises strain" as the f ailure criterion? Should maximum shear stresses also be considered?
Resnonse The relations that describe elastic plastic behavior, based on the Prandtl-Reuss flow rule, are expressed in terms of an equivalent or of fective stress, O' o (Ref.
QCS220.37I6)and an equivalent or ef fective plastic strain incronent deThe expression f or the equi 1).
as the von Mises yield criterion, F,,
so that just as yielding begins the two are equal.
An expression f or the equivalent or ef fective strain 6, also has the same f orm as the strain E. corresponding to the von Mises yield criterion IC.
- c,end r:c;d the relationship between the ef fective stress and the equivalent or ef fective strain is taken f rom the unlaxial tensile stress strain curve as explained in Ref. Q220.37(b)-1.
Since the expression f or the equivalent or ef fective strain has the same f orm as the strain corresponding to the Von Mises yield stress criterion the term generalized von Mises strain and equivalent or of fective strain or simply generalized strain have been used in CRBRP documents interchangeably. The ef fective stress and ef fective strain are def ined by the following equations:
(% -7.Y + (7 -7 [ + (7 -6,(
2 2
3 3
=5 (C,- EsY + ( Ez.- EsY + (Es-E,f"?
cs Where:
3
~
CT,,(
are the principal stresses
% E,E, are the total principal strains 3 g QCS 220.37(b)-1 Cnvuva
Page - 60 (82-0184) [8,22] #38 lt should be pointed out that Section ill, Division 1 of the ASME B&PV Code accepts the yield criterion and associated flow rule based on the energy of distortion method (von Mises), for plastic analysis of nuclear power plant components (Appendix F, Section F-1321.1.c).
The above subject is also discussed in Appendix 3.8-B, Attachment D of the PSAR (page 3.8-B.18), which justifies the use of von Mises strain as the f ail ure criterion.
The 3/8" carbon steel liner plate used in CRBRP is a very flexible structural eiwnent particularly at elevated temperatures where there is a significant reduction of the stif f ness.
Due to its flexibility the liner can undergo large oef ormations without rupture and its behavior is similar to that of a f l exi bl e membrane.
Ihus, shear f orces f rom the stud anchor are resisted primarily by menbrane f orces that develop as the liner stretches to large def ormati ons.
For this reason shear f racture is not considered a possibil ity.
Further discussion on this subject is provided in the response to question 220.36.
Reference QCS 220.37(b)-1 Mendelson, A. Plasticitv Theory and AcoIIcation. The MacMi l l an Ccrnpa ny, Nw Yor k, 1 % S.
q QCS 220.37(b)-2 Amend. 68 May 1982=--
._=
Faga - 26 (52-0184) L8,22J 738
.Ouestion CS 220.38 During the TMBDB accident scenario several modes of f ailure for concrete Internal structures are considered.
One of these is termed "section f ailure" and occurs when the moment capacity of a section is exceeded (for example, the floor of a pipeway cell). This failure mode involves large displacements (rotations) and theref ore could result in f alling the cell liner allowing additional sodium-concrete Interaction.
The applicant has provided information showing that such section failure do not occur before TNBDB requirements are met. Additional detalis should be provided to show how close these internal structures come to failure as the TNBDB scenario progresses.
Rates at which failure is approached and times that f ailure is expected should also be provided. This should be done for all critical sections. This is particularly important because no f actor of safety is applied to the section failure capacities.
Resoonse Critical sections in the RCB structures have been or are being evaluated for rates of failure progression and expected failure times.
in accordance with the TMBDB scenario described in CRBRP 3, Vol. 2 the Reactor Cavity and the three Pipeway Cells are the only cells in the Reactor Containment Building that would be exposed to sodium.
The base case scenarlo-.
considers failure of the liner to occur at certain times and tiiese failure times establish the minimum' structural Integrity requirements for the liner
) ;tc.:.
C r.:c the liner system is anchored to concrete structures, failure of these structures (sectional capacity exceeded) is considered, conservatively, to imply potential liner failure so that the minimum times for structural Integrity requirements f or the liner apply also to the supporting concrete.
i The reactor cavity and the pipeway cell structures were evaluated and found to meet the scenario requirements with substantial margin in most cases.
A plot of the time variation of thermal moment and moment capacity is given in Figure Q220.38-1 for the upper portion of the reactor cavity wall. This plot provides Information on how the thermal moment approaches the moment capacity for this particular section of the structure.
Although the margin may not necessarily be as high for other parts of the cells in question the plot canonstrales Tnat Tne capacity is not approached asymptotically in the time frame in question and the evaluations that demonstrate Integrity, at the required times, provide sufficient assurance that scenario requirements are met.
QCS 220.38-1 Amend. 68 May 1982 L
P:g2 - 61 (82-0184) [8,22] #38 The questic.: makes specific ref erence to the pipeway cell floor and for this reason a brief discussion is given for this structure. The pipeway floor consist of two layers of concrete; a bottom layer which is anchored to the reactor cavity and the surrounding vertical walls, and a top layer which is free to expand. Th floor liner is anchored to the top layer and for this reason its integrity is not af fected by the deformations of the bottom concrete layer as long as 'here is no collapse.
In any case, liner integrity is required f or only 30 hours3.472222e-4 days <br />0.00833 hours <br />4.960317e-5 weeks <br />1.1415e-5 months <br /> and the capacity of the bottom concrete layer is not exceeded before the boil dry time (132 hours0.00153 days <br />0.0367 hours <br />2.18254e-4 weeks <br />5.0226e-5 months <br />), so there is substantial margin.
I e
1 QCS 220.38-2 Amend. 68 May 1%2
MoMcur l
3000 C4 M ciry C
O
_9 b
k Y
I j
k zwo
.o 7WERM4L~]l U
.l L
l 2
~ ~<~r O
W f
/mo _
p]
al4 Aw t
T 4%
$vil al t Nl l
I o
i I
l l
0 So 79
,99
,,5
,o 7~<' M E HouMS FIGURE Q220.38-1 REACTOR CAVITY WALL - UPPER REGION i
TIME HISTORY OF THERMAL M0 MENT AND M0 MENT CAPACITY
Ouestion CS 270.40 (a)
The appi ? cant needs to determine the most likely bucklin liner.
If the shear symmetric mode is expected, g mode for the cell anchors must be evaluated to determine shear stress sta PSAR.
and provided in the Resnonse The cell liner analysis performed Indicate that the m pattern for the cell liner is that of alternate panels bost likely buckling a checkerboard pattern.
The reply to this question also describes the cri 220.36(d).
prevent plate / stud tear out failures.
a adopted to A case of symmetric buckling, where the liner bears imposes a large tensile load on the stud anchoragainst the concrete and Figures QCS220.40(a)-1 and QCS220.40(a)-2.
, has been examined.
See compression develops in the liner in equilibrium w stud, membrane Consequently, no significant shear stresses can be de stud tension.
Strain levels in the liner and stud under this mode a limits.
veloped in the liner.
re within the acceptable In a further evaluation of shear ef fects, membrane action and that punching shear develops in theit was assumed The sheer stresses in the plate were determined to plate. The shear acity of the stud.
of the ultimate tensile strength of the plate for 1/2 inche respectively diameter studs.
and 3/4 inch QCS 220.40(a)-1 Amend. 68 RhoWM hflfTL ---~
MAXIMUM GENERALISED STRAINS M=
MEMBRANE STRAIN M + B = MEMBRANE PLUS BENDING STRAIN i
MAXIMUM LINER STRAIN (in/in)
.019 (M + B)
LINE OF SYMMETRY
.011 (M)
[.18"
.05" 0.0"
.05"
.10"
.15" y
1 STUD STRAIN (in/in:
.10" g- -
.021 (M + B)
N l
/
N l
/
l
.021 (M)
\\
l
\\l/
l
\\
STUD ANCHOP
.15 g
N N.15"
\\
l LINER DEFLECTI
[
I N
['1go.
INFLUENCE LIN
.20" l
/
\\
\\
h
/
.05"
\\
/
'- 0.0" I
\\,
\\,/
l
.25" l
i
__g l
(__
7 g
y N
.05" N
\\
\\
l
\\
l,/ /
\\
/
\\
l
- 25"
.25"
.20"
.15"
.10" 4 Spaces @ 1.875" = 7.5" LINE OF SYMMETRY l
CELL LINER - SYMMETRIC BUCKLING; TENSION IN STUD FIGURE Q220. 40 (e)-1 l
0220.40(a)-2
LINER DEFLECTED SHAPE UNDEFLECTED
.T.S.)
SHAPE
[-----
_f_f u
I, *..'
'~,* ', '. -
STUD ':-
s.
gypy
/
STRUCTURAL FACE OF INSULATING M 77 CONCRETE 77 7 CONCRETE INTERFACE CELL LINER - DEFLECTED SIIAPE - SYMMETRIC BUCKLING FIGURE O220. 4G (a)-2 0220. 40 (a)-3 c-.
Paga - 34 (82-0184) L8,22J #38 Ouestion CS 220.40 (b)
Evaluate the ef fect of pre-existing cracks and of cracks generated during the life of the plant; discuss possible propagation of these cracks in the liner.
Resoonse The propagation of pre-existing cracks in the cell liners is not mechanistically possible. The cell liner system is designed, fabricated, installed, and Inspected as an Engineered Safety Feature. The requirement for the ESF cell liners are defined in PSAR Section 3.8-B.
The materials, f abrication, welding and construction NDE requirements identified in PSAR Appendix 3.8-B, Attachment B, preclude the incorporation of defects in the cell liner which are capable of propagation. The acceptance standards for cell liners have been extracted from ASME B&PV Code Section ill, Division 2.
The generation of cracks in the cell liner system during the plant lifetime is prevented by the design. The strain ilmitations imposed for DBA conditions preclude plastic instability. The generation end propagation of cracks by f atigue is addressed in the design requirements (PSAR Appendix 3.8-B).
A fatigue evaluation of the cell liners using ASME B&PV Code Section lil, Division 1, Section NE-3222.4d concludes that for the thermal cycles specified for the cell liners, a fatigue analysis is not reautred.
Additionally, the potenital for cell liner brittle fracture due to raidation has been investigated in accordance with the methods and Ilmits established by USNRC Re3L i u l v. y Goido 1.99 " Effects of Residual Elements Predicted Radiation Damage to Reactor Vessel Materials." In the worst case exposure condition, at the rector cavity beltline, the resultant maximum adjustment in the Nil-Ductility Transition (NDT) temperature is 10 F.
This indicates that the cell liner is not affected by neutron embrittlement.
The propagation of cracks in the cell liner is prevented from impacting the response of the cell liner under DBA sodium spills by both scheduled and unscheduled inservice inspection requirements which are identified in PSAR Appendix 3.8-B, Attachment E.
The scheduled inservice inspection program assures the continued Integrity of the overall cell liner.
Unscheduled Inservice Inspection is required in the event of a cell liner repair, small sodium spill, any indication of corrosion in excess of design limits, any thermal cycling in excess of the duty cycle or indications of excessive oxygen in-leakage.
QCS 220.40(b)-1 Amend. 68 May 1982
Pago - 69 (82-0184) [8,22] #38 pggstIon CS 220.40 (e)
'I In addition to the non-uniform temperature distribution in liner to be considered (Ref. Question 220.36e) the response to a shallow pool spill (for P I'-
instance localized in a very stif f area such as a corner) should be evaluated.
Resoonse The response of the cell liner to a shallow sodium spill, localized in the b
region of a cell liner corner, has been evaluated.
An analysis has been performed using the Wall-Floor Liner corner model shown in Figure 3A.8-3.
The
.# I' analysis assumes the cell liner floor panel, located adjacent to the corner,
- I to be heated to 1000 F while the remainder of the cell liner is maintained at 70 F.
The Induced strains in ihe liner are lower than the strains reported in Section 3A.8.3.3 for the uniformly heated liner case.
c-DC Actual conditions preclude such a confined hot area since the conductivity of the plates will result in a rapid redistribution of heat.
Us T6
-1 Tl
'l
.'ry ve-3.
i QCS 220.40(c)-1 Amend. 68 r2% RD2
Maga - 36 (82-0164) LB,22J 736 Questfon CS 220.41 The portion of containment at cells that experience pressure over 10 psig may have to be quellfled as ASME Section 111, Division 2 concrete pressure vessels because the concrete is relied upon to carry a portion of the pressure load, unless applicant can justify his position in the PSAR.
Resoonse The RCB lined cells are not a part of the Containment System.
As saf ety related reinforced concrete structures the cells are designed per ACI 349.
This code includes design requirements for pressure and other loads on saf ety related concrete structures. Theref ore the ASME Section lil, Division 2 Code is not applicable in this case.
QCS 220.41-1 Amend. 68 May 1982
P2g3 - 31 (62-0164) LB,22J FSB Ouestion CS 220.42 The applicant has performed bounding seismic response calculations by using
" soft" and " stiff" soll springs. The applicant is required to show that Intermediate values of soll stif f ness will not increase the response levels.
This could be shown by comparing modal frequencies of major contributing modes to the response spectrum.
Resoonse The attached Figure QCS 220.42-1 compares the response spectra at the operating floor for the following soll properties:
a) upper bound b) average c) lower bound The curve for the average properties f alls within the band of upper and lower bound properties for the entire frequency range under consideration.
Thus, intermediate values of soll stif f ness will not increase the response levels.
QCS 220.42-1 Amend. 68
_. _ _ MapJ982___
COMPARISON OF RESPONSE
- [
FOR AVERAGE, UPPER BOUND AND LOWER BOUND
[
SOIL PROPERTIES
[
RCB EL 818'-O" SSE N-S MODEL E
DAMPING = 3%
5-i i
1
.;i
.n 9
.\\
.,M l
L
..\\
L
^
e h
3 5-1 l1 e
zO L
5.
J g-e e,
i-
=
W J
e :-
W
=
0 E
O 4
2 W
C/)
E W
-h u)
LEGEND E-Z 4
AVERAGE
=
E UPPER BOUND h
4_f LOWER BOUND I-4 5-5 E
-==
b 4-E f f t t t t ttlttttttitif f f f f!!t ti!tttt! infmtfrin! t t t t t t t t t !s i t t titit!tttttttttf f lit!!!!!!It!! !??!!!!!'!"n!
t t t t I t lt t t tt t t ttlt!?tt'tt'
! I t
!!tt
- 2
.3 4
.6 A
I 2
3 4
6 8 10 to 30 40 FREQUENCY (HZ)
FIGURE Q220.42-1 0220.42-2
Page - 73 (82-0184) [8,22] #38 Ouestion Cs 220.43 (a)
The applicant used a finite element axisymmetric structural mmputer code to evaluate containment buckling f rom thermal stresses derived f ran other analysis. This is probably inappropriate beccuse of the short wave lengths (local wrinkling ef fects) actually expected f rom thermal stresses in the sh el l. This local buckling results must be reconsidered in the applicant's evaluation of the containment shell in the PSAR.
Resoonse Thermal Buck!!no for CRBRP Containment Vessel.
A linear elastic analysis was perf crmed to mmpute the critical buckling temperature using the finite dif f erence computer progran BOSOR 4* (Ref. QCS 220.43(a)-1). This method is considered appropriate since the short wave length (local wrinkling) ef fects are accounted f or.
BOSOR4 results showed that the thermal buckling is of a local nature and occurs in the vicinity of Elevation 816".
The mathunatical model has a cylindrical shell and a dome.
The radius of the cylindrical shell is 93' and its' thickness is 1.5" f rom El. 816 to El. 849'.
The vessel was assumed clanped at Elevation 816 and a uniform, axisymmetrical temperature rise was assumed. The critical temperature (Tcrit) was computed using bif urcation buckling analysis which is treated as an eigenvalue problen by BOSOR 4.
In the bif urcation buckling problen the eigenvalues ( ) are compted f or several circumference wave numbers (n) and the stability equation is solved f or the smallest eigenvalue.
The value of (n) corresponding to the minimum is 57, i.e., there are 57 circumferential waves at the buckling temperature. This is used to calculate the value of Tcrit.
The value obtalned Is Terit = 7080F, based on a rof erence temperature of 700F.
Theref ore, the buckl Ing temperature is 778oF.
In addition to the computer calcuiations, a hand calcuiation was perf armed using the method of ref erence Q220.43(a)-2, which provides a procedure to determine the critical buckling temperature for cylindrical shells with various edge conditions.
The hand calculation gave a value Terit = 10460F for a hinged and cylindrical sh el l. The method of Ref erence QCS220.43(a)-2 gave a higher critical temperature (Tcrit) for a 1-3/4 inch thick cylinder than for 1-1/2 inch.
Since BOSOR4 results are based upon a 1-1/2" thick vessel, while the revised thickness is 1-3/4", the computed Terit is on the conservative side.
The temperature that causes yield at the bottom of the Containment Vessel is 2700F, which is substantially lower than the buckling temperature.
Based on this, the critical stress f or thermal buckling can be considered to be yfeld stress.
QCS 220.43(a)-1 Amend. 68
Page - 74 (82-0184) [8,22] #38 References QCS220.43(a)-1 D. Bushnell - Stress, Stability and Vibration of Canplex Branched Shells of Revolution: Analysis and User's Manual for BOSOR 4-NASA / Langley Research Center, Hanpton, Virginia, March 1972.
QCS220.43(a)-2 D. J. Johns, " Local Circumferential Buckling of Thin Circular CyiIndrIcal ShelIs" NASA TND1510 (1%2).
- See Appendix to this response f or detail s on BOSOR4.
l i
(
i i
l 4
l i
l f
.i QCS 220.43(a)-2 Amend. 68
. _..May 1982
Page - 75 (82-0184) [8,22] #38 Ouestion CS 220.43(a) - Appendix BOSOR4 COMPUTER PROGRAM B050R4 is a computer progran f or stress, stability and vibration analysis of shel ls of revol ution. The program was developed by D. Bushnell of Lockheed Missiles and Space Company.
The computer code is based upon the linear, elastic, thin shell theory.
The structure should be axisymmetric. The progran can handle various kinds of i
wel l material s and l oadings.
Both.nechanical and thermal loads are pennitted in the analysis.
In cases involving stress analysis of a shell for non-exisymmetric loading, the program finds the Fourier series f or the loads, calculates the shell response in each hannonic to the load components with that hannonic, and superimposes the results f or all hannonics.
The progran has an option by which the stability analysis of a shell can be treated as a hif Jrcation buckling problem and mathematically it is treated as an ei genval ue probl em.
The program also handles shell vibration as an eigenvalue problem and f inds mode shapes and f requencies.
l BOSOR4 uses a finite-dif ference scheme as a numerical technique in the solution of shel l probl ens.
l I
QCS 220.43(a)-3 Anend. 68
Pagi - 74 (82-0184) [8,22] #38 Ouestion CS 220.43 (b)
The applicant must consider the torospherical wrinkling caused by Internal pressure in their containment evaluation in the PSAR.
Resoonse in some torospherical shells under internal pressure, compressive membrane stresses are developed; however, for the particular configuration of the CRBRP containment there are no such compressive stresses as shown in Table Q220.43(b)-1.
The configuration of the steel containment vessel dome is shown in Figure Q220.43(b)-1.
It consists of a 1.25" thick ellipsoidal knuckle and 1-3/16" thick spherical cap.
Tlie specified steel is SA516 Grade 70. The design internal pressure is 10 psig.
Membrane fcrces N4 and No.are shown in Figurs Q220.43(b)-2 and Table Q220.43(b)-1.
N4 = Meridional membrane f orce in lbs/In.
4 = Hoop membrane force in Ibs/in.
Sign Convention: Positive sign is tension, Negative is compression The membrane forwes clearly indicate that +he dome is subjected to tensile stresses only under the Internal pressure.
Hence, stability of the dome is not a concern under the Internal pressure and torospherical wrinkling will not occur.
l l
l l
i l
l I
l l
QCS 220.43(b)-1 Amend. 68 m
p S ( RCB i
I o
i
..O i
o a
a d
p o*
=
..O o
g g
8 I
"o
,V, cn v
_a n o
a i
b d
$ $_6 k'%
0
'O n
w 2-c-
'n d
x w
tu p p
TANGENT LINE E L 899'-O"
, 80 1 y
y 48' EL 88T-10'/
N 33* he F
"o 'l' b' 3'h
,+, a
- CRANE GIRDER W
_ _d e
w-o1 0
"9
'b SHELL E L 85G' -O" STlFFENERS dD
-l 839' 4 85G' IR = D3'-O" b
i
~
i_.
O.
E L 6b? -O" 4
5 M
_J L __ EL 81G'-O" OPERATING FLOOR E L t_O7
r - -_
i n
i I
4l I
I
'I i
I j
W l
l SHELL & STIFFEMERS
=__
I MATERIAL: SA SIGGm ?O I
I I
i I
I j
r[j l
cTOP OF'4' LINER Ft-TOP T NSE R fei727-2'z-
+-~ " ' = ' - o" l
I FIGURE Q220.43(b)-1 CRSRP CONTAINMENT VESSEL 0220.43(b)-2
N $ = N 0 = 6018
- : a N$ e 6020 g
g NS = 4496 4
& = 125 f
<b
\\
N4 = 5584
& = 90 '
Ed y
El. 899 NO = 5407 o
N&
-i g
l
-N0 o
"^Tr:
2_. 22. 3 coment due to ll l
internal pressure is o
l insignificant in t'p head.
o il il o
El. 816' NPHZ, NTHETA ALONG MERIDION AT 0 DEG. (PSI)
CRBRP INTERNAL PRESSURE 10. PSI MAXIMA NC = 6038.
NO = 11659.
FIGUREQ220.43(b)-2 MEMBRANE FORCES Nt AND NO DUE TO INTERNAL PRESSURE Q220.43(b)-3
TABLEQ220.43(b)-1 MEM5RANE FORCES IN DOME DUE TO 10 PSIG INTERNAL PRESSURE
~ '
055
'OLLMN 1
2 3
l 4
IN7. PRESS.
E M OF DEAD SUM OF DEAD
+ SCAF. CLIPS
.5 PSI 10 PSI LOAD + IT PRESS.
LOAD + INT. PRESS g
LES/IN.
LBS/IN.
LES./IN.
COL 1&2 LBS/IN. COL 113 LSS/IN.
i N9 NG N9 N9 N9 NG M9 NG.
N9 NG 90
-335 259
-279 270 5554 5407
-61 4 1
5249 5576 92.5
-324 469
-279.
127 5555 2544
-603 342 5252 3013 95
-31.6 522
-230
-62 5592 1235
-596 470 5276 1767 97.5
-307 507
-230
-55 5604 1105
-587
,452 5297 i61 2 100
-299 463
-231
-65 5621 1290
-520 398 5322 1753 102.5
-293 423
-232
-73 5642 1450
-57 5 350 5349 1873 105
-255 389
,-213
-78 5557 1553
-569 311 5381 1952' 110
-275 325
-287
-90 5732 1758
-552' 236
'5457 2124
)
105 5315
'2092
-557 156 5549 2353 115
-255 251
-291 l
120
-253 196-
-256 118 5918 2351
-554 78 5550 2557 125
-252 253
-301 225 6020 4496,
-553 23 5768 4749 304 6035 6b78
-563
-325 5774' 6055 130
-251
-22
-302 135
-151
-52
-302 302 6034 6033
-553
-354 5783 5981 140
-242
-85
-302 302 6034 6034
-544
-357 5792 5949 150
-227
-139 *
-302 302 6034' 6034
-529
-44T.
5507 5595 160
-223
-124
-302
- -302 6034 6034
-525
-485 5811 5250 170
-205
-195
-302 302 6034 6035
-507
-497 5829 5840' 175
-203
-201
-302 302 6032 6037
-505
.-503 5529 5535 NOTES:
3.
Column 3 shews the distributien of membrane forces (Nd, NG) due to internal pressure in the deme.
2.
d is defined on Figure QI20.43,('6)...
2 Q220.43(b)-4
_.,.__,anv.,___.___.,. _ _ _ _ _.. _,, _ _.- - -, _,,,.,,-, -,,,. _
-en-,--
Pega - 75 (82-0184) [8,22] #38 Ouestion CS 220.43 (c)
The applicant has not evaluated buckling of large openings and buckling near local penetrations in the PSAR.
We require this analysis be included.
Resoonse Buckling has been considered at large openings and near penetrations.
All penetrations are f ully reinforced in accordance with ASME Code rules, therefore the buckling strength of the vessel will not be reduced by these penetrations. Of the two major penetrations above the operating floor, the equipment hatch was not only reinf orced in accordance with the code but buckling is prevented by a stiffening frame provided around the opening as shown in Figure QCS 220.43(c)-1.
The other major penetration is the Equipment / Personnel Airlock and the stresses around the opening will be checked against the buckling allowable.
l l
QCS 220.43(c)-1 Amend. 68
[hn 1982
i m
I
~.
'l T
??
+
l u
4 5
83 l
1~;
/r i
s I
6
~
l
/
I Ji I
/
l
~ ?
if W
e at 1o3 i 5
l i
$$ 4 l,s I
J F*
w a cr
/
i.
i 7 &.Hi I i ji +,'
=~j &
-l-5--
i m
9,..--
.z
,/
Q a
I
/
_L
".* 1
,e
~,
e n u n nini., w 4
'h',
l D.i i
e y
c l
h.
i o
/;
3
,n t
.__.m__
J l
I W.
~
l I
4
, 4, -
e,/
I T
I y
l
.,w---
2 l
i
,1 X
a l
s 9
a l
1 h
_1
/
g
-a a
a ;
a Ix y
I y.
j s
t w
)...'!
)
I 4c2% _ _1
+-t<,ij
~
- - + _ - - _ = _ - -. -1, ---
-- ~--- -
. 7. g ;
i d's i
I
's +
- i. j m,.
Ns' l
l 3
^
s i
4'l ss r
,s l l!
I E I
l lll N sN s
lll
't't i I
Ill 111
'l l
q!
s m
I M<s I
M
]
s l
l
-M llll jD1!
i
's I'l i b
?
6-
'i
$ 'e' 21-i I
4.j g
4 W
W
~
w 0220.43(c)-2
l PJge - 78 (82-0184) [8,22] #38
_Duestion CS 220.43 (d)
Provide a discussion on your analysis of the containment buckling in the region adjacent to polar crano support.
Resoonse The polar crane support consists of two ring stif f eners and 60 equally spaced vertical radial stif f eners or gussets as shown in Figures 1 Q220.43(d)-1 and Q220.43 ( d )-2.
The stresses on the chell including those fran the Containment Vessel itsel f and those dn to '..e polar crane reactions.
The polar crane reactions were provided by
's crane manuf acturer and were applied as concentrated loads by means of Fourier Series on the shell using the Kalnin's program.
Buckling was checked in accordance with the buckling criteria given in PSAR Appendix 3.8-A.
The ring stif feners were checked in accordance with new Section 7 of Appendix 3.8-A.
The vertical stif feners (gussets) were checked as column :octions per the AISC code.
The critical buckling siresses of the shell plates betwren the ring stif feners and vertical gussct plates were calculated using the equations presented in Sections 3.2.a, 3.2.b, 3.2.c and 3.2.d of the PS AR Appendix 3.8-A, for axial compression, circumferential compression, torsion and bending respectively.
The critical buckling stress f or transverse shear was made equal to 1.25 the critical shear stress f or torsion per Section 6.0 of the PSAR Appendix 3.8A.
Saf ety f actors of 1.33 for the OBE and 1.11 for the SSE were used in accordance with Table 3.8-A-2 of the PSAR Appendix 3.8A.
Based on the calculated shell stresses, critical buckling stresses and saf ety f actors, the stability was evaluated through the interaction equation in l
rev! sed Section 6.0 of the PSAR Appendix 3.8A and the results were shown to meet ine acceptance criterion.
l l
l l
l l
l QCS 220.43(d)-1 Amend. 68 I"vuSU;f9L--
r SEIII THI C ESS 5"ttL CONTAINMENT CONCRE*t
\\
00NFNt h TNT El. 974-3-3/16"
.J e g
ss.
N'
- 7.,!
/
E1.931'-5-3/16* /
J4
/
\\
s'- i
,[
oct a e coaNg ttl tEl. 999'.0"
/
S? tau GENtea*09 M
j
.t.....
f VtSSE*. 23 18 6'* 0*
S E
- v'c g 9.;G g
m o tN...,:.
C.
l 1
54 A4 WEv
~,
l r - l- - --
(
enena No rt.
l gf s-EL.916 c*
, n.sor -r m
1 A
1 y
I esNt* CSNC9t*t l
wa at38'tnica) d!
CV?tm CONC *tQ
}
wAu(12 %.cn a g
l
~
! }j }'
SANCwlCwtS t'Er. '
f*Ett L NE4 CCNTAssus.3T vt&5C.
i
-i l
EL.733'-Q*
' 1l i
- b l
- CNP CONCSE*E 9
ys==
,,,-c..rse r
- ,,,..c..
t-
/
t..*17-C'
.A
- .1/;"
D 1
g' iT i
t,
, vtt$r.
FOUNCAPCN g g er g'.f 5M.4
- WA?
4 se i
l FIGURE Q220.43(d)-1 CONTAINMENT BUILDING CROSS SECTION 0220.43(d)-2 l
R= 1116.875"
,R = 1084.0" I
EL. 890'-6" 6"
gg,
./ s
\\
21" 1-3/4" 26" 1-3/4"PL o
e 1%"
_it ss-Il
_ VERTICAL SUPPORT PLATES u
A A,
y 12" p
1-3/4" 45" i'
O SECTION A-A
'1" GUSSET PLATE i
(60 AROUND SHELL) i k
m l-3/4" PL EL. 875'-0" 2 9 -- 3 /4 "
/
o EL. 874'-10" 5'-11" V
Y FIGURE Q220.43(d)-2 SECTION THRU CRANE GIRDER SUPPORT 0220.43(d)-3
Paga 41 [8,22]#39 fasg Factor
- Local Buckling (2.00/1.67) x (1.67/1.11) = 1.8 with SSE Conf inement Construu..n Leeds (See Note 2 Table 3.8A-2)
(1) With Snow load 2.0/.5
= 1.33 (l1) WIthout SnowIosd 2.0
.67 = 1.2
- These f actors are based on the saf ety f actors given in Table 3.8A-2 of thi s Appendix.
6.
Interaction Ecuation (includina Thermal Stresses)
To avoid buckling due to concurrent thermal, seismic, pressure, dead and live loads at the operating floor, where radial thermal expansion is restricted, the f olicwing criteria (given in the f orm of Interaction equations) will be met:
(1)F (2)F (4)p (g3)y 7
+
F C
+
+
(
+
(1) critical
( 2)criti cal g4) critical
( 3)e Itical g5)F
)2 (cih I2
)
+
(
) i 1
[) critical)
(gth al lowabl e)
)
(
)
where:
F(1)
= Ccrnpressive merldlonal stresses due to axial loads p 2)
= Compressive hoop T(3)
= Membrane shear stresses due to pure torsion 6(4)
= Compressive meridional stresses due to bending-g5)
= Mebrane shear stresses due to lateral shear 6+h
= Thermal hoop membrane stresses th g
al lowabl e =.67 yleid f cr OBE Design Load Combinations
=.80 yield f or SSE Design Loed Combinations F = Factor of saf ety as def ined in Table 3.8A-2 for SSE Design Load Combi na ti ons.
For OBE Design Lead Combinations, increase the f actor of saf ety by a f actor of 1.2 geritical = Critical stress values,(as def ined in Section 2 of this Appendix, with the exception of g 5) critical which is def ined as g(5) critical = 1.25 g(3) critical.
3.8 A-10 Amend. 68 w
P3e52[8,22]#39 9
NOTE: The stresses to be used in the above interaction equation are all to be taken f or the same point of the shell under evaluation.
Points will be evaluated between elevation 816 feet and up to and including 3 Rt above 816.0 ft.
7.
Horfzontal Stiffener Reaulrements to Assure Overall Stability of a Cylindrical Shell The criteria f or sizing the horizontal stif feners, to assure overall stabil Ity, Is based on comparing a total moment of inertia,1st, for the stif f ener to the moment of inertia required by the ASME Code. The greater of the two moments of inertia will be used in the design.
It should be noted that this applies to ASME stif feners only, for non-ASNE stif feners the total moment of inertia,1st, will be used.
The f ollowing method to calculate 1st is based in part on " Fabricated Cylindrical Shells Under Canbined Axial Compressive Load and External Pressure" by C. D. Mil ler, CBI, 7/77 Revision:
1 st = l sa + ISC + l ss i st = Total manent of Ir.ertia f or the sti f f ener Isa = Moment of inertia to control axial buckling (see Section 7.1)
Isc = bbment of inertia to control circumferential buckling (see Section 7.2)
ISS = Moment of inertia to control torsional buckling (see Section 7.3) 7.1 isa - Moment of inertial to control Axial Bucklino Isa is the stif fener moment of inertia required to prevent overall Instability of the shell under axial loading.
Isa = 5.33 F t3 Ls/('Ls/ Rt)l*0 Where:
2.,.; c..' ;;.; W.ess of the shell to control axial buckling at the stif fener location Ls = Ef f ective length of the shell to be used f or sizing the stif fener R = Radius of the Vessel i
F = Saf ety f actor of 1.5 7.2 lse - Moment of inertia to Control circumferential BuckIfno isc is the stif f ener moment of inertia required to prevent overall Instabil Ity of the shell under a unif orm radial external pressure.
3.85-11 Amend. 68 M
Paga 42 [8,22]i39 c
( 1.2 ATH}
isc
=
~
(n,3 )2 2 2 2
2 2
(n -1)
(n +.5h -1) s
___6 Where:
Design length of a vessel section L
=
The radium to the centroidal axis of the combined smaller R
=
c stiffener and effective width of shell in inches Equivalent thickness of the shell required to resist the t
=
membrane hoop stresses. This thickness should be calculated, based on the criteria of this appendix, for hoop stresses with a saf ety factor for general membrane stress as found in Table 5.
Number of waves into which the shell will buckle in the n =
circumferential direction.
A/.8 where A is the f actor used to enter the applicable A
=
TH material chart in the ASME Code h=
R/L where L is the length of the cylinder between stiffeners.
See Note (1).
The value of Isc is found by selecting an integer value of n 2. 2 which maximizes the right hand side of the equation.
7.3 iss - Moment of inertia to control Buckling Caused by Shear Loads Iss is the stif fener moment of inertia required to prevent overall Instability of the shell under torsional or lateral shear loads.
Moments of inertie are calculated by using the formulas In Sec. lon 7.1 and 7.2, with "t" being defined as the thickness required to resist shear. The-higher moment of inertia is determined and defined as Iss.
Netc 't): L, is the length of the cylinder between bulkheads or stif fening elements with suf f icient stif f ness to act as bulkheads. A head can be considered as a bulkhead plus an additional length of cylinder of one-third of the depth of the head.
This means that for the CRBRP Containment Vessel, a conservative Lg, equals the length of the cylinder (899-816=83') plus 1/3 of the head.
3.8 A-11-1 Amend. 68 DRE5R
u i
Page - 79 (82-0184) [8,22] #38 5
Ouestion CS770.44 We require additional detall for the reactor vessel support ring bef ore
. s-eval uating that structure.
In particular, the means f or transferring load f rom the ring to primary concrete structure needs to be provided. The applicant should also provide their analysis of this load transfer path and the predicted saf ety margin.
Resoonser Additional details f or the Roactor Vessel Support Ring are provided in response to Question CS220.34(a).
The Reactor Vessel Support Ledge (RVSL) is a steel structure supported by, and T
embedded in, the Reactor Cavity as shown in PSAR Figure 3.8-9.
A system of horizontal ring plates, radial brackets and stif feners, and vertical cylindrical panels welded together, comprise the Reactor Vessel Ledge.
The two upper ring plates, at Elevation 800'-7-1/4" and at Elevation y
795 '-1 -1/ 4", receive the loads directly from the Reactor Vessel Support Sy stem.
The upper support ring plate receives the downward load by bearing.
The l ower support ring plate takes the upward loads through hold down bolts.
(The hoicdown bolts transf er the upwerd load to the lower support ring plate).
The base plate (El. 7 60 '-7-1/ 4" ) is embedded in concrete and is also anchored U
in the concrete by anchor bolts.
The ext eri or cy l inder ( R=20 '), together with the radial brackets and stif f eners, transf er the vertical loads f rom the support plates to the base plate.
The two Inner cylinders (at R=12'-0-1/2" and at R=13 '-11-1/2")
transf er loads and contribute to stif fen the Intersecting plates by forming box sections.
The radial brackets continue outwards (beyond the exterior cylinder at R=20')
te cen:tut: *ha radial stiffeners. These radial stif feners are enbedded in the cavity concrete and act as shear keys f or horizontal loads.
Thus the vertical loads, upward and downward, are f irst taken by the upper and lower plates respectively.
These are then transferred to the base plate by (1) the radial plates and (2) the exterior vertical cylinder.
The downward verti cal loads are transferred to the concrete by the base plate by bearing.
The upward vertical loads are transf erred by tension In the base plate anchor bolts and shear in the concrete above the base plate, f
1 r
QCS220.44-1 Amend 69 1
Pag) - 60 (82 0184) [8,22] #38 The horizontal loads (selsmic) are transf erred through the upper horizontal plates, and the radial brackets to the stif fener plates embedded in the concrete acting as shear keys.
The load is transf erred to the concrete by bearing against the stif f ener plates.
The 1/2 inch radial gap between the exterior cylinder and the concrete, allows f or f ree thennal expansion of the ledge under DBA conditions.
The most severe loading on the ledge is the Sl@DB load.
(see Response to Question CS220.34(b)). Project assessments show that the support rings can accommodate the f uli SfEDB lced.
QCS220.44-2 Amendo @9
Page - 30 lb2-0104) LB,22J Fod Ouestion CS 220.45 From your presentation on seismic analysis it appears that in your mathematical model, you used the 2-D finite element analysis to derive the spring constants for soll and used damping values on the basis of half space theory. The staff has reservations in such an analysis approach and your justification for such an approach is requested.
Resoonse The use of half-space theory is an accepted method to determine spring and damping constants for soil-structure Interaction, if, Instead of using the theoretical equations, the half-space is represented by finite elements, the calculated spring constants will be the same (or close) as the theoretical values.
Since the material under the CRBRP foundation mat is not strictly a half-space because of the inclined layers of slitstone and limestone, a static finite element approach was used. The purpose was to define an equivalent half-space. The rock under the foundation is not considered sufficiently rigid to jusfify a fixed base analysis.
Although a two-dimensional finite element analysis was used, a correction f actor for three-dimensional ef f ects was introduced as Indicated in Section 3.7.1.6 of the PSAR (Page 3.7-3b).
The spring constants calculated in this manner were compared with the values calculated using 1he half-space equations (using siltstone as uniform material tr.d:r 'hc f undation) and the agreement between the two sets of values was satisfactory.
The damping values f or the soll-structure Interaction were calculated with the theoretical half-space equations. This was further justified by the satisf actory agreement between the finite element calculated springs and half-space equations.
In addition, the analyses were performed for upper bound, lower bound and average rock properties and the results for the three cases were enveloped.
This would account for any uncertainties in properties and methods of calculations.
QCS 220.45-1 Amend. 68 May 1982
_ _ _ _ _ _ _ _ _ _