ML20082G556
| ML20082G556 | |
| Person / Time | |
|---|---|
| Site: | Fort Saint Vrain  |
| Issue date: | 10/26/1983 |
| From: | Bennett D, Fugelso E LOS ALAMOS NATIONAL LABORATORY |
| To: | Office of Nuclear Reactor Regulation |
| Shared Package | |
| ML20082G536 | List: |
| References | |
| CON-FIN-A-7258 NUDOCS 8311300246 | |
| Download: ML20082G556 (47) | |
Text
ni THERMAL STRESS ANALYSIS OF SELECTED FORT ST. VRAIN HOT SPOTS Eric Fugelso and Deborah Bennett Advanced Engineering Technology Los Alamos National Laboratory '
Los Alamos, New Mexico October 26, 1983 Division of Licensing--NRR, NRC .
.NRC FIN A-7258 4
Review of Selected Fort St. Vrain Issues 8311300246 831118 PDR ADOCK 05000267 p PDR
4 r
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NOTICE 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, or the results of such 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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1 THERMAL STRESS ANALYSIS OF SELECTED FORT ST. VRAIN HOT SPOTS Erik Fugelso and Deborah Bennett, Q-13 Advanced Engineering Technology Los Alamos National Laboratory Los Alamos, New Mexico 1.0 Summary Under normal operating conditions at the Fort St. Vrain (FSV) Nuclear Generating Station, seven hot spot areas have been identified in the liner cooling system of the Prestressed Concrete Reactor Vessel (PCRV). Each of the hot spot areas is located at the interface between the PCRV concrete and the steel liner encasing the PCRV interior and its penetrations, and are generally the result of a higher than anticipated heat load or a discontinuity in the protective thermal barrier. The temperature data for the seven hot spot areas have been reviewed, and thermal analyses performed to determine the temper-ature distribution in each of the areas. The objective of this study is to calculate the thermally-induced stresses in the PCRV regions where concrete temperatures exceed the Final Safety Analysis Report (FSAR) hot spot criteria, utilizing the thermal analyses in the previously submitted Los Alamos report,III and to compare the calculated thermal stresses to ASME and FSAR limiting values.
Structural calculations were performed using a recommended evaluation cri-terion on four of the most severely affected hot spots. The criterion employs an initial static thermal stress calculation, using temperature dependent elastic constants, in determining the stresses and displacements associated with a hot spot region. The resultant stresses are then compared to temperature-dependent compressive and tensile concrete limiting stresses.
We have concluded that none of the seven hot spot regions identified causes stresses to exceed ASME stress limits. Specifically, feye hot spots, namely those with concrete temperatures in excess of 300 F, were evaluated in detail. Compressive stresses generated in all these cases were signifi-cantly less than the compressive ultimate strain. In two of the cases, the
o peak value of the tensile stress exceeded the FSAR limit, but not the ASf1E limiting values, while in the other two cases evaluated, the maximum tensile stresr:a. were significantly below both ASME and FSAR limits. The remaining three cases had less severe temperatue excursions and their thermal stresses should be even less severe. Therefore, liner cooling system hot spots presently generate no serious stress concentrations in the concrete liner and/or the PCRV of the Fort St. Vrain Nuclear Generating Station.
2.0 Background
Bennett OI analyzed the seven identified PCRV liner hot spots to deter-mine the localized temperature distribution and the extent or volume of con-crete exceeding the hot spot temperature limits specified in the FSAR. The seven regions studied are listed below and are identified in Fig.1.
o the core support floor region beneath the core barrel support struc-ture; o the loop divider baffle intersection with the PCRV bottom liner; o the sidewalls and the lower access penetration; o the helium purification train crossover pipe and penetration inter-section in the PCRV top head; o the core outlet thermocouple penetrations in the PCRV sidewall; o the refueling penetrations in the PCRV top head; o the peripheral seal at the edge of the lower floor; o the steam generator penetration intersection with the PCRV bottom head.
PCRV Thermal Environment The safe operation of the PCRV is dependent on strict control of its thermal environment at all times. This is accomplished by protecting the PCRV from the principal source of potentially damaging heat, the primary helium inlet coolant circulating within the vessel. The PCRV is protected by a thermal barrier lining the interior of the PCRV and many of its penetrations, and the PCRV liner cooling system. The thermal barrier consists of ceramic
m
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o insulation blankets compressed against the helium side of the PCRY liner by steel plates. Heat passing into the liner is removed by the liner cooling system, which is composed of two independent cooling loops, each loop consist-ing of water-cooled tubes (1 in.2 in area) continuously welded to the con-crete side of the liner in an alternate tube / alternate loop manner.
Under nomal operating conditions, the two means of thermal control (in-sulation and water cooling) operate to maintain the three criteria defining the design temperature distribution in the PCRV such that:
o the bulk concrete has an effective temperature of 130 F.
o the local concrete temperature is a maximum of 150 F, midway be-tween cooling tubes at the liner / concrete interface, and 0
o the liner cooling water differential temperature is 20 F.
The FSAR defines the existence of a hot spot when the local maximum con-crete temperature is 2000F under normal operating conditions, and 250 F with only one cooling loop in operation. This definition is further expanded by Amendment 29 of the FSAR such that, in the core support floor under the 0
barrel support pad, the cooling water differential temperature may be 40 F, 0
- and the local maximum concrete temperature may reach 250 F under normal operating conditions, and 360 F with one cooling loop in operation. The FSAR clearly states that these extended temperature limits are applicable to the core support floor periphery only.
The heat load to the liner cooling system varies with reactor power, is dominated by the heat transfer across the themal barrier, and is a direct function of the differential temperature between the inlet gas temperature (u750 F at 100% power) and the liner cooling system average water temper-ature (wl100F). A much smaller fraction of the heat load is contributed by bypass flow of the primary coolant within the insulation and around local internal equipment such as the lower floor peripheral seal.
The liner hot spot areas are generally caused by a discontinuity in the thermal barrier, providing a lower resistm ce path for the heat to the liner /-
concrete. Structural attachments to the liner also provide a heat conduction I
1
r path to the liner. Higher than expected heat fluxes can also be caused by bypass or impinging helium flow patterns, or by unsuitable cooling tube con-figurations .
3.0 Analysis Bennett UI modelled each of the seven PCRV liner cooling system hot spots identified, using the TSAAS finite element code. The basic results of the thermal analyses are listed in Table I, and include the peak temperatures and the volume of concrete affected for each hot spot area.
Based on the thermal calculations, four of the hot spots in the liner cooling system containing concrete temperatures exceeding 300 F were se-lected for further analysis using the ABAQUS I2I thermo-mechanical finite element code to investigate the structural effects of the increased temper-atures and the resultant integrity of the concrete. The four regions selected are:
- 1) the core support floor below the core barrel,
- 2) the intersection of the loop divider baffle with the concrete wall,
- 3) the helium purification crossover train penetration, and
- 4) a typical thermocouple penetration.
Only the core support floor involves a region where significant structural loads accompany the additional thermal stresses. The three remaining hot spots are geometrically similar to the thermocouple penetration configuratio1, but with lower concrete temperatures. Therefore, the thermally generated stresses were assumed correspondingly lower and not investigated in detail.
3.1 Material Properties Concrete material properties are known to be functions of temperature.
The modulus of elasticity, compressive strength, thermal conductivity, thermal expansion and Poisson's ratio are concrete properties that experience varying amounts of degradation with elevated temperature.
+ . .
TABLE I PCRV LINER HOT SPOT IDENTIFICATION Predicted Local Max.
Concrete Temp. Depth of Apparent Cause 100% Power Penetration in. Vol . , Ft3 Core support Bypass flow Between Tubes floor under core barrel 3500F < 0. 5 ( 0.4 support pad. 2100 at T/C Location Loop divider Discontinuity in 350 < 3.0 < 1.2 baf fle thermal barrier producing thermal short to liner.
Helium purifica- Insufficient cooling 330 1.0 0.6 tion Train at intersection of c rossover HTFA penetration penetration crossover pipe.
Thermocouple Discontinunity in 345 2.5 0.5/
penetration thermal barrier Penetration produciag thermal short to liner. .
Refueling Flow pattern 290 1.6 < 0.03 impinging on penetration liner.
Steam generator Insufficient cooling 273 2.2 x 5.6 < 0.02 at intersection of penetration and bottom head.
Peripheral seal Discontinuity in 220 0.25 < 0.1 thermal barrier producing thermal short to liner.
C
- o. .
In complying with elevated temperature degradation, the concrete proper-L ties have been modelled in the following manner. The modulus of elasticity at room temperature (70 F) is 5.0 x 106 psi, and linearly decreases to 80% of that value at 200 0F, with an additional linear decay to 50% at 500 F. The uniaxial ultimate compressive stress is modelled to decay linearly from its
- room temperature value of 6000 psi to 80% of the room temperature value at 2000 F, and assumes a constant value at tcmoeratures any higher. Poisson's ratio is constant at 0.167. The thermal conductivity (k = 1,0 Btu /ft hr F) and the linear thermal expansion coefficient (a = 5.3 x 10-6jF o ) are considered constant throughout the range 70 F-400 F. The properties of the carbon steel liner are considered constant throughout this temperature range.
Much research has been done on how elevated temperature affects concrete properties, but Davis I3I concludes from his published work that for tempera-tures up tc dOO F, the deterioration in concrete structural properties is ordinarily tolerable. Thus a rationale exists that temperatures in excess of 250 F may be tolerable if each case is examined individually for stress magni tudes .
3.2 Boundary Conditions Using slight refinements to the finite element meshes in the previous study, the thermally-induced stresses in the concrete and in the steel liner were calculated using the finite element ccde ABAQUS. Appropriate heat transfer coefficients and heat flux parameters were applied to the axisym-metric and plane strain models. A sliding' interface is modelled between the PCRV concrete and the steel liner, thereby allowing compressive stresses to be transmitted across the interface, but not tensile stresses.
l 3.3 Calculations Estimates on the allowable increase in stress-caused by an elevation in l concrete temperature can be found from the elastic equation ;
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.- ~ ~ , _ _ . _ - . _ _ _ _ - , _ , - _ - . - - _ _ _ _ . . - - - . _ - - - . _ - - _ -
r Em T U
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using two limiting sets of material properties and comparing with the allow-able compressive stress. The upper bound on the stress occurs when material properties are specified from American Concrete Institute (ACI) and ASliE Boiler And Pressure Vessel standards (Ref. 4 and 5) at room temperature, and a lower bound is produced by using concrete material properties from the FSV Final Safety Analysis Report (FSAR) as given in Ref. 6, degraded by elevated temperature. The differences in material property values are shown in Table II.
TABLE II BOUlJDIl1G MATERIAL PROPERTY VALUES Property Upper Bound Lower Bound Temperature, T 700F 3500F flodulus of Elasticity, E 5.0 x 106 psi 3.25 x 106 psi Poisson's Ratio, V 0.20 0.167 Thermal Expansion, a 5.1 x 10-6 /oF 5.3 x 10-6 /oF Concrete Compressive 6000 psi 4800 psi Strength , fc '
The temperature difference considered is the difference between the allow-able hot spot temperature of 200 F and the general 350 F temperature of the four most severe hot spots, resulting in aT = 150 F. Substituting the temperature difference and the material properties into Eq. (1) indicates an Upper bound a = 2i00 psi and Lower boundc = 1500 psi These upper and lower limits should cover the range of increased stresses ex-pected by a temperature elevation of 150 F.
o, .
The FSAR (Section 5.4.4 and E.1-1) gives the minimum compressive strength of.the FSV concrete as fc ' = 6000 psi at prestressing. The FSAP. and the
- ASME Boiler and Pressure Vessel code give the limit design stress, or the combination of primary and secondary stresses, under normal conditions as FSAR ASME Primary Compressive .f cc = 0.6 C fc ' f cc = 0.6 C f cua Stress, psi Primary Tensile f ct =6 f I = 7.5 f cua c ct Stress, psi where f' = minimum compressive strength c
f
- primary compressive stress cc f
cy,
= ultimate compressive stress as measured f
ct
- principal tensile stress
, C = multiaxial compression coefficient The measured compressive strength at room temperature for the FSV concrete is 8300 psi (10) Thus the maximum principal compressive stresses become 3600 and 4980 psi for the FSAR and ASME limits, respectively at room temperature, wi th C = 1.0. Using temperature-degraded strength values at 350 F we get maximum principal compressive stresses of 2880 and 3980 psi, respectively.
The maximum principal tensile stresses are 230 psi and 685 psi for the room temperature FSAR and ASME limits, with values of 210 psi and 610 psi for the elevated-temperature tensile stresses at 350 F.
These bour. ding calculations are introduced to emphasize the following point. If the original design is ' based on the fact that the temperatures in the concrete are below or at the FSAR specifications, the additional secondary stresses that must be accommodated are of the order of 0.3-0.35 cf '. This value falls below FSAR and ASME code stress limits, and thus the combination of the normal design stresses and the secondary thermal stresses developed by an excess of 150 F should remain within tolerable limits.
-g-4.0 Results The results of the ABAQUS finite element calculations for the four se- .
lected hot spot regions are presented in this section. As mentioned pre-viously, Ref. I had identified the hot spot regions, had developed mathemati-cal models of each region suitable for further analysis, and had conducted parametric studies of the temperature distributions. In the following sec-tions, we will use these results in combination with refined basic models for each hot spot.
Core Support Floor
~
One of the most serious hot spot areas is on the upper, outer edge of the core support fitor, beneath the core barrel support pad, as shown schemati-cally in Fig. 2, and is caused primarily by a bypass flow between the core support pad and the core support floor liner. The detailed corner of the core support floor is shown with the core barrel resting 'on the core support pad in Fi g. 3.
Model boundary conditions shown in Fig. 4 indicate the modelled heat transfer coefficients, the cooling tubes and the concrete area affected. The t finite element mesh used in the calculations is shown in Fig. 5. Slip along and separation of the steel liner and the concrete is provided. Vertical dis-placement only is allowed at the left edge and radial displacement only is allowed along the bottom, since this mathematical model is a very small por-tion of a massive structural component.
The calculated temperature profiles are shown in Fig. 6. Temperatures U
exceeding 250 F exist in the concrete between the third ar.1 fourth cooling tubes, the fourth and fifth, and the fifth and sixth to a depth of about 0
1/4 in. The peak temperature in the concrete is above 350 F between the fourth and fifth cooling tubes at the steel-concrete interface.
Figure 7 shows the maximum principal stress contours in the concrete due to thermal stresses only. The contours indicate that 'he maximum principal
l .
stresses vary from a 200 psi tensile (+) stress at the concrete / liner inter-face between the fifth anc sixth cooling tubes, to a 200 psi compressive ( i stress at the concrete / liner interface between the fourth and fifth cooling tubes.
The minimum principal stress contours are shown in Fig. 8. All minimum principal stresses are compressive (-), and range from 200 psi in the bulk concrete between the fourth and fifth cooling tubes, to 850 psi at the concrete / liner interface between coolings tubes two and three.
In this configuration, the steel liner on the side of the core support floor tries to separate from the concrete. The liner on top of the core sup-port periphery bends due to differential expansion and slips relative to the concrete / liner interface, thus transmitting very minimal compressive stress to the concrete. The resulting maximum principal tensile stress (+200 psi) gen-erated is well below the high-temperature tensile limit of 610 psi (ASME) and the naximum principal compressive stress is less than 0.18c f ' at the ele-vated temperature, using fc ' = 4800 psi.
The combined effect of primary stresses and secondary themal stress is considered by applying a vertical load of 70 psi that corresponds to the dis-tributed weight of the core barrel and support structure. In this model, no other primary loads are applied to the upper or outer edge of the core support floor. The upper, outer periphery of the core support floor is restrained to radial movement and slip is again allowed between the liner and the concrete
- interface.
Under these conditions, the maximum principal stress contours in Fig. 9 show a slightly greater depth for the 200 psi tensile stresses in the concrete between the fifth cnd sixth cooling tubes. Likewise, Fig.10 indicates a sim-ilar minimum principal stress distribution with a maximum compressive stress of -925 psi, or 0.19 f ' . In general, the location and values of the ,naxi-mum and minimum principal stresses are very much the same as in the previous analysis with no consideration of primary loads.
Loop Divider Baffle The loop divider baffle divides in half the plenum between the lower floor and the PCRV bottom head. A hot spot is generated beneath the intersection of the loop divider baffle and the PCRV, and arises from the heat conducted along the steel baffle structure and from the gap convective effects between the thermal barrier and the baffle support. Figure 11 schematically shows the arrangement of the baffle seal, support bracket and support flange relative to the liner / concrete interface.
Figure 12 shows the modelled intersection with appropriate thermal boun-dary conditions. Fig.13 shows the finite element mesh associated with the structural model, using one half of the structure for syrnetry reasons. Slip is provided for between the liner and the concrete interface. No external structural loads are placed on this assembly.
Figure 14 shows the temperature field in the baffle bearing plate, support bracket, support flange, liner, and concrete. Temperatures in the concrete exceed 350 F at a point directly beneath the baffle and continue to exceed 250U F at a depth of 2 inches.
The maximum principal stresses (Fig.15) are only 40 psi compressive (-)
in the concrete directly beneath the loop divider support structure, with a ninimal peak tensile stress of 40 psi (+) an inch deeper into the concrete.
Minimum principal stresses (Fig.16) do not exceed 650 psi compressive (-) in the same location. Therefore, the maximum tensile stresses are well below the required tensile limits, while the maximum compressive stresses are only slightly greater than 0.3 fc '*
Helium Purification Crossover Train A crossover pipe connects the high temperature filter absorber (HFTA) and the front end cooler of the helium purification trains in the PCRV top head.
A hot spot region has been identified in the concrete at the intersection of the crossover pipe and the HTFA penetration, as modelled with the thermal boundary conditions in Fig.17.
The finite element mesh used for this intersection is shown in Fig.18.
Slip is allowed between the liner and the concrete. The interior surface of the penetration or pipe is not structurally loaded, and pipe deformation is allowed only in the radial direction.
The temperature contours at the intersection are plotted in Fig.19. The 0
maximum temperature is about 300 F, and occurs in a small triangular ring surrounding the crossover pipe intersection with the wall of the HTFA penetration.
The maximum principal stress contours shown in Fig. 20 indicate a region of tensile stresses on the order of 600 psi (+) in the bulk concrete around the intersection. The stresses become compressive (-3000 psi) closer to the interface of the pipe liner and the concrete cavity. It is apparent that the outward thermal expansion of the steel liner retards the inward thermal expansion of the concrete cavity, resulting in a compressive stress at the interface.
The minimum principal stresses are shown in Fig. 21, where the region of concrete in 500 psi (+) tension is significantly reduced. However, the stress contours show more dramatically the compressive effect of the outward growth of the pipe liner on the concrete cavity. In moving closer to the concrete / liner interface, the minimum principal stresses become compressive with values reaching as high as 5100 psi (-) at the interface in the pipe cavi ty. However, the compressive stresses at the liner / concrete interface are confined.to a very narrow boundary layer in the concrete around the crossover pipe, and fall dramatically to approximately 20% of the surface value in less than 1/8 inch of concrete depth. In the angular corner around the HTFA penetration, a similar-set of compressive stress contours with lower values exists, but the boundary layer thickness is about 1/2 in.
The peak compressive stresses require further discussion, in that their values appear to exceed ASi1E limits. Determining the directions and values of the principal stresses at the pipe liner / concrete cavity interface indicates that the maximum principal stress (-4000 psi) is radially compressive, that
)
the minimum principal stress is a compressive hoop stress (-5100 psi) and the intermediate principal stress is a . axial compression (-4000 ps. i. It is known that the ultimate compressive strength in concrete (fc ') increases rapidly with lateral confinement (Ref. 7, 8 and 9), as is caused by the ex-pansion growth of the steel liner impinging on the concrete cavity. Using these compressive stresses, the triaxial factor C is calculated from Ref. 4 as C = 2.0. Thus fcc = 5750 psi and, thus, the compression in the " boundary layer" is below the limit.The maximum tensile stress at the PCRV remain within the AS!1E limit of 610 psi, but exceeds the FSAR limit of + 210 psi. The stress calculation was redone with a free surface introduced along a line where the tensile stress is at a maximum. All the readjusted stressec within the concrete were lower than without a crack. The concrete may thus be broken; however, it carries no structural load, and is confined and restrainea by the steel liner, thus the thermal and bulk properties remain intact. We expect that no effect on any reactor or PCRV parameters will be made by this possible break.
Themocouple Penetration Seven penetrations through the PCRV sidewall permit access for the themo-couple instrumentation measuring the core outlet helium temperatures. A hot spot region was identified in le region of the intersection of the thermo-couple penetrations with the inside of the PCRV sidewall, as shown with the typical section in Fig. 22.
The finite element mesh for the penetration intersection is shown in Fig. 23, and uses the thermal boundary conditions indicated in the previous figure. The slip condition is again imposed at the interface of the penetra-tion liner and the PCRV liner.
The temperatures in the concrete are shown in Fig. 24. The peak temper-ature is about 3000F at the corner of the pipe and PCRV intersection.
s
The maximum principal stresses are shown in Fig. 25, and indicate a region surrounding the penetration that exhibits tensile behavior on the order of 500 psi (+). Imitating the stress behavior of the crossover pipe, the tensile stresses become compressive at the liner / concrete interface with a value of 1500 psi (-). Similar recalculation of the structural behavior in this case, as in the previous case, with an introduced crack, shows no significant change in maximum or minimum stress. It is therefore concluded that the concrete in this approximately 1/2 inch boundary layer is significantly below the ultimate compressive capability if the confined concrete.
The minimum principal stress contours in Fig. 26 also show a drastically reduced area in tension (+500 psi) with a quickly developed boundary layer area in compression (-3600 psi) at the interface. These results imply that the maximum and minimum stresses generated have the same qualitative and quan-titative characteristics as described for the previous calculation for the crossover pipe. In this case, the smallest compressive principal stress is 1600 psi (-). The triaxial compression factor (Ref. 4) is 1.6.
Conclusions Four of the more severe hot spots in the PORY of the Fort St. Vrain Nuclear Power Generating Station, identified and thoroughly described in a previous study, were analyzed for thermal stresses, utilizing the ABAQUS fi-nite element code. These hot spots were all characterized by concrete tem-0 peratures in excess of 300 F. Generally, the hot spot regions are not structural members, but one of the hot spots, the core support floor, was in a region where structural load bearing capacity was an important consideration.
In all of the ABAQUS calculations, the thermally generated stresses caused by increasing the concrete temperature some 150 F over the hot spot temper-ature limit were well below the ultimate stress for the existing state of tri-axial stress and temperature, as defined in ASME and FSAR codes. Table III summarizes the calculated stress states in each of the four hot spot areas.
l .
Because of material degradation from elevated temperature effects, the pos-sibility of incipient microcracking in the concrete is present in all of the cases calculated, but should be of minor concern with respect to overall re-actor operation and safety.
The four cases studied fell into two main categories. The first de,cribes the thermal stresses in the core support floor and the loop divider baffle, while the second characterizes the helium purification crossover train and the thermocouple penetration.
In the first category, the steel liner overlies the concrete body, which is essentially in a state of plane strain. The liner expands and slips along the interface transmitting little or no compression to the concrete below.
The concrete is thus allowed to expand and relieve the compressive stresses generated in a thermal hot spot; in the case of the core support floor, to about 850 to 900 psi and, in the loop divider baffle case, to even lower val ues. The tensile stresses developed about the cooling tube cut-outs are quite small. The values of the maximum compressive stresses, expressed as a 0
fraction of the compressive ultimate stress at 350 F, are about 0.2 cf ' 0" smaller, and in the core support floor, when combined with the structural load, remain below 0.22 fc'
- The second category of cases is characterized by a bulk concrete body with a cylindrical hole. In the case of the helium purification train, and the thermocouple penetration, both the cylindrical cavity and the free surfaces have a steel liner, with the hot spot in the concrete located near the junc-ture of the two liners. The compressive stresses within the concrete readjust under themal expansion, but are confined by the pressure generated in the thermal expansion of the steel liners. As a result, the stresses in the con-crete are quite low everywhere, except for a very narrow boundary layer, on the order of 1/8 in. - 1/4 in., near the cavity face and near the plane liner. All the principal stresses in this volume are compressive, diminishing quickly with distance from the free surface. In the layer near the cavity, the typical maximum stress is about -4500 psi and the typical minimum stress
~. .
(always the hoop stress) is about -5500 psi for the HFTA and -2500 psi and
-3600 psi, respectively, for the TCP. These values fall below -1000 psi at 1/4 in depth. The most compressive stresses, when considered in isolation, exceed the ASME and FSAR standards. However, the standards do not adequately describe the increase in concrete ultimate strength with confining pressure.
When we consider the effects of triaxial compression on the compressive strength of concrete and compare the compressive stress with this increased strength, the concrete is well within safe limits. However, the maximum ten-sile stress obtained in both cases is very nearly at the ASME limit and ex-ceeds the FSAR limit. The volume of concrete thus affected i; very small, most probably is elastic (since it is below the ASME limit), is in a non-structural area, and we therefore conclude that no concrete failure will occur. In addition, calculations with added crack at the region of larges',
tensile stress indicates that all readjusted stresses are well below all applicable limits.
Thus, the stress concentrations from the four hot spots with the most severe temperatures are below any rational failure criteria. Since the tem-peratures in the three other hot spots were smaller, it is concluded that the stresses in those cases will be less severe than the cases calculated. Since these stresses are all within the elastic limits, more complicated thermal loading histories, such as cyclic cooling and heating, should have negligible effect.
e
+ -
TABLE III SIGNIFICANT THERMAL AND THERMAL STRESS FEATURES OF THE FORT ST. YRAIN HOT SPOTS Yolume Exceeding Maximum Temperature Maximum Minimum Allowable Temperature Limit Principal Principal Compression Location (OF) (ft 3) Stress (psi) Stress (psi) f cc (psi)
Core barrel /
core support floor 350 0.4 200 -850 -2880 Loop divider baf fle 350 1.2 40 -650 -2880 Helium Purification Train Crossover 330 0.1 600 -5100 -5750 Core Outlet Thennocouple Penetration 31 0 0.5 500 -3600 -4600
References
- 1. Bennett, D., " Evaluating the Severity and Consequences of the PCRV Linear Hot Spots at the Fort St. Vrain huclear Generating Station," Los Alamos report, TN-FSV-2,1982.
- 2. "ABAQUS: A General Purpose Finite Element Code," Hibbitt, Karlsson and Sorenson, Providence, Rhode Island,1982.
- 3. Davis, H. S. , " Effects of High-lemperature Exposure on Concrete,"
Materials Research and Standards, Vol . 7, No.10, pp 452-469 (1967).
- 4. " Code Requirements for Nuclear Safety Related Concrete Structures (ACI 349-76) and Commentary on Code Requirements for Nuclear Safety Related Concrete Structures (ACI 349-76)," American Concrete Institute, Detroit, Michigan,1976.
- 5. ASME Boiler and Pressure Vessel Code,Section III, Division 2, " Nuclear Power Plant Components," 1980.
- 6. " Fort St. Vrain Final Safety Analysis Report," Public Service Co. of Colorado,1983.
- 7. Mills, L. L. and Zimmerman, R. M. " Compressive Strength of Plain Concrete Under Multiaxial Loading Conditions," ACI Journal, j7, 802-807 (1970).
- 8. Gardner, N. J. " Triaxial Behavior of Concrete," ACI Journal, j6,136-146, (1969).
- 9. Cedolin, L. , Crutzen, Y.R.J. , and Del Poli, S. , " Triaxial Stress-Strain Relationship for Concrete," Journal EMD 103, 423-439 (1977).
- 10. Cheung, K.C., "FSV-Core Support Floor Stress Analysis," General Atomic Co. , VEC:729:78,1978.
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- s
. i gyg /
- '..f. ^ 1.. ^
w ' / 2*2 C "
&M - I ama -n M N P.DDDsE mmaf8I a
Fig. 2. Schematic of the core support floor hot spot.
\ .
l CORE BARREL SEAL p vg y .
CORE BARREL
/
mlY/
Q'A NN\ Ns_NN s\ g f e.
\_ .\k \
r CORE SUPPORT STRUCTURE
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. cr.c;;;;;;;.e .;;:ee .w.-:%- 4 :. w..:..w: 1 :a.:.r..c ,
i,e,i, ,0. . ' . . i .e s Fig. 3. Core support floor periphery.
l aus.
l h = 150 - 230 BTU /hr ft27 support pad m -
Adiabatic
, ,, ,r r--"
.I- .I f
concrete Adiabatic Fig. 4 Sectional taodel of the core support floor periphery.
< o i i
// ; i l\\
h l I i l l l l h* 3! UI i O '
U [_.L i
l i
Cl i i i t i I i '
I i g i I i I l l l I !
L.
II i 4 II '
iI. IIJ l g
i h.
4 1-
,. t t i e 6 I t t t g
I h i r.
n 7
F Fig. 5. Finite element mesh of the core support floor.
K~
2**
l l 5**
t.
l!
',,l ,.
g m.g : . 2 2:
^ ^
z 4 i M'D'tf6WU C E
O Q .
k rt
, *I h- i f r! ST v:Als pts. t 5
i i-3 Fi g . 6. Contours of temperature in core support floor hot spots.
& //////
IMAGE EVALUATION / 44, 4
I/$/ TEST TARGET (MT-3) ((/j\NNXy k//77 %+ 4- e
'M> (44 '["
f 4 i.o gna na
. 8 EE I.I
!!lE
!' Ele
{ l.8 1.25 1.4 1.6
< - 150mm >
< 6" >
- < % + /4
- k,V h/ ?' sib lO o
4+4*
w .
l
.. . , , , o s . s. m I *.E 3 8 63 3 63 4 C3 4 '5 3 8 *s.m t est t K.E I a3 9 EJ B
.e ==
i V . ~ u) y f -
e-s - o. -
[ - .
s -
- l y 1 i
- I
!.; ri, ST, vaatn P t!!, !
e, . -.
's f
! ?
Fig. 7. Contours of maximum principal stress in the concrete for the core support floor hot spot (thermal stresses only).
l
E'O' "
!, 7."
% *E..
V.'
am n =
.:. . =
u in.=
si S W- , Qi
)
l ;7
.i o
a i
il I
e i
rT. ST. VRalN P909. I wa ms l l!
iw -
i
! I l
Fi g. 8. Contours of minimum principal stress in the concrete for the core support floor hot spot (thermal stresses
( o nl y) .
t I
l
\
l
e .
e e
St. et tC % TTTI g I .8. AJ s t *4K 3 3 M.3 S ec 3 4 Mac 3
% *E .3 8 al.3 % alt
? .E 3 0 ter. 3 m3 i.i =3 n ise 3 j
=
L/ '
i
_}"/ j _{
w i !(
l-f
' d k .
rT, ST. vmRIN PROS. ;
[ ' ****"
2 l l3 l l*
s l
l Fig. 9. Contours of maximum principal stress in the concrete i
for the core support floor hot spot with structural
! load.
i
~l: T "
! Te" I Yi l "l,,*e ..
l l i ,
2.".
! Ez f
4 l
1
.a .l '
'I I
! -l i
[i. ri. ST. vRR[N PROS. !
y wi -i s
l'.
)
Fi g . 10. Contours of minimum principal stress in the concrete for the core support floor hot spot with structural load.
PCRV CONCRETE LINER SUPPORT FLANGE
- ~
-[' . BEARING PLATE i'
ft" m ,
- .( ,
, 1 l'
(y 1 l \
,e-
_ c ,
() -
.9
@ ug SEAL
- -s
".b'.*
SUPPORT BRACKET PARTIAL PENETRATION WELD (Region B)
FILLET WELD (Region A)
Fig. 11. Loop divider baffle seal cross section.
. _ _ - . - - - - - ,, , , . . ..,,_._,7 -- .,- . ,-,- -
. . _ . . . _ -, _,._._,_.--y
f LODP DIVIDER BRFFLE
. t . I f . f f . . t h -
~
4 W- pO h = 150 btu /hr/f t F 2
s .' h = 70 btu /hr/ft f -
r
- a. -
, d- ,<s -
l -
l .
~
- Adiabatic 4- .
4~ .
- I .
i . . .
- -1. a. 4. s. : .
Fi g . 12. Cross-sectional model of the loop divider ba f fl e .
l l
y +-e , - .v.,,. - .---%, -..--,,..,,---,..--_._,,,,.,-.ec,- -
-- _.,,,gnr- - - - - - -
e m --,a -ww ,,,,m, _--,-
I. -
? -
i i
!' M I
( l Iset i ;i l I 111 l 41 !
11 l l
iii ! .ii 6 iit't! Il I itii; ..li ~
I l l . w l
... l .
)
5 s
, u lit ti... i '
j .
- iiti'!, , I l
- iisi .! t t t r
.t'i ;t.; i g l6i,;i.t, i 6 I. t. ! ' + . i e j l g
_ 11. .
.. , i j i ii1 'i r g ,
l
.- ii. i i -
8 d.... .. i d.
d b,ii! ; I c a i i l ? - - , i l :i --
- 3. . . . , . ... =. :. .
- 1
,, e a. -
g l
! J
< \
l
'a !
l i
F,
.c l.
l l
l Fi g . 13. Finite element mesh of the loop divider ba f fl e.
l l
i t
f l
l l
l t
mm t.t.* suLa $"
i R.3 8 ET.E 3 IV).c /
e E.E 1 W.E S E.E O e:N. N 9 M.r 80 BC.C it TW:.E , F, 88 et.c ,
fv?.C
- 33 ts M.E r j
11 T.C 14 E.C <A
(
k[
ii
\
P E
+
li!
P 1
l-ik a a
b '
E L rT. ST. 79:N Barrt,t ws e i T.
et i
I i
Fig. 14. Contours of temperature for the loop divider baffle hot spot.
Al n-
..,.% ,,=,,
\l4.
i.e. u i l t m.z a .:= =
3 -tm.m 4 4 1
- s .:
8 *J e t ,
, e.x 8 c.2 9 IJC.E 10 lE.E
- i. m.=
l L
et i
a l
E E
r /
( /
l iI 31
\Y' I J t 3
l
- s *
- r. i
$ C . ST. M;!N Brirr,.c
, , , ~ ~ .
3 I:
l . l, i a-w i
i Fi g . 15. Contours of maximum principal stresses in I the concrete for the loop divider baffle hot spot.
l .
l
. .. mm B.8. suL.2 4 -EnE.E 3 *M 2 3
- AEI E 4 *182.2 5 *let 2 0 *inz z
<1.1 0 *K . 2 9 RT 10 EE.E
. .am.
I' I
F E-F T
k *l 1' i ;
i
! a i
a r7, 57, veggy garrLE wa l$ m' .
3
.f k -
Fi g . 16. Contours of minimum principal stress in the concrete for the loop divider baffle hot spot
~
v .,w... .
I
. . Pipe Centerline . .
d- 2 0
.: h = 150 to 200 Btu /hr/f t F T 650 to 750 F ,
9 e'o :
- ~
~
.:- r ,
I[
r, . ;
a :
,.4- - = - ,
' - .- .- b dk
- O
%s 4-t
.e a.
o .
u ,
m .
~
e .
- a.
Adiabatic g f
)
f ,
J. .
d
- d. d T" .
L lt SL Ik L & L F i g . 17. Sectional model of the helium purification crossover train pipe.
1
'-----^-=a.~-..m. . . _ , ,_ __.
e 6 9 l
I lll
, l. l' f i . .
t l !
! i ili i r.
a l, no ti iii i
f I
i l i i .'
' l 1 li r
1 k' +ll.
il l' i N N
- :: . r N i I
l j i*(
! l ! \ ,
E. , G, o.
Y, !
- i. i i -1, ME_!W TRelW
[f
'a L
P Fi g . 18. Finite element mesh for the helium purification Cr0SSover train,
s .
?
lI Eu
' (
1: !)
- ;;2 i l q : :: 9 /
I
- == ai i /
ii m .=
- LJ g
.,a
\ !
. ti: s 'N .
,'\si-\ \
]p i i \ s\
IN\\\ ' i
\
! ss, N'g g 'k1 I
( $\.g \ i ,<3 :
i_t '
) HEllUM TR9fN 9
h Fi g . 19. Contours of temperature in the helium purification crossover train hot spot. -
- - - - - -- ' - - - - - - -m c. w,,- ycw - * - - - , - - - - - - . - -, .- --
l 1
I l
l l
l l
n l
f m, S-i = =
l =tHE.5
=== b i a -es W l 3 *u.m 4 *M.a
.l
- ltE W
-- f r == ]
t J
\ ti l
I '
1 .
(
k O
i- '
l 8 l h il I i 8.
( HCLlUt1 TRAIN g e- -i v
d l I l
Fig. 20. Contours of maximum principal stress in the concrete for the helium purification crossover train hot spot.
i l
t i ~ -. -
-e s..te.
m
.. se~ t ceg .
t =fME .E 2 -eta E p 3 m.E 4 *M.2 5 *!ts :D !
e Ee p r is.=
id-i-
'I
- . . i a
I 9
s b
j 8
~
4 '
h l
B j j s I.
, HCLILM TRAIN 3
se e n===r i e
W d
I Fig. 21. Contours of minimum principal stress in the concrete l for the helium purification crossover train hot spot.
l l
l
- --. ,.,.,._ , --, , . . - . . - - . - - , , . , . _ , . - , , . - - ~ ~ , , , , - . , - _ _ _ .
/
Adiabatic - ?
e b
15.1 I
l i
h = 10 ETV/hr ft2 'F Fig. 22. Sectional model of a core outlet thermocouple penetration.
O ,
o i
I l
l l illlI !!l
!- ! iiiii lll 4-
! ! l i!!lllll. l i: l I'iil iii
!! I !!'ll ll !!
! i! ; !" i I: i l li i I!! ! l !i g
! !' idl'll!! !
! 11 iil l til l
,! li i I!i l l i I i l -
I ! li!iI I i I l
h ! t i:1 i ' II'j j li i l'i i i i i ! !
P ff li i i:' ! ! i i ! i 19 ! l 1: Ii i !lI -
2 . ; i .: .
't I -
W;, . .
g i.
e-
- t j Tm, in I
e.
- i. l
.- . r.
o--
. . ~ .
,e e. p,_.;.
- o::re.
.s..
.r
, I a
- f. l 7 f Fig. 23. Finite element mesh for the thermocouple penetration.
i
, , , --ny .,-w--. - , - , - - , -~~-r--e +* rvn"
'"v"" **"4'" M * "*""~~ "' ~
1 I
W, l t
s.s. u
! IE.2 a : ..x ,
j 3 Jac.x .
t 3 .* .2
- 1 E.2 I e ex , ?
9 SE.E i i e w.x -
1 I i mz e 8
. 'z.x 2
, .. se.x l 'l l l i
- 1
- l i l
\
1 /
f '
a I l
r .4 \v[aa
' ~~%us:
f.l.M f
'I
'l8 s e -
(=ct* -
. i i fg -ro -- , 2 e pggg7r::[g:;
re- .. e .
e l m
f" l+.
5 I
Fig. 24 Contours of temperature in the thermocouple hot spot.
1 l
l
_ _ _ _ _ _ . . . _ - - _ _ _ _ _ _ _ _ - _ . ._ ._ _ .___ _ 1
O,
- 5 i
1 I
=s. = m seem e 4
!, . =I. l i
- M. W
. . I 1 e -en.m I r au ,
I 3
I
- I I I I
J i 5
=
5
, l i h#
s /
M.D I ,
,E TNF C PEN TRATION B
W m
5 Fig. 25. Contours of maximum principal stress in the concrete for the thermocouple penetration hot spot.
r~ '
3
. o, .
F i
1
, 1 I
meL set asvut m j iA m,4 5 N '
q I
$.~ E
$ .tB E d s - '
r uma , , i
, < i 1
l
, l i
f g <
4 , i l n
[ , 1
- s. .
q l 11 Ni i i I
- d THCRM000JpLC PCNCTRATION a l
-i i 3
l I
Fig. 26. Contours of minimum principal stress in the concrete for the thermocouple penetration hot spot.
. - - - - . . . . . ~ . . - - - , - - - - . - , - - - . -
.- , - - - . . . - - . . , - - . . . , - - - - - - - - . , - - . - - - - , - - ~ , . - . , --
1 I
i HEY EVERYBODY! THANKSGIVING IS NEXT WEEK! I l
Following Region IV tradition, we will be celebrating on Wedne Jay, !
November 23, 1983, in the conference room. You are all invited to bring some type of delicious food to the conference room for everyone else to share in. If it is a dish that needs heating or needs to be kept cold, the best time to take it to the conference room would be about 11:15 because many of us will eat our lunch there that day.
Other foods may be placed there first thing in the morning and devoured by all participants all day long at their desks. Volunteers are needed to bring paper plates and napkins and plastic cups and utensils.
Bringing these supplies will exempt you from bringing food. Please let Denise (Ext. 279) know if you plan to bring some of the sup? lies.
HAPPY, HAPPY TURKEY DAY TO EVERYONE!!
e g .0 L