ML20136F694

Draft Evaluation of Concrete Creep & Stress Redistribution in Fort St Vrain Pcrv Following Rupture of Prestressing Tendons

ML20136F694
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Site:	Fort Saint Vrain
Issue date:	10/31/1984
From:	Fugelso E LOS ALAMOS NATIONAL LABORATORY
To:	John Miller Office of Nuclear Reactor Regulation
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Evaluation of Concrete Creep and Stress Redistribution in the Fort St. Vrain PCRV Following Rupture of Prestressing Tendons WRC Fin No. A-7258 October 31, 1984 Los Alamos National Laboratory University of California Los Alamos New Mexico 87545 E. Fugelso, Q-13 C. A. Anderson, Q-13 Responsible NR Individual and Division J. Miller / ORB 3 Prepared for the U.S. Nuclear Regulatory Comission Washington, D. C.

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I DISCLAIMER This report was prepared as an account of work sponsored by an agency of the United States Government. Neither the United States Government nor any agency thereof, or any of their employees, makes any warranty, expressed or implied, or assumes any legal liability or responsibility for any third party's use, of any information, apparatus, product or process disclosed in this report or represents that its use by such third party would not infringe privately owned rights.

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ABSTRACT

,4 During a routine lift-off force test of the prestressing tendons in the prestressed concrete reactor vessel (PCRV) in the Fort St. Vrain, Colorado reactor, broken strands in several tendons were observed. These pretensioned tendons apply suffi-cient compression in the concrete to balance or exceed the cir-cumferential and vertical tension in the concrete that results.

from the internal pressure. A combined analytical and numerical study was undertaken to evaluate the evolution of these stresses, both to the initial prestressing and to subsequent partial and total rupture of these tendons. At the stress levels anticipated in the concrete, and for the anticipated operating life span of the PCRV, the concrete behavior can be modeled as a linear visco-elastic solid with the creep strain varying proportionally with the logarithm of time at constant stress throughout the projected reactor lifetime.

l A one-dimensional model of a long concrete column of rec-I I

tangular cross-section with an embedded prestressing tendon along the length was used to evaluate the concrete and steel stresses as well as the hold-down and lift-off forces. These were e talu-ated for the intact tendons and the degraded tendons. The degree of tendon degradation is described through the ratio of the i

number of unbroken strands to the original number of strands.

l Initial time of rupture was varied from the time of initial pre-stressing to 400 days after emplacement. The formulation led to an integral equation, which was solved numerically. The hold-down forces decayed approximately with the logarithm of time and for both the extreme observed degradation (21 broken strands) i and for a more extren case (40 broken strands) the hold-down force still exceeded the minimum safety design requirements.

In addition, several finite element calculations, using the finite element code NONSAP-C, were made to evaluate complete l

tendon failure in a 600 sector of the Fort St. Vrain PCRV.

This code has an extensive material library of constitutive rela-l

< tions to model the various properties of concrete, together with l

a ' specialized element model to simulate prestressing tendons.

l Two rows of vertical and are row of circumferential tendons were incorporated in the model as a baseline calculation, the tendons were prestressed to 704 of the ultimate and an internal pressure of 775 psi was applied (this pressure is the internal pressure of the helium coolant in the '4TGR) and the creep of the concrete and slow decay of the tendon stresses were evaluated out to 30,000 days. Then, three cases wherein one tendon was removed at one day were evaluated. First the middle vertical tendon' in the outer row and in line with the outer buttress was removed.

Second, an inner vertical tendon opposite the thinnest portion of the PCRV wall was removed. Finally, an inner layer circum-ferential tendon at midheight was removed. Stress redistribu-tions at 300 days after ruptures were calculated and shifts of the remaining tendon loads to accommodate the broken tendon were calculated. Regions of local tensile and shear stress in the j

concrete portion of the PCRV were identified and related to overall structural integrity..

With all tendons present, the mean vertical stress was about

-760 psi, the radial stress decreased from the applied internal pressure of -705 to about -1200 psi at the ring of circumferential tendons and the tangential stress ranged from -2400 psi at the inner wall to about -2200 psi at the same place.

Removal of a vertical tendon reduced the mean axial stress by about +40 psi, the local tangential stress by -10 psi and did not materf ally affect the radial stress. Removal of a circumferential tendon reduced the mean tangential stress by +30 psi and the local axial stress by -80 psi. The vertical hold-down force from zero days through 30,000 days decreased linearly and remained above the prescribed safety limit, as did the circumferential hold-down force.

Comparison of the analytical solution and a small finitt element problem simulating the analytical problem was made to verify the viscoelastic creep models and the tendon element in the NONSAP-C code. Excellent agreement for stresses, strains and hold-down forces was obtained.

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FOREWORD This technical evaluation report is part of the technical assistance program, " Review of Selected Fort St. Vrain Issues", FIN No. A-7258, and is supplied to the U.S. Nuclear Regulatory Commission, Office of Huclear Reactor Regulation, by Los Alamos Itational Laboratory.

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Chapter I I.

Introduction 1.1 Statement of the Problem During a routine inspection of the lift-off forces of the prestressing tendons in the Fort. St. Vrain reactor's prestressed concrete reactor vessel (PCRV),

several of the tendons were observed to be partially degraded in that up to twenty one of the individual strands (of 169 total strands per tendon) had failed. Since these prestressing tendons carry a large tension, thereby placing the surrounding concrete in compression, the evaluation of the concrete stresses and the subsequent adjustment of these stresses to the degraded tendons, needs to be accomplished to asses:: the continued structural integrity and functional capability of the PCRV.

Group Q-13 of the Los Alamos National Laboratory therefore undertook a systematic investigation of the evaluation of the concrete stresses in response

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to changing tendon configuration. The initial level of the stress in the concrete' is such that the major feature of the concrete deformation is slow creep in compression.

The approach to the desired engineering evaluation of the structural integrity l

is twofold. First, an analytical model of the tendon and concrete stresses with prestressing of the tendon, subsequent elastic behavior of the tendon and visco-elassic creep of the concrete is developed. The concrete creep behavior is genera 112ed from the long time creep tests at constant stress on samples of the

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concrete,in place. Tendon degradation is moleled by reduction of the effective cross-sectio'n area of the tendon proportiona'l to the Naction of broken strands.

i Time evaluation of the hold-down and lift-of f force

..e then evaluated through solution of the resulting integral equation. The second' step is the evaluation of the stresses in a sector of the PCRV wi)h a number of initially prestressed tendons.

The concrete stresses with all tendons intact is then compared with the stresses with selected tendons degraded, missing or removed at some time after initial pre-stressing. Solution of this problem is effected through the finite element code NONSAP-C, utilizing, in particular, the viscoelastic constitutive model for concrete creep and the tendon element model in the code. Connection between the analytical formulation and the finite element calculations is provided by finite element solu-tion of the analytical model.

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With all tendons present, the mean vertical stress was about -760 psi, the radial stress decreased from the applied internal pressure of -705 to about -1200 psi at the ring of circumferential tendons and the tangential stress ranged from

-2400 psi at the inner wall to about -2200 psi at the same place. Removal of a vertical tendon reduced the mean axial stress by about +40 psi, the local tangential stress by -10. psi and did not materially affect the radial stress! Removal of a circumferential tendon reduced the mean tangential stress by +30 psi and the local axial stress by -10 psi. The vertical hold-down force from zero days through 30,000 days decreased linearly and remained above the prescribed safety limit, as did the circumferential hold-down force.

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Chapter II I.

Introduction The time evolution of the stresses in the prestressing tendons and surrounding concrete was accomplished for a simplified, one-dimensional model af a typical Fort St. Vrain PCRV tendon system. This study was done to investigate the decay of an intact tendon system and a degraded system wherein a small portion of the tendon strands had been broken at some time during the structure's history. The concrete is modeled as a viscoelastic mediurn defined by creep tests at constant stress.

Decay of the tendon stress, concrete stress, hold-down force and lif t-off force were generated for a temperature range representative of the FSV operating environ-ment and for a number of degraded tendon conditions encompassing those observed during routine inspections.

Figure 1 shows the schematic diagram of this problem. A steel tendon, of total cross-section area A, and initial length, L,, is essentially stretched to an

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initial tension, e. This tendon is attached to base plates, which compress the g

concrete surrounding the tendon. The initial cross-section area of the concrete column is A and its initial length before compression is L. The initial com-c g

pressive stress in the concrete is o. We consider only the untaxial stress g

component and uniaxial defor: nation in the tendon and concrete column.

Stress equilibrium requires that l

A,c, + A ogg=0 through the history of the defomation.

Equality of the lengths of the stretched steel tendon and the compressed l

< concrete column requires L, fl + s e (1 + c )

I=L g

c where e, and c are the strains in the stee?

• erd o and concrete column, c

respectively.

The steel is assumed to be linearly elastic throughout o,. E,c, i -_

'. '. while the concrete is assumed to be viscoelastic.

Its properties are defined through a creep test at constant stress. Appendix E of the updated FSAR (Ref. 2) yields the following form for the creep data for a time interval from t = 0 days to t = 300 days.

= [,

1 + e in (t + 1) e g C

Where e is a constant dependent only on temperature.

We assume this form of the creep equation will be valid for the range 0 1t i 11000 days,(- 30' years). The values for a were determined for the data presented in the FSAR, (ignoring the initial 2-8 day rapid variation), and, using 6

E, = 5 x 10 psi, we found a = 0.12 for T = 75*F and a = 0.40 for T = 150*F.

Generalizing the strain data at constant stress to that of a variable stress-time history through the Duhamel integral, we have:

g I (t - C) og (C)dC

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1 e (t) = p j

g c j o

where j(t) = 6(t) +

gg with a(t) denoting the Dirac delta functional.

Eliminating o, o and c, from the four equations, the following Volterra g

g integral equation of the second kind is found for the concrete strain e (C)dt g

Acg (t)

+B

+ [C + D in (t + 1)] = o

~~

t t+1 h

trith A,E,L ~

c A=

• I AEL cCs 1 l l

c4,E,L c

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A E L, gg

^s s c

C=

-1

, and D = oC The steel stress can be written in terms of the concrete strain as e, = E, (1+c)-I c

the hold-down force, F, is given by F = A, e, The area of the steel tendon, A, is the product of the number of strands-in the tendon, N, times the ares, 6, of each individual strand.

3 A, = N a,

.Kach strand is 1/4" in diawter and, initially, there are 169 strands per tendon.

The steel, with all strands in place, is stretched to 0.7 of its untaxial yield strength,' which is taken as 240,000 psi. The concrete is compressed to an initial compression, at t = 0, of -750 psi.

Thus A/A,=224 g

and L,fLg = 0.99428202 5-

The integral equation tas solvsd numerically for N 169 and the temperature of 75'F and 150*F. The hold-down forces for these two cases are shown in Figure 2.

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We also solved these equations for the case where the initial pre-tensioning is applied, and then M individual strands break at t = 0,, just after the initial load is applied and equilibrated. These are also shown in Figure 2, for M = 10 and M = 20. Also one case was evaluated when M = 0 for 0 1 t g 365 diiys, at which time 20 strands'(aM = 20) were allowed to break. This was done for T 150*F and this curve is also shown on Figure 2.

From these solutions, it is seen that the stress decays almost linearly with log time for a > 0 where the load is applied at t = 0 and decreases faster with increasing e.

For a = 0, the elastic solution is ge'nerated and there is no stress relaxation. When the condition M > 0 is applied at t = 0, the stress drops to a lower value and then the subsequent time behavior is as above. When M is changed at a later time, the stress relaxation curve changes rapidly, within a few days, from the initial relaxation curve to the other. In these calculations. F falls in 30 years from an initial value of 1394 kips to a low of 1130 kips for the case

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M = 20 and T = 150*F, which is well above the minimum requirement of 980 kips.

The hold-down force for several tendons has been monitored for about 10 years, l

and there is reasonable agreement between the long term trend of the data and those calculations. The measured hold-down force (Ref.1) (measured for six vertical tendons through load cells from t = 100 days to t = 4000 days) missed slightly the theoretical curve using the design data listed above.

A second set of calculations was done using the measured data instead of the l

initfal design data for the steel and concrete. The initial yield stress of the steel tendons was measured and found to be between 252000 and 256000 psi. We then used 250000 pst instead of 240000 psi and took the initial stress in the steel tendons to be 0.7 of that value. Re-examination of the concrete creep data in i

Appendix E of the FSAR (Ref. 2) yielded values of a = 0.25 and a = 0.50 for i

T = 75'F and 150*F, respectively, where t > 20 days. The rapid initial strain during the first 2-8 days was incorporated into the initial elestic strain and thus 6

the effective initial concrete modulus was reduced to E ~. 4 x 10 psi. Figure c

3 shows the time evolution of the hold-do'wn force, using these values, and the comparison with the load cell data (Ref.1). The fit between the theory and j

measured values is much better.

The current lift-off force can be calculated in terms of the current. hold down force, by assuming that only elastic forces are applicable durig a short time 6-

Itft-off test and applying just enough axial tension to the steel (and thus axial

' strain) to reduce the current concrete compression to zero. Thus f = F (1 + s )/ 1-F/AE gg where f is the lift-off force and F is the hold-down force and e=AE,j[AE s

cg For this problem F/A E - 10~4 gg a = 0.026B for the design parameters, and

= 0.0321 for the measured parameters.

Thus, f = 1.027 F (design parameters) or r

1.032 F (measured parameters).

The relationship between the lift-off force and the number of broken strands, M, was determined by linear regression analysis to the data (Ref.1) which yielded t

f = 1425 - 6.95M + 30 kips.

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Chapter 171 III. The Simple Finite Element Model.

An elementary test problem using the finite element code, NONSAP-C, (Ref. 3) tras formulated to evaluate the combined viscoelastic creep model for-concrete and the prestressing tendons. NONSAP C is a three-dimensional finite element code, derived from the NONSAP code.

It is specifically designed to evaluate reinforced

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concrete structures, with a number of concrete constitutive models specifically fomulated for evaluation of various pertinent concrete properties. The main property of the concrete needed here is creep at low stress levels.

In NONSAP-C the creep compliance at constant stress is represented as a sum of exponentials, N

~ ' '*P "Y'i The maximum value of W allowed in the program is far. By choosing og = o, 4

3 E,. E, and v$ = 10 v, 74 1073, 73 10073, v2 = 1073, a good approximation 3

to the logarithmic time fit, described in the previous fit is available.

If 7 3

3 days, the fit is valid for 1 day 5 y g 60 000 days. The values of E, and E c

were obtained from the fit of the Ft. St. Vrain concrete creep data (Ref. 2),

described in Section 2

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The analytical problem of the previous chapter was simulated by a finite element structure with ten 20-node bricks of concrete in a five-high by two-wide column. A steel plate cap of two elements simulated the confinement. A single i

tendon element from the fixed base to the steel cap and passing through the mid-side nodes of the common concrete element was pretensioned. The finite element deformations and stresses were evaluated over a time period from 1 day to 30 000 days after pretensioning in logarithmic steps of one-half decade. Dynamic inertial tems were neglected. Figure 4 shows the axial hold-down force in the tendon versus time.

The time decay of the tendon hold down force is very similar in shape to that calculated from the analytical model. Slight differences in the peak axial. stress 1.

above and below the mean are due to the three-dimensional nature'of this problem and the welded approximation to the steel plate-concrete interface.

The numerical solution to the finite-element simulation of the analytical rodel is sufficiently close to the previous analytical solution in magnitude of stress and axial force, and in its time evolution, that we may confidently use the fit parameters in the NONSAP-C Dirichlet expansion for the creep for_the more complex models of the reactor containment wall described in the next chapter, e

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Chapter IV FINITE ELEMENT ANALYSIS OF A SECTOR OF THE FORT ST. VRAIN PCRV CONTAINMENT WITH PRESTRESSED TENDONS 4.1 Finite Element Representation In this section, we will look at the stresses and tendon forces in a 60* sector of the outside wall of the containment vessel of the Fort St. Vrain reactor. Figure 5 shows the mesh used to mock up the PCRV wall. F,igure 6 shows the horizontal cross-section of the mesh for the containment wall for this sector. From memory limitations of the program, the two inner rows of vertical prestressing tendons and

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the third row tendon are included in this model, and the 210 circumferential tendons are compressed into 5 mock-tendons at a radius of 255 in., which corresponds to the inner row of the three actual layers of these tendons. Each of these circumferen.

tial tendons mocks up 42 of the actual tendons, which is accomplished by using a cross-section and 42 times the actual value. The tendon locations are marked on the figure. The computational model has 1116 nodes,190 20-node bricks, 20 tendon elements, 3049 degrees-of-freedom and the stiffness matrix has a band-width of 733.

The NON5AP-C program was run on the CRAY-XMP and averaged 9 minutes and 30 seconds per the step.

A baseline calculation, described below, was made for en intact tendon struc-ture. This calculation will serve as the point of departure. In this calculation all of the tendons were pretensioned to 70% of the ultinate stress of the steel, which is 168 000 psi. An internal pressure of 705 psi, which is the operating pressure of the helium coolant in this HTGR was applied. The stresses were evalu-ated from one day to 30 000 days in logarithmic steps of one-half decade. Figures 7, 8 and 9 show th2 stresses on the mid-height plane at 300 days. The induced i

axial stress is approximate by uniform, about -750 pst between the inner wall and the ring of circumferential prestress', and is smaller outside. The radial stress varies from the internal pressure of 705 psi applied at the inner wall to a value j

of about -1200 psi at the inner row of the circumferential prestress'. Outside this ring, the radial stress in this calculation goes rapidly to a tensile value and then returns to zero at the outside wall, with a concentration near the. buttress._

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and cylindrical shell junction. This region of tensile stress is an artifact of the finite element tendon mockup and the very limited number of circumferential tendons in the model.

In actuality, the 210 tendons are distributed over three rings of 70 tendons each with the outer ring coincident with the exterior boundary on the innter buttress boundary. The radial stresses between the inner tendon ring and the exterior tendon ring will vary approximately linearly from its. peak compres-J sive value to zero. The decay of the average vertical hold-down force are shown in Figure 10 for the center vertical tendon in the second row.

Three calculations show the effects of removing three different tendons at the time of prestressing. The first calculation removes the center vertical tendon in the second row. The second removes the center vertical tendon nearest the thinnest part of the wall. The third calculation removes one inner circumferential tendon at mid height.

The stress contour plots for the three cases with a tendon removed are not plotted as the significant differences do not show up well on the scale as plotted. The differences are best illustrated by the several plots along the-stated arcs or lines.

The main results for each of these cases are best shown through comparisors of the three stress components, viz., the axial, radial and tangential stresses along certain arcs. Figures 11 through 13 show these components along an are from 0* to 60* at mid height through the outer ring of Gaussian integration points'in the second radial row of elements. This arc is shown in Figure 6.

On each graph, the three cases with one tendon missing are labeled I, II and III in the order described above". Figures 14 through 16 and 17 through 19 show the radial dependence at mid-height at o. 30* and e = 60*, respectively. The former is the line from the inner w'all through the middle of the buttress and the latter passes through the thinnest section of the wall.

When one vertical tendon is removed, there are two basic effects. The first is that the mean axiol compressive stress throughout the cross-section is reduced by approximately

[H)(a,AjA ) where e, is the stress in the tendon and A, g

and A, are the cross-section areas of a steel tendon and the concrete, respective-ly. When one vertical tendon is removed the mean axial stress changes by +40 psi, and the tangential stress by -10 psi. When a circumferential tendon is removed, the mean tangential stress within the circumferential tendon radius - changes by +30 psi and the axial stress by -10 psi. The radial stress distribution is only slightly changed.

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Chap 2er V CONCLUSIONS AND REC 06HENDATIONS j

This report represents a finite element calculation of the foncrete creep in a model of the Fort St. Vrain PCRV after prestressing and the subsequent stress redistribution when selected tendons are degraded or removed. The complexity of the finite element mesh used in these calculations was a compromise between exact physical detail and short computational time. Modelling of the complete PCRV with all of its prestressing tendons would have resulted in inordinately long compu-

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tational times, so fifteen vertical and five circumferential prestressing tendons were included in a 60* sector of the PCRV wall.

The calculations of the creep and stress redistribution were accomplished for l

four cases. These were creep of the model with all tendons present, and creep of the model with one selected tendon missing. These latter three cases eliminated, in turn, the vertical tendon at 30* in the second row, the leftmost vertical tendon in the inner row, and one mid-height circumferential tendon in the inner ring.

The viscoelastic creep model in NONSAP-C is a limited constitutive model in which only the untaxial compressive Young's modulus is allowed to vary with time.

Poisson's ratio remains constant throughout the deformational process. Further, the creep rate varies linearly with stress.

In view of the Ifmited creep test data on the Fort St. Vrain concrete specimens, which is limited to unf axial stress versus uni, axial strain data over a very limited compressive stress range, no further sophistication in the triaxial viscoelastic modelling is warranted.

The upshot of the Ifmited modelling accuracy imposed by computer memory and time 1{mitations and the restricted viscoelastic creep model in the program (and further justified by the limited experimental data available) is that the numerical values of the calculated stresses are not to be interpreted as an exact representa-tion of the actual stresses in the structure. Rather, the orders of magnitude, their trends, and, in particular, the differences between the successive calcula-tions, are significant.

The most significant departure in the structural stresses due to the limited numer of tendons included in the model is the tangential stress component. The solution without any tendons present indicates that the maximum value of-o, is I-

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abDut +1850 psi at the inner wall. With the mockup of the circumferential tendons.

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in the model, an approximate uniform compression of -3700 psi is added.

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Comparison of the unprestressed creep and creep with all prestressing tendons in place show that the axial and the circumferential stress are altered by the impo.

sition of an approximately uniform compression given by i

- A, l

Ao =

gog while the lateral stress increment will vary from zero to approximately (-1/4N)(ao),

dependent on the overall boundary conditions.

l The radial stress is not affected significantly. See the discussion in the I

previous section for the adequacy of the radial stress component computation beyond l'

the circumferential tendon ring.

If one vertical tendon is removed, the axial stress redistribution has two l

I components. The mean axial compression is reduced by approximately (1/N)(ao) and i

there is an additional local reduction of the same order confined to one vertical tendon spacing. The circumferential compression is reduced by the. same amount locally, and the radial stress is only slightly affected if at all.

Calculations for the one-dimensional defomation of a concrete column pre-stressed by a steel tendon yielded estimates of the time evaluation of the hold-down and lift-off forces in the column. The steel was assumed elastic and the j

j concrete was assumed viscoelastic with a constant log time creep law under constant l

stress. The slowly time varying defomation and stress were evaluated numerically

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from the resulting Volterra integral equation. Two sets of material parameters f

were used in this evaluation, first, the design parameters of the steel and concrete j

and, then, the measured parameters as reported in the updated FSAR. The hold down l

force decreases with time after the initial application of the prestressing load; i

t this decay, for the first 20,000 days is approximately linear with log (t + 1) and j

is dependent on the number of broken, strands. The hold-down force decreases approx-imately linearly with the number of broken strands in the range up to 20.

One case was evaluated to investigate the later breakage of tendon strands; f

initially, no broken strands were present at load application, then 20 strands were j

j broken at t 365 days. The hold-down force time history shifted rapidly from the I

M = 0 line to the M 20 line and by t 400' days, the subsequent equal deformation was indistinguishable from the M. 20 at t = 0.

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In all cases, the prGdicted hold-dswn forces stayed well above 980 kips for the range of time O to 20,000 days, for the number of broken tendon strands from M. o to 20 and for the range of material parameters and temperatures (75* to 150*F) considered.

The main conclusions of this evaluation are' in two parts. First, we state the conclusions for the partial degradation of an isolated prestressing tendon. Then we examine the complete failure of one tendon in a region of multiple tendon.

The hold-down force and the related concrete stress decreases proportionally with the fraction of broken tendon strands up to spproximately one quarter of the strands broken. This decrease is independent of the time the strands break. The hold-down force decreases linearly with the logarithm of time due to creep. For the extreme combination of concrete creep at 150*F and tendon degradation over the lifetime of the structure, the prediction is that the hold-down force at the end of the lifetime will be higher than the minimum safety requirements.

When we consider complete degradation or total failure of one tendon, we must consider the adjacent tendons. At the time of tendon failure there will be a uniform readjustment of the compressive stress. This mean change in the fraction of the compressive stress is the hold-down force for the one tendon divided by the total cross-section area.

In addition, there is a slight local tension added in a region near the removed tendon, whose typical size is a circle with radius equal to the original tendon spacing. The magnitude of this localized stress change is 20 to 50% of the mean stress change. Over the lifetime of the structure, the change in the remaining hold-down forces is the same as if all tendons were present, and t $.this change in the stress state will still be in the safe range.

'Conbining these two sets of conclusions, either partial or complete degradation of one,prestressing tendon will slightly reduce the mean compressive stress in the same direction in the concrete. In addition, a very slight local efffect is noticed.

This stress and the hold-down force will remain within structurally safe bounds.

ACKNOWLEDGEMENT Mary Marshall, 0-13, painstakingly created the NONSAP C mesh and performed the code runs on the PCRV wall structure.

REFERENCES 1

1.

Letter: Warembourg, D. W., (Manager, Nuclear Production), Fort St. Vrain Nuclear Generating Station, to Collins, John (Regional Administrator)

U.S.

Nuclear Regulatory Commission, Region IV, April 12, 1984, 2.

Fort St. Vrain Nuclear Generating Station, Up' dated Final Safety Analysis Report, Vol. 6, Appendix E-20.

3.

Anderson, C. A., Smith, P. D., and Carruthers, L. M., "NONSAP-C: A Non-linear Stress Analysis Program for Concrete Containments Under Static, Dynamic, and Long-Term Loadings," Los Alamos LA-7496-MS, Rev.1, NUREG/CR-0416,(1982).

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STEEL CAP d

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a s

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A E'

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C, C

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Fig. 1.

Schematic Diagram of the Concrete Column With a Steel Prestressing Tendon.

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

t*

s.

1 r

I2 1400' 1

r=<F M=0 I

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T = 15( F M=0 T = 75' F M -10 j

1300 -

1

'E 5

T =150*F M = 10 U

M -20 AT 305 DAYS cc O*

T=#F M = 20 1

2g o 1200 T = 150* F M = 20 I

O J

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j 1100 l

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t i

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a e

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10 100 1000 10,000 i

l TIME + 1 ~ DAYS l

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Calculated Hold-Down Force versus Time at 75'F and 150*F for 0, 10 and 20 broken strands.

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MEASURED AVG lEMTRAPOLATED) 1 r

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MEASURED + 14 N

N aL N

VM17 i

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M4 NN N

w N

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CALC (= Uo67 N

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y 4* = -750 pel

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N N

N s

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