ML21278A133
Text
CALVERT CLIFFS UFSAR 5.1-1 Rev. 47 STRUCTURES 5.1 CONTAINMENT STRUCTURE 5.1.1 DESIGN BASIS General plans at various elevations and sections through the Containment Structure interior are shown in Figures 1-6 through 1-15 and show the general arrangement of various equipment such as reactor, steam generators, pressurizer and reactor coolant pumps. The support and anchorage details of these components are shown in Figures 5-11 through 5-15. At an Elevation 177'0", a landing platform, to the ladder on the vent stack provides an access route to the top of the Containment Structure. A steel ladder, provided inside the Containment Structure provides an access up to the polar crane. The Containment Structure is a Seismic Category I structure and is designed for all loading combinations described in Section 5A.3.
The Containment Structure completely encloses the Reactor Coolant System (RCS) to minimize release of radioactive material to the environment should a serious failure of the RCS occur. The structure provides adequate biological shielding for both normal operation and accident situations. The Containment Structure is designed for maximum of 0.16%/day leakage by weight of the original content of air at a design pressure of 50 psig and a concrete surface temperature of 276F.
The principal design basis for the structure is that it be capable of withstanding the internal pressure resulting from a loss-of-coolant accident (LOCA) with no loss of integrity. In this event, the total energy contained in the water of the RCS is assumed to be released into the Containment Structure through a break in the reactor coolant piping. Subsequent pressure behavior is determined by the building volume, engineered safety features, and the combined influence of energy sources and heat sinks.
Energy is available for release into the Containment Structure from the following sources:
RCS Stored Heat Reactor Stored Heat Reactor Decay Heat Metal-Water Reactions The energy release and the containment pressure transient curves are shown in Section 14.20.
The design of the engineered safety features systems and their operation is discussed more fully in Chapter 6; only their relation to the basis of Containment Structure design is discussed below. The engineered safety features systems are provided to limit the consequences of an accident. Their energy removal capabilities limit the internal pressure so that Containment Structure design limits are not exceeded and the potential for release of fission products is minimized.
The safety injection systems inject borated water into the reactor vessel to remove core decay heat and to minimize metal-water reactions and the associated release of heat and fission products. Flashed primary coolant, RCS sensible heat, and core decay heat transferred to the Containment Structure are removed by the Containment Cooling System which is comprised of two subsystems; the containment spray subsystem and the air recirculation subsystem.
The containment spray subsystem reduces pressure in the containment by condensing the Containment Structure steam and removing heat from the containment atmosphere by recirculation of the spray water through the shutdown cooling heat exchangers.
CALVERT CLIFFS UFSAR 5.1-2 Rev. 47 The air recirculation subsystem reduces pressure and removes heat directly from the Containment Structure atmosphere to the Service Water System with recirculating fans and cooling coils.
5.1.2 DESIGN CRITERIA 5.1.2.1 General Description The Containment Structure houses the RCS. Its purpose is to contain any accidental release of radioactivity from the RCS. It is designated as a Seismic Category I structure.
The basic design criteria are that the integrity of the liner plate be guaranteed under all loading conditions and the structure shall have a low-strain elastic response such that its behavior will be predictable under all design loadings.
The structure consists of a post-tensioned reinforced concrete cylinder and dome connected to and supported by a massive reinforced concrete foundation slab as shown in Figure 5-1. The entire interior surface of the structure was lined with a 1/4" thick welded American Society for Testing and Materials (ASTM) A36 steel plate to assure a high degree of leak tightness. Numerous mechanical and electrical systems penetrate the Containment Structure wall through welded steel penetrations as shown in Figures 5-2 and 5-3. The penetrations and access openings were designed, fabricated, inspected, and installed in accordance with the American Society of Mechanical Engineers (ASME), Boiler and Pressure Vessel (B&PV) Code,Section III, Class B.
Principal dimensions of the Containment Structure are:
Inside Diameter 130' Inside Height (including Dome) 181-2/3' Vertical Wall Thickness 3-3/4' Dome thickness 3-1/4' Foundation Slab Thickness 10' Liner Plate Thickness 1/4" Internal Free Volume 1,989,000 ft3 The Containment Structure is shown in Figures 1-6 through 1-16.
In the concept of a post-tensioned Containment Structure, the internal pressure load is balanced by the application of an opposing external force on the structure.
Sufficient post-tensioning was used on the cylinder and dome to more than balance the internal pressure so that a margin of external pressure exists beyond that required to resist the LOCA pressure. See Appendix 5E for an evaluation that reduced the original containment minimum design prestress. Nominal, bonded reinforcing steel was also provided to distribute strains due to shrinkage and temperature. Additional bonded reinforcing steel was used at penetrations and discontinuities to resist local moments and shears.
The internal pressure loads on the foundation slab are resisted by both the external bearing pressure due to dead load and the strength of the reinforced concrete slab. Thus, post-tensioning was not required to exert an external pressure for this portion of the structure.
CALVERT CLIFFS UFSAR 5.1-3 Rev. 47 The post-tensioning system consists of:
- a.
Three groups of 68 dome tendons oriented at 60 to each other for a total of 204 tendons anchored at the vertical face of the dome ring girder.
- b.
Two hundred four vertical tendons anchored at the top surface of the ring girder and at the bottom of the base slab.
- c.
Six groups of 78 hoop tendons, each enclosing 120 of arc, for a total of 468 tendons anchored at the 6 vertical buttresses.
Each tendon consists of approximately 90 1/4" diameter wires with button-headed BBRV-type anchorages, furnished by the Prescon Corporation. The tendons are housed in spiral wrapped, corrugated, thin-wall, carbon steel sheathing. After fabrication, the tendon was shop dipped in a petrolatum corrosion protection material, bagged and shipped. After installation, the tendon sheathing was filled with a corrosion preventive grease (Viconorust 2090P). The ends of all tendons were covered with pressure-tight, grease-filled caps for corrosion protection. All the vertical tendons for each Unit received new corrosion preventive grease between 1997 and the end of 2002. Some original vertical tendons for each Unit were restressed or replaced with a new tendon between 2001 and 2002. See Appendix 5E for details.
American Society for Testing and Materials A615, Grade 60 reinforcing steel, mechanically spliced as needed with B-and T-series CADWELDS, was used throughout the foundation slab and around the large penetrations. The same type of steel was used for the bonded reinforcing throughout the cylinder and dome as crack control reinforcing and at areas of discontinuities to provide an additional margin of elastic strain capability.
The 1/4"-thick liner plate was attached to the concrete by means of an angle grid system stitch welded to the liner plate and embedded in the concrete. The details of the anchoring system are provided in Figure 5-1. The frequent anchoring is designed to prevent significant distortion of the liner plate during accident conditions and to insure that the liner maintains its leak-tight integrity. The design of the liner anchoring system also considers the various erection tolerances and their effect on its performance. The liner plate was protected from corrosion on the inside with 3 mils of inorganic zinc primer topped with 6 mils of an organic epoxy up to Elevation 75'0", and 3 mils of an inorganic topcoat above that elevation.
There is no paint on the side in contact with concrete.
The aggregate used in the structure produced an excellent high-strength, dense, sound concrete. The 28-day design strengths were 5000 psi for the shell and 4000 psi for the foundation slab.
Personnel and equipment access to the structure is provided by a two-door personnel lock with double seals on both doors and by a 19'0" clear diameter, double gasketed, single-door equipment hatch as shown in Figure 5-3. A two-door emergency personnel escape lock is also provided. These locks and hatch were designed and fabricated from A516, Grade 70 firebox quality steel made to A300 specification and Charpy V-notch impact tested to 0F in accordance with the ASME, B&PV Code,Section III. All piping penetrations furnished adhered to the same requirements.
A containment outage door on the exterior of each Containment Structure at the equipment hatch opening serves as a substitute for the equipment hatch when setting containment closure during Modes 5 and 6 conditions. The containment
CALVERT CLIFFS UFSAR 5.1-4 Rev. 47 outage door was designed, fabricated, examined, inspected and tested to the requirements of the 1995 Edition of ASME Section VIII, Rules for Construction of Pressure Vessels, Division 1. However, the containment outage door cannot be credited for severe weather. If Emergency Response Plan Implementation Procedure 3.0, Preparing for Severe Weather, is implemented and requires containment closure to be established, then the equipment hatch must be used.
Structural brackets provided for the Containment Structure polar crane runway were fabricated from ASTM A36 steel shapes and ASTM A516, Grade 70 insert plates (Figure 5-1). Like the penetration assemblies, structural brackets and thickened plates were shop fabricated, stress relieved and shipped to the job site for welding to the 1/4" liner plate.
The strength of the Containment Structure at working stress and overall yielding was compared to various loading combinations to assure safety. The Containment Structure was examined with respect to strength, the nature and the amount of cracking, the magnitude of deformation, and the extent of corrosion to assure proper performance. The structure was designed and constructed in accordance with design criteria based upon American Concrete Institute (ACI) 318-63, ACI 301-66, and ASME, B&PV Code, Sections III, VIII, and IX to meet the performance and strength requirements prior to prestressing, at transfer of prestress, under sustained prestress, at design loads, and at yield loads.
The structure was originally analyzed using Bechtel's Finite Element Program for Cracking Analysis CE 316-4, for individual and various combinations of loading cases of dead load, live load, prestress, temperature and pressure. The computer output included direct stresses, shear stresses, principal stresses and displacements of each nodal point. See Appendix 5E for an evaluation that reduced the original containment minimum design prestress.
Stress plots which showed the total stresses from appropriate combinations of loading cases were made and areas of high stress were identified. The modulus of elasticity was corrected to account for the nonlinear stress-strain relationship at high compression in these areas and stresses were recomputed.
In order to consider creep deformation, the modulus of elasticity of concrete under sustained loads such as dead load and prestress was differentiated from the modulus of elasticity of concrete under instantaneous loads such as internal pressure and earthquake loads.
The forces and shears were added over the cross-section and the total moment, axial force and shear were determined. From these values, the straight-line elastic stresses were computed and compared to the allowable values. The ACI 318-63 design methods and allowable stresses were used for concrete and prestressed and unprestressed reinforcing steel except as noted in the design criteria.
It is the intent of the criteria to provide a structure of unquestionable integrity that will meet the postulated design conditions with a low strain elastic response. The Calvert Cliffs Containment Structure meets these criteria because: (See Appendix 5E for an evaluation that reduced the original containment minimum design prestress.)
CALVERT CLIFFS UFSAR 5.1-5 Rev. 47
- a.
The design criteria are in general based on the proven stress and strain to meet the ACI or ASME codes. Departures from or additions to these codes have been made in the following manner:
- 1.
The environmental conditions of severity of load cycling, weather, corrosion conditions, maintenance, and inspection for this structure have been compared and evaluated with those for code structures to determine the appropriateness of the modifications.
- 2.
The consultant firm of T.Y. Lin, Kulka, Yang and Associate was retained on earlier projects to assist in the development of design criteria. In addition to assisting with the criteria submitted in the Preliminary Safety Analysis Report, they were involved in the review of design methods to assure that the criteria were implemented as intended.
- 3.
Dr. Alan H. Mattock of the University of Washington was retained on earlier projects to assist in developing the proper design criteria for combined shear, bending and axial load.
- 4.
All criteria, specifications and details relating to liner plate and penetrations and corrosion protection have been referred to Bechtel's Metallurgy and Quality Control Department. This department maintains a staff to advise the corporation on problems of welding, quality control, metallurgy and corrosion protection.
- 5.
The design of the Calvert Cliffs Containment Structure was continually reviewed as the criteria were revised for successive license applications.
- b.
The primary membrane integrity of the structure is provided by the unbonded post-tensioning tendons, each one of which is stressed to 80%
of ultimate strength during installation and performs at approximately 50 to 60% during the life of the structure. Thus, the main strength elements are individually proof-tested prior to operation of the plant.
- c.
Eight-hundred-seventy-six such post-tensioning elements have been provided, 204 in the dome, and 204 vertical and 468 hoop tendons in the cylinder. Any three adjacent tendons in any of these groups can be lost without significantly affecting the strength of the structure due to the load redistribution capabilities of the shell structure. The bonded reinforcing steel provides for crack control assures that this redistribution capability exists.
- d.
The unbonded tendons are continuous from anchorage to anchorage, being deflected around penetrations and isolated from secondary strains of the shell. Thus, the membrane integrity of the shell can be assured regardless of conditions of high local strains.
- e.
The unbonded tendons exist in the structure at a slightly ever-decreasing stress due to relaxation of the tendon and creep of the concrete and, even during pressurization, are subject to a stress change of very small magnitude (2% to 3% of ultimate strength). Thus, the main structural system is never subjected to large changes in load, even during accident conditions.
- f.
The concrete portion of the structure, similar to the tendons, was subject to the highest state of stress during the initial post-tensioning. During pressurization, it is subject to a large change in load (or state of stress) but the change is, in general, a decrease in load. The large membrane compressive forces are diminished and/or replaced by relatively small radial pressures and stresses.
CALVERT CLIFFS UFSAR 5.1-6 Rev. 47
- g.
The deformations of the structure during plant operation, or due to accident conditions, are relatively minor due to the low-strain behavior of the concrete. The largest deformations occurred at the time of initial post-tensioning and shortly thereafter, prior to operation. This low strain behavior and the inherent strength of the structure permit the anchoring of all piping penetrating the structure directly to the shell. Such details (Figure 5-2) eliminate the use of expansion bellow seals and significantly reduce the likelihood of leaks developing at the penetrations. The exception to this is when the fuel transfer tube is in use, requiring use of the transfer tube bellows (Section 5.1.4.4.d).
5.1.2.2 Loads Prior to prestressing, the structure was designed as a conventionally reinforced concrete structure. It is designed for dead load, live loads and a reduced-wind load. Allowable stresses are computed in accordance with ACI 318-63.
Loads at Transfer of Prestress See Appendix 5E for an evaluation that reduced the original containment minimum design prestress.
The Containment Structure is checked for prestress loads and the stresses compared with those allowed by ACI 318-63 with the following exceptions:
ACI 318-63, Chapter 26, allows concrete stress of 0.60 at initial transfer. In order to limit creep deformations, the membrane compression stress is limited to 0.30 whereas in combination with flexural compression the maximum allowable stress will be limited to 0.60 per ACI 318-63.
For local stress concentrations with nonlinear stress distribution as predicted by the finite element analysis, 0.75 is permitted when local bonded reinforcing is included to distribute and control the localized strains. These high local stresses are present in every structure but they are seldom identified because of simplifications made in design analysis. These high stresses are allowed because they occur in a very small percentage of the cross-section, are confined by material at lower stress and would have to be considerably greater than the values allowed before significant local plastic yielding would result. Bonded reinforcing was added to distribute and control these local strains.
Membrane tension and flexural tension are permitted provided they do not jeopardize the integrity of the liner plate. Membrane tension is permitted to occur during the post-tensioning sequence but will be limited to
. When there is flexural tension but no membrane tension, the section is designed in accordance with the ACI Code, Section 2605(a). The stress in the liner plate due to combined membrane tension and flexural tension is limited to 0.5 fy.
Shear criteria are in accordance with the ACI 318-63 Code, Chapter 26, as modified by the equations in the structural yielding subsection of this section using a load factor of 1.5 for shear loads.
Loads Under Sustained Prestress See Appendix 5E for an evaluation that reduced the original containment minimum design prestress.
The conditions for design and the allowable stresses for this case are the same as above except that the allowable tensile stress in unprestressed reinforcing is
CALVERT CLIFFS UFSAR 5.1-7 Rev. 47 limited to 0.5 fy. ACI 318-63 limits the concrete compression to 0.45 for sustained prestress load. Values of 0.30 and 0.60 are used as described above, which bracket the ACI allowable value. However, with these same limits for concrete stress at transfer of prestress, the stresses under sustained load are reduced due to creep, shrinkage, relaxation, and possible tendon wire breakage.
See Appendix 5E for a discussion on possible tendon wire breakage.
At Design Loads This loading case is the basic "working stress" design. The Containment Structure is designed for the following loading cases:
- a.
D + F + L (Construction case)
- b.
D + F + L + To + E (Operating case)
- c.
D + F + L + P + TA (Design incident case)
- d.
D + F + L + 1.15P (Test case)
- e.
D + F + L + Ts + E (Prolonged shutdown case)
D
= Dead Load L
= Appropriate Live Load F
= Appropriate Prestressing Load P
= Design Pressure To = Thermal Loads Due to Operating Temperature TA = Thermal Loads Corresponding to Pressure P E
= Operating Basis Earthquake (OBE) of 0.08g Ts = Thermal Loads Due to Transient Wall Temperatures Over a Prolonged Shutdown (20F outside face, 50F inside face)
Sufficient prestressing is provided in the cylindrical and dome portions of the vessel to eliminate membrane tensile stress (tensile stress across the entire wall thickness) under design loads. Flexural tensile cracking is permitted but is controlled by bonded reinforcing steel.
Under the design loads, the same performance limits given for loads at transfer of prestress apply with the following exceptions:
- a.
If the net membrane compression is below 100 psi, it is neglected and a cracked section is assumed in the computation of flexural bonded reinforcing steel. The allowable tensile stress in bonded reinforcing is 0.5 fy.
- b.
When the maximum flexural stress does not exceed 6 and the extent of the tension zone is not more than 1/3 the depth of the section, bonded reinforcing steel is provided to carry the entire tension in the tension block.
Otherwise, the bonded reinforcing steel is designed assuming a cracked section. When the bending moment tension is additive to the thermal tension, the allowable tensile stress in the bonded reinforcing steel is 0.5 fy minus the stress in reinforcing due to the thermal gradient as determined in accordance with the method of ACI-505.
- c.
The problem of shear and diagonal tension in a prestressed concrete structure should be considered in two parts: membrane principal tension and flexural principal tension. Since sufficient prestressing is used to eliminate membrane tensile stress, membrane principal tension is not critical at design loads. Membrane principal tension due to combined
CALVERT CLIFFS UFSAR 5.1-8 Rev. 47 membrane tension and membrane shear is considered in the next subsection.
Flexural principal tension is the tension associated with bending in planes perpendicular to the surface of the shell and shear stress normal to the shell (radial shear stress). The present provisions of ACI 318-63, Chapter 26 for shear are adequate for design purposes with proper modifications as discussed in the next subsection using a load factor 1.5 for shear loads.
Crack control in the concrete is accomplished by adhering to the ACI-ASCE Code Committee standards for the use of reinforcing steel. These criteria are based upon a recommendation of the Prestressed Concrete Institute and are as follows:
0.25 percent reinforcing shall be provided at the tension face for small members 0.20 percent for medium size members 0.15 percent for large members A minimum of 0.20% bonded steel reinforcing is provided in two perpendicular directions on the exterior faces of the wall and dome for proper crack control.
The liner plate is attached on the inside faces of the wall and dome. Since, in general, there is no tensile stress due to temperature on the inside faces, bonded reinforcing steel is not necessary there.
Loads Necessary to Cause Structural Yielding The structure is checked for the factored loads and load combinations that will cause structural yielding.
The load factors are the ratio by which loads will be multiplied for design purposes to assure that the load/deformation behavior of the structure is one of elastic, low-strain behavior. The load factor approach was used in this design as a means of making a rational evaluation of the isolated factors which must be considered in assuring an adequate safety margin for the structure. This approach permits the designer to place the greatest conservatism on those loads most subject to variation and which most directly control the overall safety of the structure. It also places minimum emphasis on the fixed gravity loads and maximum emphasis on accident and earthquake or wind loads. The final design of the structure satisfies the load combinations and factors shown in Appendix 5A.
The load combinations, considering load factors referenced above, are less than the yield strength of the structure. The yield strength of the structure is defined as the upper limit of elastic behavior of the effective load carrying structural materials.
For steels (both prestressed and unprestressed) this limit is taken to be the guaranteed minimum yield given in the appropriate ASTM specification. For concrete, it is the ultimate values of shear (as a measure of diagonal tension) and bond per ACI 318-63 and the 28-day ultimate compressive strength for concrete in flexure (). The ultimate strength assumptions of the ACI Code for concrete beams in flexure are not allowed; that is, the concrete stress is not allowed to go beyond yield.
The maximum strain due to secondary moments, membrane loads and local loads exclusive of thermal loads is limited to that corresponding to the ultimate stress divided by the modulus of elasticity (l) and a straight-line distribution from there to the neutral axis assumed.
CALVERT CLIFFS UFSAR 5.1-9 Rev. 47 For the loads combined with thermal loads the peak strain is limited to 0.003 in./in.
For concrete membrane compression, the yield strength is assumed to be 0.85 to allow for local irregularities, in accordance with the ACI approach. The reinforcing steel forming part of the load carrying system is allowed to go to, but not to exceed, yield as is allowed for ACI ultimate strength design.
A further definition of yielding is the deformation of the structure which causes strains in the steel liner plate to exceed 0.005 in./in. The yielding of unprestressed reinforcing steel is allowed, either in tension or compression, if the above restrictions are not violated. Yielding of the prestressed tendons is not allowed under any circumstances.
Principal concrete tension due to combined membrane tension and membrane shear, excluding flexural tension due to bending moments or thermal gradients, is limited to 3. Principal concrete tension due to combined membrane tension, membrane shear, and flexural tension due to bending moments or thermal gradients is limited to 6. When the principal concrete tension exceeds the limit of 6, bonded reinforcing steel is provided in the following manner:
- a.
Thermal Flexural Tension - Bonded reinforcing steel is provided in accordance with the methods of ACI-505. The minimum area of steel provided is 0.20% in each direction.
- b.
Bending Moment Tension - Sufficient bonded reinforcing steel is provided to resist the bending moment on the basis of cracked section theory using the yield stresses stated above with the following exception: When the bending moment tension is additive to the thermal tension, the allowable tensile stress in the reinforcing steel is fy minus the stress in reinforcing due to the thermal gradient, as determined in accordance with the methods of ACI-505.
Shear stress limits and shear reinforcing for radial shear are in accordance with ACI 318-63, Chapter 26 with the following exceptions:
Formula 26-12 of the Code was replaced by:
= + [
1] + 1 where
= [1.75 0.036
+ 4.0]
but not less than 0.6 for p' 0.003.
For p' >0.003, the value of K shall be zero.
=
[6 + + + ]
fpe = Compressive stress in concrete due to prestress applied normal to the cross-section after all losses (including the stress due to any secondary moment) at the extreme fiber of the section at which tension stresses are caused by live loads.
fn
= Stress due to axial applied loads (fn shall be negative for tension stress and positive for compression stress).
CALVERT CLIFFS UFSAR 5.1-10 Rev. 47 fi
= Stress due to initial loads at the extreme fiber of a section at which tension stresses are caused by applied loads including the stress due to any secondary moment (fi shall be negative for tension stress and positive for compression stress).
n
=
475
, constant in value of K above p' =
ratio of compression steel area to area concrete V
= Shear at the section under consideration due to the applied loads.
M' = Moment at a distance d/2 from the section under consideration, measured in the direction of decreasing moment, due to applied loads.
Vi = Shear due to initial loads (positive when initial shear is in the same direction as the shear due to applied loads).
The lower limit placed by ACI 318-63 on Vci of 1.7 b'd is not applied.
Formula 26 -13 of the Code was replaced by:
= 3.5()1/2 [1 + +
3.5 ]
1/2
+
Where Vp = radial shear component of effective prestress due to tendon curvature at the section considered, and the term fn is as defined above. All other notations are in accordance with ACI 318-63, Chapter 26. It should be noted that this formula is based on the tests and work done by Dr. A. H. Mattock of the University of Washington, and has been included in ACI 318-77, Section 11.5.2.
When the above-mentioned equations show that allowable shear in concrete is zero, radial horizontal shear ties are provided to resist all the calculated shear.
Other Design Loads The Containment Structure shell is designed for the following loads:
- a.
Dead load
- b.
Prestress forces
- c.
Live load including allowances for piping, ductwork and cable trays
- d.
Wind, including tornado
- e.
- f.
Thermal expansion of pipes attached to the Containment Structure wall
- g.
Uplift due to buoyant forces Transients resulting from the LOCA and other lesser incidents are presented in Chapter 14 and serve as the basis for the Containment Structure design pressure of 50 psig and a design concrete surface temperature of 276F.
The external design pressure of the Containment Structure shell is 3 psig. This value is approximately 0.5 psig beyond the maximum external pressure that could be developed if the Containment Structure were sealed during a period of low barometric pressure and high temperature and, subsequently, the Containment Structure atmosphere was cooled with a concurrent rise in barometric pressure.
Vacuum breakers are not provided.
CALVERT CLIFFS UFSAR 5.1-11 Rev. 47 5.1.2.3 Equipment Supports
- a.
Reactor Vessel Supports
- 1.
Restrain the vessel to maintain the integrity of emergency core cooling systems and to prevent the rupture of additional primary pipes should LOCA occur due to single pipe rupture;
- 2.
Permit slow radial thermal expansion of the vessel under normal operation; and
- 3.
Restrain the vessel against seismic and LOCA jet forces.
- b.
Steam Generator Supports
- 1.
Restrain the vessel to prevent simultaneous rupture of the primary coolant pipe, and the steam or the feedwater pipes;
- 2.
Permit slow thermal growth of the loops and the vessel; and
- 3.
Restrain all motion under seismic or LOCA loads.
Calvert Cliffs Nuclear Power Plant Units 1 and 2 are approved for leak-before-break based on References 6 and 8, and compliance with Regulatory Guide 1.45 for leak detection as documented in UFSAR Section 4.3.1. As a result of the application of leak-before-break, the mechanical/structural loads associated with the dynamic effects of a large break LOCA in the RCS hot leg or cold legs are no longer considered part of the plant design basis. Accordingly, leak-before-break is credited in the design of the steam generator sliding base supports for the replacement steam generators.
- c.
Pressurizer Support
- 1.
Support normal operating loads; and
- 2.
Restrain the vessel under seismic loads.
- d.
Main Coolant Pumps Supports
- 1.
Permit slow thermal movements of the pump; and
- 2.
Restrain the pump under seismic loads.
- e.
Safety Injection Tank Support
- 1.
Support normal operating loads; and
- 2.
Restrain the tank under seismic loads.
Materials for Equipment Supports
- a.
Reactor Vessel
- 1.
Plate material ASTM A302, Gr B
- 2.
Structural shapes ASTM A441
- 3.
- 4.
Welding electrodes ASTM A233, E 7018
- b.
- 1.
Plate material ASTM A302, Gr B
- 2.
Anchor bolts ASTM A490
- 3.
Welding electrodes ASTM A233, E 7018
CALVERT CLIFFS UFSAR 5.1-12 Rev. 47
- c.
Pressurizer
- 1.
- d.
Main Coolant Pump
- 1.
Structural shapes ASTM A36
- 2.
Machine bolts ASTM A325
- 3.
Steel plate (S. hangers)
ASTM A515, Gr 65
- 4.
Pipe section (S. hangers)
ASTM A53
- 5.
Machined rod (S. hangers)
AISI C-1015
- 6.
Spring coils (S. hangers)
AISI 4161 H
- e.
Safety Injection Tank
- 1.
Anchor bolts ASTM A354, Gr BC Allowable Stresses and Strain for Equipment Supports Excluding Supports for Piping and Vessels:
- a.
Stresses under normal operating loads and the OBE are within the applicable code allowable limits.
- b.
Stresses under combined normal, Safe Shutdown Earthquake (SSE) and accident loads are kept below the yield strength of the material.
- c.
Allowable local and average strains are determined by the maximum permissible deformation for the support.
5.1.3 CONTAINMENT STRUCTURE DESIGN ANALYSIS The analysis for the Containment Structure shell falls into two parts, axisymmetric and non-axisymmetric. The axisymmetric analysis is performed through the use of a finite element computer program for the individual loading cases of dead load, live load, temperature, prestress and pressure, as described in Section 5.1.3.1. The axisymmetric finite element representation of the Containment Structure shell does not include the buttresses, penetrations, brackets, and anchors. These items of configuration, the lateral loads due to earthquakes or winds, and various concentrated loads are considered in the non-axisymmetric analysis described in Section 5.1.3.2.
This section includes only analytical techniques, references and design philosophy. The results of these analyses are shown in Section 5.1.4. The design criteria and analysis have been reviewed by Bechtel's consultants, T.Y. Lin, Kulka, Yang and Associate.
5.1.3.1 Axisymmetric Techniques The finite element technique is a general method of structural analysis in which the continuous structure is replaced by a system of elements (members) connected at a finite number of nodal points (joints). Conventional analysis of frames and trusses can be considered an example of the finite element method. In the application of the method to an axisymmetric solid structure such as the Containment Structure shell, the continuous structure is replaced by a system of rings of quadrilateral cross-section which are interconnected along circumferential joints. Based on energy principles, sets of work equilibrium equations are formed in which the radial and axial displacements at the circumferential joints are the
CALVERT CLIFFS UFSAR 5.1-13 Rev. 47 unknowns. The results of the solution of this set of equations are the deformations of the structure under the given loading conditions. For the output, the stresses are computed knowing the strain and stiffness of each element.
The original finite element mesh used to describe the structure is shown in Figure 5-4 (see Sheets 1 and 2). See Appendix 5E for an evaluation that reduced the original containment minimum design prestress. The upper and lower portions of the structure were analyzed independently to permit the use of a greater number of elements for those areas of the structure of major concern, e.g., the ring girder area and the base of the cylinder. The finite element mesh of the base slab was extended down into the foundation to take into consideration the elastic nature of the foundation material and its effect upon the behavior of the base slab. The tendon access gallery was designed as a separate structure.
The finite element mesh for the Containment Structure does not include the interior structures. The interior structures were included in the finite element input as a lumped mass. The finite element analysis produces stresses due to axisymmetric loads. The stresses from interior structure loads and earthquake loads are superimposed on the finite element stresses. The final summation of all stresses was used to design the base slab, exterior shell and dome. The use of Bechtel's finite element computer program, CE 316-4, permitted an accurate estimate of the stress pattern at various locations of the structure. The major benefit of the program is the capability to predict shears, normal forces and moments due to internal restraint and the interaction of the foundation base slab with the subgrade.
The forces and moments were applied to all directions. The following material properties were used in the program for the various loading conditions:
Econcrete' Foundation 3.64x106 psi Econcrete' Shell 4.07x106 psi concrete (Poisson's Ratio) 0.17 concrete (Coefficient of Expansion) 0.55x10-5 Eliner 29x106 psi Fy liner 34,000 psi Esoil (Construction and Operating Case, Figure 5-4, Sheet 1) 1st Layer E = 6,200 psi 2nd Layer E = 9,600 psi 3rd Layer E = 12,000 psi Esoil (Testing and Accident Case) 1st Layer E = 9,600 psi 2nd Layer E = 14,400 psi 3rd Layer E = 18,000 psi Esoil (Factored Load-Yield Stress) 1st Layer E = 10,000 psi 2nd Layer E = 20,000 psi 3rd Layer E = 30,000 psi
CALVERT CLIFFS UFSAR 5.1-14 Rev. 47 In arriving at the above-tabulated values of E, the effect of creep is included by using the following equation for long-term loads such as thermal load, dead load and prestress:
Ecs = Eci [i/(s + i)],
where:
Ecs
= sustained modulus of elasticity of concrete, Eci
= instantaneous modulus of elasticity of concrete, i
= instantaneous strain, in./in. per psi, and s
= creep strain, in./in. per psi The thermal gradients used in the design are shown in Figure 5-5. The design pressure and concrete surface temperature of 50 psig and 276F became 75 psig and 276F at factored conditions.
The compressive stress and strain level is the highest (after the LOCA when the temperature is still relatively high, 200F, and the pressure is dropping rapidly) at the inside face of the concrete at the edge of openings and also near the liner plate anchors. Neither concentration is a result of what may be considered a real load. In the case of an opening, the real stress is a result of prestress, reduced pressure and dead load. Applying stress concentration factors to these stresses maintains the concrete stress essentially in the elastic range. When the strain and resulting stress from the thermal gradient are also multiplied by a stress concentration factor, the total strain and resulting stress will be above the linear stress range determined by a uniaxial compression test. The relatively high stress level is not of real concern due to the following:
- a.
The concrete affected is completely surrounded by either other concrete or the penetration nozzle and liner reinforcing plate. This confinement puts the concrete in triaxial compression and gives it the ability to resist forces far in excess of that indicated by a uniaxial compression test.
- b.
The high state of stress and strain exist at a very localized area and have no effect on the overall containment integrity.
However, to be conservative, reinforcing steel was placed in these areas. The penetration nozzle will also function as compressive reinforcement.
The concrete under the liner plate anchors has some limited yielding in order to get the necessary stress distribution required to resist the liner plate self-relieving loads.
By criteria, yielding was only permitted in the design of the liner plate and of the bonded reinforcing steel for the Containment Structure. Subsequent design analyses, as tabulated in Table 5-1, indicate that the stresses in the liner plate and bonded reinforcing steel will not exceed the allowable yield stresses.
The thermal loads are a result of the temperature gradient across the structure wall. In the finite element analysis, when temperatures are given at every nodal point, stresses are obtained at the center of each element.
The liner plate was handled as an integral part of the structure and was included in the finite element mesh of the Containment Structure, but having different material properties (Figure 5-4, Sheet 1).
CALVERT CLIFFS UFSAR 5.1-15 Rev. 47 Under the LOCA condition or factored load condition, cracking of the concrete at the outside face would be expected. The value of the sustained modulus of elasticity of concrete, Ecs, was used in ACI Code 505-54 to find the stresses in concrete, reinforcing steel and liner plate from the predicted design incident thermal loads and factored incident loads.
The method of determining stresses in the concrete and reinforcement required the evaluation of the stress blocks of the cross-section being analyzed.
Stress values were taken from the computer output in the case of axisymmetric loading and from analytical solutions in case of non-axisymmetric loading. Both computations were based on homogeneous materials; therefore, some adjustment was necessary to evaluate the true stress-strain conditions when cracks develop in the tensile zone of the concrete.
An equilibrium equation was written considering the tension force in the reinforcement, the compressive force in the concrete and the axial force acting on the section. In this manner, the neutral axis was shifted from the position defined by the computer analyses to a position which is a function of the amount of reinforcement, the modulus ratio, and the acting axial forces.
The thermal stresses in the containment wall are comparable to those developed in a reinforced concrete slab which is restrained from rotation. The temperature varies linearly across the slab. The concrete will crack in tension and the neutral axis will be shifted toward the compressive extreme fiber. The cracking will reduce the compression at the extreme fiber and increase the tensile stress in reinforcing steel.
The following analysis is based on the equilibrium of normal forces; therefore, any normal force acting on the section must be added to the normal forces resulting from the stress diagram. The effects of Poisson's ratio are considered assuming the reinforcement to be identical in both directions.
Stress-strain relationship in compressed region of concrete:
Ec x = x - vc y (1)
Ec y = - vc x+ y (2)
From the above equations (1) and (2):
=
+
12 (3)
=
+
12 (4)
Substituting, x = y = c and x = y = c into equations (3) and (4)
=
1 1= 1.205 (if vc = 0.17)
The reinforcement is acting in one direction, independently from the reinforcement in the perpendicular direction.
CALVERT CLIFFS UFSAR 5.1-16 Rev. 47 Example:
If Ec = 4.07x106 and Es = 29x106
=
29 1.205 4.07 = 5.9 The liner plate is acting in two directions, similar to the concrete except for the difference caused by the Poisson's ratios and elastic modulus:
=
1 1
= 1.35, = 0.26
=
1.35 29 1.205 4.07 = 7.98, = 0.17 The concrete and reinforcement stresses, due to a moment caused by a loading other than thermal, are calculated by conventional methods. The analysis assumes homogeneous concrete sections. Those concrete and reinforcing steel stresses are then added to the thermal stresses as obtained by the method described above.
Notation:
Ec Modulus of elasticity of concrete.
Es Modulus of elasticity of steel.
nL Modular ratio of liner plate-concrete.
nR Modular ratio of reinforcement-concrete.
c Concrete strain.
s Steel strain.
x Concrete strain in the X direction.
y Concrete strain in the Y direction.
vc Poisson's ratio of concrete.
vL Poisson's ratio of liner plate.
c Stress in concrete.
L Stress in liner plate.
x Stress in concrete in the X direction.
y Stress in concrete in the Y direction.
5.1.3.2 Non-axisymmetric Techniques The non-axisymmetric aspects of configuration or loading required various methods of analysis. The descriptions of the methods used, as applied to different parts of the containment, are given below.
Buttresses The buttresses and tendon anchorage zones are defined as Seismic Category I elements and were designed in accordance with the general design criteria for the Containment Structure and with the applicable provisions of ACI 318-63, Chapter 26.
The buttresses were analyzed for two effects, non-axisymmetric and anchorage zone stresses.
At each buttress, two out of three hoop tendons are spliced by being mutually anchored on the opposite faces of the buttresses; the third tendon is continuous through the buttress. (The anchors are located on 21-3/4" centers, and are
CALVERT CLIFFS UFSAR 5.1-17 Rev. 47 staggered every 7-1/4" to the opposite face of the buttress. Combined with the continuous tendon, this results in the hoop tendons being positioned at every 7-1/4" along the vertical cross-section of the wall.) Between the opposite anchorages, the compressive force exerted by the spliced tendon is twice as much as elsewhere. This value, combined with the effect of the tendon which is not spliced, will be 1.5 times the prestressing force acting outside of the buttresses.
The cross-sectional area at the buttress is about 1.5 times that of the wall, thus the hoop stresses, as well as the hoop strains and radial displacements, can be considered as being nearly constant all around the structure.
The vertical stresses and strains, caused by the vertical post-tensioning, become constant at a short distance away from the anchorages because of the stiffness of the cylindrical shell. The stresses and strains remain nearly axisymmetric despite the presence of the buttresses. The effect of the buttresses on the overall analysis is negligible when the structure is under dead load or prestressing loads.
When an increasing internal pressure acts upon the structure, combined with a thermal gradient such as at the design incident condition, the resultant forces are axisymmetric. The stiffness variation caused by the buttresses will decrease as the concrete develops cracks. The structure will then tend to shape itself to follow the direction of the acting axisymmetric resultant forces even more closely. Thus, the buttress effect is more axisymmetric at yield loads (which include factored pressure) than at design loads including pressure. This fact, combined with the design provision that alternate horizontal tendons terminate in a single buttress, indicates that the buttresses will not reduce the margins of safety available in the structure.
The analysis of the anchorage zone stresses at the buttresses has been determined to be the most critical of all the various types of anchorage areas of the shell. The local stress distribution in the immediate vicinity of the bearing plates has been derived by the following two analysis procedures:
- a.
The Guyon equivalent prism method: This method is based on experimental photoelastic results as well as on equilibrium considerations of homogeneous and continuous media. It should be noted that the relative bearing plate dimensions are considered.
- b.
In order to include biaxial stress effects, use has been made of the experimental test results presented by S. J. Taylor at the March 1967 London Conference of the Institution of Civil Engineers (Group H, Paper 49). This paper compares test results with most of the currently used approaches (such as Guyon's equivalent prism method). It also investigates the effect of the rigid trumpet welded to the bearing plate.
The Guyon method yields the following results:
Maximum compressive stress under the bearing plate, c = -2400 psi Maximum tensile stress in spalling zone, spalling = +2400 psi = - c Maximum tensile stress in bursting zones, maximum bursting = (0.04) x avg. stress = +96 psi S. J. Taylor's experimental results indicate that the anchor plate will give rise to a similar stress distribution pattern as Guyon's method; the main difference lies in
CALVERT CLIFFS UFSAR 5.1-18 Rev. 47 the fact that the central bursting zone has a tensile stress peak of twice Guyon's value:
maximum bursting = +192 psi A state of biaxial tension in the concrete will appear on the outside face under the loading case 1.05D + 1.5P + 1.0TA + 1.0F. The superposition of the corresponding state of stress with the local anchor stresses reduces the load carrying capacity of the anchorage unit and causes a reduction in the maximum tensile strain to cracking.
On the other hand, the uniform compressive state of stress (vertical prestress) applied to the anchorage zone increases the load carrying capacity of the anchorage unit, with the maximum tensile strain to cracking being increased.
The design of the buttress anchor zones considered such additional vertical stresses, leading to a state of pseudo biaxial stress, the second direction being radial through the thickness.
For the above-mentioned case, i.e., 1.05D + 1.5P + 1.0TA + 1.0F, the averaged vertical (meridional) stress component is:
fa ~ + 400 psi The compressive bearing plate stress at 10" depth below the bearing plate is:
fc ~ - 1500 psi Thus, the two values introduced in the biaxial stress envelopes proposed in S. J.
Taylor's article are:
fc/ = 1500/5000 = 0.3 fc// = 400/5000 = 0.08 These values show that failure could occur if vertical reinforcing was not provided.
In fact, the maximum allowable vertical average tensile stress according to Taylor's interaction curve is fa/ = 0.03, therefore fa = +150 psi.
The three-dimensional stress distribution in the anchor zones was analyzed in sufficient detail to permit the rational evaluation of stress concentrations. A conical wedge segment was used as the basic design element and the radial splitting tension was determined as a tangential distribution function. The summation of splitting stresses through the entire volume of the lead-in zone established the value of the splitting force. This force is a function of the a/b ratio and the cone angle and/or, a/b and h. Several different combinations of the values were analyzed and the most critical values selected. A system analysis for the vertical splitting force was carried out based on statics. The magnitude of vertical and spalling forces were also determined.
The most unfavorable loads and load combinations were considered in the analysis of the anchorage zone. Stresses based on transient thermal gradients were used in all cases where the use of a steady state gradient underestimated the stresses and strains and were superimposed on the bursting stresses determined from the triaxial stress calculations. The computed stresses are less than the ACI allowable values. The design of the concrete reinforcement is based on this conservative analysis to provide a margin of safety similar to the other components of the reactor building structure and to control cracking in the
CALVERT CLIFFS UFSAR 5.1-19 Rev. 47 anchorage zone. As a result, there is no danger of delayed rupture of the concrete under sustained load due to local overstress and microcracking.
The reinforcing details, including the method for anchoring and splicing the reinforcing, are shown on Figure 5-1.
The reinforcement required has been designed primarily to resist tensile forces, and has been located such that it will efficiently do so. The reinforcement was provided for load cases which create the maximum tensile forces and for other load cases the relevant shear forces or stresses were superimposed.
The amount of reinforcing steel was computed manually for all Seismic Category I structures, except the exterior shell and the base slab of the Containment Structure, using conventional reinforced concrete design methods for "Working Stress Design" and "Ultimate Strength Design" depending on the governing load combination.
The seismic analysis was conducted as described in the following subsection, providing values for lateral accelerations, shears, moments and displacements at specified locations. The lateral acceleration values are applied to the axisymmetric Containment Structure as non-axisymmetric static loading in Bechtel's Analysis of Axisymmetric Shell and Solid Subject to Non-Axisymmetric Static Loading, Dynamic Loading or Base Acceleration Program, CE 771. The Containment Structure is idealized as an assemblage of a series of discretized elements. Resultant shear forces, longitudinal, circular and cross moments are calculated from stresses obtained as an output from program CE 771 at 15 increments, from 0 to 180 and at 270. The combined shear forces, normal forces, and moments are applied to the specified section and, using crack section analysis and compressive and tensile stresses in concrete, liner plate interior and exterior reinforcement are determined. Allowable working case stress and yield case stress determine the required area of reinforcement in the specified section.
The possibility of the concrete breaking along shear planes was considered at the intersection of (1) the buttress with the cylinder and (2) the cylinder with the base slab.
- a.
Buttress - Cylinder Intersection An increase in the compression force at the buttress corresponds to an increase in the concrete area of the same magnitude.
- b.
Cylinder - Base Slab Intersection An analysis for the most critical radial shear conditions was performed.
The difference in shear stiffness between the shell and the buttress and the remainder of the shell was included as a shear amplification factor. The reinforcing required was less then the reinforcing provided.
The possibility of concrete breaking along a shear plane is excluded by providing ample reinforcing. In other locations, breakage along the shear plane has been excluded by the opposition of prestressing and anchor forces.
For this reason, special anchorage zone reinforcing is based on the following considerations:
- a.
Full-scale load tests and final designs of similar anchorages.
CALVERT CLIFFS UFSAR 5.1-20 Rev. 47
- b.
The post-tensioning supplier's recommendations of anchorage reinforcing requirements.
- c.
Review of the final details of the combined reinforcing on earlier projects by the consulting firm of T.Y. Lin, Kulka, Yang and Associate.
Seismic or Wind Loading Seismic loading of the structure is higher in all cases than that of tornado or wind loading. The seismic analysis was conducted in the following manner:
The loads on the Containment Structure caused by earthquake were determined by a dynamic analysis of the structure. The dynamic analysis was made on an idealized structural model of lumped masses and weightless elastic columns acting as springs.
The analysis consists of three steps:
(1) The determination of the natural frequencies of the structure and its mode shapes, (2) the response of these modes to the earthquake by the response spectrum technique, and (3) combining modal responses to obtain structural response.
The natural frequencies and mode shapes were computed using the matrix equation of motion shown below for a lumped mass system. The form of the equation is:
2[]{} = []{}, where
[K] =
matrix of stiffness coefficients including the combined effects of shear and flexure in the structure and the rotation, and horizontal translation of the base slab on soil.
[m] =
matrix of concentrated masses.
{n} =
matrix of mode shapes n
=
angular frequency of vibration in the n-th mode.
The results of this computation are the several values of n for n and mode shapes n = 1, 2, 3,.p, where p is the total number of degrees of freedom (i.e., lumped masses) in the idealized model.
The response of the structure to the specified earthquake was then computed by the response spectrum technique as follows:
- a.
Using mode frequencies and respective damping values, a response acceleration, Sn, is read from the spectrum curves for each mode.
The modal acceleration, An, is given by:
An = Sn
=1
2
=1 The acceleration per point i and per mode n is given by:
Ain = in An
CALVERT CLIFFS UFSAR 5.1-21 Rev. 47 The inertial force per point per mode, Fin' is given by:
Fin = mi An where mi is the mass lumped at point i.
- b.
Using the inertial force per point per mode, Fin, shears and moments are computed per point per mode. The "Root Mean Square" method is used for combining modal responses (shears, moments, stresses, deflections, and/or accelerations) in the response spectrum modal analysis of Seismic Category I structures. In this method, responses are combined using the square root of the sum of the squares of responses of each mode. This method was for combining all predominant modal responses including closely spaced modal frequencies. We have examined the effect of adding closely spaced modes linearly, and find that allowable stresses have not been exceeded.
- c.
Seismic and wind shears are transferred across construction joints either by friction, by bond, by shear keys or by a combination of these.
Large Openings (Equipment Hatch and Personnel Lock Opening)
The primary loads considered in the design of the equipment hatch and personnel lock opening, as in any other part of the structure, were dead load, prestress, pressure, earthquake, and thermal loads. The secondary loads considered were the following effects caused by the above primary loads:
The deflection of tendons around the opening; The curvature of the shell at the opening; and The thickening around the opening.
The primary loads listed are mainly membrane loads with the exception of the thermal loads. In addition to membrane loads, incident pressure also produces punching shear around the edge of the opening. The magnitudes of these loads for design purposes were the magnitudes at the center of the opening. These are fairly simple to establish knowing the values of hoop and vertical prestressing, incident pressure, and the geometry and location of the opening.
Secondary loads were computed by the following methods:
- a.
The membrane stress concentration factors and effect of the deflection of the tendons around the equipment hatch were analyzed for a flat plate by the finite element method. The stresses computed by conventional stress concentration factors, compared with those values found from the above-mentioned finite element computer program, demonstrated that the deflection of the tendons does not significantly affect the stress concentrations. This is a plane stress analysis and does not include the effect of the curvature of the shell. However, it gives assurance of the correctness of the assumed membrane stress pattern caused by the prestressing around the opening.
- b.
With the help of Reference 1, stress resultants around the large opening were found for various loading cases. Comparison of the results found from this reference with the results of a flat plate of uniform thickness with a circular hole, showed the effect of the cylindrical curvature on stress concentrations around the opening.
CALVERT CLIFFS UFSAR 5.1-22 Rev. 47 Normal shear forces (relative to the opening) were modified to account for the effect of twisting moments (Reference 1). These modified shear forces are called Kirschoff's shear forces. Horizontal wall ties were provided to resist a portion of these shear forces.
- c.
The effect of the thickening on the outside face around the large opening was investigated using several methods. Reference 2 was used to evaluate the effect of thickening on the stress concentration factors for membrane stress. A separate axisymmetric finite element computer analysis, for a flat plate with thickening on the outside face, was prepared to handle both axisymmetric and non-axisymmetric loads. This program predicts the effect of the concentration of hoop tendons with respect to the Containment Structure at the top and bottom of the opening.
For the analysis of the thermal stresses around the opening, the same method was used as for the other loadings. At the edge of the opening, a uniformly distributed moment, equal but opposite to the thermal moment existing on the rest of the shell, was applied and evaluated using the methods of the preceding Reference 1. The stresses were then superimposed on the stresses calculated for the other loads.
In the case of LOCA temperature, after the incident pressure has already decreased, very little or no tension develops on the outside, so thermal strains will exist without the relieving effect of the cracks. However, the liner plate will reach a high strain level and so will the concrete at the inside corner of the penetrations, thereby relieving the very high stresses, but still carrying a high moment in the state of redistribution stresses.
In the case of 1.5P (prestress fully neutralized) + 1.0TA (accident temperature), the cracked concrete, with highly strained tension reinforcement, constitutes a shell with stiffness decreased but still essentially constant in all directions. See Appendix 5E for an evaluation that reduced the original containment minimum design prestress. In order to control the increased hoop moment around the opening, the hoop reinforcement is about twice that of the radial reinforcement (Figure 5-3).
The equipment hatch opening was thus thickened for the following reasons:
- 1.
To reduce the predicted high membrane stresses around the opening;
- 2.
To facilitate tendon placement;
- 3.
To facilitate steel reinforcing placement; and
- 4.
To compensate for the reduction in the overall shell stiffness due to the opening.
Since the resultant forces on any part of the containment exterior structure produce compressive stresses in the concrete, and sufficient unprestressed reinforcement is provided to control any local cracking, no significant cracks are expected. In the absence of any significant cracks, it is believed that the concrete will provide adequate corrosion protection for the liner plate. Cathodic protection for the Containment Structure is described in Section 5.1.7. Thus, no surveillance measures are necessary to detect the corrosion of the liner plate in the Containment Structures.
The working stress method, i.e., elastic analysis, was applied to both load combinations for design loads, as well as for yield loads, using the analytical
CALVERT CLIFFS UFSAR 5.1-23 Rev. 47 procedures described above. The only difference is the higher allowable stresses under yield conditions. The various factored load combinations and capacity reduction factors are specified in Appendix 5A and were used for the yield load combinations using the working stress design method. The design assumption of straight line variation of stresses was maintained under yield conditions.
The governing design condition for the edges of the equipment hatch opening at the outside face is the LOCA. Under this condition, approximately 60% of the total bonded reinforcing steel needed at the edge of the opening at the outside face, is required to resist the thermal load.
Excluding the thermal load, the remaining stress, equivalent to approximately 40%
of the total stress including thermal, at the edge of the outside face, is the sum of the following stresses:
Normal stresses, resulting from membrane forces, including the effect of thickening, contribute approximately minus 45% (minus 18% of the total).
Flexural stresses, resulting from the moments caused by thickening on the outside face, contribute approximately 155% (62% of the total).
Normal and flexural stresses, resulting from membrane forces and moments caused by the effect of cylindrical curvature, contribute approximately minus 10% (minus 4% of total).
Penetrations Analysis of the Containment Structure penetrations was divided into three parts:
(1) the concrete shell, (2) the liner plate reinforcement and closure to the pipe, and (3) the thermal gradients and protection requirements at the high-temperature penetrations. The three parts will be discussed separately.
- a.
Concrete Shell In general, special design consideration was given to all openings in the Containment Structure. Analysis of the various openings has indicated that the degree of attention required depends upon the penetration size. Small penetrations are considered to be those with a diameter smaller that 2-1/2 times the shell thickness, i.e., approximately 8' in diameter or less. For openings of 8' diameter or less, the curvature effect of the shell is negligible (Reference 1). In general, the typical concrete wall thickness has been found to be capable of taking the imposed stresses using bonded reinforcement, and the thickness is increased only as required to provide space requirements for radially deflected tendons. The induced stresses, due to normal thermal gradients and postulated rupture conditions, distribute rapidly and are of a minor nature compared to the numerous loading conditions for which the shell must be designed. The small penetrations are analyzed as holes in a flat plate. Applied piping restraint loads due to thermal expansion or accident forces are assumed to distribute in the cylinder as stated in Reference 3. Typical details associated with these opening are indicated in Figure 5-2.
- b.
Liner Plate Closure The stress concentrations around openings in the liner plate were calculated using the theory of elasticity. The stress concentrations were then reduced by the use of a thickened plate around the opening. In the case of a penetration with no appreciable external load, shear connectors were used to maintain deformation compatibility between the liner plate
CALVERT CLIFFS UFSAR 5.1-24 Rev. 47 and the concrete. Inward displacement of the liner plate at the penetration was also controlled by shear connectors.
In case significant external operating loads are imposed upon the pipe penetration, the stress level from the external loads is limited to the design stress intensity values, Sm, given in the ASME, B&PV Code,Section III, Article 4. The stress level in the shear connectors from external loads is in accordance with American Institute of Steel Construction (AISC) Code for A-36 steel.
The combining of stresses from all effects was performed using the methods outlined in the ASME, B&PV Code,Section III, Article 4, Figure N-414. Figure 5-9 shows a typical penetration and the applied loads.
Stresses due to the effects of pipe loads, pressure loads, dead loads, and earthquake loads were calculated and the stress intensity was kept below Sm.
The stresses from the remaining effects were combined with the above-calculated stresses and the resultant stress intensity kept below Sa.
- c.
Thermal Gradient The only high temperature pipes penetrating the Containment Structure shell are the main steam and feedwater pipes, steam generator blowdown line, and the reactor coolant letdown and sampling lines. Cooling was provided to maintain the temperature in these penetrations below 150F.
Liner Plate The primary purpose of the liner plate (including welds) is to provide leak tightness integrity to the post-tensioned concrete containment. Structural integrity of the structure is provided by the post-tensioned concrete and not by the liner plate.
The design, construction, inspection, and testing of the liner plate, which acts as a membrane and is not a pressure vessel, was not covered by any recognized code or specification. All components of the liner that must resist the full design pressure, such as penetrations, were selected to meet the requirements of ASME, B&PV Code,Section III, Nuclear Vessels, Paragraph N-1211. ASTM A516, Grade 60 or 70 made to ASTM A300 is a steel which meets these requirements and thus was used as a plate material for penetrations. This material has excellent weldability characteristics.
There are no design conditions under which the liner plate is relied upon to assist the concrete in maintaining the integrity of the structure, even though the liner may provide assistance in order to maintain deformation compatibility.
Loads are transmitted to the liner plate through the anchorage system and through direct contact with the concrete and vice versa. At times, loads may also be transmitted by bond and/or friction with the concrete. These loads cause, or are caused by, liner strain. The liner was designed to withstand the predicted strains.
Possible cracking of concrete has been considered and reinforcing steel is provided to control the width and spacing of the cracks. In addition, the design is such that total structural deformation remains small during the loading conditions, and that any cracking will be orders of magnitude less than that sustained in the repeated attempts to fail the prestressed concrete reactor vessel "Model 1," and
CALVERT CLIFFS UFSAR 5.1-25 Rev. 47 even smaller than the concrete strains of overpressure tests of "Model 2" (both at Gulf-General Atomic) (References 4 and 5).
As described, the consequences of concrete cracking on structural integrity are limited by the bonded reinforcing and unbonded tendons that have been provided.
The effect of concrete cracking on the liner plate has also been considered. The anchor spacing and other features are such that the liner will sustain orders of magnitude of strain less than did the liner of Model 1 at Gulf-General Atomic (Reference 4) without tensile failure.
Liner Plate Anchors The liner plate anchors were designed to preclude failure when subjected to the worst possible loading combinations. The anchors were also designed such that, in the event of a missing or failed anchor, the total integrity of the anchorage system would not be jeopardized.
The following loading conditions were considered in the design of the anchorage system:
Prestress Internal Pressure Shrinkage and Creep of Concrete Thermal Gradients Dead Load Earthquake Loading Wind or Tornado Loadings Vacuum The following factors were also considered in the design of the anchorage system:
Initial inward curvature of the inner plate between anchors due to fabrication and erection inaccuracies.
Variation of anchor spacing.
Misalignment of liner plate seams.
Variation of plate thickness.
Variation of liner plate material yield strength.
Variation of Poisson's ratio for liner plate material.
Cracking of concrete in anchor zone.
Variation of the anchor stiffness.
The anchorage system satisfies the following conditions:
The system has sufficient strength and ductility, with energy absorbing capability sufficient to restrain the maximum force and displacement resulting from the condition where a panel with initial outward curvature is adjacent to a panel with initial inward curvature.
The system has sufficient flexural strength to resist the bending moment which would result from the above condition.
The system has sufficient strength to resist any radial pull-out forces.
CALVERT CLIFFS UFSAR 5.1-26 Rev. 47 Liner Supports In designing for structural bracket loads applied parallel to the plane of the liner plate, or loads transferred through the thickness of the liner plate, the following criteria and methods have been used:
The liner plate was thickened to reduce the predicted stress level. The thickened plate, with the corresponding thicker weld attaching the bracket to the plate, will also reduce the probability of the occurrence of a leak at this location.
For tensile loads applied perpendicular to the plane of the liner plate, sufficient anchorage is provided.
The allowable stress in the perpendicular direction was calculated using the allowable predicted strain in that direction together with the predicted stresses in the plane of the liner plate.
In setting the above criteria, the reduced strength and strain capability of the material, perpendicular to the direction of rolling, was also considered. In this case, the major stress is normal to the plane of the thickened liner plate. The allowable stresses were reduced to 75% of the allowable stress calculated above.
5.1.4 IMPLEMENTATION OF CRITERIA See Appendix 5E for an evaluation that reduced the original containment minimum design prestress.
This section documents the manner in which the design criteria were met by the designer.
Section 5.1.4.1 discusses original isostress plots and tabulations of predicted stresses for various materials. The isostress plots of the homogeneous cracked concrete structure indicate the general stress pattern for the structure as a whole, under various loading conditions. More specific documentation is made of the predicted stresses for all materials in the structure. In these tabulations, the predicted stresses are compared with the allowable to permit an easy comparison and evaluation of the adequacy of the design.
Sections 5.1.4.3 and 5.1.4.4 illustrate the actual details used in the design to implement the criteria.
5.1.4.1 Results of Analysis See Appendix 5E for an evaluation that reduced the original containment minimum design prestress.
The isostress plots, Figures 5-6 and 5-7, show the three principal stresses and the direction of the principal stresses normal to the hoop direction. The principal stresses are the most significant information about the behavior of the structure under various conditions and were a valuable aid for the final design.
The plots were prepared by a cathode-ray tube plotter. The data for plotting were taken from the stress output of the finite element computer program of the following load cases:
D + F + L (Construction Case)
D + F + L + 1.15P (Test Case)
CALVERT CLIFFS UFSAR 5.1-27 Rev. 47 D + F + L + P + TA (Design Incident Case) 1 (1.05 D + F + 1.5P + TA) (Factored Load Case)
The above axisymmetric loading conditions have been found to be governing in the design since they result in the highest stresses at various locations in the structure.
The containment stress analysis results for structural concrete and liner plate, including shear stresses, are shown here.
5.1.4.2 Prestress Losses See Appendix 5E for a discussion on possible prestress losses due to tendon wire breakage.
In accordance with ACI Code 318-63, the original design provided for prestress losses caused by the following effects:
- a.
Seating of anchorage;
- b.
Elastic shortening of concrete;
- c.
Creep of concrete;
- d.
Shrinkage of concrete;
- e.
Relaxation of prestressing steel stress; and
- f.
Frictional loss due to intended or unintended curvature in the tendons.
All of the above losses can be predicted with sufficient accuracy.
In this case, the environment of the prestress system and concrete is not appreciably different from that found in numerous bridge and building applications.
Considerable research has been done to evaluate the above items and is available to designers in assigning the allowances. Building code authorities consider it acceptable practice to develop permanent designs based on these allowances.
The following categories and values of prestress losses have been considered in the design:
Type of Loss(a)
Assumed Value Seating of Anchorage None Elastic Shortening of Concrete 2 ksi Shrinkage and Creep of Concrete for Dome tendons (test data) 146.7x10-6 in./in.
Hoop tendons (test data) 248x10-6 in./in.
Vertical tendons (test data) 137.5x10-6 in./in.
Relaxation of Prestressing Steel(b) 9.5% of 0.70 = 15.96 ksi Frictional Loss K = 0.0003, = 0.158 (a)
See Appendix 5E for a discussion on possible prestress losses due to tendon wire breakage.
(b)
See Appendix 5E for a discussion on the relaxation value of the prestressing steel used for vertical tendons replaced in 2001 through 2002 on both Units.
CALVERT CLIFFS UFSAR 5.1-28 Rev. 47 There is no allowance for the seating of the BBRV anchor since no slippage occurs in the anchor during transfer of the tendon load into the structure. Sample lift-off readings will be taken to confirm that any seating loss is negligible.
The loss of tendon stress due to elastic shortening was based on the change in the initial tendon relative to the last tendon stressed.
The value used for shrinkage and creep loss represents only that which could occur after stressing. In general, since the concrete is well aged at the time of stress, little shrinkage and creep is left to occur and add to prestress loss.
The value of relaxation loss is based on the information furnished by the tendon system vendor, The Prescon Corporation. See Appendix 5E for a discussion on the relaxation value of the prestressing steel used for vertical tendons replaced in 2001 through 2002 on both Units.
Frictional loss parameters for unintentional curvature (K) and intentional curvature
() are based on full-scale friction test data. This data indicates actual values of K = 0.0003 and = 0.125 versus the design values of K = 0.0003 and = 0.158.
Assuming that the jacking stress for the tendons is 0.80 or 192,000 psi and using the above prestress loss parameters, the following tabulation shows the magnitude of the design losses and the final effective prestress at the end of 40 years for a typical dome, hoop, and vertical tendon. See Appendix 5E for a discussion on the relaxation value of the prestressing steel used for vertical tendons replaced in 2002 through 2002 on both Units, and a discussion on the final effective prestress at the end of 60 nominal years due to license extension.
Dome(c)
(ksi)
Hoop(c)
(ksi)
Vertical(c)
(ksi)
Jacking Stress 192 192 192 Friction Loss 20.86 29.6(a) 21.7 Seating Loss 0
0 0
171.14 162.4 170.3 Elastic Loss 2.0 2.0 2.0 Creep and Shrinkage Loss 4.26 7.19 3.98 Relaxation Loss 15.96 15.96 15.96 Final Effective Prestress(b) 148.92 137.25 148.36 (a)
Average of crossing tendons.
(b)
This force does not include the effect of pressurization, which increases the prestress force.
(c)
Losses shown are for original nominal 40-year license. See Appendix 5E for a discussion of losses at the end of 60 nominal years due to the license extension.
To provide assurance of achievement of the desired level of final effective prestress and that ACI 318-63 requirements were met, a written procedure was prepared for guidance of post-tensioning work. The procedure provided nominal values for end anchor forces in terms of pressure gauge readings for calibrated jack-gauge combinations. Force measurements were made at the end anchor since that is the only practical location of such measurements.
CALVERT CLIFFS UFSAR 5.1-29 Rev. 47 The procedure required the measured temporary jacking force for a single tendon to approach, but not exceed 850 kips (0.8 ). Thus the limits set by ACI 318-63 2606(a)1, and of the prestressing system supplier, were observed. Additionally, benefits were obtained by in-place testing of the tendon to provide final assurance that the force capability exceeded that required by design. During the increase in force, measurements were required of elongation changes and force changes in order to allow documentation of compliance with ACI 318-63 2621(a). The procedure required that the prestressing steel be installed in the sheath before stressing for a sufficient time period that the temperatures of the prestressing steel and concrete reach essential equilibrium, to establish conformance with ACI 318-63 2621(e). The jacking force of 0.8 further provided for a means of equalizing the force in individual wires of a tendon to establish compliance with ACI 318-63 2621(b). The procedures required compliance with ACI 318-63 such that, if broken wires resulted from the post-tensioning sequence, compliance with Section 2621(d) was documented. Each of the above procedures contributed to assurance that the desired level of final effective prestress would be achieved.
Paragraph 2606(a)2 of ACI 318-63 refers to "tendons" rather than to an individual tendon. Further, the paragraph does not refer to the location to be considered for the determination of 0.8, for example, "temporary jacking force" referred in paragraph 2606(a)1.
Two interpretations were required; both had to consider the effect of the resultant actions on both the prestressing system and structure.
The first interpretation was that the location for measurement of the seating force, used in calculating the percentage of, was at the end anchor and just subsequent to the measurement of the "temporary jacking force" referred in ACI 2606(a)1. The advantages of this location are several. One is that it is a practical one and thus the possibility for achieving valid measurements is greater. The second is that it is the same location used for measuring the "temporary jacking force" and measurements could be made without the added complexity of additional measuring devices. The third advantage is that measurements at this location provide assurance that the calculated percentage of at seating does not anywhere exceed the maximum percentage of to which that tendon has been subjected.
Several possible cases were considered for the second interpretation so as to allow anchoring of an individual tendon without exceeding the requirement stated for "tendons" collectively in ACI 318-63 2606(a)2. One such case assumed that the anchoring force for the typical tendon was that for a tendon anchored midway through the prestressing sequence. It further assumed that the losses to be assumed were one-half of the sum of elastic losses, and of the creep, shrinkage, and relaxation predicted to occur during the entire prestressing sequence. This interpretation, however, was considered to be neither practical nor enforceable, since it resulted in changing the seating forces as the actual (as compared to the schedule) time length of the prestressing period was dictated by weather and manpower availability.
In another case, the stressing is done by jacking each tendon to the required force of 850 kips (0.8 ), and placing shims of predetermined thickness, corresponding to the calculated elongation, between the bearing plate assembly and the washer.
Proper tendon stress is assured by comparing the jack pressure and tendon elongation with previously calculated values.
CALVERT CLIFFS UFSAR 5.1-30 Rev. 47 The case adopted was to seat each tendon with a measured "pressure" reading for the jack, at "lift-off" of the end anchor, of 775 kips (between 0.72 and 0.73,).
This procedure had several advantages.
One advantage was that the force on the containment and the tendon was within the bounds of those for which it had been tested and resulted in no known detrimental effects. The second advantage was that the stressing procedure was simplified, since the stressing crews did not have to accommodate a large number of different anchoring force requirements. The third advantage was that, at the completion of stressing the last tendon, the expected losses were such that the average percentage of at the end anchors of the tendons would be less than 0.7 thus establishing compliance with ACI 318-63 2606(a)1 and 2. The fourth advantage was that the percentage loss of prestressing force was less than would be the case if the tendons were anchored in such a manner that the calculated percentage of nowhere exceeded 0.7.
The latter advantage deserves special mention since it plays a strong role in assuring that the final effective prestress equaled or exceeded the desired value.
For example, if the percentage of at the anchorage of the tendons were 0.1,
creep and shrinkage of concrete could result in the loss of almost all of the prestressing force. Assuming that the total losses due to creep, shrinkage and elastic shortening equal 0.1, then the final effective prestress would be 20% of an initial prestress equivalent to 0.5. If the initial prestress was equivalent to 0.7, the final effective prestress, neglecting relaxation for the moment, would be about 86% of the initial prestress. Clearly, the assurance (that the concrete creep and shrinkage losses have been properly accounted for) increases as the percentage of for the anchored tendon(s) increases. However, this design was committed to meeting the ACI 318-63 requirement and the anchorage force for the tendons was kept at or below 0.7 in accordance with the interpretation described.
5.1.4.3 Liner Plate The following design bases were applied to the containment liner so that it meets the specified leak-rate under LOCA conditions:
- a.
The liner is protected against damage by missiles. (Section 5.1.5.3)
- b.
The liner plate strains are limited to allowable values that have been shown to result in leak-tight vessels or pressure piping.
- c.
The liner plate is prevented from development of significant distortion.
- d.
All discontinuities and openings are properly anchored to accommodate the forces exerted by the restrained liner plate, and careful attention is paid to details of corners and connections to minimize the effects of discontinuities.
Pressure vessels, pressure piping, high pressure hydraulic tubing, and similar containers are made by cold forming, drawing, and dishing operations, where strains may approach the elongation capacity of the material. (For mild steel at failure, this elongation varies from 15 to 30%.) These forming operations result in high strains both in tension and compression. Vessels and piping components manufactured by these methods have a history of high leak-tight integrity, proving that subjecting the steel material to high strain does not affect its leak-tight integrity.
The best basis for establishing allowable liner plate strains is considered to be that portion of the ASME, B&PV Code,Section III, Nuclear Vessels, Article 4.
CALVERT CLIFFS UFSAR 5.1-31 Rev. 51 Specifically, the following sections have been adopted as guides in establishing allowable strain limits:
Paragraph N-412(m)
Thermal Stress Paragraph N-414.5 Peak Stress Intensity Table N-413 Classification of Stresses for Some Typical Cases Figure N-414 Stress Categories and Limits of Stress Intensity Figure N-415(A)
Design Fatigue Curves Paragraph N-412(n)
Operational Cycle Paragraph N-415.1 Vessels Not Requiring Analysis for Cyclic Operation American Society of Mechanical Engineers design codes require that the liner material be prevented from experiencing significant distortion due to the thermal load and that the stresses be considered from a fatigue standpoint
[Paragraph N-412(m)(2)]. The following fatigue loads were considered in the design of the liner plate:
- a.
Thermal cycling due to annual outdoor temperature variations. Daily temperature variations do not penetrate a significant distance into the concrete shell to appreciably change the average temperature of the shell relative to the liner plate. The original number of cycles for this loading is 40 cycles for the plant life of 40 years. Reference 10 provides an evaluation of the effect of thermal cycling on the liner plate for a total of 60 years. The evaluation concluded that the containment liner plate is acceptable for 60 years.
- b.
Thermal cycling due to interior temperature variations during the startup and shutdown of the reactor system. The number of cycles for this loading was assumed to be 500.
- c.
Thermal cycling due to the LOCA was taken very conservatively to occur only once during plant life. Thermal load cycles in the piping systems are somewhat isolated from the liner plate penetrations by the concentric sleeves between the pipe and the concrete. The attachment sleeve was designed in accordance with ASME, B&PV Code,Section III fatigue considerations. All penetrations were reviewed for a conservative number of cycles to be expected during plant life.
Thermal stresses in the liner plate fall into the categories considered in Article 4,Section III, Nuclear Vessels of the ASME, B&PV Code. The allowable stress in Figure N-415(A) is for alternating stress intensity for carbon steels and temperatures not exceeding 700F. In addition, the ASME code further requires that significant distortion of the material be prevented.
In accordance with ASME, B&PV Code, Paragraph 412(m)(2), the liner plate is restrained against significant distortion by continuous angle anchors and never exceeds the temperature limitation of 700F. It also satisfies the requirements for limiting strains on the basis of fatigue consideration. A typical section showing the anchors is included in Figure 5-1.
American Society of Mechanical Engineers, B&PV Code, Paragraph 412(n),
Figure N-415(A), has been developed as a result of research, industry experience, and the proven performance of code vessels. Because of the conservative factors it contains on both stress intensity and stress cycles, and its being a part of a recognized design code, Figure N-415(A) and its appropriate limitations have been used as a basis for establishing allowable liner plate strains. Since the graph in
CALVERT CLIFFS UFSAR 5.1-32 Rev. 51 Figure N-415(A) does not extend below ten cycles, ten cycles was used for the LOCA instead of one cycle mentioned above.
Establishing an allowable strain based on ten significant thermal cycles of the LOCA condition would permit an allowable strain [from Figure N-415(A)] of approximately 2%. Maximum allowable tensile or compressive strain has been conservatively set at 0.5% (compared to 2% shown above). The maximum predicted strain in the liner plate during LOCA conditions has been found to be 0.25% compression.
At the design LOCA condition, there will be no tensile stress anywhere in the liner plate membrane. This is true both at the time of initial pressure release and under any later pressure and temperature condition. The purpose of specifying a non-destructive examination temperature requirement is to provide protection against a brittle fracture or cleavage mode of failure. However, this type of failure is precluded by the absence of tensile stresses.
No allowable compressive strain value has been set for the test condition because the value will be less than that experienced under the LOCA condition. The maximum allowable tensile strain will be 0.2% under test conditions; the predicted value is much smaller.
The maximum compressive strains are caused by LOCA pressure, thermal loading prestress, shrinkage, and creep. The maximum calculated strains do not exceed 0.0025 in./in. and the liner plate will always remain in a stable condition.
The stability of the liner plate has been studied for the loading cases and deformations to which it may reasonably be subjected. The critical loading cases that were considered included the LOCA condition and the operating condition during the winter.
Two separate solutions to the plate stability problem were made:
- a.
As a compressed panel under biaxial compression, assuming that the channel and angle stiffeners are rigid in their attachment to the prestressed concrete Containment Structure and the liner.
- b.
As a compressed panel under biaxial compression, assuming the panel to be a portion of a large cylinder with a flexible stiffener system.
Figure 5-1 illustrates the actual physical configuration of the stiffening system for the liner plate. The channels function as horizontal stiffeners and the angles as vertical stiffeners.
For the solution, an initial deflected form for the liner plate is expressed in terms of a Fourier Series of the form:
= *cos *
=
=0 where: defines the central angle in a plan view of the cylinder from a point on the circumference where there is zero radial deflection to the point on the circumference where there is maximum radial deflection; x defines the radial deflection and defines the unsupported length in the vertical direction.
CALVERT CLIFFS UFSAR 5.1-33 Rev. 51 The stability analysis based on the above assumptions indicates that the critical buckling stress of the liner will be approximately 29,000 psi.
Under normal operating conditions, the maximum compressive stress in the liner will be approximately 23,000 psi. This reflects a buckling margin of 20%. Under LOCA conditions the expected stress in the liner plate will be 35,000 psi. Under this stress the overall structural stability of the liner plate can be maintained.
Also of concern is the nature of the state of stress and its behavior at the point of attachment between the stiffeners and the liner plate. Special tests have been conducted on simulated models of the liner plate and vertical stiffener assembly to determine the shear capacity of the angle and vertical stiffener assembly in order to determine the shear capacity of the angle anchorage. In the tests, two different configurations of support were used for the simulated continuous anchor. One case simulates the expected condition that will exist in the Containment Structure.
A second case simulates the condition that might exist at an isolated location if the concrete were not in continuous contact with the anchor. Being guided in the proportioning of the liner plate stiffeners by the more critical values of shear transfer for the case of non-continuous contact will, in general, result in a margin of safety for progressive failure of anchors of approximately 2.7. The weld configuration shown in Figure 5-1 is felt to be adequate to transfer all loads that are considered in the design of the structure between the liner plate and the stiffener-anchors.
The conservative design approach of the stiffening system used in the liner plate to prevent significant distortions at accident conditions, and the stringent welding and weld inspection requirements ensure that the leak tightness of the liner plate at the LOCA condition will not change from that at the test condition.
In isolated areas the liner plate may have initial inward curvature due to construction. The anchors are designed to resist the forces and moments induced when a section of the liner plate between anchors has initial inward curvature.
The liner plate is anchored at all discontinuities to eliminate excessive strains. The forces in the liner plate at the discontinuities were evaluated and the anchors designed to resist these forces.
The liner adjacent to the penetrations is backed up by concrete. Refer to Figure 5-2 for a description of the anchoring arrangement at discontinuities and typical details.
The containment penetrations behave, partly or fully depending on their size, as elements of a pressure vessel. The sizes range from the small closure pieces for small pipe penetrations to larger components such as air locks and the equipment hatch. Those portions that are not backed up by concrete (the penetrations themselves) are treated as a pressure vessel and comply with ASME, B&PV Code,Section III, Nuclear Vessels, Subsection B, and are designed in accordance with the requirements of Section VIII, Paragraphs UG-27 through UG-33 for the service conditions outlined in Section 5.1.
The pressure part to thickened liner plate transition areas are designed and reinforced in accordance with the requirements of Paragraphs UG-36 through UG-41 as a welded assembly conforming to ASME, B&PV Code,Section VIII, Paragraph UW-13.
CALVERT CLIFFS UFSAR 5.1-34 Rev. 51 The air locks to thickened liner attachment rings were analyzed by finite-element methods; the membrane meridional, membrane hoop and the average radial stresses were investigated and the adequacy of design was verified.
The liner plate is considered as a composite part of the containment shell and as such, it is investigated by utilizing the strain and deformation data obtained by the finite-element analysis of the containment shell as tabulated in Table 5-1. See Appendix 5E for later tables associated with an evaluation to reduce the original amount of required containment prestress. The ASME B&PV Code is used only as a guideline to establish allowable strain limits, and not to verify the structural and leaktight adequacy of the liner.
In general, where the liner is firmly anchored and bonded to the concrete and maintains a strain compatibility with the concrete shell, the thickening of the liner will neither reduce nor increase the stresses. However at the openings, and particularly in the transition boundaries where the pressure elements of the penetrations are welded to the liner, the increase in the thickness of the steel plate will:
- a.
reduce the possible local stress concentrations, both in the steel plate and the adjacent concrete;
- b.
increase the stiffness of the plate and hence its buckling limit; and
- c.
increase the leaktight integrity of the transition welds.
At all penetrations the liner plate is thickened to reduce stress concentrations in accordance with the ASME, B&PV Code 1968,Section III, Nuclear Vessels. The thickened portion of the liner plate is anchored to the concrete by use of anchor studs all around the penetrations. For details of the penetrations see Figure 5-9.
The sleeves, pipe cap and all welds associated with the penetrations are designed to resist all loads previously mentioned and also the prestress forces and internal design pressure.
At each location where a load is to be delivered to the walls, slab, or dome of the Containment Structure, an insert plate is provided to transmit the load through the liner. The insert plate is thickened and stiffened as required to deliver the load and to reduce stress concentrations in the liner. The insert plate is anchored to the concrete by appropriate anchors and shear connections. Typical examples of such insert plates are the polar crane brackets, and the floor beam brackets at the operating deck; typical details are shown in Figure 5-1.
5.1.4.4 Penetrations All penetrations are pressure resistant, leak-tight, welded assemblies designed, fabricated, and tested in accordance with the ASME, B&PV Code,Section III, Nuclear Vessel Code.
Types of Penetrations
- a.
Electrical Penetrations Two types of electrical penetration assemblies are used - canister and unitized header. All electrical penetration assemblies are fabricated and tested in accordance with the ASME, B&PV Code,Section III, Nuclear Vessel Code. The canister type is inserted in a nozzle of suitable diameter integral with the Containment Structure and field welded on the inside end.
The unitized header type is welded to the nozzle on the outside end. All penetration assemblies are provided with a means to pressurize for
CALVERT CLIFFS UFSAR 5.1-35 Rev. 51 monitoring of leakage. Any abnormal depressurization of an assembly is annunciated locally and in the Control Room.
There are three different types of electrical penetrations, used as follows:
- 1.
Type 1 - 15 kV medium voltage power penetration canisters. These are 30" diameter cylinders constructed from steel plate (SA516, GR 70, 0.652" thick). Stainless steel header plates are then welded into each end containing epoxy bushings welded into the header plates.
- 2.
Type 2 - Low voltage power and control penetration:
- a.
Canister design: These are constructed of 10" diameter schedule 40 seamless steel pipe (SA 106B). Conductors are sealed with glass hermetics either direct fired into header plates or as assemblies welded into header plates.
- b.
Unitized Header Design: Header is constructed of stainless steel, Type 304 (SA 240). Feedthroughs consist of conductors with resilient seal, are mounted and sealed to a header.
- 3.
Type 3 - Instrumentation, thermocouple and coaxial penetrations:
- a.
Canister design: See type 2.a above.
- b.
Unitized Header design: See type 2.b above.
Figure 5-2 shows the different types of electrical penetration assemblies.
- b.
Piping Penetrations Single barrier piping penetrations are provided for all piping passing through the containment walls. The closure of the pipe to the liner plate is accomplished with a pipe cap welded to the pipe and to the liner plate reinforcement. In the case of piping carrying hot fluid, the pipe is insulated and cooling is provided to reduce the concrete temperature to 150F.
Figure 5-2 shows typical hot and cold pipe penetrations. The modes of containment isolation are covered in Section 5.2.
The anchorage of penetration closure connecting pipes to the containment wall were designed as Seismic Category I structures to resist all forces and moments caused by a postulated pipe rupture. The design conditions include the maximum pipe reactions and pipe rupture forces.
The following design basis for typical piping penetrations was used to ensure the integrity of the liner penetration junction at the piping.
- 1.
The penetration assembly, consisting of pipe cap and the assembly welds and welds to the liner plate, utilizes full penetration welds. The assembly is anchored into the wall concrete and designed to accommodate all forces and moments due to pipe rupture and thermal expansion.
- 2.
The design basis is that the pipe penetration is the strongest point in the system when a pipe break is postulated. Pipe stops, increased pipe thickness or other means are used to attain this. Part of this basis also is that the operation of closure valves will not be impaired by any postulated pipe break.
CALVERT CLIFFS UFSAR 5.1-36 Rev. 51
- c.
Large Penetrations (Equipment and Personnel Access Hatches)
An equipment hatch opening, 19' in diameter, is provided as shown on Figure 5-3. The equipment hatch opening is covered on the inside of the Containment by a dished hatch, fabricated from welded steel, furnished with a double gasketed flange and bolted in place. The equipment hatch opening also has an outage use door at the exterior opening. The containment outage door is comprised of three assemblies: (1) a transition ring assembly welded directly to the equipment hatch nozzle steel liner plate; (2) a door frame assembly attached to the transition ring by means of bolts and locating pins that support all dead weight and seismic loads; and (3) a hinged door assembly. The frame of the containment outage door is equipped with fittings or penetrations that will allow electricity, compressed air, water, etc., to be run through the equipment hatch opening without going through the door. Flanges, cable penetrations, and isolation valves at these penetrations will be designed to withstand containment pressure resulting from postulated boiling in the RCS, as might occur during a loss of shutdown cooling. The design load to pass through the containment outage door is a reactor coolant pump motor on its transport trailer. For larger equipment, like replacement steam generator components, the bolted-on door frame assembly and door assembly may be removed from the transition ring assembly to restore the equipment hatch opening to its full diameter. Internal lugs welded to the nozzle liner plate and designed to carry the reactions due to the axial pressure loads and seismic loads to the nozzle liner plate may be removed and replaced when the door is restored.
Two personnel locks are provided. One of these is for emergency access only. Each personnel lock is a double door, welded-steel assembly. A quick-acting equalizing valve connects the personnel lock with the interior and exterior of the Containment Structure for the purposes of equalizing pressure in the two systems when entering or leaving. Typical details of the personnel access lock are shown in Figure 5-3.
The two doors in each personnel lock are interlocked to prevent both from being opened simultaneously and to ensure that one door is completely closed before the opposite door can be opened. Remote indicating lights and annunciators situated in the Control Room indicate the operational status of the door. Provision is made to permit bypassing the door interlocking system to allow doors to be left open during plant cold shutdown. An emergency lighting and communication system operating from an external emergency supply is provided in the lock interior.
- d.
Special Penetrations
- 1.
Refueling Tube Penetration A refueling tube penetration is provided for fuel movement between the refueling pool in the Containment Structure and the spent fuel pool in the Auxiliary Building. The penetration consists of a 36" stainless steel pipe installed inside a 42" pipe sleeve. The inner pipe acts as the refueling tube and is fitted with a gate valve in the spent fuel pool and an encapsulating pipe sleeve which is welded to the refueling pool liner and sealed off from the containment with a testable double O-ring blind flange in the refueling pool. This arrangement prevents leakage through the refueling tube in the event of a LOCA. The 42" pipe sleeve is welded to the containment liner.
CALVERT CLIFFS UFSAR 5.1-37 Rev. 47 Bellows expansion joints are provided on the transfer tube to compensate for any differential movement between the tube and the building structures. Figure 9-14 is a drawing of the refueling tube installation. The design basis for each expansion joint is listed below:
AUX. BUILDING FUEL POOL EXPANSION JOINT CONTAINMENT REFUELING POOL EXPANSION JOINT External design pressure 28 psig External design pressure 29 psig Internal design pressure 28 psig Internal design pressure 29 psig Design temperature 150F Design temperature 273F Lateral movement 1-1/2" Lateral movement 1/2" Axial movement 1-1/2" (expansion or contraction)
Displacements were selected to accommodate the maximum differential building settlements. The expansion joint in the refueling pool may be visually inspected from the outside with a periscope when the refueling pool is empty. Repair would probably involve removing the 54" OD pipe that forms the containment boundary. The outside of the expansion joint in the spent fuel pool may be visually inspected by draining the pool or by remote means. A test connection on the bellows provides a means of testing bellows integrity. Repair would require draining the spent fuel pool.
- 2.
Containment Supply and Exhaust Purge Penetrations The ventilation system purge penetrations are equipped with a testable double O-ring blind flange in the penetration room and a tight-seating butterfly valve inside containment used for isolation purposes. The blind flange is used to provide containment integrity during Modes 1-4.
The valve is manually opened for containment purging in Modes 5 and 6 as described in Section 9.8.2.2.
- 3.
Containment Vent Penetration This system is equipped with two valves to be used for isolation purposes. These valves are opened by a handswitch in the Control Room to vent containment pressure during power operation. Although control of hydrogen in Containment following an accident is not required, this penetration may be used as a hydrogen purge.
Design of Penetrations
- a.
Design Basis Penetrations conform to the applicable sections of American Standards Association (ASA) N6.2-1965, "Safety Standard for the Design, Fabrication, and Maintenance of Steel Containment Structures for Stationary Nuclear Power Reactors" which has since been withdrawn. All personnel locks and any portion of the equipment access door extending beyond the concrete shell conform in all respects to the requirements of ASME, B&PV Code,Section III, Code for Nuclear Vessels. All future penetrations will conform to ASME, B&PV Code,Section III, Division 1 for fluted head analysis, and Division 2 for those integrated with steel in concrete openings.
CALVERT CLIFFS UFSAR 5.1-38 Rev. 47 Each line which penetrates the containment and contains high-pressure or high-temperature fluids (steam, feedwater, and steam generator blowdown) passes through a structural steel sleeve mounted on the containment wall.
This sleeve acts as a positive stop to prevent whipping associated with fracture of a line containing high internal energy and thereby prevents damage to the penetration and breaching of the containment.
Further protection of each line, necessary to preclude pipe rupture between the penetration and the first valve, is accomplished by shortening the exposed length of pipe and installing the first valve as close as possible to the internal or external wall of the structure, dependent upon valve operating and maintenance clearances. Design bases which apply to the provision of automatic and manual isolation valves in the penetration lines are contained in Section 5.2.
All penetrations, except the equipment hatch and emergency personnel lock, are inside the Auxiliary Building; therefore, the temperature of Auxiliary Building penetration material will not fall below 30F. Using the assumption of +60F temperature inside the containment and outside air temperature of +20F, the temperature of the equipment hatch and emergency personnel door will not be below +30F, during startup, operation, or cooldown of the reactor. An investigation was made to determine necessity and/or feasibility of protecting the exterior surface of these penetrations. None was required. Two temperature indicators are provided to monitor the temperature inside the containment. A technical specification is not considered to be necessary for the above reasons.
- b.
High-Temperature Penetrations The main high-temperature piping consists of two penetrations for feedwater, two penetrations for main steam, two penetrations for steam generator blowdown, one for the reactor coolant letdown line, and one for the reactor coolant sampling line. These have a maximum operating temperature range between 435F and 653F. Thermal insulation is provided on the outside diameter of each line and separate coolant circulation, with instrumentation suitable for flow monitoring, is provided in the air gap between the insulation and the penetration liner sleeve. The combination of insulation and coolant circulation is designed to restrict the maximum temperature in the concrete to 150F.
For the condition of loss of penetration coolant circulation, the maximum steady state temperature in the concrete will be 300F at the penetration surface and decreases to 120F at a maximum radial depth of 48" in the containment wall. Actual peak temperatures in the penetrations resulting from LOCA conditions are expected to subside within six hours. A maximum temperature of 390F may be tolerated for 120 days (Reference 5) without appreciable loss of concrete strength.
The basis for limiting strains in the penetration steel is the ASME, B&PV Code,Section III, Nuclear Vessels, Article 4, 1965, and therefore, the penetration structural and leak-tightness integrity will be maintained. Local heating of the concrete immediately around the penetration will develop compressive stress in the concrete adjacent to the penetration and a negligible amount of tensile stress over a large area. The mild steel
CALVERT CLIFFS UFSAR 5.1-39 Rev. 47 reinforcing added around penetrations will distribute local compressive stresses for overall structural integrity.
- c.
Penetration Materials The material for the penetrations, including the personnel and equipment access hatches together with the mechanical and electrical penetrations, is carbon steel and conforms with the requirements of the ASME, B&PV Code, Nuclear Vessel Code. As required by the Nuclear Vessel Code, the penetration materials which form the pressure boundary meet the necessary Charpy V-notch impact values at a temperature 30F below the lowest service temperature.
- 1.
Piping Penetration Materials Materials specifications are listed below:
Piping Penetration Material Specification Penetration Sleeve ASTM A155 Penetration Reinforcing Rings ASTM A516 Penetration Sleeve Reinforcing ASTM A516 Bar Anchoring Rings and Plates ASTM A516 Rolled Shapes ASTM A36 (nonpressure boundaries)
- 2.
Electrical Penetration Materials The penetration sleeves to accommodate the electrical penetration assemblies are Schedule 40 carbon steel pipe, except where otherwise noted.
- 3.
Access Penetration Materials The equipment hatch and personnel access locks materials are all ASTM A516 made to ASTM A300.
Installation of Penetrations The qualification of welding procedures and welders is in accordance with Section IX, "Welding Qualifications" of the ASME, B&PV Code. The repair of defective welds is in accordance with paragraph N-528 of Section III, "Nuclear Vessels" and Section VIII of the Code.
Testability of Penetrations Only the following penetrations are ASME, B&PV Code,Section III, Class B penetrations as classified in the Atomic Energy Commission (AEC) Technical Safety Guide, "Reactor Containment Leakage Testing and Surveillance Requirements," December 15, 1966. As Class B penetrations, they are subject to individual periodic leak-rate tests separate from the integrated Class A tests.
- a.
Equipment Hatch
- b.
Personnel Lock
- c.
Emergency Personnel Escape Lock
- d.
Refueling Tube
- e.
Purge Line Inlet and Outlet
CALVERT CLIFFS UFSAR 5.1-40 Rev. 47 5.1.5 INTERIOR STRUCTURE For the original compartment designs, the occurrence of a LOCA was postulated to result from the rupture of the RCS piping, including the main loop piping, either within the reactor cavity or the SG compartments of the containment. For the current compartment designs, breaks are not postulated to occur in the main loop piping based on the Leak-Before-Break Evaluation (References 6 and 7). The design of the Babcock & Wilcox, Canada replacement steam generators also incorporates the application of leak-before-break, which included dynamic analysis of steam generator internals (U-tubes, divider plate, lattice grids, and shroud) for the effects of the most limiting postulated break (12 pressurizer surge line break). American Society of Mechanical Engineers Code-reconciled design and methods were used to install the replacement steam generators to ensure the equivalency of the replacement steam generators-to-RCS piping connections to the original steam generator-to-RCS piping connections, and thus the continued validity of the assumptions made in References 6 and 7. Since only main loop piping is present in the reactor cavity, no LOCA is currently postulated to occur in this compartment. However, smaller bore piping connected to the RCS is present in both steam generator compartments and the pressurizer compartment. The following discussion regarding a LOCA in the reactor compartment is provided as historical design information only.
5.1.5.1 Design Basis Design of the containment interior structures evolves around four basic systems:
The primary coolant system, the turbine steam system, the engineered safety system, and the fuel handling system.
The structures which house or support the basic systems are designed to sustain the factored loads of Appendix 5A.
The design bases applied are:
- a.
The structures will sustain all operating dead and live loads, thermal loads, and design seismic loads, without exceeding code allowable stresses.
- b.
Loads and deformations resulting from a LOCA failure and its associated effects in any one of the basic systems will be sustained and restricted such that propagation of the failure to any other system is prohibited. In addition, a failure in one loop of the Nuclear Steam Supply System will be restricted such that propagation of the failure to the other loop is prohibited.
Loss-of-coolant accident loads and associated effects include:
- a.
Thrust loads resulting from rapid mass release from a pipe break in any system;
- b.
Pressure buildup in locally confined areas such as the reactor cavity or the secondary shielding compartments; and
- c.
Jet forces resulting from the impingement of the escaping mass upon adjacent structures. The following method was employed to compute the jet forces:
- 1.
Rupture may occur anywhere on portions of the main coolant pipes that are near to the structure or component under investigation. The jet force due to a double-ended break acts along the pipe axis, which may change its direction. The jet force due to a slot break may act in any radial direction normal to the pipe axis.
CALVERT CLIFFS UFSAR 5.1-41 Rev. 47
- 2.
Both the double-ended and the slot break expose an area equal to the inside cross-sectional area of the pipe ruptured. The length of a slot break is equal to twice the diameter of the ruptured pipe.
- 3.
The magnitude of the jet force is computed by the following equation, and does not change with distance to target:
F = 2.0 x Operating Pressure x Inside Cross-Sectional Area of Pipe
- 4.
The area of the jet plume diverges at a half angle of 10 from the pipe opening. The jet force is evenly distributed to the area of plume at any distance to obtain the impinging pressure on the structure.
- 5.
A spectrum of all possible breaks are considered for a particular structure, and the ones that produce maximum stresses in the structure are used for design.
- d.
Erosion effects of jet spray.
- e.
Pipe whipping following a break in the pipe.
- f.
Rapid rise in ambient temperature to 276F and accompanying rise in ambient pressure to 50 psig. All Containment Structures have been evaluated for the revised maximum vapor temperature referenced in Section 14.20.
- g.
Missiles as described in Section 5.1.5.3.
The containment interior structures were analyzed by utilizing a lumped parameter model which treated the exterior shell, base slab, foundation media and the interior structures as a coupled system, and included details of the interior structures. The ground response spectra, as shown in Section 2.6.3, were used to determine the shears, moments, forces and displacements in the various structural, components, expected to be induced by the SSE and OBEs. The method employed is basically the analysis through the response of the normal modes to spectral accelerations, and is outlined in Section 5.1.3.2 as part of the "Containment Structure Design Analysis." Seismic induced shears, moments and forces are then multiplied with the load factors for Seismic Category I structures as indicated in Section 5A.3.1.
When checking each component, particular attention has been paid to tall and slender, or tall and cantilevered portions for possible further amplification. Such components were investigated by utilizing the floor response spectra.
Seismic analysis for the interior structures was done using procedures outlined in Section 5.1.3.2.
Where concrete structures such as the primary shield wall are subjected to sustained internal heat buildup, mechanical cooling devices are included to keep the internal temperatures below 150F. Localized concrete temperatures up to 200F are acceptable.
The total heat generation rate in the primary shield includes contributions from the following radiation sources: fission neutrons and gammas; core secondary gammas; secondary gammas from the core shroud, core support barrel, the reactor vessel, and the intermediate water regions; and secondary gammas in the concrete primary shield. Secondary gamma sources due to neutron scattering and capture were based on neutron distributions calculated using methods described in Chapter 3. The computer program QAD-P5 incorporates a point-kernel numerical integration method for the gamma radiation and a point-kernel moments method for each of the source contributions listed above and the heating rates due to the several contributions summed. The temperature gradient of the primary shield wall resulting from radiation-generated heat is combined with the
CALVERT CLIFFS UFSAR 5.1-42 Rev. 47 temperature gradient resulting from convective heat from the reactor to form a combined temperature gradient in the primary shield wall. The combined gradient was used in the thermal stress analysis and the resulting stresses in the wall were determined by using the design methods established by ACI 505-54. These stresses were combined with the stresses resulting from the analysis of loading combinations as described in Section 5.1.5.1. The critical stresses were then used to evaluate required reinforcing in both directions. All stresses are within the allowable as established Appendix 5A.
5.1.5.2 Design Loads and Materials
- a.
Design Loads:
The reactor cavity wall, the steam generator compartment walls, and containment interior structures are all Seismic Category I structures, the final design of which satisfied the most severe of the load combination equations of Appendix 5A.
- b.
Materials:
The following materials have been used in the construction of the Containment Structure interior:
Concrete
= 5000 psi @ 28 days Rebar A615, Gr 60 Plate Steel A441 & A533 Structural Steel A36 & A441 Anchor Bolts A325, A354 & A490 5.1.5.3 Missile Protection Inside Containment High pressure RCS components which could be a source of missiles are suitably screened, either by the concrete shield wall enclosing the reactor coolant loops, by the concrete operating floor or by special missile shields, to block any passage of missiles to the containment walls. Potential missile sources are oriented so that the missile will be intercepted by the shields and structures provided. A shield is provided over the control rod drive mechanism to block any missiles generated from postulated fracture of the nozzles.
All internally-generated missiles inside the Containment Structure are listed in Table 5-2. For all other internally-generated missiles outside the Containment Structure, see Section 5.3.1.1. All Seismic Category I structures or parts of structures have been checked for impact of missiles. The modified Petry formula was used for checking the effects of missiles on the structure at the point of impact. Even under extreme assumptions, missiles do not penetrate the concrete barriers completely.
Missile protection inside the Containment Structure is provided to comply with the following criteria:
- a.
The Containment Structure and liner are protected from loss-of-function due to damage by such missiles as might be generated in a LOCA for break sizes up to and including the double-ended severance of a reactor coolant pipe.
- b.
The engineered safety features and components required to maintain containment integrity are protected against loss-of-function due to damage by such missiles.
CALVERT CLIFFS UFSAR 5.1-43 Rev. 47 Missile protection necessary to meet the above requirements was implemented using the following methods:
- a.
Components of the RCS were examined to identify and to classify missiles according to size, shape, and kinetic energy for purpose of analyzing their effects.
- b.
Missile velocities were calculated considering both fluid and mechanical driving forces which can act during missile generation.
- c.
The RCS is surrounded by reinforced concrete and steel structures designed to withstand the forces associated with the double-ended rupture of a reactor coolant pipe and designed to stop missiles.
- d.
The structural design of the missile shielding takes into account impact loads and is based upon the state of the art of missile penetration calculational techniques.
The types of missiles for which protection is provided are:
- a.
Valve stems;
- b.
Valve bonnets;
- c.
Instrument thimbles; and
- d.
Various types and sizes of nuts and bolts.
Protection is not provided for certain types of missiles for which postulated accidents are considered incredible because of the material characteristics, inspections, quality control during fabrication, and conservative design as applied to the particular component. Included in this category are missiles caused by massive, rapid failure of the reactor vessel, steam generator, pressurizer, and main coolant pump flywheels and casings.
It is not expected that the polar crane will become or generate an internal missile.
The crane is anchored with seismic lugs and the crane and its components are designed for the OBE and SSE. Additionally, design features and administrative controls are in place to minimize the probability of a load drop from the polar crane. A description of our means of controlling heavy loads is presented in Section 5.7.
5.1.5.4 Thermal Gradients Ventilation and cooling systems maintain thermal gradients at a level low enough to have very small structural effects on the concrete walls. These effects are, nevertheless, considered in the design.
Reactor Cavity Wall Energy is deposited in the concrete of the reactor cavity wall by nuclear radiation emanating from the surface of the reactor vessel. In the equilibrium condition, assuming no axial leakage, all the deposited energy is removed either by the air-cooling system at the inner surface of the cavity wall or by convection from the outer surface. In order to find the temperature distribution within the concrete, an analytical expression for the energy deposition distribution is first derived. This expression is substituted in the heat diffusion equation which is solved for temperature using the temperature boundary conditions at the inner and outer face of the concrete.
CALVERT CLIFFS UFSAR 5.1-44 Rev. 47 Steam Generator Compartments The ventilation system provided eliminates temperature gradients across the secondary shield walls, across the refueling pool walls and across the operating floor.
5.1.5.5 Differential Pressures Generally, the occurrence of a LOCA is postulated to result from a rupture of the primary system piping either within the reactor cavity or the steam generator compartments of the containment. A pipe rupture of this type, within a compartment, results in the expulsion of high enthalpy water that flows out of the ruptured pipe, flashing partly to steam. As the pressure builds up within the compartment, the steam-air-water mixture flows through openings into the main containment. The maximum pressure differential will depend on the number and shape of the openings leading between the compartments, the volume of each compartment, and the blowdown rate from the broken pipe.
Differential pressure analyses are made with the Bechtel computer program COPRA, which calculates the transient pressure responses of two containment compartments during a LOCA. COPRA calculates a mass and energy balance of the two-phase, two-component steam-water-air mixture as primary coolant enters the compartment during the LOCA and exits through vents and openings into the main containment building. There is no provision in the code for heat transfer to structures or for engineered safety features. These options generally have a negligible effect on compartment pressures for the short time following the rupture within which peak differential pressures occur.
Because of this short time interval between the LOCA and the peak compartment pressure differential, each case analyzed is divided into very small time intervals, generally equal to or less than 0.1 msec. The thermodynamic equations are solved for each time advancement.
The steam, water and air throughout each compartment are in thermal equilibrium at all times. Water, steam, and air entering a compartment are mixed homogeneously and instantaneously; no accumulation of water occurs on the walls or in the sump.
A COPRA Summary Document which lists and discusses the program assumptions, is contained in Reference 9. This document is on file with the Nuclear Regulatory Commission (NRC).
The discussion of the results of the analyses performed for each interior containment compartment is contained in FSAR [Final Safety Analysis Report]
Section 14.16.5, as revised by FSAR Amendment No. 29, Revision 18 (currently Updated Final Safety Analysis Report Section 14.20).
The compartment walls have been designed for jet impingement forces and differential pressures, both assumed to occur simultaneously with the SSE.
This design has the following conservative considerations:
- a.
A dynamic load factor of two is applied to the maximum jet load. (The jet load is based on sudden severance of the primary loop piping nearest to the wall.)
- b.
The differential pressure and jet impingements are considered to be in phase.
CALVERT CLIFFS UFSAR 5.1-45 Rev. 47 As dictated by design, any transient response of the containment compartment walls is expected to contain the most severe accident postulated.
Reactor Cavity There are two types of openings for flow out of the reactor cavity into the main containment volume:
- a.
Openings around the main coolant pipes - 40.0 ft2
- b.
Openings around the reactor vessel seal ledge - 173 ft2 The current plant design includes a permanent refueling pool seal that further restricts the opening around the reactor vessel. This analysis is provided as historical design information only (Reference 7).
The areas around the main coolant pipes extend along the annular space between the pipes and the walls of the pipe tunnels. A calculated nozzle flow coefficient of 0.61, based on the geometry of these flow spaces, is utilized. The opening around the refueling seal ledge is treated as an orifice. Flow coefficients for orifice geometries are supplied automatically by the COPRA computer code.
Immediately after a pipe rupture, the volume available to the expanding steam within the reactor cavity is the free volume below the seal ledge.
Both guillotine and slot primary pipe failures have been considered:
- a.
For the guillotine break, the coolant pipes are partially restrained by the tunnel walls. Restraining the primary pipes in this manner gives a flow area of 0.61 ft2 for failure of the 42" line, which results in a cavity peak differential pressure of 23 psi.
- b.
The slot longitudinal break has a length of twice the inside pipe diameter and an area equal to the pipe cross-sectional area. The flow from the part of the break within the pipe tunnel is divided between the reactor cavity and the steam generator compartment, while the flow from the part of the break that is between the primary shield wall and the nozzle of the vessel enters the reactor cavity only. The total flow of mass and energy into the reactor cavity is approximately 27% of the flow from a 42" single-ended rupture.
As the reactor cavity pressure builds up, the flow leaving the pipe within the pipe tunnel will enter the steam generator compartment; however, credit for this effect is conservatively ignored to simplify the problem. The peak differential pressure resulting from the slot failure is 31 psi.
The following table summarizes the input parameters used for the computer analysis of reactor cavity pressurization:
Main containment free volume 2.0x106 ft3 Reactor cavity free volume 5.135x103 ft3 Nozzle-type relief area (around piping) 40.0 ft2 Nozzle coefficient 0.61 Orifice-type relief area (around seal ledge) 173.0 ft2 Initial containment temperature 110F Initial containment pressure 14.7 psia Initial containment humidity 50%
CALVERT CLIFFS UFSAR 5.1-46 Rev. 47 Note that this analysis was redone for the neutron shield with no significant differences.
Steam Generator Compartments For the steam generator compartments, only the double-ended guillotine break in the hot leg is considered since it provides the largest rate of energy and mass release. The relief areas are divided into three classes: long, sharp-edged nozzles, sharp-edged orifices and well-rounded orifices. An orifice coefficient of 0.61 is applied to the long nozzles and a coefficient of 0.97 is applied to the well-rounded orifices; coefficients for sharp-edged orifices are supplied by the computer. The compartment volumes and total flow areas are listed below:
Compartment Volume Long Sharp
-edged Nozzles Sharp
-edged Openings Well Rounded Openings P
East (a)53,980 ft3 356 ft2 770 ft2 1,072 ft2 16.4 psi West 51,500 ft3 182 ft2 414 ft2 1,072 ft2 18.4 psi (a)
East volume is greater than west because east includes pressurizer as well as one steam generator and two reactor coolant pumps.
The initial conditions and the main compartment volume are the same as those used in the reactor cavity.
5.1.5.6 Design Bases Temperature and Pressure Differentials Safety factors were not determined or established for pressure and temperature acting alone. The design and analysis methods as applied to the interior structures, accounted for and investigated the effects of temperature and pressure in conjunction with the other loads occurring in those conditions as stated in Appendix 5A. The associated individual safety margins can best be examined by:
discussing the contribution of these loads to the total stress intensities, or to the total ultimate load carrying capacity of the structure or component under the two main categories of design loading conditions.
The two main design loadings conditions and the serviceability required of the interior structures in each case are that:
- a.
They provide support during plant operation and prevent the occurrence of a LOCA from:
- 1.
Normal operating loads (including dead loads)
- 2.
Design live loads
- 3.
Hydrostatic loads
- 4.
Thermal loads
- 5.
OBE seismic loads
- b.
They mitigate its consequences, should a LOCA occur, by protecting the containment and all engineered safety features systems from:
- 1.
Blowdown forces (including jet thrust and impingement loads)
- 2.
Whipping pipes
- 3.
Missiles
- 4.
Differential pressures
CALVERT CLIFFS UFSAR 5.1-47 Rev. 47
- 5.
SSE seismic loads
- 6.
Loads a.1, and others, if applicable The safety factor of the structure is the ratio of allowable stresses to the computed stresses in case (a) above, and the ratio of the ultimate load resisting capacity of the structure to the total applied loads in case (b).
In general, this ratio is not the same for all components of the interior structure and varies with both the modes of stress or loading and the hazard associated with the structure or component thereof (i.e., some of the components that were governed by radiological shielding requirements have capacities much higher than the design loads; uniaxial, flexural normal, shearing torsional, tensile and compressive stresses were treated differently). Further, the loads above were combined per Appendix 5A load combination equations which have an inherent factor of safety defined by their load coefficients. No credit is taken for this safety factor.
For the operating loads, the minimum safety margin is as determined from ACI 318-63 code for reinforced concrete, and AISC Specification, 1963, for structural steel. The contribution of the operating temperatures is negligible in all components except the reactor cavity wall, for which additional reinforcing steel was furnished for the thermal stresses in accordance with the requirements of ACI 505-59. The accident condition controlled the design of most of the interior structures, including the reactor cavity wall, the secondary shield walls which house the RCS, and all major vessel support structures. Where stresses due to blowdown forces were in the same direction and thus combined with the stresses due to other loads including temperature and pressure, the share of the former ranged from 60 to 80%, thus resulting in an increase of 2.5 to 5 times the inherent safety factor for the latter loads acting alone. Where temperature and pressure stresses acted alone, the safety factors inherent in the load combination equations were increased by 3.0 and 1.5 in the operating and accident cases respectively, for both reinforcing and structural steel. In all cases, the safety margin for concrete stresses were kept higher than steel in order to preclude brittle modes of failure.
The calculated differential pressures and temperature gradients, and those that have been used as the design bases of the major interior structures are listed below:
P (Design)
P (Calculated)
T (Design)
T (Calculated)
East Steam Generator Compartment Walls 28 psi 16.4 psi None None West Steam Generator Compartment Walls 28 psi 18.4 psi None None Reactor Cavity Wall 96 psi 31.0 psi(a) 30F/7' 6.7F/7' (a)
The current plant design includes a permanent refueling pool seal which further restricts the opening around the reactor vessel. This analysis result is provided as historical design information only (Reference 7).
5.1.6 LIGHTNING PROTECTION OF CONTAINMENT Lightning protection is provided over the dome of the containment vessel and consists of the following:
CALVERT CLIFFS UFSAR 5.1-48 Rev. 47 One 1/2"x24" copper air terminal is placed in the center of the dome. Eight equally spaced 1/2"x25" air terminals are installed on a 20'-diameter circumference. These are connected to closed loop of 2/0 AWG cable. Four cables of like size connect the loop with the center air terminal. Another 16 equally spaced 1/2"x24" air terminals are installed on 40' circumference. These are connected to a closed loop of 2/0 AWG cable. Another 20 equally spaced 1/2"x24" air terminals are installed on a 69'3" radius circumference.
These are connected to a closed loop of 2/0 AWG cable. Four equally spaced 1/2"x24" air terminals connect the three loops. Four 2/0 AWG down conductors connect the outer loop to the ground grid. Ground guards are provided to protect down conductors to an approximate height of 8' above grade.
5.1.7 CATHODIC PROTECTION The containment cathodic protection system and related environmental or galvanic corrosion protection schemes are explained in the following discussion.
Heavy waterproofing (40 mils thickness) membrane is used underneath the containment base slab, within the 4"-thick mud mat and around the cylindrical shell up to finish grade. This membrane is protected to full height by 1/2" (minimum) thick asphalt-impregnated fiber board prior to backfilling. Remotely-located shallow anodes are used in cathodic protection system. Anodes placed approximately 100' from the membrane could protect the structure regardless of the total cross-section of apertures and would provide protection as required for pipes and tank bottoms. Zinc and/or copper sulphate permanent reference electrodes are located in the soil to enable measurement of protective gradients around the foundation. This scheme is based on the concept that the containment building foundation steel is formed with an inhibitive concrete cover, a heavy waterproofing membrane and a relatively low chloride/oxygen environment. Steel embedded in or covered by concrete in the containment sub-structure will be protected by the waterproof membrane. The system is conservatively designed for a 40-year life, derating manufacturer's recommendations for anodes by approximately 50%.
The surface of the exposed liner plate was protected by an initial cleaning by wire brushing or sandblasting, and then coated with Amercoat zinc paint or its equivalent at the fabrication plant. After erection, all exposed areas adjacent to field welds were cleaned or reprimed, and the entire exposed face was painted with a finish coat. The outside of the liner is protected by a minimum of 3-1/4' of prestressed concrete, which is cast against the liner offering a very high degree of corrosion protection.
The ACI Building Code Requirements for Reinforced Concrete (ACI 318-63),
Section 808, Concrete Protection For Reinforcing, have proved to be effective for preventing chemical corrosion of concrete reinforcement. The reinforcing concrete protection exceeds Section 808 requirements by 50% or more. Tendon sheathing has a minimum of 11" of concrete cover at any location in the Containment Structure.
5.1.8 LEAKAGE MONITORING SYSTEM No continuous leakage monitoring system is provided.
The barrier to leakage in the Containment Structure is the 1/4" steel liner plate. All penetrations were continuously welded to the liner plate before the concrete in which they are embedded was placed. These penetrations, shown on Figures 5-2
CALVERT CLIFFS UFSAR 5.1-49 Rev. 47 and 5-3, became an integral part of the liner and were so designed, installed, and tested.
The steel liner plate is securely attached to and is an integral part of, the prestressed concrete Containment Structure which is conservatively designed and rigorously analyzed for the extreme loading conditions of the postulated LOCA, as well as for all other types of anticipated loading conditions. Thorough controls were maintained over the quality of all materials and workmanship during all stages of fabrication and erection of the liner plate and penetrations, and during construction of the entire Containment Structure.
The comprehensive program for preoperational testing, inspection, and post-operational surveillance is described in detail in Section 5.5 and is summarized in the following paragraphs.
During construction, the entire length of every seam weld in the liner plate was leak tested. Individual penetration assemblies were shop tested. Welded connections between penetration assemblies and the liner plate were individually leak tested after installation. Following completion of construction, the entire Containment Structure, the liner and all its penetrations were tested at 115% of the design pressure to establish structural integrity. The initial leak rate test of the entire structure was conducted at 50 and 100% of the calculated peak pressure to demonstrate vapor tightness, and to establish a reference for periodic leak testing for the life of the plant.
Penetrations such as the permanent equipment hatch opening and personnel access locks cannot be opened except by deliberate action. The personnel access locks are interlocked and provided with alarm devices so that the Containment Structure cannot be breached unintentionally. The liner plate over the foundation slab is protected by cover concrete. Wherever access to the liner plate welds is not possible, means are provided so that they can be tested for leakage. The liner plate is protected against corrosion by suitable coatings and by cathodic protection. Walls and floors for biological and missile shielding also provide protection for the liner plate.
Once the adequacy of the liner plate was established initially, there is no reason during the life of the plant to anticipate progressive deterioration which would reduce the effectiveness of the liner as a vapor barrier. Inside the Containment Structure, the atmosphere is subject to a high degree of temperature control. The outside of the liner is protected by prestressed concrete which is resistant to all weather conditions.
Inspection on a periodic basis, as necessary, may be conducted in all accessible areas during full power operation. Biological shielding is provided to reduce direct radiation from the RCS to acceptable limits. A visual inspection of the Containment Structure, from the inside and outside, will be conducted during regularly scheduled unit shutdowns.
All penetrations, except the following, are located in groups, and a penetration room is located at each group.
- a.
Equipment Hatch Opening
- b.
Personnel Access Lock
- c.
Emergency Personnel Access Lock
- d.
Refueling Tube
CALVERT CLIFFS UFSAR 5.1-50 Rev. 47
- e.
Purge Line Inlet and Outlet
- f.
Containment Emergency Sump Any leakage that might occur from these penetrations will be collected and exhausted through vents or drains. In this manner, leakage which might occur from these groups of penetrations will be isolated from leakage which might occur through the Containment Structure, itself.
Provisions are made such that these penetrations may be pressure tested for leakage during normal operation.
The containment normal sump penetration is another exception that is not located in a penetration room. However, there are no provisions made to collect and exhaust leakage from this penetration or to enable pressure testing during normal operation. This penetration is further described in Section 5.2.2.
Within the penetration rooms, provisions are made such that individual penetration assemblies with resilient seals may be pressure tested for leakage during normal operation. This degree of control over leakage through penetrations greatly reduces the probability of undetected leakage for the Containment Structure as a whole.
A considerable background of operating experience is being accumulated on containment and penetrations. Full advantage of this knowledge is taken in all phases of design, fabrication, installation, inspection, testing, and operation.
For the foregoing reasons, it has been concluded that a continuous leakage monitoring system is unnecessary. Since there is no such system provided, there can be no misoperation or malfunction, which in itself might constitute a hazard.
The steel-lined Containment Structure is self-sufficient, and other than valves and hatch doors, there are no operating parts. The containment boundary is extended only by specific penetrations which are further described and tabulated in Section 5.2, "Isolation System."
5.
1.9 REFERENCES
- 1.
A.C. Eringen and A.K. Naghdi, "State of Stress in a Circular Cylindrical Shell with a Circular Hole"
- 2.
Levy, Samuel, A.E. McPherson, and F.C. Smith, "Reinforcement of a Small Circular Hole in a Plane Sheet Under Tension," Journal of Applied Mechanics, June 1948
- 3.
K.R. Wickman, A.G. Hopper, and J.L. Mershon, "Local Stress in Spherical and Cylindrical Shells due to External Loadings," Welding Research Council Bulletin No. 107, August 1965
- 4.
HTGR and Laboratory Staff, Prestressed Concrete Reactor Vessel, Model 1, GA7097
- 5.
Advanced HTGR Staff, Prestressed Concrete Reactor Vessel, Model 2, (GA7150)
- 6.
CEN-367-A, Leak Before Break Evaluation of Primary Coolant Piping in Combustion Engineering Designed NSSS, February 1991
- 7.
Letter from Mr. D. G. McDonald (NRC) to Mr. R. E. Denton (BGE), dated February 3, 1994, Installation of a Neutron Shield/Pool Seal at the Calvert Cliffs Nuclear Power Plant, Unit Nos. 1 and 2 (TAC Nos. M87176 and M87177)
CALVERT CLIFFS UFSAR 5.1-51 Rev. 51
- 8.
Letter from NRC to Mr. E. C. Sterling, III, Chairman CEOG, dated October 30, 1990, Acceptance for Referencing Topical Report CEN-367, Leak-Before-Break Evaluation of Primary Coolant Loop Piping in Combustion Engineering Designed Nuclear Steam Supply Systems
- 9.
Containment Pressure Analysis, NS731 TN, December 1968
- 10.
Letter from C. H. Cruse (BGE) to Document Control Desk (NRC), dated April 2, 1999, First Annual Amendment to Application for License Renewal
CALVERT CLIFFS UFSAR 5.1-52 Rev. 47 TABLE 5-1 STRESS ANALYSIS RESULTS(8)
STRUCTURAL DATA LOCATION CONCRETE REINFORCING STEEL
- psi t - in Pm - %
Ph - %
A 5000 39
.260
.094 B
5000 39
.619
.214 C
5000 55.6
.159
.150 D
5000 55.6
.253
.232 E
5000 138.12
.060
.060 F
5000 138.12
.286
.112 G
5000 45
.316
.185 H
5000 45
.741
.289 J
5000 45
.235
.235 K
5000 45
.235
.235 L
5000 70.2
.315
.247 M
5000 70.2
.570
.247 N
4000 120
.223
.185 O
4000 120
.447
.479 P
4000 120
.286
.222 Q
4000 120 1.144
.514 KEY ELEVATION (Showing location of reference sections)
CALVERT CLIFFS UFSAR 5.1-53 Rev. 47 TABLE 5-1 STRESS ANALYSIS RESULTS(8)
ALLOWABLE STRESSES WORKING STRESS DESIGN YIELD STRESS DESIGN SHELL CONCRETE:
fa = 1500 psi fa = a (fc) = (0.85) (5000) = 4,250 psi fce = 3000 psi fce = ce (fc) = (0.90) (5000) = 4,500 psi BASE CONCRETE:
fce = 1800 psi fa = a (fc) = (0.85) (4000) = 3400 psi fce = ce (fc)= (0.90) (4000) = 3600 psi STEEL: A615, GR 60 fs = 30,000 psi fs = (fy) = (0.90) (60,000) = 54,000 psi
CALVERT CLIFFS UFSAR 5.1-54 Rev. 47 TABLE 5-1 STRESS ANALYSIS RESULTS(8)
D + F + L (Stresses in psi) Construction Case I MERIDIONAL HOOP SHEAR SECTION e
OUTSIDE e
INSIDE a
AXIAL e
OUTSIDE e
INSIDE a
AXIAL vci vcw Shell A - B
-827
-1630
-1256
-814
-1609
-1207
-23 105 687 C - D
+243
-1630
-609
+77
-734
-250
+32 45 452 E - F
-248
-784
-464
+136
-863
-330
+15 43 360 G - H
-1
-26
-618
-55
-1154
-613
-40 104 513 J - K
-1
-8
-664
-575
-1645
-1120
-8 69 427 L - M 2
-27
-499
-481
-645
-583 6
98 359 Base N - O 113
-251
-41 565
-426
-29 32 138 157 P - Q 560
-971
-69 565
-470
-12 44 135 123
CALVERT CLIFFS UFSAR 5.1-55 Rev. 47 TABLE 5-1 STRESS ANALYSIS RESULTS(8)
CONTAINMENT STRUCTURE -
SUMMARY
OF CONCRETE AND REINFORCING STEEL STRESSES (psi)
CONCRETE COMPUTED COMPUTED VS.
ALLOWABLE SECTION LOAD CASE em eh am ah
II D+F+L+1.15P
-997
-935
-591
-542
-6 0.332 0.394 0.017 III D+F+L+TO+E
-2141
-2363
-1307
-1324 20 0.788 0.883 0.056 A-B IV D+F+L+TA+P
-2990
-2985
-545
-510 7
0.993 0.363 0.020 V
1.05D+F+1.5P+ TA
-2610
-2213
-215
-61 2
0.580 0.051 0.003 VI 1.05D+F+1.25P+ TA+1.25E
-2048
-2205
-652
-519 24 0.490 0.153 0.040 VII D+F+P+ TA+E'
-2247
-2392
-804
-732 17 0.532 0.189 0.030 II D+F+L+1.15P
-814
-573
-240 50 68 0.271 0.160 0.200 III D+F+L+TO+E
-2210
-1367
-645
-293 65 0.737 0.430 0.180 C-D IV D+F+L+TA+P
-2820
-2539
-224
-35 83 0.940 0.149 0.230 V
1.05D+F+1.5P+ TA
-1970
-2058
-94 28 128 0.457 0.022 0.220 VI 1.05D+F+1.25P+ TA+1.25E
-2444
-2183
-145 25 171 0.543 0.034 0.280 VII D+F+P+ TA+E'
-2848
-2564
-226
-39 125 0.633 0.053 0.210 II D+F+L+1.15P
-357
-709
-293
-235 15 0.236 0.195 0.040 III D+F+L+TO+E
-232
-1409
-471
-293 15 0.469 0.314 0.040 E-F IV D+F+L+ TA+P
-346
-2984
-269
-133 7
0.995 0.179 0.020 V
1.05D+F+1.5P+ TA
-451
-2626
-197 42 23 0.584 0.046 0.040 VI 1.05D+F+1.25P+ TA+1.25E
-374
-2843
-231
-25 20 0.632 0.054 0.033 VII D+F+P+ TA+E'
-349
-3014
-272
-135 11 0.670 0.064 0.018
CALVERT CLIFFS UFSAR 5.1-56 Rev. 47 TABLE 5-1 STRESS ANALYSIS RESULTS(8)
CONTAINMENT STRUCTURE -
SUMMARY
OF CONCRETE AND REINFORCING STEEL STRESSES (psi)
REINFORCING STEEL COMPUTED COMPUTED VS.
ALLOWABLE SECTION LOAD CASE m
h
II D+F+L+1.15P
-269
-16 0.009 0.001 III D+F+L+TO+E 2098 694 0.070 0.023 A-B IV D+F+L+TA+P
+3590 4253 0.120 0.142 V
1.05D+F+1.5P+TA 4760 4512 0.088 0.084 VI 1.05D+F+1.25P+TA+1.25E 14440 28664 0.267 0.531 VII D+F+P+TA+E' 9888 16483 0.183 0.305 II D+F+L+1.15P 1750 4786 0.058 0.160 III D+F+L+TO+E 6282 10485 0.209 0.349 C-D IV D+F+L+TA+P 4040 15967 0.134 0.532 V
1.05D+F+1.5P+TA 147 19229 0.003 0.356 VI 1.05D+F+1.25P+TA+1.25E 1475 18794 0.027 0.348 VII D+F+P+TA+E' 4084 16127 0.076 0.298 II D+F+L+1.15P 124 2148 0.004 0.072 III D+F+L+TO+E
-276 5701 0.01 0.190 E-F IV D+F+L+TA+P 1110 8667 0.037 0.289 V
1.05D+F+1.5P+TA 4080 10259 0.076 0.190 VI 1.05D+F+1.25P+TA+1.25E 2343 10036 0.0434 0.186 VII D+F+P+TA+E' 1121 8754 0.021 0.162
CALVERT CLIFFS UFSAR 5.1-57 Rev. 47 TABLE 5-1 STRESS ANALYSIS RESULTS(8)
CONTAINMENT STRUCTURE -
SUMMARY
OF CONCRETE AND REINFORCING STEEL STRESSES (psi)
CONCRETE COMPUTED COMPUTED VS.
ALLOWABLE SECTION LOAD CASE em eh am ah
II D+F+L+1.15P
-69
-746
-154
-224
-13 0.249 0.149 0.037 III D+F+L+TO+E
-1415
-1770
-683
-793 61 0.590 0.529 0.170 G-H IV D+F+L+TA+P
-101
-2827
-68
-195
-11 0.942 0.130 0.310 V
1.05D+F+1.5P+TA
-96
-1766 114 23
-66 0.392 0.027 0.110 VI 1.05D+F+1.25P+TA+1.25E
-1453
-1496
-162
-221
-20 0.332 0.052 0.033 VII D+F+P+TA+E'
-1529
-1683
-194
-364 1
0.374 0.086 0.002 II D+F+L+1.15P
-60
-653
-201
-133 1
0.218 0.134 0.003 III D+F+L+TO+E
-1821
-2244
-732
-1209
-9 0.747 0.806 0.250 J-K IV D+F+L+TA+P
-67
-2629
-111
-99
-2 0.876 0.074 0.006 V
1.05D+F+1.5P+TA
-96
-2828 169 420 2
0.628 0.099 0.003 VI 1.05D+F+1.25P+TA+1.25E
-1488
-957
-209
-75 5
0.331 0.049 0.008 VII D+F+P+TA+E'
-1697
-1669
-311
-447
-3 0.377 0.105 0.005 II D+F+L+1.15P
+10
-848
-277
-718 1
0.283 0.479 0.003 III D+F+L+TO+E
-621
-2012
-494
-635
-2 0.671 0.423 0.006 L-M IV D+F+L+TA+P
-29
-2914
-254
-731 147 0.971 0.487 0.415 V
1.05D+F+1.5P+TA
-84
-2973
-263
-465 163 0.661 0.109 0.271 VI 1.05D+F+1.25P+TA+1.25E
-1633
-3510
-197
-914 14 0.780 0.215 0.023 VII D+F+P+TA+E'
-1427
-3360
-296
-870
-6 0.745 0.205 0.010
CALVERT CLIFFS UFSAR 5.1-58 Rev. 47 TABLE 5-1 STRESS ANALYSIS RESULTS(8)
CONTAINMENT STRUCTURE -
SUMMARY
OF CONCRETE AND REINFORCING STEEL STRESSES (psi)
REINFORCING STEEL COMPUTED COMPUTED VS.
ALLOWABLE SECTION LOAD CASE m
h
II D+F+L+1.15P 2884 2284 0.096 0.076 III D+F+L+TO+E 2616 3332 0.087 0.111 G-H IV D+F+L+TA+P 16649 10293 0.555 0.343 V
1.05D+F+1.5P+TA 7381 15437 0.136 0.285 VI 1.05D+F+1.25P+TA+1.25E 23724 28024 0.440 0.520 VII D+F+P+TA+E' 23595 20489 0.437 0.379 II D+F+L+1.15P 2368 2882 0.080 0.096 III D+F+L+TO+E 5283 351 0.176 0.012 J-K IV D+F+L+TA+P 11580 11410 0.386 0.380 V
1.05D+F+1.5P+TA 8720 10523 0.161 0.195 VI 1.05D+F+1.25P+TA+1.25E 29043 26964 0.537 0.499 VII D+F+P+TA+E' 25474 14396 0.471 0.266 II D+F+L+1.15P 22007
-5482 0.734 0.183 III D+F+L+TO+E 3366 1812 0.112 0.0604 L-M IV D+F+L+TA+P 11216 3168 0.374 0.105 V
1.05D+F+1.5P+TA 13369 3063 0.247 0.057 VI 1.05D+F+1.25P+TA+1.25E 19995 9888 0.370 0.183 VII D+F+P+TA+E' 10770 9651 0.199 0.178
CALVERT CLIFFS UFSAR 5.1-59 Rev. 47 TABLE 5-1 STRESS ANALYSIS RESULTS(8)
CONTAINMENT STRUCTURE -
SUMMARY
OF CONCRETE AND REINFORCING STEEL STRESSES (psi)
CONCRETE COMPUTED COMPUTED VS.
ALLOWABLE SECTION LOAD CASE em eh am ah
II D+F+L+1.15P 69
-362 68 76 127 0.201 0.600 III D+F+L+TO+E
-911
-1204
-64
-17 48 0.669 0.220 N-O IV D+F+L+TA+P
-207
-594 26 58 114 0.330 Note 0.540 V
1.05D+F+1.5P+TA 7
-704 20 86 149 0.196 (a) 0.413 VI 1.05D+F+1.25P+TA+1.25E
-861
-1203 54 52 164 0.334 0.454 VII D+F+P+TA+E'
-1098
-1198 25 52 154 0.333 0.427 II D+F+L+1.15P
-1338
-473 56 113 74 0.743 0.350 III D+F+L+TO+E
-1562
-1540
-43
-63 43 0.868 0.202 P-Q IV D+F+L+TA+P
-1697
-599
-45 121 65 0.943 Note 0.306 V
1.05D+F+1.5P+TA
-2253
-829
-31 68 135 0.626 (a) 0.374 VI 1.05D+F+1.25P+TA+1.25E
-2115
-1086
-15 75 74 0.588 0.205 VII D+F+P+TA+E'
-1621
-1251
-42 46 79 0.450 0.220
CALVERT CLIFFS UFSAR 5.1-60 Rev. 47 TABLE 5-1 STRESS ANALYSIS RESULTS(8)
CONTAINMENT STRUCTURE -
SUMMARY
OF CONCRETE AND REINFORCING STEEL STRESSES (psi)
REINFORCING STEEL COMPUTED COMPUTED VS.
ALLOWABLE SECTION LOAD CASE m
h
II D+F+L+1.15P 256 20957 0.009 0.700 III D+F+L+TO+E 19168 26123 0.640 0.871 N-O IV D+F+L+TA+P 13172 26362 0.440 0.878 V
1.05D+F+1.5P+TA 16596 35106 0.307 0.650 VI 1.05D+F+1.25P+TA+1.25E 30315 41218 0.561 0.763 VII D+F+P+TA+E' 29196 40450 0.542 0.749 II D+F+L+1.15P 26933 24818 0.897 0.827 III D+F+L+TO+E 24970 30500 0.832 1.000 P-Q IV D+F+L+TA+P 26824 30012 0.894 1.000 V
1.05D+F+1.5P+TA 36983 42126 0.684 0.780 VI 1.05D+F+1.25P+TA+1.25E 35862 40468 0.664 0.749 VII D+F+P+TA+E' 29984 41033 0.555 0.759 NOTE (a):
fa = 0.3 not applicable to base slab.
CALVERT CLIFFS UFSAR 5.1-61 Rev. 47 TABLE 5-1 STRESS ANALYSIS RESULTS(8)
CONTAINMENT STRUCTURE -
SUMMARY
OF CONCRETE AND REINFORCING STEEL STRESSES (psi)
CONCRETE REINFORCING STEEL SECTION LOAD CASE em eh m
h A-B D + F + L + To
-1860
-1845
-18
-190 111 D + F + L + To + E
-2141
-2363 20 2098 694 D + F + L + Ts
-1400
-1374
-23
-4470
-4419 D + F + L + Ts + E
-1466
-1526
-10
-10516
-10246 1.05D + F + L + 1.25P + TA
-653
-465 23 3700 4198 1.05D + F + L + 1.25P + TA + 1.25E
-2048
-2205 24 11731 25506 D + F + L + P + TA
-897
-794 7
3590 4253 D + F + L + P + TA + E'
-2247
-2393 17 9889 16483 C-D D + F + L + To 7
1 43
-12,200
-6074 D + E + L + To + E
-2281
-1393 65 8139 7912 D + F + L + Ts 137 220 32
-7880
-1838 D + F + L + Ts + E
-2283
-1392
-4 8101 7903 1.05D + F + L + 1.25P + TA 7
1 114
-4040 7353 1.05D + F + L + 1.25P + TA + 1.25E
-2006
-946 171 14042 10321 D + F + L + P + TA 8
1 83
-4340
-1033 D + F + L + P + TA + E'
-1545
-1209 125 9390 13818 E-F D + L + F + To
-242
-1266 10
-1396 5701 D + L + F + To + E
-836
-1542 15
-728 10234 D + F + L + Ts
-225
-681 15
-993 1337 D + F + L + Ts + E
-732
-723 30
-1135 901 1.05D + F + L + 1.25P + TA 321
-653 13
-1052 7427 1.05D + F + L + 1.25P + TA + 1.25E
-193
-2132 20
-750 52514 D + F + L + P + TA 262
-711 7
-809 9653 D + F + L + P + TA + E'
-336
-2012 11
-1691 49431
CALVERT CLIFFS UFSAR 5.1-62 Rev. 47 TABLE 5-1 STRESS ANALYSIS RESULTS(8)
CONTAINMENT STRUCTURE -
SUMMARY
OF CONCRETE AND REINFORCING STEEL STRESSES (psi)
CONCRETE REINFORCING STEEL SECTION LOAD CASE em eh m
h G-H D + F + L + To
-1321
-1522 41 2622 3475 D + F + L + To + E
-1415
-1770 61 2616 3332 D + F + L + Ts
-816
-1022
-40
-1187
-1466 D + F + L + Ts + E
-783
-1095 18 4114 4984 1.05D + F + L + 1.25P + TA
-2584
-2440
-32 14927 12730 1.05D + F + L + 1.25P + TA + 1.25E
-1452
-1495
-20 23724 28024 D + F + L + P + TA
-2660
-2712
-11 14972
-10960 D + F + L + P + TA + E'
-1529
-1682 1
23595 20489 J-K D + F + L + To
-1595
-2115
-6 4228 476 D + E + L + To + E
-1821
-2244
-9 5283 351 D + F + L + Ts
-927
-1407
-8
-511
-3931 D + F + L + Ts + E
-990
-1453 29 3960
-8747 1.05D + F + L + 1.25P + TA
-1029
-1648 57
-511 17727 1.05D + F + L + 1.25P + TA + 1.25E
-989
-957 5
3960 26965 D + L + F + TA
-2655
-2628
-2 11580 11410 D + L + F + TA + E'
-1696
-1669
-3 25473 14396 L-M D + F + L + To
-304 392
-15
-2422 3821 D + F + L + To + E
-621
-2012
-2 3366 1812 D + F + L + Ts
-1039
-614 6
961
-3751 D + F + L + Ts + E
-1117
-788
-16 833
-10188 1.05D + F + L + 1.25P + TA
-2005
-2980 165 37136 2914 1.05D + F + L + 1.25P + TA + 1.25E
-1633
-3510 14 19995 9888 D + F + L + P + TA
-1845
-2914 147 28111 3168 D + F + L + P + TA + E'
-1427
-3360
-6 10770 9651
CALVERT CLIFFS UFSAR 5.1-63 Rev. 47 TABLE 5-1 STRESS ANALYSIS RESULTS(8)
CONTAINMENT STRUCTURE -
SUMMARY
OF CONCRETE AND REINFORCING STEEL STRESSES (psi)
CONCRETE REINFORCING STEEL SECTION LOAD CASE em eh m
h N-O D + F + L + To
-708
-696 24 12251 15290 D + F + L + To + E
-910
-1070 48 19168 26123 D + F + L + Ts
-537
-492 32 8742 5101 D + F + L + Ts + E
-930
-1080 33 20719 27297 1.05D + F + L + 1.25P + TA
-484
-728 98 22724 30046 1.05D + F + L + 1.25P + TA + 1.25E
-861
-1203 164 30315 41218 D + F + L + P + TA
-763
-743 114 25316 28528 D + F + L + P + TA + E'
-1098
-1198 154 29196 40450 P-Q D + F + L + To
-1251
-688 22 17161 15582 D + F + L + To + E
-1562 1355 43 24970 30500 D + F + L + Ts
-911
-446 44 8506 3814 D + F + L + Ts + E
-1165
-630 60 17945 13720 1.05D + F + L + 1.25P + TA
-1821
-670 44 32542 30141 1.05D + F + L + 1.25P + TA + 1.25E
-2115
-1086 74 35862 40468 D + F + L + P + TA
-1697
-599 65 26824 30012 D + F + L + P + TA + E'
-1621
-1251 79 29984 41033
CALVERT CLIFFS UFSAR 5.1-64 Rev. 47 TABLE 5-1 STRESS ANALYSIS RESULTS(8)
NOTES:
- 1.
Loading Cases I, II, III, & IV are working stress analysis whereas Loading Cases V, VI, & VII are yield stress analysis.
- 2.
For notation see next page.
- 3.
All concrete extreme fiber stresses sc are for the inside surface. Outside surfaces stresses are indicated by ( ). The stresses listed are the controlling stresses for that section.
- 4.
Computed vs. allowable ratios for Cases V, VI, and VII include appropriate factors, e.g.,f a/fa = a/a(fc).
- 5.
Allowable shear stresses include stirrups whenever applicable.
- 6.
Deviations in allowable stresses are in accordance with Section 5.1.2.2.
- 7.
Reinforcing steel is Type A615, GR 60.
- 8.
See Appendix 5E for later tables associated with an evaluation to reduce the original amount of required containment prestress.
NOTATION D
Dead load F
Prestress P
LOCA pressure load E
E' Earthquake (SSE)
TA Accident temperature fc Concrete compressive strength at age of 28 days fy Steel re-bar yield stress fa Allowable concrete axial stress fce Allowable concrete axial & flexure stress y
Allowable concrete shear stress including stirrups if applicable fs Allowable steel stress a
Nominal membrane stress e
Combined axial & flexure nominal stress Actual shear stress h
Subscript indicating hoop direction m
Subscript indicating meridional direction ph Hoop steel percentage pm Meridional steel percentage
CALVERT CLIFFS UFSAR 5.1-65 Rev. 47 TABLE 5-1 STRESS ANALYSIS RESULTS(8)
+
Tensile stresses Compressive stresses Cracked section analysis ci Section 5.1.2.2 (loads necessary to cause structural yielding) cw Section 5.1.2.2 (loads necessary to cause structural yielding)
CALVERT CLIFFS UFSAR 5.1-66 Rev. 47 TABLE 5-2 ASSUMED INTERNALLY GENERATED MISSILES IN CONTAINMENT STRUCTURE ITEM AND LOCATION WEIGHT MASS/
AREA SOURCE SHAPE IMPACT VELOCITY POINT OF IMPACT lb/in2 (b,c) ft/sec I.
REACTOR VESSEL a) Closure Head Nut 116
.0034 1
Cylindrical 35.5 Overhead Shielding Slab b) Closure Head Nut and Stud 710
.0477 1
Cylindrical 21.2 Overhead Shielding Slab c) Instrument Assembly 335
.0445a 2
Rod 49.4 Overhead Shielding Slab d) Instrument 165
.0219a 2
Flat Rod 211 Overhead Shielding Slab II.
STEAM GENERATOR a) Primary Manway Cover 1000 6.57a 2
Disk 156.6 Steam Generator Walls b) Primary Manway Cover Stud and Nut 24 7.64 1
Rod 33 Steam Generator Walls c) Secondary Handhole Cover 150 3.76a 2
Disk 181 Steam Generator Walls d) Secondary Handhole Cover Stud and Nut 4
4.04 1
Rod 17.9 Steam Generator Walls e) Secondary Manway Stud 4.67
.0098 1
Rod 506.7 Pressurizer Shield Wall f) Inspection Port Cover 30 1.72 2
Disk 668 Steam Generator Walls g) Inspection Port Cover Stud and Nut 1.33 3.02 1
Rod 1295 Steam Generator Walls III.
PRESSURIZER a) Safety Valve & Flange 550/800
.4540 2
Cylindrical 97.2/80.6 Pressurizer Shield Wall b) Valve Flange Bolt 3.75
.0079 1
Rod 16.5 Pressurizer Shield Wall c) Relief Valve & Flange 150
.0311a 2
Cylindrical 150 Pressurizer Shield Wall d) Manway Cover 680
.0150a 2
Disk 369 Pressurizer Shield Wall e) Upper Temp. Element 2.75
.0052a 2
Rod 81.8 Pressurizer Shield Wall f) Lower Temp. Element 3
.0056a 2
Rod 80.2 Pressurizer Shield Wall g) Safety Valve 200
.1648 2
Cylindrical 158 Pressurizer Shield Wall IV.
CONTROL ROD DRIVER 1100
.0365 2
Cylindrical 58 Overhead Missile Shield V.
MAIN PUMP & PIPE a) Temp. Nozzle & Rtd 11.1
.0209a 2
Cylindrical 81.3 Secondary Shield Wall VI.
SURGE & SPRAY LINES a) Thermal Well & Rtd 1.75
.0033a 2
Cylindrical 105 Secondary Shield Wall
CALVERT CLIFFS UFSAR 5.1-67 Rev. 47 TABLE 5-2 ASSUMED INTERNALLY GENERATED MISSILES IN CONTAINMENT STRUCTURE ITEM AND LOCATION WEIGHT MASS/
AREA SOURCE SHAPE IMPACT VELOCITY POINT OF IMPACT lb/in2 (b,c) ft/sec VII.
SHUTDOWN COOLING a) Line Valve Stem 85
.0450 2
Cylindrical 50.4 Secondary Shield Wall a
Cross-section used is projected edge area (dia x thk) all other areas are minimum projected area.
b Stored mechanical strain energy converted to kinetic energy by separation of stud.
c Hydrostatic pressure force converted to fluid jet force by separation of mechanical restraint.