ML20107H755

Review of Waterford III Basemat Analysis

ML20107H755
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Site:	Waterford
Issue date:	07/18/1984
From:	BROOKHAVEN NATIONAL LABORATORY
To:	NRC
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4 REVIEW OF WATERFORD III BASEMAT ANALYSIS Structural Analysis Division Department of Nuclear Energy Brookhaven National Laboratory Upton,_NY 11973 9

July 18, 1984 l

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8502270265 840820 PDR FOIA GARDEB4-455 PDR

9 TABLE OF CONTENTS .

4 Page No.

ii ABSTRACT. . . . . . . . . . . . . . . . .- . . . . . . . . . . . .

1 INTRODUCTION. . . . . . . . . . . . . . . . . . . . . . . . . . .

. . . . . . . . . . . . . . . 2

GENERAL COMMENT

S. . . .

. . . . . . . . . 4 STRUCTURAL ANALYSIS TOPICS REVIEWED . .

. . . . . . . . . . . 4

1. Dead Loads. (D). . . . . . . .

2. East West Cracks Internal to the Shield Wall. . . . . . IZ

3. Buoyancy Forces (B) . . . . . . . . . . . . . . . . . . 13 Variable Springs Used For the Foundation Modulus. . . . 14 4.

Vertical Earthquaka Effects . . . . . . . . . . . . . 14 5.

16

6. Side SoiT Pressure. . . . . . . . . . . . . . . . . .

Boundary Constraints. J . ... . . . . . . 17 7.

18 8 Hnite Element Mash and Its Effects . . . . . . . . . .

Average Vertical Shear. . . . . . . . . . . . . . . . 20 9.

10. Punching Shear. . . . . . . . . . . . . . . . . . . ' 20

11. Stresses Resulting From Pouring Adjacent Mat Blocks . . 21

12. Effect of Sidewall Loads on Basemat Capacity. . . . . . 21

13. Shear Margins For Mat Areas Located Between Column 22 Lines (9M to 12A) and (Rt to Q ) . . . . . . . . . . . .

14. Vertical Wall Cracking. . . . . . . . . . . . . . . . 24 25 C0hCLUSIONS AND REC 0leENDAT10NS . . . . . . . . . . . . . . . . .

A-1 APPENDIX A: LIST OF CONTRIBUTORS . . . . . . . . . . . . . . . .

B.1 APPENDIX B: STRESSES INDUCED WHILE POURING BLOCKS. . . . . . . .

C-1 APPENDIX C: EFFECT OF SIDEWALL LOADS ON BASEMAT CAPACITY . . . .

t A85 TRACT' '

The Structural Analysis Division of the Department of Nuclear Energy at ' .

BNL undertook a review and evaluation of the Waterford III basemat.

Based upon a review of the detailed finite element analyses performed by Harstead Engineering Associates on behalf of the applicant, together with approximate analyses developed by BNL, it is concluded that the observed cracks developed on the top surface of the mat during the construction phase and were most probably caused by differential settlement induced i

by the dead loads acting alone or by dead loads, acting on a mit already cracked by normal themal and/or shrinkage effects. For this latter case, the bending induced in the mat by the dead loads would cause these cracks to open and become larger. "In the latter stages of construction, when the backfill was in place and the water table restored to its*

~

natural level, additional loadings caused by the side wall soil and water pressures offset these effects.

! All of the approximate check calculations performed by BNL confinn these conclusions and, together with engineering judgment, lead to the conclu-sion that the safety margins in the design of the mat are adequate. The BNL calculations also indicate that the development of cracks in the basemat due to diagonal tension, either from t,he applied dead loads or from any potential differential settlements between blocks during con-struction, is unlikely. The cracks that have appeared in the vertical walls do not alter thes conclusions and do not appear to present a significant safety issue. Nonetheless, it is reconnended that some b

- ii -

B' s

detailed confimatory calculations be performed, although it is not4 -

anticipated that these analyses will lead to any substantial differences in the results. In addition, it is recommended that a surveillance .

program be initiated to monitor tne cracks.

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E INTRODUCTION e At.the request of the NRC Staff., the Structural Analysis Division of the Department of Nuclear Energy at BNL undertook a review and evaluation ~

of the HEA Waterford III mat analysis d tad in Harstead Engineering i

Associates (HEA) Reports, Mos. 8304-1 and 8304-2 Both reports are i

entitled " Analysis of Cracks and Water Seepage in Foundation Mat."

c Report 8304-1 is dated September 19.1983 while Report 8304-2 is dated l

October 17 ,1983. Major topics addressed in the first report are:

(1) Engineering criteria used in the design, site preparation and construction of the Nuclear Plant Island Structura basemat.

(2) Discussion of cracking and Teakage in the basemat.

]

(3) Laboratory tests on basemat water and leakage samples.

I (4) Stability calculations for the containment structure. ~

T.he second report concentrates on the finita element analysis and its results. Specifically , it describes:

(1) The geometric criteria and finite element idealization.

(2) The magnitude and distribution of the loads.

! (3) The final computer results in tems of moments and shear versus the resistance capacity of the mat structure.

Supplemental information to these reports was obtained at meetings ,

held in Bethesda, MD, on March 21. March 26,, and July 3,1984; at the Waterford plant site in Louisiana on March 27 ,1984; and at Ebasco head-At the close of quarters in New York City on April 4 and July' 2,1984.

the E8ASCO meeting on April 4,1984, a complete listing of the HEA computer run was made available to BNL.

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I.

The BNL affort's were originally concentrated on a review of the -

results presented in HEA Report No. 8304-2 and on the supplemental ;

This infomation contained in the computer run given to us by HEA.

computer run contains the ninet design load cases and their various combi-nations. The input / output printout consists of roughly two thousand Selected portions were reviewed in detail, while pages of information.

the remaining sections were reviewed in lesser detail.

As a result of further discussions with the NRC Staff, the BNL workscope was expanded to include the following additionai topics:

a) East-West cracks internal to the shield wall b) Average vertical shear c) Punching shear Stresses resulting from pouring adjacent mat blocks ,

d) e)

Effect of sidewall loads-on basemat capacity Shear margin for mat areas located between column lines f)

(9M to 12A) and (R to Qt) g) Cracks in the vertical walls of structures placed on the basemat.

Coments regarding this work are given in the sections that follow.

GENERAL COPMENTS Basically, the HEA report concludes that bending moments will l-i produce tension on the bottom surface of the ' mat in the final as-built condition (backfill in place and water table restored to its natural For this c'o5dition, including required seismic loading, it elevation).

Furthermore, the shear margin is shown that the design is conservative.

r t-

,,,n,ms- - - - , ,,n-,m_.---an-w,-,,-,-.,-.---,,._-,-,,wnn-+-------,-m---- ,,-,,~ ~ , , - -

-,w--

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o~wn- m--., --,+.--,-_.,,.vn

(shear capacity vs. the shear produced by load combinations) is concluded -

by HEA to be adequate although a few elements were fou',.6 to be close to the design capacity. Accordingly, the cracking of the top surface is

' attributed by HEA only to " benign" causes such as. shrinkage, differential soil settlement, and temperature changes.

i Based on the discussions held with EBASCO and HEA, and on the review l

of data given to BNL., it is our judgment that the bottom reinforcement as well as the mat shear capacity is adequate. The statement that the cracking of the top surface is attributable to " benign" causes, however, was not analytically demonstrated by HEA. In the BNL review of the reports and data, an attempt was made to ascertain the reas3ns for existing crack patterns that appee5around the outside of the reactor '

shield building as. depicted in Figure 0-1, Appendix D, of HEA Report

' Other effects influencing the structural behavior and safety No. 8304-2.

were also investigated Specifically ,the structural analysis topics

, ~ reviewed in more detail include:

(1) Dead loads and their effects.

(2) East-West cracks internal to the shield wall.

(3) Buoyancy forces and their effects.

Variable springs used for the foundation modulus.

(4)

(5) Vertical earthquake effects.

(6) The side soil pressures.

The boundary constraint conditions used for the mat.

(7)

(8) Finite element mosh size and its effects.

(9) Average vertical shear.

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(10) Punching shear.

1 (11) Stresses resulting from pouring adjacent mat blocks.

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.(12) Effect of sidewali loads on basemet capacity. .

I (13) Shear margins for mat areas located between column lines (9M l

J to.12A) and (R to Qy ).

1 I (14) Cracks in vertical walls of structures placed on the base mat.

STRUCTURAL ANALYSIS TOPICS REVIEWED I 1. Dead Loads (D)

As mentioned EBASCO in-its, diset:ssion,. and HEA in its reports.

In have not shown analytically the cause of the top surface cracks.

reviewing the HEA computer outputs, it was found that element moments

, ~

These indivi-and shears for individual loadings. ara explicitly given.

' dual loadings are factored together at the end of the computation to Since individLal provide the overall combined loading calculations.

loading conditions were explicitly given, BNL extracted certain informa-

' tion from these computer runs to provide an assessment of the contribu-Thus, for tion of the various loading conditions to calculated stresses.

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the case involving dead loads only, a number of elements in the regions L

where the principal surface cracks appear exhibit moments (positive in I

\

l sign) that can produce tension and thus create cracking on the top surface.

and This situation is shown in Table 1 which gives moment data (Mx , My M,y) for elements under various load conditions (dead (D), bou I and normal side pressure) in the regions which generally correspond to the areas where the pri,ncipal surface cracks appear (compare Fig. 1 to Fig. 2).

i

TABLE 1 i

Mx (kip-f t/f t) My(kip-ft/ft) Mxy(kip-ft/ft) Normal Side* Pressure

' ' t - ,

1 Mx Mx Mxy i

0 8 0 8 ELDEht 0 8

-196 93

-$74 19 7 116 - 31 -294 79 4 37 ' -242 11 3 106 - 25 -663 -392 212 655 595 207 91 48 -2 19 -416 - 76 205 -412 217 -296 50

' s. 211 -605 * - 81 15 -319 -193 99 -136 136 66 i 5 207 64 168 172 -170 39 - 12 -347 -489

• 'I 441 -105 531 -130 -274 -258 117

-1193 357 T 436 -719 269

-159 158 - 60 '26 -730 -347 27 j .

438 269 142

- 55 -653 -339 -127 210 88 248 j

~

u "T 447 665 59 87 569 72 -143 28 -361 -420 24

- 49 M 204 193 75 -354 -771 l

350 32 898 - 24 -241

-247 30 i 208 236 39 - 21 -574

-676 260 -995 -171 -486 61 l 203

-705 310 332 - 65 I 426 -542 157 .

l I

l -133 '81 154 - 36 259 62 148 -

75 0 18 5 71 531 i 253 5 41 10

' 255 30 58 670 , .

NOTE! D - Dead Load l 1 $$ i 87 8 86 24 611 ~

i $ 5 252 412

• 41 69 9 Sil 26 44 12 8 - Bouyancy 162 = 23 h' 2$4 5 251 37 - 15 + M causes tension

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- 38 15 - 81 "i 2S7 320 57

- 29 - 6 at the top surface

- 26 16

"' 248 255 29

- 64 28 of the inat, 80 87 118 267 -236 - 82 32 - M causes tension at 59 434 10 j

269 -173 - the bottom surface j

of the mat x = moment along north-

-635 313 - 30 12 south orientation

-314 137

- 29 of mat 4 19 ,

270

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42 -201 102 1  : t! '401 -180

-304 118 -130 17 8 44 - 19 of mat j & .' 414 -200 93 440 41 - 17 - 15 xy = twisting moment i 'i- 4 17 428 - 32 98 - 18 kip - 1000 lbs.

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1A fig. I t.ocation of Elements Listed in Table 1.

D sM t

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l Prior to discussing the results shown in Table 1, it is necessary ,

In particular. -

to consider the strength characteristics of the base mat.

the banding moments causing tensile stresses in the top of the sat are '

of interest. As may be seen from Figure 3 ,all of the north-south reinforcement for the top of the mat consists' of No.11 bars placed six The top east-west reinforcement is also inches on center (#11 9 6").

fil 9 6" except irt the vicinity of the containment'where #11912" are added.

Table Z gives approximate values of the bending moment required for the two reinforcement patterns utilized,in the top of the mat, (a) to l

cause the steeT ta reach an allowable stress of 24 ksi (ACI Code working stress capacity), (b) to crack the concrete. (c) to yield the top f

!, reinforcement, and (d) to reach the ultimate moment capacity of the section.

TA8LE 2

' TOP

, TOP REINFORCEMENT

, REINFORCEMENT BENDING. MOMENT

1. 11 9 6* #11 9 6" + #11 9 12" REQUIRED TO:

820 kips-ft/ft 1230 kips-ft/ft Reach working stress f capacity (in steel) 1640 1640 Crack the concrete 2040 i

Yield top reinforcement 1360 1480 , 2220 Reach ultimate . ,

capacity 4

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Table 2. indicates that be'iding moments of, lass than 1640 kip-ft/ft"

^ _

would not cause cracks to occur in a section of the basemat which had no existing cracks. However,'if some cracks already existed in the met (as '

undoubtedly.would be the case because of temperature and shrinkage ,

effects), one would expect the existing cracks to become more pronounced

' when the bending moment approached the working stress capacity of the top steel (820 kip-ft/ft for the more lightly reinforced section, and 1230 kip-ft/ft for the mere heavily reinforced section).

Returning to the data shown in Tabla 1., it is to be noted that the moments in severai elements exceed the working stress capacity. For -

example, for element 208, the dead load (D) moments M g and M y are Thus, as respectively equal to 350 and 895 kip-ft/ft and are positive.

mentioned previously, the top surface of the met in this area .is in tension. The maximum principal moment is a function of Mx , My, and Mxy and its computed value is close to 1000 kip-ft/ft. This moment exceeds l

the working stress capacity and thus would be expected to expand any Similarly, any existing shrinkage and thermal cracks in this element.

such pre-existing concrete cracks would become more significant under the dead load ccndition in other elements located in the three areas where major cracking in the basemat outside the shield wall was observed, i.e. , elements 447,212 ,204,253, 255,269, 257, 417, and 404 (see Figure 2).

e

\

O Thus, the cracks on the top surface of the mat outside the shield _

wall would be expected to have occurred after construction of the super-structure, but before placement of the backfili. It should be noted that there was, in fact, no period in which the superstructure was fully completed before the backfill was pTaced and before the ground water was allowed to rise and exert a buoyancy effect, i.e., the condition modeled by the " dead Toad" calculations. However, based on information provided -

in discussions with EBASCO and HEA personnel, there was a period before i

dewatering was stopped and beforet the backfill was placed.when a substan-1 tial portion of the superstructure was in place; the shield wall was virtually completed and many of the side walls and internal structures were in place.* Thus, while there is no explicit computation of loading 4

conditions at this point in the conktruction of the facility, the dead i

load portion of the HEA finite element analysis provides a reasongble

! simulation of the actual loading to pennit a conclusion as to the probable cause of the surface cracking.

i In view of the comments made in section 8 of this report, regarding l

the finite element grid size used by HEA and EBASCO in their analysis and the effect of this grid size on the accuracy of the analysis, an This strip approximate analysis of a. strip of the mat was made by BNL.

was taken at the center of the reactor building in the N-S direction l'

1 In this analysis, the [nat was considered to be i with a width of 22 ft.

infinitely stiff (a conservative assumption) and subjected to the dead loads taks.n from the HEA~ computer input. The maximum moment for this case occurs close to the center of the reactor anc results in tension I

l

. ~ .

on the top surface of the mat. idhile this analysis admittedly is-  ;

conservative, it supports the previous conclusion that cracking coeld I have occurred during construction due to tensile bending stresses in the. '

top of the mat. Similar results would be found at the other cracked' l sections shown shaded in Fig. I.

In sunnary,, the cracking on the top surface of the sat is most 4 probably caused by dead loads acting on elements already cracked due to nomal thermal and shrinkage effects. The dead load moments would enhance previously existing small and most likeTy unobservable cracks, .

causing them to become Targer and observable.

t*

2. East-West Cracks Internal to the Shield hil As shown in Fig. 2, crack patterns were also noted in March 1977, internal to the shield wall. At that time, based on information provided 3

by EBASCO, the shield wall was partially constructed up to elevation 187' and the steel containment was supported on temporary footings. Other walls or structures on the mat eitne were not yet constructed or were only partially constructed. Since the computer dead load calculations refer to the mat with all existing structures, which of course has different dead weight loads and different spatial distribution of these f

loads than the partially completed case (for example, the massive l

concrete fill on which the containment rests, plus all equipment in I

containment, would not be included in a model of the 1977 configuration),

it is not possible to utilize the HEA computer results to explain the l

l l

l  ;;

1977 cracks.

It is noted, however, that the additional top reinforce- 4 ments (i.e.. # 11 9 12" as'shown in Fig. 3) are essentially located in areas under the shield wall and are placed in an east-west direction. ~

Thus, if cracking should occur, the preferred direction would be parallel to the direction of the heavier reinforcement, or east-west in orienta-tion.

This is indeed the direction of the cracks, and these cracks,' in all likelihood, had the same origins as the cracks outside the shield wall (dead loads acting in conjunction with themal and shrinkage effects).

4 The additional east-west direction top reinforcements will also cause prevailing cracks in elements located outside the shield wall circle, directly east and west to the shield wall (i.e., those shown I shaded in Fig.1 irt areas R-P-2M-iA and R-P1-12A-9M) to be oriented in an This is indeed the pattern indicated in Fig, 2. In east-west direction.

~

contrast ;since there is no additional top reinforcement in the elements shown shaded in Fig.11ocated bets een sections T2-R-12M-7FH (in the northeast sector of the mat), the prevailing cracks do not necessarily have to be oriented in the east-west direction.

3. Buoyancy Forces (B) 1 The moment results from our analysis (Table 1) show that these forces when acting alone would mostly cause tensile stress on the top l

The moments causing these stresses are tabulated in

' surface of the mat.

Table 1 under the column heading B for groups of elements in the identi-As can be seen,these moments are not fied cracked regions of the mat.

as severe as those due to dead weight.. By superposition they could in 5-

some cases contribute to higher tensile stresses and thus result in: -

further. cracking in some a'reas of the top surface of the mat.

s

~.,

4. Variable Springs Used for the Foundation Modulus Moments and shears developed in the basemat were computed using the concept of the Winkler foundation; namely, the soil is represented as a series of relatively uniform independent springs. The stiffness of the springs is obtained from approximate analyses which are based on generalized analytical solutions available for rigid mats on the surface of elastic soils. The actual design of the mat was based on a series of iterative computer runs in which the soil stiffness was varied until the computed contact pressures under the mat were fairly uniform and equal to the overburden stress at the elevation- of the foundation mat.

This approach appears to be reasonable when assessing the final stress

~

conditions. Long term consolidation effects can be anticipated to cause effective redistribution of loads and cause the mat t'a behave in a flexible manner. However, during the initial loading stages this approach is not applicable since load redistribution is continuously .

taking place.

5. Vertical Earthquake Effects Vertical earthquake effect was not discussed in the HEA reports.

However, from the finite element analysis printout and conversation with HEA engineers, it was stated by HEA and EBASCO that this effect was included in the load cabination cases by specifying an additional factor of 0.067, which was then applied to the dead and equipment load case.

i- * .

i*

From these discussions and our review, it is not clear to BNL whether an -

.' amplification factor due to vertical mat frequency was or was not used.

s in order to obtain a rough estimate of this effect, the north-south direction of the mat was simulated by a beam on four, teen elastic supports.

The total weight of the mat, the superstructure, the equipment, etc., as well as the spring constants ,were the same as those used by Ebasco and HEA in their computer run.

The natural frequencies obtained from this analysis are shown below in TabTe J.

Table 3 Natural Frequencies of Simulated Mat Frequency ,

Moda (Cycles /sec)

Number _

1 .4557E+01 2 .5308E+01 3 .5753E+01 I 4 .5923E+0L 5

.6210E+01 6 .7035E+01 7 .8007E+01 8 .1058E+C2 9

.1295E+02

.1769E+02.

10 11

.2009E+02

.2461E+02 12

.3248E+02 13

.3752E+02 14 ..

As can be seen from Table 3, the frequencies vary from 4.56 to 37.52 cps (i.e. 4557E+Oi to .3752E+02). Using Regulatory Guide 1.60, for the 5% damping case, it is found that amplification factors for these

'l , I s For the first seven frequencies frequencies will vary from 3.0 to 1.0.  :

4

- shown in Table 2, the amp 1'ification factors will be less than 3.0 but s

i i above 2.60.

From the review it seems that the vertical amplification .

It should be realized, I factor used'by HEA was 1.34, which is below 2.60.

however, that not all response parameters (moments, shears, etc.) are f Moreover, the equally aftected by the natural frequencies of the mat.

Hence.,to apply an frequencies were obtained from a simplified model.

overall amplification factor of even 2.5 to all response parameters is I

not reasonable.

The flexibtTity of the sat generally will result in i

some Tocal increases. in the computed seismic moments at some particular locations. Where this increase occurs is hard to ascertain withcut l Since the effects are performing a very detailed dynamic analysis.

localized, we believe that they should not grea,tly influence the gross However, a proper dynamic analysis t

resultant forces acting on the mat.

~

should be.-perfonned to verify the stresses which may be expected to 4,

result.

l i 6. Side Soil Pressure ~

According to the STARDYNE computer results obtained from HEA, the l ,

normal side soil pressures produce large moments that are opposite to those caused by the dead loads, as shown in Table I where moments ef ~

elements located in one of the cracked regions outside of the shield building (TZ-R-12M-7FH) are compared. The total moments in some cases In other regions, there (e.g., elements 447 and 208) become quite small.

is a reversal of the total bending moment, resulting in tension on the l

bottom surface and compression on the top. This compression would tend i

?

c

- L7 -

to close the cracks on the upper surface. Thus, it appears that pide soil pressure is an important load case for the mat design.

1 For the static or normal operating condition, the lateral pressures are based on the at-rest stress condition and are uniform around the periphery of the structure. For'the seismic problems the pressures are computed to approximately account for relative movements between the structure and the soil. On one side, the structure will move away from the lateral soil (active side) and reduce the pressures, while the opposite will occur on the other side (passive side). Tite actual computations by EBASCO made use of actual site soit properties to arrive No at the soil pressures, rather than the standard Rankine analyses.

dynamic effects on either the lateral soil or pore pressures were included. The sensitivity of the calculated responses to these effects* .

are currently unknown. However, approximate estimates of these dynamic effects made by BNL indicate that the total lateral load should change by no more than 15 per cent. Nonetheless, a proper analysis including dynamic effects should be performed to verify this result.

7. Boundary Constraints For equilibrium calculations no special consideration need be made for the vertical case since the soil springs prevent unbounded structural

~

motion. However, the same cannot be said for the horizontal case since Rather, the soil springs are not used to represent the soil reactions.

lateral soil forces are'[iirectly input to the model. To prevent unbounded rigid body motion, artificial lateral . constraints must be imposed on the

,--a .. . _- m L -

d model. The constraints are depicted in Fig. 4. The nodes shown circled were constrained from moveinent in the y (east-west) direction, while those described by "x" were constrained in the x (north-south) direction.

As is cosmonly practiced in finite element applications, the constraints are placed in a manner that they do not overly affect the static and dynamic response caTculations. From the output presented in the EBASCO The stresses caused by and HEA reports, this effect was not evaluated.

the artificial boundaries should be calculated and compared with those presented.

8. Finite Element Mesh and its Effects In generaT. finite element models for plate structures require at least four eTements between supports tar obtain reasonable results on stress computations. The models used by both EBASCO ar.d HEA violate l ~

this " rule of thumb" in the vicinity of the shield wall. The signifi-I cance of this effect is demonstrated in Figure D-3 of Report No. 8304-2 l

which presents a. plot of moment taken through the center of the slab.

The computed moments in adjacent elements 193 , 194 and 455 are -3800, The elements used in the HEA analysis are constant

-2500 and +400K.

curvature elements so that the computed moments will be constant within each element.

The steep moment gradient between the elements indicates that a finer mesh would be adv1 sable. A similar effect was also noted when investigating the elements forming the junction between the lateral earth retaining walls and the base mat. Finally, in order to obtain a betterapproximationoflheshearandbendingmomentswithinanelement (with less oscillation about the true' solution), quadrilateral elements are reconneoded for the mat analysis.

r.. - . . - - ..

ooe< ' m

l. .t .

(. t

(. a. t

![ N, ,\'\N,\,N,\,\N, ,!.!.si

.. , , s .

W o , ,_

_ \ s s X-

. s ,

'\- s' , s u L  ;

, , s .

\s \, \,\, if y h

~ s= , , s s .

, , , s _= y

=

,x s u$ .

V je M%%M  %~%%%%

,

e

%NA ANANANN- \%% iV

• aX . X

/-\,,<N/-Nf- -

- ,. .n%*L A

=

%X \'NX NX NX"

\ANAN \f%%% c i

= Xy, -

1. e%,y, y, N

a X%%'k& -=XX  %  %

>= AAA-a u v 8/OA

=

>dguvu A A

%2%s%%-2%y% >s(X%X 1

8%%M -

e  %%%% %%%%%%  %%%%

• 8  %%%X XX%%%%%%%X; _g.

L -

9. Average Vertical Shear ,

Several elements in the Ebasco/HEA analysis indicate local areis However, shear failure where al10wable shear stresses are exceeded ,

should not be associated with local exceedance of an allowable shear stress, because the loads are distributed across the entire potential failure plane. 'All of the ACI code shear requirements are based on this approach. Average vertical shear stresses (i.e. diagonal tension) were computed by BNL in the base mat for two sections across the mat; one section is in the. E.-W direction and the other in the N-S direction.

These sections were chosen to include those elements which indicated high shear stresses in the HEA analysis and, where actual cracking was noted.

The highest average shear stress computed for any design load combination is 50 psi. The allowable shear stress for this case, in accordance with ,

Chapter 11. of ACI Standard 318-77, is 107 psi (M(. Thus, a ,

safety factor of greater than two '.is available to prevent shear failure under the ' design load combination.

10. Punching Shear Another potential failure surface in the base mat considered by BNL is a punching shear section located a distance of d/2 outside the reactor The shield wall, as recossiended in Chapter 11 of ACI Standard 318-77.

peak value of shear stress due to both SSE overturning moments and normal operating loads (plus proper load factors) wen close to but always less than the allowable design shear (%).

11. Stresses Resulting From pouring Adjacent Mat Blocks  :

BNL has explored the question of whether diagonal tension cracks may To .

have occurred during the process of pouring adjacent mat blocks.

determine if such cracking could have occurred an approximate analysis

w. 5 made, as set forth in Appendix L The adjacent blocks are assumed to The rest on foundation springs which' represent the soil flexibility.

second block poured was assumed to harden instantaneously thereby over-estimating the shear Toad carried by the first block. due to relative settlement of the two blocks. As indicated in Appendix B, the resulting stresses were found to be sufficiently small so that neither diagonal tension stress, nor bending tensile stress alone, would be expected to cause cracking Moreover, the likelihood of moment cracking was signifi-K These conclusions cantly greater than the Tikelihood' for shear cracking.

are valid evert for the case with soft spots in the foundation soiles, i.e.,

i where the satT modulus under one block is one-half that of the soil modulus under the adjacent block It should be noted that, according to EBASCO, soil settlement at the The site was found to be instantaneous based on actual measured data.

concrete has almost no strength for the first eight to twelve hours and therefore even the small str isses calculated in Appendix B are unlikely.

l

12. Effect of Sidewall Loads on Basemat Capacity Under normal operating conditions the loads acting on the side walls produce an average' compressive stress in the base mat of about 50 psi. When seismic loads are included in this computation, the average 5~

..c._,_.ww-w--- r-wmvw w-v 'w-ywew-p-mw-u-'*--wwww w we re e sw--*-se'w- *iW- ww---'

r . - .

compressive stress in the base mat is reduced, but is still about 38 psi. ,

These compressive stresses, provide additional shear strength which have not been included in evaluating the capacity of the mat to carry diagonal tension stresses. It should be noted, as indicated in section 9 of this report, that the highest average shear stress developed. in the base mat is only 50 psi. If this shear stress is combined with the 38 psi average compressive stress one finds that the diagonal tensile stress in the concrete is reduced to 34 psi. It is unlikely that this shear stress could cause a shear (diagonaT tension) failure given the 107 psi shear capacity. This analysis is presented in Appendix C.

13. Shear Margins For Mat Areas Located Between '

Column Lines 9M to 12A

, and K to Q7.

In response- to a. request by the NRC Staff, EBASCO provided an estimate of peak diagonal tensile stress irt a region bounded by column lines 9M-12M-R-Q1. (Figure 1). EBASCO indicated that the average diagonal tensile stress, for the SSE case, in elements 410 , 413, 414 and 419 was A meeting was 210 kips /ft, as compared to a capacity of 274 kips /ft.

The following held at EBASCO on July 2,1984 to review these data.

conclusions were reached at that meeting:

i  :

(a) The EBASCO estimates of 210 kips /ft was overly conservative, for the following reasons. Two shears are associated with each of the elements. The first (F g) acts on a plane lying in the north-south direction, and the seco,nd (Fp)actsonaplanelyingintheeast-west

' direction'. The computer output gives these results in local element

_ _ _ _ . .wm_.y.--- . . , _ _ _ . _ . _ - . _ - , , - - - ~ . . - .

i. . .

- -n.

coordinates rather than in global coordinates. EBASCO used the mitximum ,

, although these maximum values were not acting value of either Fy or '

How-in the same plane, for each element to obtain the average shear. .

ever, if the appropriate values of F, and tyr (f.e., those acting in the same plane) are combined, it was found that the average value of F,, is is 106 kips /ft. As may be seen ,

132 kips /ft and the average value of Fg these values are considerably less than 210 kips /ft, and are less than one-half of the shear capacity of 274 kips /ft.

(b) The above values represent the shear at a point. If an average shear is calculated along an east-west line running between column lines R and Q from the containment to the exterior of the mat using the r,2A and EBASCO (ESI) computer runs, the following results are found: ~

HEA ESI 103 k/ft 127k/ft DBE (with Toad combinations) 66 k/ft 65 k/ft Normal (with load combinations) l l

i Once again it may be seen that the shear stresses are much less than In addition, based on one-half of the shear capacity of 274 kips /ft.

the discussion in section 11 of this report, it can be estimated that the maximum additional. shears that can be developed from differential settlement of the base met, even when postulating a gross difference (2:1) in soil stiffness under adjacent blocks, are calculated to be on ,

the order of about 16 kips /ft. Thus, the developed shear stresses will still be small as compared to the shear capacity of the mat.

l

• It should be noted, as discussed in section 8 of this report', that ,

~

large triangular finite e1ements were used.by the applicant to model the mat and its associated structures. The use of these elements produces .

sharp variations in computed moments and shears from element to element.

Because of these variations, BNL's evaluation looked at average values for these forces, derived from several sets of adjacent elements, in order to arrive at representative values.

14. Vertical Wall Cracking The shield. wall is. very stiff es compared to the basemat because of its wall thickness and circular geometry. It is therefore unlikely that the differential settTement of the basemat could have developed the cracks in the shield walt. In our opicion, these cracks must have been caused by thermal and shrinkage effects which occurred after the concrete placement.-

Cracks have also been observed in other vertical walls such as those at the cooling tower. These walls are not as stiff as the shield wall since they are plane. Therefore, it would be possible for these cracks to have been caused by the differential movements of the basemat in addition to themal and shrinkage effects. It is our opinion that the cracks in these walls occurred during construction when the basemat was Now that the long subjected to its largest differential settlements.

term settlements have stabilized, these cracks are not expected to grow.

A more refined analysis, considering the actual configuration of ~

the plant during various stages of construction, would provide the ' i 4

l quantitative basis for determining the origin of the vertical wall cracks. However, it is conclude'd that these cracks do not appear to raise a significant safety issue.

CONCLUSIONS AND RECOMENDATIONS (a) ~The Waterford plant is, primarily a box-like concrete structure supported on a 12-foot thick continuous concrete mat which houses all Class 1 structures. The plant island is supported by rela-tively soft overconsolidated soils. To minimize long term settle-ment effects, the foundation mat was designed on the floating foundation principle. The ave' rage contact pressure developed by i

the weight of the structure is made approximately equal to the existing intergranular stresses developed by the weight of the soil E

Thus, 4

overburden at the level of the bottom of the foundation mat.

net changes in soil stresses due to construction and corresponding l

settlements can be anticipated to be relatively small.

(b) In reviewing the information , reports,and computer outputs

\

! supplied to BNL by EBASCO, HEA, and LP&L, it is concluded that l

normal engineering practice and procedures for the analysis of nuclear power plant structures were ehloyed.

o (c) Accepting the information supplied to BNL pertaining to loadjngs, -

geometries of,the structures, material properties and finite element mesh data, it is our judgment that: I.

(i) the bottom reinforcement as well as the shear capacity of the base mat are adequate for the loads considered.

(ii) the computed dead weight output data can be used to explain the pattern of cracking that has appeared on the top surface of the mat. The cracks that appear probably occurred after construction of such of the superstructure but before place-ment of alt of the backft1T and restoration of the ground water ta its natural levek'. Growth of the cracks would then have been constrained by subsequent backfili soil and wa,ter pressures. -

(iii) The cracks that have ppeared in the vertical walls of structures placed on the base mat do not affect the conclusions regarding- the strength of the base mat and do not appear to present a significant safety issue. ,

(d) It is recommended that a surveillance program be instituted to monitor the cracks, water leakage and chNiical content of the water on a regular basis.

. y. 1 (e) BNL has reviewed the information and analyses provided by EBASCO,  ;

)

HEA, and LP&L'. Those analyses could be refined in the following areas:

(t) dynamic coupling between the reactor building and the base mat for seismic stresses resulting from the vertical earthquake input (see section 5);,

(ii) dynamic effects of laterai soiT/ water loadings (see section 6);,

(iii) artificial boundary constraints in finite elements models (see section 7);

(iv) fineness of base mat element mesh (see section 8);

(v) origin of cracks in the vertical walls (see section 14). ,

Based upon our approximate calculations together witir engineering Judgment, we do not anticipate that=the refinement of these analyses -

will lead to major changes in calculated stress levels; nonetheless.

it is reconnended that the detailed confirmatory calculations mentioned above be p'erformed. For all of these reasons it is our conclusion ' hat the safety margins in the design of the base mat are adequate.

O O

e

APPENDIX A .

i LIST OF CONTRIBUTORS Listed below in alphabeticaT order are the names of the contribu-l tors to this report: l Costantino, C. J.

Miller, C.. A. .

Philippacopoulos, A. J.

Reich ,M.

Sharma ,S i

' Wang,.P. C.  ;

1 o

I e

O O

O y-

' - - - - - - ~ - - _ . _ _ _ _ _ _ _ _ _ _ . . _ _ . _ _

Appendix B ,

Stresses Induced While Pouring Blocks

[

e 9

T e

• OO 9

• e E?

l 1

l l

A question has been raised concerning the stresses which coul.d have ~

l Therespionse l been introduced when the basemat blocks were being poured.

of two adjacent blocks during construction are considered.

The first It is block is'taYen to be in place when the second block is placed.

l also assumed that the concrete in the second block hardens immediately so that it can transmit loads to the first block. The subgrade modulus under the two blocks is assumed to be different so that the effect of A sketch of the problem to be

• soft spots in the soiT can be considered.

considered is shown in Fig. BL a

/

i 9

K K 1, 2-Fig.B1 Construction of Two Adjacent Blocks l

When the first block is poured it settles an amount, O =y W/K 1 .

The second block is then poured. If the concrete is conservatively assumed to harden before the soil settlement can occur, the second block l

B-2 will introduce additional loadings on the first block. The new deforma-  :

tion caused by the weight of the second block is shown on Fig. B2.

s *- i m s m a nn. ora ~~ .-

2 DEFORr1ED SHAPE 5_ m Fig.B2 Deformed Shape of Blocks .

The loads acting on the blocksy then be determined by natitiplying the deformations by the foundation moduli. These loads a're shown an Fig. B3. .

l W

l .. . .

3~.-- - ,_

K (A+ E L)

. 2 2 K4 K (a +6 - 5 L) '

22 1 1 2 Kfo+Ai,

- 11 2 EigB3 Loads A,cting on Blocks l

i

. . . , - - - , - - . - - ,e

. . ~ .

B-3 I

Force and moment equilibrium allow the two unknown displacements '(p2' -

to be calculated. The results are.

O2 * " EI

• IIll + 1# M )3/K1 .

/ = 12. W/LL Ky(1+14A+A2)]

where,A = Kg/Ky Once the displacements are known the loads on the blocks may be evaluated and beam shears and bending moments may be computed. This is Peak values of shear done for foundatiort moduli ratios of 1 and. 0.5.

and moment are tabuTated in Table St.

5 fable El Shear and Moments in Blocks During Construction Maximum Required f'c (psi)

Foundation to Prevent Moduli Ratio Shear Moment (Kips /ft) (Kip-ft/ft) Shear Bending Tension

( )

15 15 1 11 101 275 31 113 0.5 16 For the design concrete strength of 4000 psi, the shear capacity of the concrete section is 274 kips /ft. As may be seen this is much larger than Bending the peak shears that could be caused during construction.

l cracks will occur in the. concrete when the peak concrete tensile stress I

[

B-4 reaches the modulus of rupture. For the concrete design strength

• this _

will occur at a bending moment of 1640 kip-ft/ft. It may be seen that l

l the peak moments are closer to the value required to cause a bending .

crack than the peak shears are to the value required to cause a diagonal tension crack. ,

The concrete will not have attained its finai strength at the time when these stresses occur. The last two columns in Table B1 list the required concrete compressive strength to prevent shear and moment failures. Two conclusions may be drawn from these data. First, even for ra.ther dramatic variations in foundation moduli, only a minimal concrete strength is required. to present either a shear or moment crack. Second, if a crack were to develop it would most likely be a bending crack.

The above analysis is based on the assumption that the concrets hardens before soil settlement occurs. If this were not so,the wet i The concrete would fill the void volume created by soil settlement.

concrete block would then be supported on the soil rather than " hanging" from the.other block. Figure B4 shows the concrete strength gain during the first day. As may be seen concrete will have no strength until about 8 hours9.259259e-5 days <br />0.00222 hours <br />1.322751e-5 weeks <br />3.044e-6 months <br />. By this time all of the soil settlement would have occurred and the second concrete block would not induce any loads on the first block. ,

I -

?

B-5 Fig. 84 From: " Concrete" by S. Mindess,J. T. Young, _

Prentice Hall 7 s

Chen a cemens 4s 80 .

- sooo

- 7000 so -

VME

) -.

/ e

? 40

/,/ mb

~

n -

h a 30 - .

• • • ** / ,/ -

- }s. .

/

\ - /

/

/

/

/

"fp 3

k 8 20

~

j

~

,/ ,/ O

/ / - 200o

/ /

, Tyne / / Tyne

- to -

m/ / I ~ ' ,

/ ./

l.

t /t >

A/ t i g 0 t 7 days 28 days 3h ah I day j 15 min Th

• Twneof curing

- el concretes made wkh rap id .herd -

. Figurer 3.13. Soongth L*,

• • . Consmsemen. ening cemenes. (Adapted frora W. Perenchio, in Nene Meteriale in Cescreer ed. S. P'. Shah. (Jneversley of tainess at Chicago Orcle. .

.Chicage.1972. p.12-VI.)

placement and have the advantage of better water resistance. But the very rapid strength gain of the cement suggests many.other applications in which the properties of'a portland cement are desired: pavement and bridge-deck repair, precasting operations. shotereteing, and slip forming.

It is unfortunate that regulated-set cement is not currently available in the U.S., but the interesting properties of the, cement will no doubt ensureits, reappearance.

VME Cement la the production of*VHE cement, calcium sulfate is added to the raw 5 is formed in the rotary kiln. This is the same mix so that C.A3 compound that is present in Type K . expansive cemen l ._ . . - - - . . _ _ _ _ _ . _ _ . _ _ _ . . . . _ . . . . . _ . . . _ . . , , . _ . . - . _ . . . , . _ _ . _ . _ . _ . _ _ _ . _ _ _ . . - , _ . . - _ . . _ , , - - . . . . - . _ _ _ . . , ~ .

l l

Appendix C Effect of Sidewail I. cads On Basemat Capacity m

,2 C Q

e O

O e

• 9 f

{

Soil pressure loads act on the sideweils and these loads intr'oduce . ,

compressive stresses in the slab of he basemat. This compressive stress will assist in resisting the diagonal tension stresses which _

occur in'tHE sT&b. The significance of this effect is discussed in this Appendix.

Table C1 lists the horizontal loads which act on the sidewalls due to the various load combinations. These Toads were determined directly from the HEA/Ebasco computer printouts Table C1.

Total Force Acting on'the Wall Surface (kips)

Wall #1 #Z #3 ' #4 Load Case 36619 36441 50942 50522 Case 4: Normal Soil Pressure 27061 110657 50684 50377 Case 8: SSE & Soil (North,to South) 26907 50684 50377 Case 10: SSE & Soil (South to North) 111051

/ ,i.. &l,1e'

.. Jn -

/ - -

(. <'- -

/l4 _-

1 .-

t

7. < sm

- endl *4- <;

l

, -u..s- -i

~"'

.. -------------- --- ---- -----.-- ---------------------------- -------------------- a

C-2 I

An elevation of the structure parallel to the long direction of the -

basemat is shown on Fig. C1.

t n .

n '

kec 54

?'. P P d$ (

T8~ / 5!Ed!6BG!#G!#G!s#G!####G!####ss, ,

k ,,

,r .

4 Fig.CI Estimated Side Loads Orr Wall _

The forces (P) are taken as the forces shown on Table CI and acting on The soil pressure'is assumed to have a triangular walls #2 and #4.

variation as shown so that the resultant force (P) acts at the thi,rd point on the wall. Since the wali- is buried about 54', the resultant force acts at a point 18' up the wall from the bottom of the basemat.

8 The stresses caused by this loading in the cross section are shown on Fig. C2.

d P3*

. ###/#/####/ #

275" 5IFigC2 Cross Section of Basemat

4

. . , i o . * *-

• C-3 The basemat is analyzed as a beam structure. The cross section shown c in -

Fig. C2 has the following properties:

~

Cross sectional area = 355Z square feet Centroid at 7.91.' above the bottom of the mat Moment of Inertia = 247300 feet Stresses are then computed as: <

f = P/A Mr/I Therefore at the top of the wall,,

f = P/355Z + P (18-7.91) (54-7.91) / 247300 The stress at the top of the slab is, f = P.3552. - P (18-7.91) (12-7.91) / 247300 ts The stress at the bottom of the slab it, f

bs = P.3552 - P (18-7.91) (7.91) / 247300 i

I

! The resultant stresses for the Case 4 loads (Nonnal soil pressure) are:

i fg = 541 psi f = 112 psi ,,

ts f = -11 psi bs l

t

1. - , - - - - - , - . _ , - , - - _,__,___y.,_ ._ _

,y

C :.

> The stresses for Case #8 (SSE in N-5) are: ,

z

' f = 465'ps! ,

tw f = 84 psi i.

3 f = -8 psi bs t-The average stresses in the slab for these two load cases are 51 psi and l

The average shear in the basemat for the vertical 38 psi respectively.

If this shear shear Toadings (see section 9) was found to be 50 psi. '

stress is. combined with the 38 psf average compressive stress one finds that the diagonal tensile stress in the concrete is reduced to 34 psf.

It is unlikely that this. stress could cause a shear (diagonal tension) failure given the 107 psi shear capacity.

l e

i O

i

- , - ,..,..-,,,-,.,,.n.- -

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