ML19305C316
| ML19305C316 | |
| Person / Time | |
|---|---|
| Site: | Trojan File:Portland General Electric icon.png |
| Issue date: | 03/17/1980 |
| From: | Broehl D PORTLAND GENERAL ELECTRIC CO. |
| To: | Schwencer A Office of Nuclear Reactor Regulation |
| References | |
| TAC-07551, TAC-11299, TAC-12369, TAC-7551, NUDOCS 8003260520 | |
| Download: ML19305C316 (47) | |
Text
.-.
a March 17, 1980 Trojan Nuclear Plant Docket 50-344 License NPF-1 Director of Nuclear Regulation ATTN: Mr A.
Schwencer, Chief Operating Reactors Branch #1 Division of Operating Reactors U.S. Nuclear Regulatory Commission Washington, D.C.
20555
Dear Sir:
The March 7, 1980, meeting between the NRC Staff, Licensee, Bechtel Power Corporation and consultants was held for the purpose of discussing and resolving structural considerations with regard to the proposed modifications to the Troian Control Building which were expressed in the NRC Staff's SER as being unresolved. As a result of the meeting, Licensee agreed to provide additional information with regard to these unreso.'"9d structural considerations. Such information has been developed and is provided in the enclosed submittal.
This letter summarizes Licensee's principal conclusions with respect to the information developed pursuant to these struc-tural considerations.
Further, as discussed between Licensee and the NRC Staff, certain additional structural improvements which add to the seismic capability of the existing structures will be incorporated into the modifications.
These improvements are described in the testimony being filed by Licensee today in the Trojan Control Building proceeding.
While we believe that evaluations performed to date are conservative and in keeping with industry standards, neverthe-less in accordance with the NRC Staff's requests we have performed additional analyses taking into account the postu-lated effects of such considerations as gross bending, dead load reductions (due to creep and shrinkage), relative dis-placement between panels due to slipping and elimination of resistance provided by beam-column connections.
Even with these. additional conservatisms, we have concluded that the
/900/
S
/!/
8 oose coq
G Mr.
A. Schwencer March 17, 1980 Page 2_of 2 modified Complex overall has an available capacity at least equal to that required to satisfy the 1.4 load factor OBE criteria.
In addition to conventional structural analysis considerations, the foregoing effects have been evaluated with regard to structural capacity, displacements and floor response spectra. While we do not concur that these effects are realistic structural considerations, we have demonstrated the necessary assurances that the modified Comple x provides the required margin in accordance with OBE criteria specified in the FSAR.
In light of the discussions at the March 7 meeting, we believe that the Staff will now have sufficient inform'a-tion available to it to enable resolution of the open items.
We appreciate the Staff's cooperation in meeting with us in San Francisco and in expediting its review of this material in anticipation of the forthcoming hearing.
Sincerely, cc: Service List 45F
0 1
i a
Supplemental Information Requested at Meeting of March 7, 1980 Q. 1.
Page 1 of 18 Capacity Consideration a)
Address diagonal shear cracking capacity as a function of l
aspect ratio.
b)
Examine additional walls to determine potential effect of gross bending on capacity / force ratios.
c)
Quantify the effect of consideration of embedded columns as equivalent reinforcing on stiffness; also the effect of potential reduced dead load.
d)
Assuming that iotal base shear remains constant, predict the relative distribution among the shear elements with gross bending and potential dead load reduction considered.
e)
Verify that double curvature flexural crg.acity is still the controlling capacity including consideration of diagonal tension.
f l
ED-1 l
l
Supplemental Information Requested at Meeting of March 7, 1980 Answer 1 (a & e) Page 2 of 18 For evaluating the capacities of the wallt af the modified complex, three distinct behavioral modes are postulated.
(a) flexural mode (b) sliding mode (c) diagonal tension (shear) mode Equations for each of these behavioral modes are described below with appropriate references from where they have been taken, unless they are directly developed.
For the purpose of comparison of capacities corresponding to the above modes, four different thickness of composite walls of the Complex are considered:
20",
23",
32" and 48".
- Also, for purposes of the example calculations only, the example wall panels are assumed to have horizontal and vertical rein-forcing steel ir. the block portions only.
With reinforcing steel in the concrete core, the capacities of the wall sections will increase accordingly.
For purposes of the example calculations only, floor beams are assumed as W24x68 and the steel columns as W14x142 because these sizes represent the typical steel sections.
Two typical values of aspect ratio for the Complex walls are considered; first, where the column spacing is 31'-0" and the aspect ratio 0.45 and second where the column spacing is 19'-3" and the correspon-ding aspect ratio is 0.73.
Figures 1-1 through 1-8 show the calculated capacities for the example wall panels corres-ponding to each of the possible behavior modes mentioned above.
ED-1
B Supplemental Information Requested at Meeting of Merch 7, 1980 Answer 1 (a & e) Page 3 of 18 Notations:
2 A
= area of masonry portion of wall section, inches B
A
= area of concrete including cell grout and core, g
2 inches A
=AB+A y
g A
= cross sectional area of a steel column, taken as sc 41.8 square inches for a typical W14x142 column fh
= specified compressive strength of concrete, psi f
= computed compressive strength of composite sec-g tion, psi fh
= compressive strength of block, taken as 1500 pai for both high density (4100 psi) and standard weight block (2700 psi) f
= specified yield strength of reinforcing steel, y
equal to 40000 psi f
= specified yield strength of column steel, equal ys to 36000 psi h
= height of wall panel, taken as 168 inches (14 ft)-
as a typical distance between floor slab and underside of the beam above l
= width of wall panel, inches y
N
= total number of columns crossing the shear plane e
V n staar force resistance cf the frictional component b
l of beam-column connect.on ED-1 l
Supplemental Information Requested at Meeting of March 7, 1980 Answer 1 (a & e)
Page 4 of 18 V,
= shear force resistance developed by shear-friction of reinforcing steel V
=V3+V i
s vf
= nominal shear streus capacity of wall panel, psi t
= thickness of wall u
= coefficient of shear friction
= vertical reinforcing steel ratio py ph
= horizontal reinforcing steel ratio o,
= wall compressive stresc due to normal l'oad, psi Flexural Mode (i)
Double Curvature The procedure for determination of capacity in a wall producing double curvature mode of behavior has been des-cribed in Section 3.4.2.2 of PGE-1020 and the equation developed therein is as follows:
1" (0.93 p f
+ 0.94o )
v
=
f h
vy o
(ii)
Single Curvature The alternative procedure for calculating the horizontal shear capacities of the walls was presented in the response f
ED-1 l
j l
Supplemental Information Requested at Meeting of March 7, 1960 Answer 1 (a & e)
Page 5 of 18 dated July 6, 1979 to NRC Staff Question 43, dated May 18, 1979.
The method described in that response was based on a single curvature model of a wall panel with restraining moments provided only at the bottom of the panel by the vertical reinforcing steel and the effective panel normal load together with those provided by the the vertical shears given by the beam-column connection and the horizontal rein-forcing steel across the two vertical edges.
The equation for the shear capacity is as given below:
1 V1 W
(0.465 p f
+ 0.47 o )+
v
=
f h
vy o
ht Sliding Mode The method for determining sliding resistance at a wall slab interface of the complex walls has been presented in the res-ponse dated June 29, 1979 to NRC Staff Question 11(a), dated May 18, 1979.
The total sliding resistance was developed as i
the summation of the resistance offered by the embedded steel columns, with the shear strength of the steel taken as f /V3, y
and the shear-friction given by the normal force and the ver-tical reinforcing steel crossing the shear plane.
The equation 1
for the sliding resistance is as given below:
ED-1
Supplemental Information Requested at Meeting of March 7, 1980 Answer 1 (a & e)
Page 6 of 18 A
f N -1.5 se s
c v = po + 1.4 A f
+
f o
ay lt N -1 y
c where the coefficient of friction, y,
for the normal force is taken as the weighted average based on the relative areas of bearing composed of mortar bed joints and concrete in cell and core.
In the absence of any code-specified value, the coefficient of friction for the mortar bed joints is taken as 0.75 which is a lower bound value used in Reference 1.
The numerator in the multiplying factor for the column resistance, (N
- 1.5), represents the number of columns available to c
resist sliding in the entire wall section and is equal to the total number of columns crossing the shear plane minus one end column and half the other end column.
The term (N
- 1) 1" c
the denominator represents the number of panels.
Therefore, A
A A
f N -1.5 B
sc v
=(0.75
+ 1.4 9
) o + 1.4 p f
+
4"O s
e A
A o
y y lt N -1 w
w w
c For the example walls, N is taken as 4.
e Diagonal Tension (Shear) Mode
_cor evaluating the lateral load capacity of a wall panel, the ACI deep beam equation is used which gives the shear stress vc as, ED-1
4 Supplemental Information Requested at Meeting of March 7, 1980 Answer 1 (a & e)
Page 7 of 18 M
(3.5 - 2.5
) 2'. /f v
=
c lV V g y
Where M and V are the moment and shear force at the section.
For.a cantilever shear panel, M = Vh (3.5 - 2.5 h) 2 ff v
=
c 1,
V g The compressive strength of the grouted masonry composite wall section is taken as a weighted average of masonry and concrete.
b + /f b
C # '. /f '
=A/f Vg V b A y c A
y y
Accounting for the normal stress, a on the wall section o,
the resulting shear capacity, v *, is e
(v 2)
(2 )
v*
+
=
c c
The horizontal and vertical reinforcing steel provide additional contribution to the diagonal tension strength.
Test results from specimens with height to width ratio of 0.5 (Reference 2) and height to width ratio of 1.17 (though ED-1
Supplemental Information Requested at Meeting of March 7, 1980 Answer 1 (a & e)
Page 8 of 18 titled as 1.0 in Reference 3) indicate that the cracking plane engages both the horizontal and vertical reinforcing and therefore, they may be considered as equally effective in providing the shear resistance.
Thus the final equation for the diagonal tension mode is:
h
=v*+0.5(p +
p f
v f
c v
h l
y y
It is important to note that all the wall sections considered in these examples have an unreinforced ecre.
In the actual Complex walls, especially in the Control Building, most of the wall panels of the shear walls have reinforcing steel in the concrete core and, therefore, their shear capacities for all the predicted modes will be increased accordingly.
From the calculated capacities of the example walls, shown in the attached figures, the following observations can be made:
o The shear' capacity given by the diagonal tension mode is the upper bound shear, for all the wall thicknesses j
and for both the aspect ratios present in the Complex, I
in the range of normal stress considered in the Complex walls.
As can be seen from the figures, the diagonal tension (shear) capacities of the wall panels with aspect ratio of 0.73 fall below 300 psi in some cases.
In actual wall evaluation, the calculated values have been taken ED-1
Supplemental Information Requested at Meeting of March 7, 1980 Answer 1 (a & e)
Page 9 of 18 rath'er than the prescribed limit of 300 psi stated in Section 3.4.2.1 of PGE 1020 when the calculated values are less than 300 psi.
o For the walls with an aspect ratio of 0.45, the shear capacity corresponding to the double curvature
-mode provides the lower bound, at least in the low normal-load range.
However, as the wall thickness increases, the sliding mode becomes more critical.
o For the walls with an aspect ratio of v.73, double curvature mode provides the controlling shear capa-city in the range of normal load generally present in the Complex walla.
The examples developed herein have been provided in order to respond to the NRC staff questions.
The shear walls in the Complex have been evaluated to assure that these capacities, based on the applicable lower bound behavior mode, exceed the load demand.
ED-1
Supplemental Information Requested at Meeting of March 7, 1980 Answer 1 (a & e)
Page 10 of 18
References:
4 1.
Hatzinikolas, M., Longworth, J.,
and Warwaruk, J., " Eval-uation of Tensile Bond and Shear Bond of Masonry by Means of Centrifugal Force", Alberta Masonry Institute, Edmonton, Al be rta.
2.
- Hidalgo, P.A., Mayes, R.L., McNiven, H.D.,
and Clough, R.W.,
" Cyclic Loading Tests of Masonry Single Tiers - Volume 3 -
Height to Width Ratio of 0.5," Report No. UCB/EERC-79/12, College of Engineering, Univ. of California, Berkely, Cal-ifornia, May 1979.
3.
Chen, S.J., Hidalgo, P.A., Mayes, R.L.,
- Clough, R.W.,
and
- McNiven, H.D., " Cyclic Loading Tests of Masonry Single Piers - Volume 2 - Height to Width Ratio of 1,"
Report No.
UCB/EERC - 78/28, College of Engineering, Univ. of California, Berkeley, California, December 1978.
ED-1 l
1 1
o 400-a90"*1 300-I c'
- e4 inc M
\\
@@3 w
200-w c.:
L*
r%
vs s-Y G
C
'ih N*
E
,9 r"o ta E*<r:
M E* l.
100' 20 Inches Thich Fall go (Block Reinforcement Only) 3 =.00183 j
c i
Aspect Ratio = 0.45 1
I 0
50 100 150 I
NORt!AL LOAD o (psi) g l
i rigure 1-1
400 '
c0"*l pa 300-Cd sS'* ole
.O E
~
w 200 -
t.1 s*
e%
e.
m c:
=
s LO 4
s
/
Z 100 -
23 Inches Thick Wall (Block Reinforcement Only) h =. 0159 O
Aspect Ratio = 0.45 i
I i
l l
L 0
50 100 150 NOR?iAL LOAD c (psi) o r
rigure 1-2 i
400-l x
300 -
.O e
4 e
tn3 d
co
- g
./
u 200 65 tr.
D D
u, c:
6 v2 W
E*$c 100 -
32 Inches Thick Wall o
.00115 o
=
(Block Reinforceraent Only)
Oh =.00161 Aspect Ratio = 0.45 0
50 100 150 NOR'mL LOAD o (psi) g Figure 1-3
400-e 300 -
c#q@
E 10 7
200 -
.o g,%*
u) va Mc:
E*
vs h
cr.
b
=
UJ ca Ed6 7;
100 -
48 Inches Thick Wall (Block Reinforce:*ent Only)
O
=.00108 h
Aspect Ratio - 0.45 0
5s 100 150 NO'c i LOAD o (psi) g rigure 1-4
400 --
300 -
etensiO" con l a
c.-
Tqatgr Dia 1* CC
, g# e,ind c.%
C
~
m0 200 --
to E
a:
N d'
c:
4 6
c
=
e tia s
y rc 9
E-6.
E 20 Inch Thick I?all 100-ao (Block Reinforcement Only) ph =.00183 Ascect Ratio = 0.73 O
50 100 150 i
l NOPJ1AL LOAD c (psi) g l
Figure 1-5
400 --
300 '
5 C
D' e
6 en$
200-
- n v2 VJ c.;
c:
E*
o<0 cd 9
so 0
.9 na E*60 100 -
23 Inch Thick Wall o
(Block Reinforcement Only) ph =.00159 Aspect Ratio = 0.73 l
l l
0 50 100 150 NORMAL LOAD o (psi) g Figure 1-6
i i
i 400" 300-4 Dia c
Y.
y,*
-~
b
'y
.c.9
_a sib 200 '
c 02 s..r:
E4e c:
c.*u 5
4 N
C4
.s e
0*
32 Inch Thick ifall f
100~
9 a"
o
.00115
=
y (Block Reinforcement Only)
.00161 Oh=
Ascect Ratio = 0.73 0
50 100 150 NORT@L LOAD o (psi) g Figure 1-7
400 "
300-na a
-piac
~l eM S
s4 200 -
v:
tr.
Y e,
. pt 0
II CY C
i S"
E u:
R V*
4 I.
H 100 -
4*
g a
p 48 Inch Thick !"all e
s 0
0 9
O
=
.00076 y
(Block Reinforcement Only) h =.00108 0
1 i
Aspect Ratio 0.73 l
O 50 100 150 NOPliAL LOAD o (psi) g i-Figure 1-8 l
l
Supplemental Information Requested at Meeting of March 7, 1980 Answer 1(b)
Page 11 of 18 This investigation is an extension of the effects discussed in Attachment 1 to Licensee's submittal dated February 13, 1980 where the possible load shift along the wall and total stiffness of the wall was examined relative to the influence of gross bending.
The OBE loads from the linear elastic solution for 2% damping have been used in this investigation.
If the gross bending effect should develop, the departure from the elastic solution would be gradual as the magnitude of the response increases.
Associated with this departure would be an increase in damping which would tend to reduce the significance of the effect to less than that estimated
-here.
The investigation has been extended to include the lower two stories of R, N, and 41 walls in the Control Building.
The lower stories were selected since the gross bending effect is a maximum at the base of the structure and decreases with distance above the base.
These three walls represent a wide range of dead load conditions and reinforcing steel ratios.
The general approach is to reevaluate the shear force distri-bution and total stiffness of the walls using the stiffness reduction factors in Appendix B of PGE 1020.
In determining the stiffness reduction factors, the influence of gross bending on the vertical forces on the various wall sections
-has been'directly included.
The vertical loads reported in
)
Licensee's response to NRC Staff Question 23 of October 2, 4
1979 are, therefore, increased or decreased depending upon DX-15 i
i t
1
Supplemental Information Requested at Meeting of March 7, 1980 Answer 1(b)
Page 12 of 18 the gross bending effect.
The reinforcing steel ratios used in this investigation also conservatively neglect the influence of the beam-column connections.
The effect of possible reduction in the dead load on the wall sections due to creep and shrinkage are treated separately as reported in response to 1(c).
If gross bending effects were to develop, the results of the analysis for the R wall indicate a shift in the shear load toward the new structural elements for the inertia loads directed north.
This shift was 21% and 49% at elevations 45-61 and 61-77 respectively.
As ir.dicated in response to Question 4, the margin for the new concrete is 50% and 87%,
respectively, for this loading condition.
The capacities in the existing wall sections are reduced due to the reduction in dead load but the shear load due to this effect is also reduced.
The capacity to force ratio for some of the existing wall sections on the tension side of the Complex would be less than 1.4 when comparing the ultimate capacity to the This indicates there may be some redistribution OBE forces.
of forces along the length of the wall, which would be an acceptable condition since behavior of the wall sections is ductile and the displacement of a wall section which may l
experience yielding is controlled by other sections of the same wall which would not experience yielding.
Also, the concept of force redistribution among the wall panels is consistent with the behavior of other concrete structures.
The strain in the reinforcing steel would thus be limited to i
DX-15
1 Supplemental Information Requested at Meeting of March 7, 1980 Answer 1(b)
Page 13 of 18 only 1.5 or 2.0 times the yield strain.
Since the ultimate strain in the reinforcing steel for uniaxial monotonic loading is greater than 15%, which is approximately 100 times the yield strain, the reinforcing steel can sustain the much smaller strain of 1.5 to 2.0 times the yield strain for the postulated 50 peak OBE stress cycles.
The. smallest capacity to force ratio found in any individua-1 panel resulting from this analysis would be 1.18 which occurs in the R wall, second story between column lines 51 and 55.
As described above, this capacity to force ratio indicates the reinforcing steel has yielded at the edge of the wall section and load. redistribution takes place within the wall panels.
This condition is reflected in the increase in the shear in the new structural elements at this elevation.
The reason the redistribution is the greatest at this location is because of two thick wall sections (35 and 41 inches) which have no core reinforcement.
The maximum increase in shear of 19% in the new concrete along the N wall occurred at elevation 45-61.
The increase is much smaller because the cores have reinforcing steel in this wall.
The wall also has an overall capacity to force ratio greater than 1.4.
l Bounding conditions were investigated in order to obtain an assessment of the maximum gross bending effects, if they were to develop.
Results show that locally, capacity (ultimate) to force (OBE) ratios could be reduced below 1.4 which would DX-15 4
Supplemental Information Requested at Meeting of March 7, 1980 Answer 1(b)
Page 14 of 18 imply some amount of additional yielding in the reinforcing steel, and hence some redistribution of loads throughout the shear wall system.
Since the structural performance of the shear wall system is ductile, the structure has the capability for such load redistributions without detrimental effects.
As long as the overall structure capacity to force ratio demonstrates the criteria margin, 'as is the t::ase with consideration of potential gross bending effects, local reductions in margin are not of significance and the desired performance of the building is obtained.
DX-15 l
l
o Supplemental Information Requested at Meeting of March 7, 1980 Answer 1(c)
Page 16 of 18 on the same bases as the stiffness reduction factors, of the existing wall sections has been reduced due to the reduction in the vertical force.
Associated with the reduction in capacity is a reduction in the shear force attracted by the existing wall panels.
This reduction in shear forces in the existing wall section results from the stiffness reduction factors which produce the increase in shear force in the new' concrete elements.
The ultimate capacity to-OBE force ratio for some of the existing wall sections reduce to 1.4.
This results in capacity to OBE force for the total length of the walls being greater than 1.4.
After observing the nature of the redistribution and the conditions in these walls that cause the maximum redistribution, it is apparent that the cases considered offer a broad enough representation of the conditions throughout the plant to conclude that the structure has the capability of accomodating the redistrib-ution and still have an ultimate capacity to OBE force ratio of 1.4.
l l
DX-15
Supplemental Information Requested at Meeting of March 7, 1980 Answer 1(d)
Page 17 of 18 The effect of creep and shrinkage of the existing walls is to reduce the dead load carried by the wall panels and correspon-dingly increase the dead load in the adjacent steel columns.
One of the major contributions of the dead load is to resist the seismic overturning tendency or gross bending.
The dead load is equally as effective in this regard whether it is actually carried by the walls or by the eclumns.
Since the sum total of dead load in the wall panel and the adjacent columns remains constant, this amount of vertical load will continue to resist the vertical tension load generated by the gross bending as predicted by the STARDYNE analysis and will not be influenced by creep and shrinkage.
The vertical forces on the various wall sections predicted by the finite element model provides the demand in the various wall sections but this demand does not have to be resisted only by the reinforcing steel in the panel in conjunction with the vertical load on the panel being reduced by the creep and shrinkage effects.
The dead load that has shifted to the embedded columns is also concurrently effective in resisting the gross bending effect.
In this evaluation, therefore, the effective dead load on the walls for purposes of both the capacity and the stiffness evaluation is the net result of the direct dead load, reduced for the vertical component of the earthquake, and gross bending effects.
This reduca; *he effects of gross bending DX-15 i
l
--Supplemental Information Requested at Meeting of March 7, 1980 Answer 1(d)
Page 18 of 18 and potential dead load reduction to the same conditions as considered in the response to 1(b).
I 2
e DX-15
Supplemental Information Requested at Meeting of March 7, 1980 02 Page 1 of 4 Response Spectra Determine how the above considerations affect the floor spectra.
Answer:
The influence of a possible reduction in the dead load due to creep and shrinkage and the influence of excluding the beam-column connections from the effective reinforcing steel ratios have been estimated by recalculating the stiffness re-duction factors which'in turn were used to predict the change in frequency.
In determining the new effective reinforcing steel ratios, the influence of the beam-column connections were neglected in all walls.
To conservatively include the effect of creep and shrinkage on the dead load in the east-west walls, all dead load was neglected above elevation 77 and below elevation 77 only one-fourth of the dead load used previously was applied.
For the north-south walls, all dead load was neglected above elevation 93 and below elevation 93 one half of the dead load used previously was applied.
This would account for creep, and for a shrinkage strain of 140 j
micro-inches per inch and the encasement effect on the beams.
t Using the corresponding stiffness reduction factors, the decrease in stiffness for the R and N walls was 21.7 and 17.2 percent.
For these two walls, the change in frequency is about 10%.
Only the stiffnesses of the R and N walls changed due to these effects since the stiffness of the other north-l' DY-3 l-i
supplemental Information Requested at Meeting of March 7, 1980 Answer 2 Page 2 of 4 south walls in the finite element model were based on no dead load and did not include the effect of beam-column connec-tions.
This indicates the stiffness reduction for the total Complex is less than that experienced by the R and N walls.
It is estimated that the frequency shift for the Complex is between 7 and 9%.
For curve broadening, 8% has been used.
A similar evaluation in the east-west direction indicates that the effect of these considerations on stiffness change is somewhat less; however, for curve broadening,the same amount i.e.,
8% has been used for this direction also.
The influence of gross bending on the frequency is obtained from the analysis described in response to questions 1(b) and 1(d).
In this analysis, the reduction in stiffness resulting from the reevaluated stiffness reduction factors is obtained by comparing the effective stiffness obtained from the gross bending analysis described in questions 1(b) and 1(d) with that i
used in the finite element model.
A frequency shift of 4.5%.
is indicated.
As indicated in the response to question 1(d),
the loss in wall's dead load due to creep and shrinkage phenomena loses its relevancy with respect to the wall's stiffness, as well as capacity, when simultaneous effects of j
gross bending are also considered.
However, for curve broadening, these effects will be conservatively considered to be additive.
In the event of vertical slip along the embedded columns, the horizontal deflection will increase.
The' frequency shift due DY-3
Supplemental Information Requested at Meeting of March 7, 1980 Answer 2 Page 3 of 4 to this apparent softening of the structure has also been in-cluded.
The finite element model predicts a deflection at the roof of the Control Building due to the north-south component of the earthquake of 0.10 inch.
Including the effect of re-duced dead load and adjusted reinforcing steel ratios (8%
decrease in frequency) and gross bending (4.5% decrease in frequency),' the deflection at the roof would increase to 0.13 inch (0.10/(0.92x0.95)2).
Using this horizontal deflection and the increased deflection due to the vertical slip, the decrease in frequency due to vertical slip is estimated to be 6%.
The broadening of the floor response spectra on the low fre-quency side of the peaks from the above effects is combined with the 16.6% frequency shift due to 50 full stress cycles to account for multiple OBE's.
This gives a total of 31%
(1 - 0.92xO.955x0.94x0.834 = 0.31).
In addition to this, the curve is broadened + 10% to account for other parametric vari-ations described in Appendix B of PGE 1020.
The largest frequency shift occurs in the fundamental mode for north-south motion but the same broadening will be applied to all peaks including those for east-west motion.
This estimate of frequency shift is considered conservative since it was based on a pseudo-static approach and material properties from pseudo-static tests.
In the dynamic response, the actual frequency shift is less than that calculated by DY-3
Supplemental Information Requested at Meeting of March 7, 1980 Answer 2 Page 4 of 4 using static methods because the strain rate effects tend to increase the stiffness and since the frequency is changing with the major load cycle.
Therefore the peak will not be able to develop to the full magnitude considered in this idealized linear elastic model.
In addition, if all these departures from the idealized linear elastic model were to take place, the damping of such a response would be higher than the 2% used in this analysis.
l DY-3
Supplemental Information Requested at Meeting of March 7, 1980 Q.3. Page 1 of'2 vertical shear Along the steel plate, take the frictional component of beam-column connection and the bolt friction only, with-out shear-friction of rebars in the portion of the wall behind the steel plate, and determine the extent to which the vertical shear force can be resisted without slip.
Answer:
The vertical shear force produced by the factored OBE at the wall-column interface at column line R-46 can be resisted without any vertical slip.
The vertical shear capacity of the wall at the embedded column is calculated using shear transfer mechanisms that do not require relative vertical displacements.
As shown in Figure 3-1, the summation of the capacities of these mechanisms exceeds 1.4 OBE forces at all levels.
The mechanisms considered are the beam-column connections, the friction-capacity developed by the posttensioned bolts, and the block and concrete core shear capacities.
The block and concrete core capacities are calculated at the column face.
The capacity of the block is taken as 75 pai, and the capacity of concrete is taken as 2W/fh.
These values are the nominal shear strengths of the materials permitted by the applicable codes.
These capacity values ED-2 wv
Supplemental Information Requested at Meeting of March 7, 1980 Answer 3 Page 2 of 2 are conservative when used in conjunction with the factored OBE load.
Between elevation 117' and elevation 61' the effective block and concrete core area is reduced by the area occupied by the column flange.
This conservatively ignores the bond between the column flange and the concrete core.
Below elevation 61', no reduction is taken because the shear studs added to the column and the new continuous horizontal rein-forcement will not allow the column to separate from the surrounding concrete.
i ED-2
EL. ll7 320
\\
\\
\\s\\
\\
990 7;of(;450 EL.98
\\
\\
N EL.77 1260 1760 2760 N
s
\\
EL.G:
1990 q790s $730
\\
\\
s
\\
EL.45 2880
\\.\\ 449o taso Force (OBE)
-- -- Capa ci ty
-- - -For c e (1.4 OBR)
All forces in kips Fig. 3-1 Vertical Shear Forces (N-S OBE O.159) and Capacities
- at Column R-46.
5
Supplemental Information Requested at Meeting of March 7, 1980 04.
Page 1 of 4 New Concrete Elements Examine the potential increase in load in the new walls re-sulting from reduction in stiffness in existing wal'Is because of reduction in vertical loads due to both dead load loss and gross bending, along with any possible relative displace-ment between the panels.
Answer The new concrete elements were evaluated for the possible increase in the in-plane shear forces due to the potential influence of relative vertical displacement, dead load re-duction and gross bending.
The effect of these phenomena would be to reduce the stiffness of the existing walls and thus the frequency of the Complex.
The apparent softening of the structure will not increase the seismic in-plane shear forces for the Complex shear walls.
In the event of relative vertical displacement at the inter-faces of the wall panels and embedded columns, the relative stiffnesses of individual walls and panels are not predicted to vary substantially.
Since the distribution of lateral loads is proportional to the relative stiffness of the walls and panels, no significant change in load is expected on the new concrete elements due to the relative panel vertical displacements.
DY-6
Supplemental Information Requested at Meeting of March 7, 1980 Answer 4 Page 2 of 4 The influence of the reduction in dead load due to creep, and a shrinkage strain of 140 x 10-6 in/in on the existing wall elements, and gross bending on the Complex would be to in-crease the demand on the new concrete elements.
The capacity to force ratios for these effects in the new concrete elements in the R and N walls are presented in Tables 4-1 and 4-2.
DY-6
Supplemental Information Requested at Meeting of March 7, 1980 Answer 4 Page 3 of 4 Table 4-1 Capacity / Force Ratios Considering Potential Dead Load Reduction Ratio of Ultimate Capacity Wall Elevation to Factored OBE Forces N
45'-65' 1.34 65'-77' 1.37 R
45'-65' 1.48 65'-77' 1.96*
- Includes additional reinforcing steel relative to the design values reported in Licensee's responses on February 13, 1980 to NRC Staff Requests during the week of January 28, 1980.
DY-6 r
J Supplemental Information Requested at Meeting of March 7, 1980 Answer 4 Page 4 of 4 Table 4-2 Capacity / Force Ratios Considering Potential Gross Sending Ratio of Ultimate Capacity Wall Elevation to Factored OBE Forces N
45'-65' l.39 65'-77' l.55 R
45'-65' l.50 65'-77' l.87 DY !
--e m
w
,m
-w 1
r,e-+--
r w
n w
ee
Supplemental Information Requested At Meeting of March 7, 1980 Q. 5 Page 1 of 3 Displacements With consideration of the results from the evaluations in 1 above:
Recalculate inter-structure displacements given in Paragraph 9 of the W. H. White affidavit in the motion for summary dispositon of CFSP 22; review the factor 4.3 given to estimate hypothetical displacements for factored OBE as reported in the July 10 responses to Staff's May 18, 1979 Question 44; and address the effect.of increased displacement, if any, on previously calculated rebar strains.
Answer:
As a result of stiffness reduction due to fifty full st.
cycles to account for multiple OBE's, vertical slip, gross bending, dead load reductions and the exclusion of beam column connections from the effective reinforcing steel ratios, the displacements of the Control Building would increase by a factor of 2.1 relative to these predicted by the STARDYNE analysis.
The displacements of the Control Building for an SSE loading condition, as given in paragraph 9 of the affidavit in the motion for summary disposition, would also increase by a j
a factor of 2.1 and are shown in Table 5-1.
The displacements
{
for the Turbine Building remain unchanged.
The gaps existing between the two structures exceed the sum of their displace-ments, and considerable margin exists for the SSE event.
EE-1 l
I
Supplemental Information Requested At Meeting of March 7, 1980 Answer 5 Page 2 of 3 The July 10, 1979 response to the Staff's question 44 indi-cated a multiplying factor of 4.3 to the linear elsstic STARDYNE predicted displacements based on the stiffness reduction of the most heavily stressed panel to arrive at a hypothetical displacement corresponding to a factored OBE case.
We feel that prediction of the structural displacement for factored OBE load is neither an explicit nor an implicit-criteria of either the FSAR or the present regulatory positions.
Furthermore, additional consideration of all the postulated events causing structural nonlinearity should not be the basis to address the factored OBE loading as regards the displacement.
The maximum rebar strain, considering gross bending, reduc-tions in dead load due to creep and shrinkage strains and the exclusion of beam column connectinns from the effective reinforcing steel ratios, is estimated to be between 1.5 to 2.0 times yield strain.
For an eight inch block height, this strain level may produce a crack width of approximately 20 mils.
In considering the above effects with a three-component SSE earthquake, the maximum reinforcing steel strain is estimated to be about 3 times yield strain.
EE-1
Supplemental Information Requested At Meeting of March 7, 1980 TABLE 5-1 Control Building _and Turbine Building Relative Displacements at Plate, and Gaps Provided Maximum EW Displacements (Inch) SSE.25g El. 93' El. 69' Turbine Building (TB) @ 5% damping 0.9 0.15 Control Building (CB) @ 5% damping 0.1 0.065 ABS Combination TB & CB 1.0 0.215 Gap provided after modifications 2.0 2.5 are made Maximum NS Displacements (Inch) SSE.25g El. 93' Turbine Building (TB) @ 5% damping 1.25 Control Building (CB) @ 5% damping 0.180 ABS Combination TB & CB 1.430 Gap Provided at plate 4.0 EE-1 i
Supplemental Information Requested at Meeting of March 7, 1980 0 6 Page 1 of 4 Determine whether any of the above considerations impact previous evaluations with respect to maintaining seismic capability during exposure of columns as part of the modifica-tion program.
Answer:
In determining the~ seismic capability of the complex while columns are exposed, the vertical shear forces and vertical shear capacities are the important items.
The impact of the items from Question 1 is expected to be minor.
The overall lateral loads on the Complex are unchanged and the actual distribution of the loads among walls will remain approxi-mately the same.
The vertical shear at the corners (41-R and 41-N) results from wall 41 acting as a flange in resisting the north-south component of the earthquake.
Since this behavior will remain unchanged as a result of any of the effects considered in Question 1, the vertical shear at the exposed columns will remain approximately the same as pre-dicted by the elastic finite element model.
The previous calculations of capacity versus demand are adequate but the error band is difficult to estimate.
In order to eliminate concern in this area, a revised construction sequence has been developed which will increase the capacity at the various stages of construction.
This revised sequence has adequate margin to account for the uncertainties.
In developing this construction sequence, the various modes DO-17
Supplemental Information Requested at Meeting of March 7, 1980 Answer 6 Page 2 of 4 of behavior such as double curvature, single curvature, sliding, etc were considered in the evaluation of structural capacity.
The revised construction sequence is as follows:
Construction will begin with the new wall along column line N'
between column lines 41 and 46.
The construction will consist of excavation for and constructing a new footing, and exposing embedded steel columns at column lines 41 and 46; this will allow the horizontal reinforcing steel to be connected along column lines 41 and 46.
The new concrete wall N' will be connected to the existing wall south of column line 46 by drilling holes and grouting reinforcing steel into the existing portion of the wall.
The new wall will also be attached at el. 65 to the existing slab.
In conjunction with the construction of the N' wall, other work between el. 45'-65' may be performed subject to the restrictions indicated in Item 1.
Item 1:
Along column line 41 between el. 45 and 65, restrictions are placed on the exposing of the embedded steel columns.
The embedded columns at column line R and N along 41 may be exposed at the same time, but not at the same time as the column at 41 and O.
When the embedded steel column at 41-0 is to be exposed, the column at 41-R and 41-N DQ-17
Supplemental Information Requested at Meeting of March 7, 1980 Answer 6 Page 3 of 4 must be encased.
If exposing the embedded steel columns at columns 41-R and 41-N from el. 45 to 65 precedes exposing the embedded steel columns at 41-0, the embedded steel columns at 41-0 will not be exposed until the new concrete at either N-41 or R-41 has reached a compressive strength of 2,000 psi.
If.the construction at column 41-0 precedes exposing the embedded steel columns at column line 41-R or 41-N, these embedded steel columns will not be exposed until the concrete at 41-0 has reached a compressive strength of 2,000 psi.
Item 2:
Before the embedded steel columns at el. 65 to 77 along column lines R and N may be exposed, the compression strength of the new concrete at column lines R and N between el. 45 and 65 must be up to 2,000 psi and the new wall at N' must have reached full design strength.
The two columns at el. 65 to 77 along column line R may be exposed at the same time, but they may not be exposed along with the corresponding embedded columns along column line N.
The same limitations hold for exposing the embedded steel columns along column line N.
If the construction between el. 65 and 77 are done along column line R first, then before exposing the columns (the embedded columns) along column line N, the strength of the concrete placed in column line R DO-17
1 Supplemental Information Requested at Meeting of March 7, 1980 Answer 6 Pags 4 of 4 between el. 65 and 77 must be equal to 2,000 psi.
The corresponding restriction on exposing the columns in R would apply if column line N were constructed first between el. 65 and 77.
Item 3:
The construction along colu:.in line 55 at O and at N' between el. 45 and 61 can be done individually at any time but would not be done simultaneot' sly.
)
Item 4:
Before the embedded column 41-R is exposed between elevation 77 and 85, the new concrete below elevation 77 along the R line must have a compressive strength of 2000 psi.
This column section will not be exposed while the columns on the N line are exposed between elevation 65 and 77.
The columns on the N line must be encased in the existing wall or new concrete having a compression strength of 2000 psi before 41-R can be exposed between elevation 77 and 85.
i l
DQ-17 l
Supplemental Information Requested at Meeting of March 7, 1980 0 7 Pege 1 of 1 Address the bearing capacity of the rock foundation to with-stand potential plate drop loads including dynamic ef fects.
Answer:
Licensee's response to the Staff Systems Branch Question 2, dated 8/17/79 stated that the compressive stress tr.ansmitted to the rock would be 18.3 ksf.
The allowable static bearing stress, based on Section 2.5.1.5 of the FSAR, was given as 20 ksf.
The allowable static bearing stress was conservatively established by considering only the weakest rock type un-derlying the foundation, which is the tuf fs having an un-confined compressive strength of 1225 psi (176 ksf).
This is a conservative bearing pressure with respect to the i
foundation conditions.
The P and S wave velocities mea-sured near foundation grade indicate a much stronger rock than one having 1,225 psi compressive strength; a compres-sive strength of 3,000 to 5,000 psi is more realistic.
If a rock compressive strength of 3,000 psi were used, the allowable static bearing stress would be 3,000 x 20 ksf = 49 ksf 1,225 Therefore, if an amplification factor of 2 is used in the calculations contained in the referenced response, the resulting compressive stress in the rock would be 38.6 ksf, which is less than the static bearing stress of 49 ksf.
DV-9
Supplemental Information Requested at Meeting of March 7, 1980 Answer 6 Page 4 of 4 between el. 65 and 77 must be equa' to 2,000 psi.
The corresponding restriction on exposing the columns in R would apply if column line N were constructed first between el. 65 and 77.
Item 3:
The construction along column line 55 at O and at N' between el. 45 and 61 can be done individually at any time but would not be done simultaneously.
Item 4:
Before the embedded column 41-R is exposed between
' elevation 77 and 85, the new concrete below elevation 77 along the R line must have a compressive strength of 2000 psi.
This column section will not be exposed while the columns on the N line are exposed between elevation 65 and 77.
The columns on the N line must be encased in the existing wall or new concrete having a compression strength of 2000 psi before 41-R can be exposed between elevation 77 and 85.
DQ-17
.~ o a
Supplemental Information Requested at Meeting of March 7, 1980 Q.7 Page 1 of 1 Address the bearing capacity of the rock foundation to with-stand potential plate drop loads including dynamic ef fe cts.
Answer:
Licensee's response to the Staff Systems Branch Question 2, dated 8/17/79 stated that the compressive stress transmitted to the rock would be 18.3 ksf.
The allowable static bearing stress, based on Section 2.5.1.5 of the FSAR, was given as 20 ksf.
The allowable static bearing stress was conservatively established by considering only the weakest rock type un-derlying the foundation, which is the tuffs having an un-confined compressive strength of 1225 psi (176 ksf).
This is a conservative bearing pressure with respect to the foundation conditions.
The P and S wave velocities mea-sured near foundation grade indicate a much stronger rock than one having 1,225 psi compressive strength; a compres-sive strength of 3,000 to 5,000 psi is more realistic.
If a rock compressive strength of 3,000 psi were used, the allowable static bearing stress would be 3,000 x 20 ksf = 49 ksf 1,225 Therefore, if an amplification factor of 2 is used in the calculations contained in the referenced response, the resulting compressive stress in the rock would be 38.6 ksf, which is less than the static bearing stress of 49 ksf.
DV-9