LR-N17-0034, Salem Generating Station, Units 1 & 2, Revision 29 to Updated Final Safety Analysis Report, Section 3.8, Design of Category I Structures

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Salem Generating Station, Units 1 & 2, Revision 29 to Updated Final Safety Analysis Report, Section 3.8, Design of Category I Structures
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3.8 DESIGN OF CATEGORY I STRUCTURES 3.8.1 Containment Structure 3.8.1.1 General Description For arrangement of containment structures, the patterns of reinforcements, and the layout for liner, see Figures 3.8-1, 3.8-3, 3.8-8 and Plant Drawings 208900, 201102, 201105, 201108, 201175, 201181 and 201131. The reactor containment structure is a reinforced concrete vertical right cylinder with a flat base and a hemispherical dome. A welded steel liner with a minimum thickness of 1/4 inch is attached to the inside face of the concrete shell to ensure a high degree of leak tightness. The design objective of the containment structure is to contain all radioactive material which might be released from the core following a loss

-of-coolant accident (LOCA). The structure serves as both a biological shield and a pressure container.

3.8-1 S GS-UFSAR Revision 27 November 25, 2013

The underground portion of the containment structure is waterproofed in order to avoid seepage of ground water through cracks in the concrete. The waterproofing consists of an impervious membrane which is placed under the mat and on the outside of the walls. The Ethylene Propylene Diene Monomers (by Uniroyal, Inc.) membrane will not tear in handling or placing of backfill against it. The installation of the membrane is described in Section 3.8.2.6.8.4. The basic structural elements considered in the design of the containment structure are the base slab, side walls, and dome acting as one structure under all possible loading conditions. The liner is anchored to the concrete shell by means of anchors so that it forms an integral part of the entire composite structure under all loadings. The reinforcing in the structure will have an elastic response to all loads with limited maximum strains to ensure the integrity of the steel liner. The lower portions of the cylindrical liner are insulated to avoid buckling of the liner due to restricted radial growth when subjected to a rise in temperature. The reinforcement patterns of the base mat are shown on Plant Drawings 201102 and 201105. The reinforcement patterns of the cylindrical wall are shown on Figure 3. 8-3. Drawing 201108. The reinforcement patterns of the dome are shown on Plant The containment structure is inherently safe with regard to common hazards such as fire, flood, and electric storm. The thick concrete walls are invulnerable to fire and only an insignificant amount of combustible material, such as lubricating oil in pump and motor bearings, is present in the containment. A lightning protection system is installed on the containment dome to protect against electrical storm damage. The dead weight of the structure is a minimum of 3. 0 times the buoyancy force that may be exerted on the structure if the ground water level is considered to be at a grade which is 3.5 feet higher than the normal ground water table. SGS-UFSAR In case of a hypothetical hurricane 3.8-2 Revision 27 November 25, 2013 flooding to a height of 20.9 feet above grade, the dead weight will be a minimum of 1.6 times the buoyant force. Therefore, the highest water conditions in the river will present no hazard to the flotation of the containment.

Internal structures consist of equipment supports, polar crane gantry, shielding, reactor cavity and canal for fuel transfer, miscellaneous concrete and steel for floors and stairs.

A 3-foot thick concrete ring wall serving as a partial radiation shield surrounds the Reactor Coolant System (RCS) components and supports the polar

-type reactor containment crane. A 3 to 5-foot thick reinforced concrete floor covers the RCS compartments. Removable concrete plugs are provided to permit crane access to the reactor coolant pumps. The four steam generators, pressurizer, and various pipes penetrate the floor. Stairs provide access to

the areas below the floor.

The refueling canal connects the reactor cavity with the fuel transport tube to the spent fuel pool. The floor and walls of the canal are concrete, with walls and shielding water providing the equivalent of 6 feet of concrete. The floor is 4.5-feet thick. The concrete walls and floor are lined with 1/4-inch thick stainless steel plate. The linings provide a membrane that is resistant to

abrasion and damage during fuel handling operations.

The containment characteristics used to determine the containment structural heat sinks considered in the containment accident analysis are shown it Tables

15.4-20 and 15.4

-21. 3.8.1.2 Design Codes

The Containment Building has been designed under the following codes:

1. Building Code Requirements for Reinforced Concrete, ACI 318-63.
2. AISC Manual of Steel Construction, 6th Edition or later edition, as applicable.

3.8-3 SGS-UFSAR Revision 29 January 30, 2017

3. ASME Boiler and Pressure Vessel Code, section III, section VIII, and Section IX (Applicable portions) -1968. 3.8.1.3 Design Loads and Loading Combinations The following loads are considered to act upon the containment structure creating stresses within the component parts: 1. Dead load The dead load consists of the weight of the complete structure as shown in the construction drawing. To provide for variations in the assumed dead load, the coefficient for dead load components is adjusted by +/-5 percent as indicated in the various cases of loading combinations. 2. Live load Live load consists of snow or construction loads on the dome and also the weight of major components or equipment in the containment. A construction load of 50 pounds per square foot, which is more severe than the snow load, is used in dome design. 3. Internal Pressure SGS-UFSAR The internal pressure transient used for the containment design and its variation with time is shown on the pressure-temperature transient curve, Figure 3.8-11. For the free volume of 2,620,000 cubic feet within the containment, the design pressure is 47 psig. This pressure transient is more severe than those calculated for various LOCAs and Main Steam Line Breaks (MSLB) which are presented in Section 15. 3.8-4 Revision 17 October 16, 1998 --
4. Thermal Thermal expansion stresses due to an internal temperature increase caused by a LOCA have been considered. This temperature and its variation with time is shown on the pressure-temperature transient curve, Figure 3.8-11. The maximum temperature at the uninsulated section of the liner under accident conditions is 246°F. For the 1.25 times and 1.50 times design pressure loading conditions given in Section 3.8.1.4.1, the corresponding liner temperature will be 285°F and 306°F, respectively. The pressure-temperature transient curves for these loading conditions are shown on Figures 3.8-12 and 3.8-13, respectively. The maximum operating temperature is 120°F. For the Main Steam Line Breaks (MSLB), Figure 15.4-100 provides the containment pressure and temperature transients for the limiting temperature case. The governing peak temperature is 351.3°F. 5. Buoyancy Uplift due to buoyant forces created by the displacement of ground water by the structure has been considered. Computations are based on normal ground water being at grade level and flood water at 20.9 feet above grade during a hypothetical hurricane. 6. Seismic Load SGS-UFSAR The site seismology and ground response spectra are described in Section 2. Seismic design criteria for structures and equipment are described in Sections 3.7 and 3.8.1.4.2. 3.8-5 Revision 17 October 16, 1998
7. Wind Load A wind load of 30 pounds per square foot, equivalent to 108 mph, was applied to structures and found to be less critical than the operational Basis Earthquake (OBE) load. 8. Tornado The Reactor Containment, Fuel Handling, and Auxiliary Buildings have been checked to withstand a tornado loading based on a peripheral wind velocity of 300 miles per hour and a translational velocity of 60 mph. Simultaneous with wind loading, an atmospheric pressure drop of 3 psig for all Class I structures has been considered. The shape factor, c, for the dome is 0.4 and for the cylinder, 0.5. No gust factor is applied. For additional information on tornado loadings, see Section 3.3.2. 9. Test Pressure The test pressure for the containment structure is 115 percent of the design pressure or 54 psig. 10. Negative Pressure . SGS-UFSAR Loading from an internal negative pressure of 3. 5 psig has been considered. A pressure of this magnitude would result from the combined effects of: cooling of the containment volume 70°F below the temperature at which 3.8-6 Revision 6 February 15, 1987

--the containment was sealed, a rise in external barometric pressure of 1 psi, and burning up of hydrogen evolved in an accident conditi.on. The load combinations utilized to determine the required limiting capacity of any structural element in the containment structure have been computed as follows: Case A Operating plus DBA + O.OSD + 1.5P + 1.0 (T + TL) + l.OB Case B Operating plus DBA plus OBE + O.OSD + 1.25P + 1.0 (T' + TL') + 1. 25E + 1 . OB Case C Operating plus DBA plus DBE C+ 1. OD + 0. OSD + 1. OP + 1.0 (T" + TL") + l.OE I + 1.0B Case D Operating plus Tornado + O.OSD + l.lOW + l.OB + l.OPb -t Case E Operating plus DBE + o.osD + I.OT + t.OE' + I.OB Case F Testing + O.OSD + 1.15P + l.OB Symbols used in these formulae are defined as follows: c Required load capacity of section. D Dead load of structure and equipment loads p Accident pressure load as shown on pressure-temperature transient curves. 3.8-7 SGS-UFSAR Revision 6 February 15) 1987 T = TL = T' = TL' = T' = TL" = T' = E = E' = B = = = SGS-UFSAR Load due to maximum temperature gradient through the concrete shell and mat, based upon temperatures associated with 1.5 times accident pressure. Load exerted by the liner based upon temperatures associated with 1.5 times accident pressure. Load due to maximum temperature gradient through the concrete shell and mat based upon temperatures associated with 1.25 times accident pressure. Load exerted by the liner based upon temperatures associated with 1.25 times accident pressure. Load due to maximum temperature gradient through the concrete shell and mat based upon temperature associated with the accident pressure. Load exerted by the liner based upon temperature associated with the accident pressure. Load due to operating temperature gradient through the steel liner, concrete shell, and mat. Load resulting from assumed OBE or wind, whichever is greater. Load resulting from assumed Design Basis Earthquake (DBE) Load resulting from buoyancy effect of ground water. Wind load due to tornado. Bursting pressure loading associated with a tornado. 3.8-8 Revision 6 February 15' 1987 The load factor approach is being 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. In the case of the containment structure, therefore, this approach places minimum emphasis on the fixed gravity loads and maximum emphasis on accident and earthquake or wind loads. The extent to which equilibrium checks of external loads against internal stresses have been made are as follows: Equilibrium checks of external loads against internal stresses have been conducted with a finite element computer program developed specifically for axisymmetric structures under non-symmetric loading by Conrad Associates. The required ultimate load capacity for any structural component of the Containment Building was established by utilizing the following load combination relationship: (a) C l.OD + 0.05D + l.SP + l.OT + l.OB (b) C l.OD + O.OSD + 1.25P + l.OT' + 1.25E + l.OB (c) c 1. OD + 0 . OSD + 1. OP + 1. OT" + 1. OE I + 1. OB (d) C = l.OD + O.OSD + l.lOWt + l.OPb + I.OB (e) C = l.OD + O.OSD + l.OT + l.OE' + l.OB Symbols used in these formulae are defined as follows: c D = SGS-UFSAR Required load capacity of section. Dead load of structure and equipment loads. 3.8-9 Revision 6 February 15, 1987 p ::: T = T' = T" = T I = E = E' = = = B = Accident pressure loads as shown on pressure temperature transient curves. Load the due to maximum temperature steel liner, concrete shell, gradient through and mat, based upon temperatures accident pressure. associated with 1.5 times Load due to maximum temperature gradient through the steel liner, concrete shell, and mat, based upon temperature associated with 1.25 times accident pressure. Load the upon due to maximum temperature gradient through steel liner, concrete shell, and mat, based temperature associated with the accident pressure. Load due to operating temperature gradient through the steel liner, concrete shell, and mat. Load resulting from OBE or wind, whichever is the greater. Load resulting from DBE. Wind load due to tornado. Bursting pressure associated with a tornado. Load resulting from buoyancy effect of ground water. Load combination a assumed that the containment will have the capacity to withstand loadings at least 50 percent greater than that calculated for the postulated LOCA alone. 3.8-10 SGS-UFSAR Revision 6 February 15, 1987 Load combination b assumed that the containment will have the capacity to withstand loadings at least 25 percent greater than that calculated for the postulated LOCA with a coincident OBE. Load combination c assumed that the containment will have the capacity to withstand loadings at least as great as those calculated for the postulated LOCA with a coincident assumed DBE. Load combination d assumed that the containment will have the capacity to withstand tornado winds and associated external pressure drop loadings. Load combination e combines the thermal gradient associated with normal operating conditions with the DBE. The resulting combination produces the maximum compressive stresses in the liner. The horizontal and vertical components of earthquake loads are considered to act simultaneously on the Containment Building. Resultant stresses from both components of loading are added directly with the other loads in the combination. Since the horizontal component of earthquake loading is non-symmetrical, producing tension on one side of the containment vessel and compression on of earthquake combinations. the other, both the positive and negative values stress resultants were considered in the load The combination producing the most critical stress was used in the design. The tornado and tornado generated missile analyses are provided in Section 3.8.1.4. The load combination Case D specifies tornado loads combined with operating loads. The tornado load (Wt) includes the static forces produced by the 360 mph maximum wind velocity t a 3 psi negative pressure and the structural response to the missile impact. 3. 8-11 SGS-UFSAR Revision 6 February 15, 1987 The stresses on any structural member produced by the effective pressure transformed from the tornado wind, the impact of the missile, and also differential pressure were superimposed to obtain the most critical total stress, provided the induced stress from these three components are in the same direction. When one of the components induced an opposite stress, thereby reducing the total stress in the member, it was neglected. In other words, all six loading combinations listed in the Standard Review Plan (SRP) have been considered with factors of 1 instead of 0.5 for Wp in combinations iv and vi and also have taken into account stress directions as stated previously. Hydrostatic loadings from the hurricane condition were applied to the structures to check their stability. The procedures used by our consultant (Dames and Moore) for transferring the static and dynamic flood effects to load were as delineated in the U. S. Army Coastal Engineering Research Center Technical Report No. 4. Total head, including wave effects, was considered to investigate the lateral and overturning effects. Containment flooding for fuel recovery was not a design consideration. The load combinations utilized in the design of the containment and other Category I structures were equivalent to or more than those outlined in the SRP. The following tabulations provide a combinations utilized with the SRP criteria. 3.8-12 SGS-UFSAR comparison of load Revision 6 February 15, 1987 Test CONCRETE CONTAINMENT STRUCTURE SRP (D+L) + Pt + Tt Salem

  • D +/- O.OSD + Pt + Tt Construction Not Critical Normal Not Critical Extreme Environmental (1) Extreme Environmental (2) Abnormal Abnormal/Service Environmental Abnormal/Extreme Environmental SRP Salem SRP Salem (D+L) + T0 + Wt + R0 + Pv
  • D +/- O.OSD + l.lWt + B + Pb (D+L) + T + E + R + P 0 0 v
  • D +/- O.OSD + T' I I + E' + B SRP (D+L) + l.SP + T + R a a a Salem * (D +/- O.OSD + 1.5P + (T + TL) + B a SRP (D+L) + 1.25P + T + 1.25E + R + y + y a a a r m Salem
  • D +/- O.OSD + 1.25P + (T1 + TL') +L25E+B SRP (D+L) + P + T + E' + R + Y + Y a a a r m Salem
  • D +/- O.OSD + P + (T1 1 + TL' ')+E'+B
  • See preceding pages of this Section for identification of Salem symbols. See Although the R and Y SRP Section 3. 8. 3 for SRP symbols. forces are not listed in the overall structural analysis load combination formulae, the local effects under piping load, jet load, and missile impingement were taken into account. 3.8-13 SGS-UFSAR Revision 6 February 15, 1987 (2) SRP Salem (2b) SRP Salem (3) SRP Salem ( 4) SRP Salem (5) SRP Salem (6) SRP Salem INTERNAL CONCRETE STRUCTURES 1.4D + 1.71 = 1.9E Not Critical 0.75 (1.4D + 1.7L + 1.9E = 1.7 T + 1.7R) 0 0 Less Critical Than (5) D + L + T + R + E' 0 0 Less Critical Than (6) D + L + T + R + 1.5P a a D + L + T + R + l.SP a a a D + L + T + R + 1. 25P + 1. 25E a a a + (Yr + Yj + Ym) D + L + T + 1.25E a + R a + 1.25P + (Yr + Yj + Ym) D + L + T + R + P + (Yr + Yj + Ym) + E' a a a D + L + T + R + P + (Yr + Yj + Ym) + E' a a The buoyancy effect of ground water has been included in the assessment of the sliding and overturning potential of the Containment Building and all other Category I structures. The buoyancy effect will reduce the dead weight and thus reduce the factors of safety against sliding and overturning. To include the buoyancy effect in assessing the sliding and overturning potential is the more conservative and correct approach. However, the maximum hurricane, flood, and earthquake are not postulated to occur simultaneously. The safety against sliding, overturning, and flotation for the Containment Building and all other Category I structures under all loading combinations are within the limits set by SRP 3.8.5. 3.8-14 SGS-UFSAR Revision 6 February 15, 1987 3.8.1.4 Design and Analysis Procedures The containment structure has been analyzed to determine stresses, moments, shear, and deflections due to the static and dynamic loads. 3.8.1.4.1 Static Analysis The containment structure has been analyzed and designed for all loading conditions combined with load factors as outlined in Section 3.8.1.3. Mathematically, the dome and cylinder are treated as thin-walled shell structures which result in a membrane analysis. Since the thickness of the dome and cylinder is small in comparison with the radius of curvature (cylinder 1/15.5, dome 1/20), the stress due to pressure and wind or earthquake can be calculated by assuming that they are uniformly distributed across the thickness. In general, membrane stresses are carried by the reinforcement. Some are carried by the steel liner, but none by the concrete unless they are compressive stresses. Manual analysis of the containment structure, based on "Theory of Plates and shells," by Timoshenko and Woinowsky-Krieger (1) and "Theory of Elasticity," by Timoshenko and Goodier (2), have been performed to obtain shears, moments, and stresses within the structure as the basis of our preliminary design for reinforcements and _J.iner plate. An independent three-dimensional axisymmetric modal analysis using the finite element method was made by Conrad Associates (3) to ascertain that the design of the containment structure was adequate. 3.8-15 SGS-UFSAR Revision 6 February 15, 1987 The manual shell analyses calculations and the "Conrad Associates" design review report are submitted separately. The design includes the consideration of both primary and secondary stresses. The design limit for tension members (i.e., the capacity required for the design load) is based upon the yield stress of the reinforcing steel. The load factors used in the design primarily provide for a safety margin on the load assumptions. The capacity reduction factor "0" is provided for the possibility that small adverse variations in material strengths, workmanship, dimensions, and control, while individually within required tolerances and the limits of good practice, occasionally may combine to result in under capacity. For tension members, the factor "0" is established as 0.95. The factor "0" is 0.90 for flexure and 0.85 for diagonal tension, bond, and anchorage. For the liner steel the factor "0" is 0.95 for tension, 0.90 for compression and shear. The detailed design has been reviewed by Conrad Associates' finite element computer program to verify its safety. Stress values for rebars and liner plates at various locations for all loading combinations involving LOCA are given in Tables 3.8-1 through 3.8-10. The designation of main reinforcement pattern for the containment structure is shown on Figures 3.8-14 and 3.8-15. seismic reinforcing consists of diagonal bars at 45° to the horizontal plane each way, extended from mat to the lower portion of the dome. They are designed to resist the lateral shear under earthquake such that the horizontal component per foot of diagonals will be equal to the maximum value of the shear flow. Although, in the cylinder, the liner and the concrete have some capacity available to resist the seismic shears, no credit was taken for the capacity. Dowel action of the main bars was also neglected. The containment structure has also been evaluated for increase in design loads due to the postulated MSLBs. The evaluation shows that for the design of the containment structures LOCA is the governing condition. 3.8-16 SGS-UFSAR Revision 17 October 16, 1998 Wall Stresses The stresses in the wall reinforcement from the independent check listed in Tables 3.8-1 through 3.8-8 are all under yield pointt except the only location where the diagonal bars are critically stressed is at Elevation 84 feet under load combination (c). Howevert as stated by Conrad Associates (3), the stress indicated at that location as 60.79 ksi was obtained neglecting all contributions of the main meridional and hoop reinforcement to the seismic shear-resisting capacity of the containment wall. An inspection of the stresses incurred by the main reinforcement as a result of forces other than the seismic shear indicates that these bars are markedly understressed in this zone of the containment shell. Thus the stress value of 60.79 ksi in the diagonal reinforcing bars resulting for seismic shear is overestimated. The discontinuity stresses are accounted for in the design. The moments and shears are computed by equating the deformations and angular rotations of the two parts of the structure at the point of juncture and solving for the resulted discontinuity stresses. The total stresses are obtained by adding the discontinuity stresses to the membrane stresses. The moments and shears at the base of the containment wall are determined on the basis of the rigidity of the resulting cracked section, with the steel on the inner face in tension and concrete on the outer face in compression. The compressive stress in the concrete is checked to ascertain that it is less than a. 75 fc. The tension bars are checked to ascertain that the stresses are not more than 0.90 fy. The shears are carried by hooked diagonal radial bars and no reliance is made on the concrete. Additional diagonal radial bars inclined in a direction normal to the shear diagonal bars will be placed in the wall to take care of diagonal tension. In the stress analysis, uncracked section for concrete is found to be more critical in creating secondary bending stresses in the areas of discontinuity. This conservative assumption was used by Conrad Associates to check the design in such areas. 3. 8-17 SGS-UFSAR Revision 6 February 15, 1987 I The deformation of the containment is larger if cracked section property is employed. The values obtained from this approach are being used in calculating the relative displacement between the buildings for clearance, assuring that adequate clearance has been provided. The working stress check under operating conditions has been found to be at very low level. The maximum concrete compressive stress under dead load, operating thermal load, and OBE is 835 psi, while the maximum stress in the reinforcement for the same loading combination is 6540 psi. The concentric dome ring was conservatively designed as a tension ring subjected to uniform pull around the periphery. Two sets of l-inch diaphragm plates are used to transfer the tension through the ring to the meridional reinforcements merging at the peak of the dome. The stress level in the cylindrical tube is minimal. The ring plate is made of ASTM A516 Grade-70 pressure vessel quality, ultrasonic testing per ASTM A-435 except with 100 percent coverage. All conforms to Section III of the ASME Boiler and Pressure Vessel Code. The ring is physically connected to the meridional reinforcement and the liner. Liner Plate The maximum tensile stress in the liner plate under the test condition is 30.9 ksi, below the minimum yield point of 32 ksi. This is the preoperational artificial pressure test without the accompanied temperature rise. This case induces the higher tensile stress in the liner plate than the design basis accident 3.8-18 SGS-UFSAR Revision 20 May 6, 2003 condition. Under other cases of critical loading combinations involving LOCA the maximum tensile stress in the liner plate is 27.5 ksi, with 14 percent extra safety margin. The maximum interaction coefficient for biaxial compression and shear in the liner plate under critical load combinations involving LOCA is 0.902 with approximately 10 percent extra safety margin. The listed stresses in Tables 3.8-1 through 3.8-9 have already taken account of the capacity reduction factors, 0. In other words, the stresses have been divided by the appropriate 0, 0.95 for tension, 0.90 for flexure, and 0.85 for shear, etc. The combined biaxial compression and shear in the liner plate have been examined by the following interaction formula: where: a , o X Z hoop and meridional stresses in liner plates oxo' azO = maximum allowable stress in hoop and meridional direction (critical buckling stress or the yield stress) T T 0 shear stress in the liner plate maximum allowable shear stress The resulting interaction coefficients for Operating, LOCA, and Test conditions 1 are listed in Table 3.8-10. The containment liner has also been evaluated for the increased containment temperature of 351.3°F and the concurrent pressure due to the postulated MSLBs. The evaluation shows that the liner in the uninsulated portion tends to yield locally at EL.l20 '-0"; however, the total design forces at this local section can 3.8-19 SGS-UFSAR Revision 17 october 16, 1998 be carried by the containment reinforcing steel alone, without using the liner as a strength element. The corresponding strains in the liner at this section are low relative to the allowable liner strain values specified in Table CC-3720-1 of 1995 ASME B&PV Code,Section III, Division 2 (Reference 6) for maintaining leaktight integrity of the liner. Thus, both strength and leaktight integrity of the containment are assured. s. B. Batdorf and M. Stein in their paper "Critical Combinations of Shear and Direct Stress for Simply Supported Rectangular Flat Plates" (NACA Technical Note 1223, 1947), obtained the critical stress combination for the case of shear and simultaneous uniaxial compressive stresses as: (t/tO) 2 + a/aO = 1 For biaxial compression, Timoshenko and Gere in their "Theory of Elastic Stability" defined the allowable biaxial compression in the form of: Modifying The Batdorf and Stein expression to include the biaxial effect we have used the following equation to check the interaction stability: (x + y) (t/t0)2 + 1 (xO +yO) The edge condition was assumed to be simple supported which is more conservative. 3.8-20 SGS-UFSAR Revision 17 October 16, 1998 Base Mat In designing the base mat, the slab is considered to be a circular plate of constant thickness, t. The loads are imposed upon the slab by the exterior cylinder wall, the central circular crane wall and, to a lesser degree, by the equipment. The soil reaction pressure was found in a conventional manner by treating the slab, which is 16-feet thick, as a rigid mat. The mat is then analyzed as a plate subjected to soil pressures and supported by a circular wall symmetrical with respect to the center of the mat. The supporting walls are considered as either simply supporting the mat or partially fixed. The exterior cylinder wall has been considered partially fixing the mat; the crane wall is a simple support. The containment base mat is analyzed as a rigid circular plate subjected to loadings from the axisymmetric exterior cylinder wall, crane wall, interior walls, and equipment acting around an equivalent circle. The soil pressure is found in a conventional manner without the benefit of its elastic deformation. Manual analysis was based on the AC! Paper, Title No. 63-63, "Analysis of Circular and Annular Slab for Chimney Foundation," by Kuang-Han Chu and omar F. Afandi. A finite element program was used to check the rebar under five loading combinations. Since the mat is covered by a 2 to 5-foot thick concrete slab, and also the lower 34-feet of cylinder liner is insulated, the thermal effect on the mat has been neglected. The design of the base mat reinforcement has been reviewed for five load combinations at three different mat sections. The maximum radial, tangential, vertical, and shear stresses at these sections are shown on Figures 3. 8-16 through 3.8-20. The stresses shown in these figures are integrated over the thickness of the slab and transformed to forces per unit length of circumference. These forces are then distributed to the top and bottom reinforcing bars at the section under investigation. The resulting stresses in the bars are all under 30 ksi. The maximum tangential shear under DBE for the interior structure at top of reactor pit is 7600k.. The shear is transmitted through the pit wall at Elevation 76 feet and then bearing against the base mat. The unit shear is 73 psi and bearing is 42 psi, both well within allowable values. 3.8-21 SGS-UFSAR Revision 17 October 16, 1998 Five static load analyses consisting of dead load, buoyancy, internal pressure, thermal, and tornado loadings have been performed for the containment structure. The complete report by Conrad Associates {3) and manual design calculations are kept on file by Public Service Electric & Gas (PSE&G). They are summarized as follows. Dead Load Analysis Finite element model is used to perform the dead load analysis. Static secant moduli are used in representing the soil stiffness under dead load. For the vertical load, only horizontal restraints are imposed at side boundaries of the soil system. Stresses, moments, and shears at containment wall and mat are shown on Figures 3.8-21, 3.8-22, and 3.8-23. Buoyancy Analysis Normal ground water table for the site is at Elevation 96 feet. For design purpose it is considered to be at 6 inches above the plant grade level, Elevation 99 feet-6 inches. Under hurricane condition the water level could be expected to rise to Elevation 120.4 feet; however, since the direction of the hydrostatic pressure is so small it does not create a critical loading combination. The result of the buoyancy-induced stresses in the containment vessel are very small and are confined to the lower portion of the structure. No plot is given because it does not affect the design. Internal Pressure Analysis The internal pressure transients used for the containment design and its variation with time are shown on Figures 3.8-11 through 3.8-13. For the free volume of 2,620,000 cubic feet within the containment, the design pressure is 47 psig. The maximum temperature at the uninsulated section of the liner under the accident condition is 246°F. For 1.25 times and 1.5 times design pressure loading conditions, the corresponding liner temperatures are 285°F and 306°F, respectively. Static pressure loads are used in design, since the pressure increase is very gradual from the transient curve. 3.8-22 I SGS-UFSAR Revision 17 October 16, 1998 Thermal Analysis The thermal gradients in the containment wall under operating and accident conditions are shown on Figure 3. 8-24. Both loadings are analyzed for the containment structure. The analytical model employed by Conrad Associates ( 3) for finite element thermal analysis is an axisymmetric assemblage of solids of revolution. Each segment across the containment wall consists of ten elements to represent the thermal gradient through the wall thickness. Orthotropic material properties are used to represent the variable shell area in the hoop and meridional directions. Due to the one-dimensional nature of the reinforcing bars, Poisson's ratio was set equal to zero in the plane of the equivalent steel shell. For accident loading, the concrete is assumed to be totally cracked in the hoop and meridional directions, but uncracked in the radial direction. For operating loading, concrete is assumed to be uncracked. The liner plate is modeled as a thin isotropic steel shell with an elastic modulus of 28,000 ksi and a Poisson's ratio of 0.3. Between the liner plate and the concrete containment shell a thin element, 0.01-feet thick, is introduced to facilitate modeling of the discontinuity in temperature occurring at the liner-to-concrete interface under accident conditions. A fixed boundary is introduced at the foundation mat. Thermal stresses and strains are not likely to develop in the thick mat which has excellent insulating properties. 3.8-23 SGS-UFSAR Revision 6 February 15, 1987 Stresses under operating and accidental thermal loadings involving LOCA are shown on Figures 3.8-25 through 3.8-30. Tornado and Tornado Generated Missile Analysis Three tornado wind distributions were investigated in the Category I structural design as shown on Figure 3. 8-31. In combination with the static forces produced by the 360 mph maximum wind, a 3 psig atmospheric pressure drop was specified for the containment structure. Evaluations of structural adequacy against tornado wind loads and tornado missiles are given in sections 3.3.2 and 3.5.2, respectively. 3.8.1.4.2 Dynamic Analysis The containment structure seismic analysis was performed through (a) lumped mass model manual analysis, using average response spectra ground input, and (b) a finite element modal analysis, using time history ground input. The detailed report from conrad Associates (3) and the independent manual calculations are kept on file by PSE&G. The computer analysis yields a slightly higher result in accelerations, shears, and moments in comparison with the manual analysis. The most conservative results are used in design. The seismic analysis of the containment structure by the finite element method is performed by computer using a step-by-step direction integration procedure. Studies have been made to establish free field soil boundary condition. The model used in the analysis is shown on Figure 3.7-13. The El Centro ground motion of May 18, 1940, was recommended by Dames and Moore as the most appropriate motion for the site. Its 3.8-24 SGS-UFSAR Revision 17 October 16, 1998 peak horizontal acceleration was normalized to 0.10 g and 0.20 g for OBE and DBE, respectively. Two-thirds of the above-mentioned values are used for vertical ground motions, and they are considered to be acting simultaneously with the horizontal ground motion. Modified Hausner's average Figures 3.7-1 and 3.7-2, are response used for spectra, as shown on normal modal analysis. Seismic design criteria and procedures for structures are described in Section 3.7. For the DBE, a damping factor of 5 percent of critical damping is used for analysis for structure and soil. Similarly for the OBE, a damping factor of 2 percent is applied for both structure and soil. Two separate modal analyses, horizontal and vertical motions, are performed and their results superimposed. The acceleration time histories from the result of the structural seismic analysis are used for the generation of horizontal and vertical response spectra at specified floors or locations for equipment of seismic design. They are presented in the Conrad Associates' report and kept on file by PSE&G. Total accelerations, peak displacements, and the envelope of forces in the containment structure under DBE and OBE conditions are shown on Figures 3.7-3 through 3.7-12. Clearances between Category I buildings and adjacent structures are checked based on the relative displacement at various building elevations under seismic and design basis accident loadings to assure that the required separations are maintained. 3.8-25 SGS-UFSAR Revision 6 February 15, 1987 3.8.1.5 Structural Design and Acceptance Criteria The containment structure is designed to meet the following design criteria stated in the "General Design Criteria for Nuclear Power Plant Construction Permits." Reactor containment shall be provided. The containment structure shall be designed (a) to sustain without undue risk to the health and safety of the public, the initial effects of gross equipment failures such as a large reactor coolant pipe break, without loss of required integrity and (b) together with other engineered safety features as may be necessary, to retain for as long as the situation requires, the functional capability of the containment to the extent necessary to avoid undue risk to the health and safety of the public. The reactor containment structure, including access openings and penetrations, and any necessary containment heat removal systems shall be designed so that the leakage of radioactive materials from the containment structure under conditions of pressure and temperature resulting from the largest credible energy release following a LOCA, including the calculated energy from metal-water or other chemical reactions that could occur as a consequence of failure of any single active component in the Emergency Core Cooling System (ECCS), will not result in undue risk to the health and safety of the public. The containment structure design parameters are based on the following: 1. Leak tightness and testing requirements 2. Seismic requirements 3. Tornado requirements 3.8-26 SGS-UFSAR Revision 6 February 15, 1987
4. Shielding requirements s. Design basis accident requirements 6. Flood conditions due to maximum probable hurricane 7. Internal missile generation The stresses of concrete, reinforcing steel, and liner plate under various loading combinations are as described in Section 3.8.1.4. The containment integrity evaluation, including the containment pressure transients and safety margin, are presented in Section 15. 3.8.1.5.1 Fracture Prevention of Containment Pressure Boundary The containment pressure boundary parts, which do not rely on concrete structures to provide the pressure retaining capability, are constructed in accordance with the material, design, fabrication, and installation requirements of the ASME Code,Section III, 1968 Edition. The Code requirements took into consideration procedures for prevention of brittle failures and fracture propagations in containment pressure boundaries. These procedures include Charpy V notch tests of plate materials, sufficient margins in the design allowables, preheat of steel plates, and postweld heat treatment of penetration assemblies. 3.8.1.6 Materials. Quality Control, and Special Construction Techniques 3.8.1.6.1 Liner Plate A welded steel liner of thicknesses varying from 1/4 inch to 1/2 inch is anchored to the inside face of the concrete shell with 3.8-27 . SGS-UFSAR Revision 6 February 15, 1987 1/2-inch diameter studs to ensure containment leak tightness. This containment liner is designed to carry a portion of the membrane force from the different combinations of loading; however, for conservatism it is not counted on in the resistance to lateral shear. The out-of-roundness tolerance of the liner shall not exceed plus or minus 2 inches from the true diameter of 140 feet. The lower 34 feet of cylinder liner is insulated, except locally around liner penetrations and around interferences with other commodities, to prevent buckling of the liner due to restricted growth under a rise in temperature. The membrane tension and the combined stress of biaxial compression and shear in the liner plate are described in Section 3.8.1.4.1. Our computations for the liner plate indicate that there would be no inelastic buckling of the plates. Under stress, the variation in plate thickness would cause small differential movements between the liner and the concrete. Also, the shrinkage cracks in the concrete would have the same result. Soft corks are placed around the studs adjoining the liner plate to allow differential movement between the liner and the concrete. The stud anchors are designed such that their failure in shear or tension will not break the leak tight integrity of the liner plate. Tests will be made to verify this criterion. nature. This would Even if stud failure developed, it would be random in not impair the liner integrity, nor would it cause progressive failure. The design load per anchor is low, and if an anchor should fail, the load it would have carried would be easily distributed to the adjacent anchor. Tensile and shear tests were conducted on the liner plate studs. Three tensile and three shear test assemblies approximating as 3.8-28 SGS-UFSAR Revision 17 October 16, 1998 close as possible the welded studs in service, were fabricated as test specimens. The results of the tests indicate that the studs pulled away from the liner plate at a tensile stress of between 74,500 psi and 80,600 psi. Under shear loading the studs sheared off between 62,600 psi and 67,000 psi. In both failure modes the leak tight integrity of the containment liner plate was not affected. Each liner plate splice in the dome, cylinder, and mat is covered by a steel channel. The steel channels are embedded in the concrete mat. To prevent any possible shearing of the channels from the differential movement between the liner plate and the inner concrete slab, they are isolated from the concrete by 1/4 inch of asphalt impregnated expansion material, and Styrofoam all around. Where there are a large number of penetrations in one area, the thickness of the liner plate is increased from 3/8 inch to 3/4 inch for reinforcement. The original intent of the steel channels was for leak testing the liner welds. However, leak testing will be performed in accordance with 10CFR50, Appendix J, "Primary Reactor Containment Leakage Testing for Water-Cooled Power Reactors," instead of pressurizing the liner weld channels. procedures are described in Section 6.2.1. The leak testing program and The 3/4-inch knuckle plate connects the cylinder liner to the base liner. The thicker plate is used to resist buckling due to concentrated loadings from liner anchors in the base mat and also to take care of the warped surface created by the double curvature at the junction. The detail of the anchor plate is shown in Section "1-1" of Plant Drawing 201175. Tension anchors to transfer the uplift force for essential pieces of equipment to the mat are also shown in Sections "X-X" and "5-5" of Plant Drawing 201175. Where there is a shear load in combination with the tensile load, as there would be in the case of an earthquake, the shear load will be transmitted into the 2-foot thick or 5-foot 3.8-29 SGS-UFSAR Revision 27 November 25, 2013 I thick concrete slab located above the liner plate by shear lugs attached to the equipment base plates. The transfer of shear load from inner structures through the bottom liner plate of the containment is by means of the reactor well acting as a key. The inside surface of the liner plate in the cylin<.;:ler and dome is painted with catalyzed epoxy paint. Surfaces treated in this manner can be easily washed for decontamination. Keeler and Long No. 7475 and No. 7844 epoxy was specified as the protective coating based on their test-proven ability to withstand design basis accident and washdown conditions. 'l'he coating, during test, had been exposed to the cycles for design basis accident condition with satisfactory results. The spray solutions used in the tests are similar to the chemical compositions proposed for Salem but of higher concentration. The protective coating also behaves well in the steam environment, as reported by the Franklin Institute Research Laboratory. The direct jet impingement with jet temperature above 300°F will cause disintegration of the coating in that local area, which can be repaired after the accident. A train of strainer modules has been connected to the sump at the bottom of the containment for retaining any disintegrated particles and to eliminate any possibility of flow blockage . Fouling of engineered safety features as a result of the local failure of the coating is not considered possible. 3.8,1.6.2 Base Mat The design of the base mat is described in Section 3.8.1.4.1. The base mat was poured on top of lean concrete fill in circular segments, as described in Section 3.8.1.6.8.7, with only vertical 3.8-30 SGS-UFSAR Revision 23 October 17, 2007 * *
  • construction joints. The base mat is poured to a level 6 inches below the final elevation of the bottom liner plate. The backing tees are then positioned and concrete poured to a level flush with the top of the backing tees. Steam generators and reactor coolant pumps are supported by heavy welded steel frames embedded in the concrete and tied down deep into base mat by 6-inch diameter and 4-inch diameter bolts, 18 feet-6 inches long to prevent the tremendous uplift during pipe rupture accident. (See Figure 3.8-32.) 3.8.1.6.3 Cylinder Wall The design of the cylindrical wall is described in Section 3.8.1.4.1. The wall pours are made in lifts 4 to 5 feet in height. 3.8.1.6.4 Dome The design of the dome is described in Section 3.8.1.4.1. The lifts in the dome are approximately 3 to 5 feet in height and each lift is poured continuously with no joints parallel to the liner plate allowed. Near the top of the dome, terminations of the lifts are horizontal rather than normal to the liner plate. For arrangement of the dome line and the top enclosure ring detail, see Plant Drawing 201181. 3.8.1.6.5 Penetrations and Openings For a description of various containment wall penetrations, hatch openings, their details and basis for design and analysis, see Section 3.8.1.6.8.8. 3.8-31 SGS-UFSAR Revision 27 November 25, 2013 The piping penetration sleeves were fabricated to applicable portions of the ASME Boiler and Pressure Vessel Code,Section III, Nuclear Vessels, 1968, with the exception that there was no requirement made to stamp the pipe with the "N" symbol. Inspection of the piping penetrations was in accordance with the above indicated code except that hydr-ostatic test of rolled and welded pipe to ASTM AISS was not performed at the mill. The plate for this pipe, however, was ultrasonically examined and welds were completely radiographed. In addition, the sleeves were pressure tested with the containment as well as pneumatically leak tested internally. Cooling, by both free and forced convection, is provided where necessary to maintain concrete temperatures adjacent to hot pipe penetrations below 150°F. A hot pipe passing through the containment wall can transfer heat to the wall via any or all of three paths. These paths, shown on Figure 3.8-33, are: 1. Radial conduction in the pipe cap and longitudinal conduction through the expansion joint and along the penetration sleeve outside the containment (Path A). 2. Radial conduction in the pipe cap and longitudinal conduction along the penetration sleeve inside the containment (Path B). 3. Radial conduction through the insulation within the penetration (Path C). The quantity of heat transferred via Path A is inconsequential due to the high thermal resistance presented by the thin cross section of the expansion bellows. 3.8-32 SGS-UFSAR Revision 6 February 15, 1987 The quantity of heat which could be transferred via Path B is significant for some penetrations, i.e. it could cause localized containment concrete temperatures to rise above acceptable limits. As such, annular heat transfer fins (extended surfaces) are provided where necessary. These fins serve to dissipate sufficient heat to the containment atmosphere by natural convection to maintain acceptable temperature in the wall. The fins are designed to dissipate the Path B heat load without the aid of any other cooling mode. The potential heat transfer via Path C can also be significant. As the magnitude of natural heat dissipation in the containment wall is not sufficient to cause a large enough steady state temperature drop in the insulation within the penetration assembly, other means are required to remove the heat and maintain the desired concrete temperature. The heat is removed by compressed air flow in plate-type heat exchangers (coolers) installed within the penetration sleeves. Protection against loss of cooling capability is provided by both the inherent "reliability" of the free convection mode and by redundant compressed air supply lines as shown on Figure 3.8-34. It has been shown that for constant exposure of concrete to temperatures up to 150°F, the loss in strength is quite small; and for temperatures as high as 500°F to 600°F, the deterioration in structural properties is tolerable. Considering the redundancy in air supply lines, the only cause of loss of penetration cooling would be complete loss of the station air compressors, a condition which would not be permitted to persi!il long enough to cause significant localized concrete deterioration. 3.8.1.6.6 Polar Crane The polar crane is described in Section 9. 1. 3.8-33 SGS-UFSAR Revision 6 February 15, 1987 3.8.1.6.7 Missile Protection High pressure RCS equipment which could be the source of missiles is suitably shielded either by the concrete shield wall enclosing the reactor coolant and pressurizer loops or by the concrete operating floor to block any passage of missiles to the containment walls even though such postulated missiles are deemed most improbable. Protection against internally generated missiles is described in Section 3. 5. I. 3.8.1.6.8 Construction Procedures and Practices 3.8.1.6.8.1 Codes of Practice Materials and workmanship conformed to the following codes and specifications: ACI 318-63 "Building Code Requirements for Reinforced Concrete" ACI 301-66 "Specification for Structural Concrete for Buildings" ACI 613-54 "Reconunended Practice for Selecting Proportions for Concrete" ACI 614-59 "Reconunended Practice for Measuring, Mixing and Placing Concrete" ACI 347-63 "Reconunended Practice for Concrete Formwork" ACI "Manual of Concrete Inspection" -1957 3.8-34 SGS-UFSAR Revision 6 February 15, 1987 ASME Boiler and Pressure Vessel Code, 1968:

Section III "Requirements for Class B Vessels" (penetrations and hatches only)

Section VIII "Requirements Pertaining to Methods of Fabrication of Unfired Pressure Vessels"Section IX"Welding Qualifications" AISC "Manual of Steel Construction," 6th Edition or later edition, as applicable ACI 301-66, "Specifications for Structural Concrete for Buildings," together with ACI 318-63 "Building Code Requirements for Reinforced Concrete," form the basis for the PSE&G concrete specifications.

3.8.1.6.8.2 Concrete

3.8-35 SGS-UFSAR Revision 6 February 15, 1987

Preliminary Tests The PSE&G Testing Laboratory obtained samples of the aggregates to be used in the concrete for preliminary testing and approvaL Testing methods and acceptance standards were as follows: Acceptance Test Method Standards Sampling ASTM 075-59 ASTM C-33 Gradation -Sand ASTM Cl36-63 ASTM C-33 Gradation -Stone ASTM C136-63 ASTM C-33 Sodium Sulfate ASTM C88-63 ASTM C-33 Soundness Loss Angeles Abrasion -ASTM C131-66 ASTM C-33 Stone Material Finer than No. 200 Sieve ASTM C117-66 ASTM C-33 Organic Impurities -ASTM C40-66 ASTM C-33 Sand Potential Reactivity -Chemical Method ASTM C289-6S ASTM C-33 In addition, the following tests were performed to give necessary information concerning the aggregates. Test Fineness Modulus Unit Weight Specific Gravity Absorption SGS-UFSAR 3.8-36 Method ASTM C125-66 ASTM C29-60 ASTM CI27-59 ASTM C128-59 Revision 6 February 15, 1987 The coarse aggregate selected and used on the Salem Project was quarried stone, crushed and graded to meet the detail specifications. The stone, commonly known as traprock, was a basic igneous rock consisting of diabase and basalt. The quarries and crushers were located in Lamberville, Pennington, and Kingston, New Jersey. The fine aggregate selected was known locally as Dorchester sand. It was a silica sand found in bank run deposits. The sand was dredged, washed, and then graded to meet project deta i 1 specifications. The Portland Cement (Type II) used conformed to ASTM Specification C-150, latest edition. Flyash was used as an admixture in the majority of the concrete and conformed to ASTM Specification C-350-65T, except that the fineness of the flyash was in accordance with the ASTM Specification C-618-68T, which has not replaced ASTM C-350. A retarding densifier was also used as an admixture which conformed to ASTM Specification C-494, Type D. The retarder was a water reducing admixture of the hydraxylated carbolic acid type and contained no calcium chloride. Trial mixes were made by the PSE&G Testing Laboratory with the above ingredients in accordance with ACI 301-66, Section 308 -Method 2. Proportions of ingredients were determined and tests conducted in accordance with ACI 613-54, "Recommended Practice for Selecting Proportions for Concrete. 11 The concrete mixes used for construction were approved by the PSE&G Structural Engineering Division and specified in the project detail specification for concrete. The dry density of the concrete mixes used for construction exceeded 144 lbs/cu ft. All concrete mixes used in the work were fully documented. For structural concrete, the maximum allowable slump for concrete placed was 4 inches. In areas with closely SGS-UFSAR spaced reinforcing bars, the detail 3.8-37 Revision 6 February 15, 1987 specification allowed the use of a concrete mix with a coarse aggregate of 3/8 inch and a maximum slump of 5 inches. For the reactor containment wall. in the area adjacent to the equipment and personnel hatches where additional reinforcing steel was specified, a more plastic mix was designed for adequate concrete placement. In this case, the slump was increased to be between 6 and 7 inches. For fill concrete, one batch in ten could have a slump of up to 5 inches with the majority of concrete placed at 4 inches slump or less. Seven-inch slump concrete was used only in the area of the equipment crowded. and The personnel hatches where reinforcing steels are water-cement ratio was 6.25 gallons per bag; 7.5 bags of cement were used per cubic yard of concrete. One hundred pounds of flyash per cubic yard of mix was added. It is believed that the erosion resistance of this specific mix is as good as the regular low slump mix. After the form was removed, the surface of the concrete appeared to be smoother and without visible cracks, due to the workability of the mix. The flyash contains no calcium chloride to cause corrosion. The 1970 edition of "Concrete Industries Year Book" states that concrete made with flyash is more resistant to weak acids and sulfates, which cause corrosion. Batch Plant The bulk of the concrete for the project was supplied from a batch plant at the site operated by United Engineers and Constructors, Inc. Technical details of this plant are as follows: 1. Eric Strayer central mix concrete plant t rated at 240 cubic yards/hourt although the maximum rate of concrete produced was 180 cubic yards/hour 3.8-38 SGS-UFSAR Revision 6 February 15, 1987

2. Four compartment aggregate bin 3. Nine cubic yard aggregate hatcher weighing aggregates cumulatively and automatically with a dial scale 4. 1,500 barrel bin divided into two compartments for flyash and cement 5. Cement and flyash hatcher weighing cumulatively and automatically with a dial scale 6. Water and ice hatcher weighing cumulatively and automatically with a dial scale The plant provided fully automatic wet hatching for the various mixes required. The operator inserted the proper card for the mix required, set a dial for the quantity of concrete desired and the machine measured out the ingredients automatically and recorded the weight automatically. Weight measurements were also visually observed by the Quality Control Inspector at the control console on three separate 2-foot diameter indicating dials to check and confirm the recording tapes. The hatching accuracy of the weighing equipment was within +/-1 percent of the true values. The weighing equipment was calibrated prior to initial use. Standard weights were used for periodic calibrations. The calibrations were made quarterly or every 40,000 cubic yards poured, whichever occurred first. Moisture probes embedded in the aggregate binds determined moisture content and compensations were made to maintain the proper water-cement ratio. During cold weather, the temperature of the concrete was controlled by heating the mixing water and heating the aggregate bins. During hot weather, the temperature of the concrete was controlled by cooling the mixing water. During extremely hot weather, flaked ice was added to the mix. The flaked ice and water was weighed separately, but cumulatively in a compartmented weigh hopper. 3.8-39 SGS-UFSAR Revision 6 February 15, 1987 During hot weather the temperature of the concrete as placed was not more than 80°F. During cold weather, when the mean daily temperature fell below 40°F, the temperature of the concrete as placed was not less than 50°F. These procedures were in accordance with the detail concrete specification. During concrete operations, the batch plant inspector verified the mix proportions of each batch of concrete and ascertained that samples were taken and tests were made of the concrete ingredients. The batch plant inspector verified that the mixed proportions complied with those of the design mixes with the water content modified as required by measurement of surface moisture on the aggregates. The batch plant inspector also prepared a daily report to document for each batch the following: mix number, mixer cycle time, weight of each ingredient (including ice), and the batch number. Truck dispatch tickets for each batch showing the time of discharge from mixer, concrete mix number, load number, total water content, and location where used, were prepared by the inspector. Placement Distribution The majority of the fill concrete was distributed directly from the concrete batch plant to the point of placement via conveyors. The longest mn was approximately 1, 000 feet and consisted of two 250 foot belts; four 100 foot belts, and two 50 foot belts. During adverse weather conditions the conveyor belts were covered with metal hoods. To ascertain the concrete integrity from the batch plant to the point of placement, slump tests, temperature measurements, and test cylinders were made from the same batch of concrete at the beginning of the belt and at the end of the belt. These tests showed no changes in strength and no significant change in slump 3.8-40 SGS-UFSAR Revision 6 February 15, 1987 and temperature. The fill concrete was then slumped at beginning of the conveyor belts and test cylinders accompanying slump tests made at the beginning of the belt. distribution point inspector did the following: the with The 1. Visually checked each batch and estimated the slump 2. Notified pour site inspector by field telephone of the quantity of concrete conveyed from the batch plant. This was done so the pour site inspector would know when to make concrete test cylinders and perform other associated tests. For the majority of the structural concrete pours, the concrete was distributed with standard transit mix trucks which served only to transport and agitate the concrete to keep it plastic. The trucks were loaded at the batch plant from a holding hopper via a short conveyor. The distribution inspector visually checked each batch on the conveyor and estimated the slump. He also prepared a truck batch ticket showing the batch number, the time, location to be used, concrete mix code, and the total amount of water in mix. For structural pours, slump tests and test cylinders were made at the location of the pour by PSE&G Test Laboratory personnel. The PSE&G Testing Laboratory Inspector performed the following for all concrete poured: 1. When necessary to add water to truck-delivered concrete for workability, the batch ticket was checked to determine how much water (if any) could be added and assure that water was added in accordance with the following limitations: a. Maximum slump was not exceeded 3.8-41 SGS-UFSAR Revision 6 February 15, 1987
b. Total water (including that added at the pour site) was not exceeded by more than 1 gallon per yard, the amount specified in the design mix to produce a maximum allowable slump. Mixer was rotated at least 30 revolutions (at mixing speed) after addition of water. However, in no case did the total revolutions of the mixer exceed 300. c. The total water in the mix (including water added to the truck at the pour site) was shown on the Slump Test Report d. Water added at the pour site was added within 45 minutes after hatching 2. Every 20 cubic yards of concrete at each pour location was checked for slump following the procedures of ASTM Cl43-66 and results recorded. Loads with higher than allowable slump were rejected. 3. Determined the temperature of the concrete each time a slump test was made and recorded the results. Loads were rejected when temperatures required by the detail specifications were not met. 4. Made one set of concrete test cylinders for curing (6 inches by 12 inches) per ASTM C31-66 daily for each 100 cubic yards or portion thereof placed per class of concrete. A set of cylinders consists of 6 cylinders. Concrete cylinders were cured initially in accordance with Section 9 (a) of ASTM C31-66. Concrete cylinder molds conformed to the requirements of ASTM C470-65T. 5. From each load of concrete sampled for the preparation of concrete cylinders, a slump test and a temperature check was made. 3.8-42 SGS-UFSAR Revision 6 February IS, 1987
6. Prepared daily reports of field concrete poured which contained the following information: a. Date b. Location of pour (portion of structure) c. Class and quantity of concrete placed d. Number and identification of test cylinders made e. List water of concrete added at batches tested with the time, pour site, if any, slump and concrete temperature The concrete cylinders made by the PSE&G Testing Laboratory Inspector, after sufficient field curing, were transported to the Salem Job Laboratory for stripping, curing, and capping in accordance with ASTM Cl92-66. Two cylinders from each set were tested at age 7 days; three at age 28 days. If the results of the 7 and 28 day tests caused concern that the concrete did not meet specification requirements, the remaining cylinder was saved and tested at age 90 days or as directed by the Structural Engineer. Otherwise, it was discarded. Compression tests of concrete cylinders were made in accordance with ASTM C39. In addition to the compression tests, the density of the concrete was measured from the test cylinders and recorded. Concrete strength tests were evaluated by the PSE&G Structural Division, Electric Engineering Department, in accordance with ACI 214-65 and ACI 301-66, Chapter 17. If any tests for individual cylinders or group cylinders failed to reach the specified compressive strength of the concrete, the Structural Engineer was immediately notified to determine if further action would be required. 3.8-43 SGS-UFSAR Revision 6 February 15, 1987 Statistical quality control of the concrete was maintained by a computer program. This program analyzed compression test results in accordance with methods required by ACI 214, "Recommended Practices for Evaluation of Compression Test Results of Concrete." The computer results of the data analyzed included normal frequency distribution curves, standard deviations, and coefficients of variation. Placing of concrete was by bottom dump buckets, concrete pumps, or by conveyor belts. Bottom dump buckets did not exceed 3 cubic yards in size. The discharge of concrete was controlled so that concrete could be effectively consolidated around embedded items and near the forms. Vertical drops greater than 5 feet for any concrete were not permitted except where suitable equipment was provided to prevent segregation. All concrete placing equipment and methods were subjected to the approval of the Resident Structural Engineer. The surface of all construction joints were thoroughly treated to remove all laitance and loose aggregate. The construction joint surfaces in the reactor containment vessel, including all the exterior walls, were roughened to expose the coarse aggregate by cutting the surface with stiff brooms or by cutting with an air-water jet after the initial concrete set had occurred, but before the concrete had reached its final set. After cutting, the surface was washed and rinsed. Where in the opinion of the Resident Structural Engineer, the use of air-water jet or brooming as above was not advisable in a specific instance, that surface was roughened by using either hand tools or other satisfactory means to produce the requisite surface. Before placing subsequent concrete lifts, the surfaces of all construction joints were thoroughly cleaned and wetted and all 3.8-44 SGS-UFSAR Revision 6 February 15, 1987 excess water was removed. Horizontal joints were then covered with a minimum of 1/4-inch thick sand/cement grout and new concrete was then placed immediately against the fresh grout. The water-cement ratio of the grout did not exceed that of the concrete itself. Where grouting was not feasible in some areas, a bonding compound such as Colma-Fix, manufactured by Sika Chemical Corp., was used instead, on top of dry cleaned concrete. Vertical joints were wetted and slushed with a coat of neat cement immediately prior to placing the next pour. Curing and protection of freshly deposited concrete conformed to ACI 301, Chapter 12, using an absorptive material with a waterproof covering and sprinkling at intervals necessary to prevent drying for 3 days. The waterproof coverings remained in place for 7 days after the pour. Also, curing compounds, conforming to ASTM C-309, were used as required. The following select sampling and testing was made of the concrete ingredients: 1. Cement was sampled from each silo used to ascertain conformance to ASTM C-150-67 for Type II cement. The cement manufacturer also supplied certified mill reports for each silo of cement used. The storage environment effects were tested in accordance with ASTM-C-0109 and C-266. Vicat apparatus (ASTM C-191) was specified to determine the time of setting of hydraulic cement in the Preliminary Safety Analysis Report. However, it is believed that for our purpose, the Gillmore Test (ASTM C-266) is a more stringent test, in that an approximation of both initial and final set is obtained, while the Vicat Test is only addressed to initial set. For that reason, the Gillmore Test was actually conducted in the field. 3.8-45 SGS-UFSAR Revision 6 February 15, 1987
2. All concrete aggregates were delivered to the site by truck and each load was visually inspected. Also, every 250 tons received was tested for gradation and determination of fineness modulus. In addition, for sand, an organic impurity test was conducted with the gradation test. These tests were conducted per test methods and acceptance standards to ASTM C-33 with the exception that for sand gradation, the requirement for the percentage of fines was decreased. 3. Flyash from each storage bin at the source was sampled and tested in accordance with C-350-65T using acceptance standards of ASTM 618-68T, which now replaces C-350-6ST, with sampling and testing frequency as follows: a. Weekly -three composites from daily samples were checked for a carbon and surface area. b. Monthly -a composite was taken from the weekly composite or a completed chemical or physical analysis. 4. Mixing water (including ice) was checked monthly to assure that it did not contain more than 100 ppm each of chlorides, sulfides, and nitrates and that the turbidity did not exceed 2, 000 ppm of suspended solids content or 25 Formaxine Turbidity Units. Due to economic and efficiency considerations, the relatively small amounts of concrete necessary to complete the Salem Station may have been obtained from the batch plant at the Hope Creek Generating Station site. 3.8-46 SGS-UFSAR Revision 6 February 15, 1987 The mixes selected for use at Salem are approved for use in Category I (seismic) structures at Hope Creek and meet the following minimum requirements:

3.8.1.6.8.3 Reinforcing Steel Material Reinforcing steel was required by specification to conform to the following for testing methods and acceptance standards.

ASTM A-432-65 "Standard Specification for Deformed Billet Steel Bars for Concrete Reinforcement," with a Minimum Yield Strength of 60,000 psi and a Minimum Tensile Strength of 90,000 psi.

A STM A-408-Standard Specification for Special Large Size Deformed Billet Steel Bars for Concrete Reinforcement," with a Minimum Yield

Strength of 40,000 psi.

ASTM A-615 (Grade 40) "Standard Specifications for Deformed and Plain Billet-Steel Bars for Concrete Reinforcement," with a Minimum Yield

Strength of 40,000 psi and a Minimum Tensile Strength of 70,000 psi.

Reinforcement bars of the above ASTM designations also conformed to ASTM A

-305-65 "Minimum Requirements for the Deformations of Deformed Steel Bars for

Concrete Reinforcement." In addition to the ASTM requirement as to chemical composition of A-432 bars, the specification required that the carbon and manganese content did not exceed

0.45 percent and 1.30 percent, 3.8-47 SGS-UFSAR Revision 6 February 15, 1987

respectively, in order to assure better bending properties. Also, the specification required that 148 and 188 bars be subjected to 90 degree bend tests using a pin with a diameter eight times the diameter of the bar being bent, to check ductility. Certification of physical properties and chemical content of each heat of reinforcing steel delivered to the jobsite was required from the steel supplier. In addition "users' tests" were performed by a testing laboratory to confirm compliance with physical requirements and verification of mill test results. Two specimens were taken for each 25 tons or less of the full heat of steel. No sample was selected from the end 12 inches of any bar. The test was performed to determine yield point, ultimate strength, and percentage elongation. If test results did not meet specification requirements, the heat of steel was resamp1ed, this time selecting four specimens instead of the two required originally. If any specimen of the second sampling failed to meet the requirements of the specification, the entire heat was rejected. At the jobsite, reinforcing steel was kept separated by size, heat, and area to be used. At the fabricator's shop and storage area, the reinforcing steel was kept identified by size and heat. Also, when loaded for shipment from the mill, the bars were properly bundled by size, heat, and tagged with the manufacturer's identification number. Reinforcing steel for the dome, cylindrical walls, and base mat of the containment was high-strength deformed billet steel bars conforming to A8TM A-432-65. For the internal concrete of the Reactor Containment vessel, the majority of the reinforcing steel required was ASTM A-15-65, and for the large bars, ASTM A-408-65. In isolated cases, the drawings called for ASTM-A-432-65 for the internal structure. 3.8-48 8G8-UF8AR Revision 6 February 15, 1987

--Placing Placing Chapter Chapter 5 8 of reinforcing steel conformed to the requirements of of ACI 301, "Structural Concrete for Buildings, 11 and of ACI 318, "Building Code Requirements for Reinforced Concrete." No tack welding to A-432 reinforcing bars was allowed. Splices All splices of main load carrying reinforcing steel in the Reactor Containment shell were made by the cadweld process using type "T" sleeves to develop the minimum ultimate tensile strength specified by the ASTM for the grade of the bar being spliced. To ensure the integrity of the cadweld splices, the detail specification required random sampling of splices in the field. The selected splices were removed and tested to the minimum tensile strength of the bar being spliced. In some cases, the drawings required bar sizes No. 11 and smaller to be spliced by the cadweld process. In a few instances, the drawings specified other than type "T" sleeves which were required for the splicing of reinforcing to special sections. A type "B" sleeve was used to join main load carrying reinforcing bars to structural steel in order to develop the same minimum ultimate tensile strength of the bar. The detail specification required the average value of all cadweld splices tested to equal or exceed the specified minimum ultimate tensile strength of the ASTM grade of bar being spliced. In addition, no more than 5 percent of the splices tested had an ultimate strength less than 85 percent of that specified by the ASTM for the grade of bar being spliced. If any of the foregoing requirements were not satisfied, production was halted until the cause and extent of the defective splices was determined. 3.8-49 SGS-UFSAR Revision 6 February 15, 1987 Quality control of the splices was maintained by three independent procedures as follows: 1. Each crew in a program including which was by the Erica Products of Ohio. Prior to the splicing of the reinforcing bars, each operator or crew three splices for each of the positions used in production \..Jork. These samples were then tested to assure conformance with the specifications. 2. Visual inspection of every splice was made by a Quality Control The inspectors to this job attended the same progr.am as the An manual containing the recommendations of the manufacturer was issued to guide the inspector in his judgment of a satisfactory Any splices judged to be in doubt as to integrity were cut out and replaced. 3. Test splices were made by having 3-foot splices produced in sequence with the production bars. These splices were tensile tested for each crew as follows: one of the first 10 , three of the next 100 and two of the next and units of 100 In one production was randomly cut out and tested for every 100 made by each crew. Should any splice tested fail at a value less than the tensile strength required for the bar, then the splice made by the same splicer immediately preceding or following the substandard splice was cut out and tested. If this second test splice did not meet the requirements all work by this splicer was stopped and five adjacent made by the splicer were cut out and tested. If any of these an evaluation was made during which time the crew SGS-UFSAR discontinued 3.8-50 Revision 25 October 26, 2010 Also, the man who made these splices was required to requalify before performing any further production splices. Should the five splices meet the test requirements, the process was considered to be in control. In addition to the above requirements, the following procedures were used to assure acceptable splices: 1. The splice sleeve, powder, and mo] ds were stored in a clean dry area with adequate protection from the elements to prevent absorption of moisture. 2. Each splice sleeve was visually examined immediately prior to use to ensure the absence of rust and other foreign material on the inside surface. 3. The molds were preheated to drive off moisture when the molds were cold. 4. Bar ends to be spliced were previously square cut. The ends of the bars were brushed to remove mill scale, rust, and other foreign material to ensure cleanliness, and then heated. 5. A permanent point was marked from the end of each bar for a reference point to confirm that the bar ends were properly centered in the splice sleeve. 6. Before the splice sleeve was placed into final position, the bar ends were examined to ensure that the surface was free from moisture. If moisture was present, the bar ends were heated until dry. 7. Special attention was given to maintaining the alignment of sleeve and guide tube to ensure a proper fill. 3.8-51 SGS-UFSAR Revision 6 February 15, 1987

8. The splice sleeve was preheated after the materials and equipment were in position. 9. Completed splices were visually inspected to assure that fill was within acceptable limits at both ends of the splice sleeve and at the top hole in the center of the splice. The following documentation and records were maintained: 1. Mill test reports of the material furnished 2. Record of cadwelder qualification 3. Record of visual inspection of splices 4. Drawings showing splice locations 5. Record of tensile tests of splices Where accessibility of limited space precluded the use of the cadweld processes, the specifications permitted splicing by butt welding. These cases constituted less than 1 percent of the total number of splices made. Welding was performed in accordance with AWS Specification D-12.1, with double "V" groove butt joints in the horizontal position and single "V" groove butt joints in the vertical position. Welding was performed by the shielded arc processes using low hydrogen stick electrodes. Qualification of the welding procedures was made in accordance with the philosophy and intent of Section IX of the ASHE code. Full section tests were made on Grade 60, 188 reinforcing steel bars. These weld tests indicated tensile, yield, and elongation values in excess of minimum requirements of the bar material. The welding procedures required pre-heated temperatures to 325°F +/-25°F and interpassed temperatures of 300°F to 500°F. The filler 3.8-52 SGS-UFSAR Revision 6 February 15, 1987 material conformed to AWS Specification No. A 5.5-69, E-100 18-02 and certified mill test reports were required. The completed welds were post heat treated at 1050°F to 1100°F for 15 minutes and then were wrapped with a protective blanket of insulating material to avoid rapid cooling. All welds were 100 percent visually and radiographically examined. Radiographic examination was in accordance with UW 51 of the ASHE Boiler and Pressure Vessel Code,Section VIII. 3.8.1.6.8.4 Waterproofing Membrane To waterproof the subgrade exterior walls and foundations, a rubber waterproof membrane was installed under all foundations and was extended vertically up to 6 inches below yard grade. The horizontal waterproofing membrane was 1/16-inch thick Ethylene Propylene Diene Monomers (EPDM rubber). The waterproofing membrane used on vertical surfaces was 3/64-inch thick nylon reinforced Ethylene Propylene Diene Monomers (Nylon Fabric Inserted EPDM). The specification detailed the properties of the material and the ASTM test methods. The specification required that mill tests by the manufacturer for conformance be witnessed by PSE&G Testing Laboratory personnel. In addition, the manufacturer was required to supply certificates of compliance of the material with the specification. Also, users' tests were made by PSE&G on a random basis. 3.8.1.6.8.5 Compaction of Fill The detail specification required compacted fill be installed under the Class I storage tanks to 98 percent and area adjacent to Class I structure to 95 percent of the maximum dry density attainable by the AASHO Specification T-180-61 method of compaction. The specification also required that all fill material conform to Type I (all classes) or Type 4 (Classes A and B) of the New Jersey State Highway Department's "Standard 3.8-53 SGS-UFSAR Revision 6 February 15, 1987 Specification for Road and Bridge Construction, except that the amount of fines (particles passing the No. 200 sieve) could not exceed 15 percent by weight. The suitability of materials for use in the construction of compacted fill was made by a representative of PSE&G' s soils consultant (Dames and Moore). Before the installation of any compacted fill material was started and during the first week of work, a Dames and Moore representative supervised and reviewed the compaction process and the testing performed by the PSE&G Laboratory. The specification required that the fill material be compacted in 8-inch layers, and, before subsequent lifts were placed, an inplace density test was made by the PSE&G Testing Laboratory personnel in accordance with AASHO T191-64 (ASME Specification D1566-64, Density of Soil in Place by the Sand Cone Method). Where test results indicated that the required density was not attained, the materials were reconditioned and recompacted to the required density. The preparation of the Optimum Moisture -Dry Density Curves was done by the PSE&G Testing Laboratory and Dames and Moore in accordance with AASHO T180-61 (ASTM Specification D1557-64). Also, particle size analysis of soils was in accordance with AASHO T-88 (ASTM D422-63). United Engineers and Constructor's Quality Control inspected the compaction operation and coordinated the testing requirements. All reports, tests, and other pertinent information were documented at the site by United Engineers and Constructor's Quality Control. 3.8. 1.6.8.6 Liner Plate General The fabrication and erection of the Reactor Containment liner, personnel locks, equipment hatch, and liner plate attachments were 3.8-54 SGS-UFSAR Revision 6 February 15, 1987
  • *
  • performed by CB and I. PSE&G or its agent reviewed all test results and monitored all work. Material The detail specification for the Reactor Containment* liner required that the steel for the main shell, including the dome, cylindrical walls, and the bottom, be low carbon/high manganese steel with fine grain structure, meeting. ASTM Specification A442-66, Grade 60. In addition, the liner material was* impact tested in accordance with the 1968 edition, ASME Boiler and Pressure Vessel Code, Section Ill, Paragraph Nl2ll, at a temperature JO*F below the minimum service temperature of so*F. Mill test reports certifying the physical and chemical properties of the plate delivered to the jobsite were submitted. The Reactor Containment liner was fabricated and erected in accordance with Part UY, "Requirements for Unfired Pressure Vessels Fabricated By Welding,"Section VIII of the ASME Boiler and Pressure *. <assel Code, 1968 edition, and PSE&G Detail Specification No. 68-7123. Where any conflict was evident, the PSE&G Specification was followed. The qualification of all welders and welding procedures was performed in accordance with Part A,Section IX, of the ASME Boiler and Pressure Vessel Code, 1968 edition. To assure good welds the following were required as a minimum: 1. Thorough cleaning of weld preparations 2. Removal of slag from previous passes 3. Proper control of welding current and polarity 3.8-55 SGS*UFSAR Revision 13 June 12, 1994
4. Make certain welding materials and base materials were dry before welding. Heating of base material in the vicinity of weld when temperature wa$ below 70*F 5. Shielding welding arcs from winds and drafts Evaluation of porosity in spot radiography was .in accordance with. the standards of Appendix 4 of Section VIII, ASME Boiler and Pressure Vessel Code, 1968 edition. Standards for field welding. were in accordance with the requirements of Section VIII of the ASME Boiler and Pressure Vessel Code, 1968 edition. The liquid penetrant inspection of the liner plate welds was in accordance with Appendix 8, "Methods of Liquid Penetrant Examination,"Section VIII of the ASME Boiler and Pressure Vessel Code, 1968 edition. Inspection of the liner seam welds was accomplished as follows: 1. A trained inspector responsible for welding Quality Control inspected every weld. 2. For the bottom liner plates, liquid penetrant and/or magnetic particle inspections of 2 percent of the weld seams was performed. In addition, the first 10 feet of weld made by each welder was also liquid penetrant and/ or magnetic particle inspected. 3. All the liner bottom plate welds were 100 percent vacuum box tested to 5 psi pressure differential with atmospheric pressure. 4. The liner plate seam welds in the cylindrical walls and dome were 2 percent radiographed for each welder and positioned in accordance with UW52. In addition, the 3.8-56 SGS-UFSAR Revision 6 February 15, 1987 first 10 feet of weld made by each welder was 100 percent radiographed. The following preliminary tests were made during the liner erection using the test channels: 1. All welds were covered by channels and zoned after which a strength test was performed by applying 54 psig air pressure to the channels in a zone for a period of 15 minutes. The exposed welded joints were given a soap test for leaks. If bubbles indicated a leak, the leak was repaired and the zone retested. 2. The zones of channels were then retested to a pressure of 4 7 psig with a 20-percent, by weight, freon-air mixture. The entire run of the channel to plate weld was traversed* with a halogen leak detector. If a leak was detected, repairs were made and a retest performed. 3. In addition, the zone of channels was held at the 47 psig air pressure for a period of 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br />. When pressure drop exceeded the standards, zones were repaired and retested. 1. Overall out of roundness below El. 130 feet: +/- 2 inches -above El. 130 feet: +/- 4 inches 2. Overall deviations out of plurnbness -l/500 of the height. 3.8-57 SGS-UFSAR Revision 6 February 15, 1987 Equipment Hatches, Personnel Locks, and Penetration Sleeves The detail fication required that the equipment hatches, except as noted '-' for the outage equipment hatch, and personnel locks be made from ASTM A-516, Grade 60, conforming to ASTM A-300 requirements. In addition, the steel was normalized by heating to 1,700°F (+/- 50°) and cooling in still air and "V" notch tested to a minimum of 15 foot pounds at -4 0°F in accordance with ASTM fication A-370-65. For the material on the interior bulkhead which is not subjected to low temperatures, the material was required to pass a Charpy "V" Notch test of 15 foot pounds at +20°F, in accordance with ASTM Specification A-370-65. The OEH conforms to ASME Code,Section VIII material requirements. Access hatches, except as noted, were fabricated and tested at the fabrication The personnel locks and any portion of the equipment access door extending beyond the concrete shell conforms to the requirements of the 1968 edition of the ASME,Section III for Class B Vessels. The chambers were completely fabricated in the shop and the weldments were inspected by the Hartford Insurance Group. The chambers were not stamped as a Class B pressure vessel, but were covered by an equivalent to a Manufacturer Partial Data Form N-lA for the personnel locks and covered by a Manufacturer Partial Data Form for the equipment hatches. The user and his authorized representative rr.oni tared test at the fabrication shop and audited all records. OEH conforms to the requirements of ASME code,Section VIII. The piping penetration sleeves are carbon steel and are described in Section 3.8.i.6.8.10. The electrical penetration sleeves are of Schedule 80 caroon steel. The piping and electrical penetration sleeves were welded to the liner assemblies at the fabrication shop in accordance with ASME,Section III, Class B vessels, except for stamping. The weldments were inspected in the The Hartford Insurance Group and are covered by a Manufacturer Partial Data Form. Strengtn and leak tests were performed in the shop on all access hatches. Leak test of the penetration sleeve welded channel was also performed in the shop. 3.8-58 SGS-UFSAR Revision 18 April 26, 2000 The shop leak test procedures were conducted in accordance with Appendix A of ANS 7. 60 "Proposed Standard for I,eak Rate Testing of Reactor Containment Structure for Nuclear Reactor." Proof tests were applied to chambers to pressurize the necessary areas to 54 psig. The pressure was maintained a sufficient time to allow soap bubbles and Freon sniff tests of all welds and mating surfaces. Any leaks found were repaired and retested. The repair of defective welds was in accordance with paragraph N-528 of Section III, "Nuclear Vessels t" of the ASME Boiler and Pressure Vessel Code, 1968 edition. Attachments to Steel Liner Nelson Studs Nelson studs were welded to the plating as shown on the drawings. Each welder, at the beginning of each day, attached at least one test stud which was tested by bending the stud approximately 45 degrees toward the plate to demonstrate the integrity of the weld. If failure occurred in the weld, the welding procedure or technique was correctedt and two successive studs successfully welded and tested before further studs were attached to the liner plate. These test studs were allowed to remain in place but were not considered as part of the regular stud pattern required by the design. All studs on which a full 360-degree weld was not obtained were removed and replaced with a new stud. Before welding a new stud, where a defective one han been removed, the area was ground flush and smooth. Stiffeners were welded to the exterior of the dome. stiffeners conformed to ASTM A-36 material. These 3.8-59 SGS-UFSAR Revision 6 February 15, 1987 3.8.1.6.8.7 Construction Construction at the site started with clearing and the installation of the first stage of a Dewatering System. The first stage consisted of installing a Peripheral Ejector System at existing grade (Elevation 99+ 1, PS datum) utilizing approximately 230 ejector wells, each 90-feet deep and double 12-inch header system connecting the wells to three diesel driven pumps. Piezometers were installed adjacent to the area to be excavated to determine the ground water level at all times. Daily readings were taken and recorded to ascertain that the ground water level remained well below the excavation level. When the Dewatering System started drawing the ground water down, an open excavation sufficiently large to accommodate all major structures was started. At an approximate depth of 23 feet below existing encircled grade, a cellular cofferdam was constructed which the excavation for all the major or Class I building structures. The cofferdam consists of 24 circular cells, 60 feet in diameter, with connecting arcs, and "toed" into the Vincetown Strata approximately 10 feet. The area within the cofferdam was then excavated to Elevation 43 feet (PS datum) and a second stage and separate Dewatering System was installed. This system consisted of approximately 140 well points, each 24-feet deep with an 8-inch header connecting the wells to one self-priming centrifugal pump. Later, this piping system was encased in lean concrete fill, abandoned in place and filled under pressure with grout. The total Dewatering Systems pumped between 1, 000 to 1,400 gpm continuously. Then the final stage of excavation within the cofferdam area proceeded to the Vincetown Strata. During this final stage of excavation, a Consulting Soils Engineer was at the site on a continuous basis. The Soils Engineer, the Resident Engineer, and a Quality Control Supervisor visually inspected the bottom of the excavation to verify that the excavation had reached the top of Vincetown Formation prior to 3.8-60 SGS-UFSAR Revision 6 February 15, 1987 placing any lean concrete. Prior to the completion of the excavation, at approximately Elevation 45 feet, 15 exploratory borings were drilled through the remaining Kirkwood Formation and into the underlying Vincetown to verify the original study borings. These additional borings showed no measurable differences from the study borings. In addition, after the Vincetown Strata had been exposed, six test borings were drilled in the excavated area into the underlying Vincetown Strata to verify and ensure that the foundation mat was, in fact, directly supported on the Vincetown Formation. Four of these borings were drilled under the Unit 2 Reactor Containment; two were drilled under the Unit 1 Reactor Containment; all borings penetrated a minimum of 20 feet into the underlying Vincetown Formation. During cold weather the exposed Vincetown was protected from freezing with insulated blankets; also, in some areas a 2 to 3-foot thick earth "blanket" was used for protection from both frost and construction traffic. The lean concrete poured over the Vincetown Strata had an overall average strength of 3509 psi with a 16.45-percent coefficient of variation as verified statistically by computer. The concrete was placed via conveyors directly from the batch plant to the point of deposit, under the quality control procedures stated in previous sections. After the lean concrete was poured and screened to the bottom level of the foundation mat, a 1/16-inch thick rubber, ethylene Propylene Diene Monomer waterproofing membrane was installed. A 1/8-inch thick hard board was installed over the membrane and then a 3-inch thick concrete protection course was installed. Later, the waterproofing membrane was extended vertically up the foundation walls with 3/64-inch thick nylon reinforced rubber and protected with 1/8-inch thick hardboard. The reactor containment base mats for Units 1 and 2 were poured in 6 segments and 8 segments, respectively. Vertical construction joints were constructed with expanded wire mesh. No horizontal joints were permitted. 3.8-61 SGS-UFSAR Revision 6 February 15, 1987 The base mat was poured to a level 6 inches below the final elevation of the bottom liner plate. The backing "T's" were then leveled to final position and concrete was poured flush with the top of the backing "T' s". Then the bottom liner plate was installed and the welds over the backing 11T' s" made. The knuckle plate was installed and the cylindrical portion of the liner was erected to Elevation 120 feet (PS datum). Reinforcing steel for the outer walls were then placed. The exterior concrete containment wall was constructed in 5-foot lifts during the construction of the interior concrete. All welds were checked for compliance with requirements. the approved weld inspection and test The reactor containment interior concrete was built on the mat liner. On the completion of the interior concrete structure to Elevation 130 feet (PS datum), the polar crane was then erected. Concrete in the exterior wall was poured in uniform 5-foot lifts around the entire circumference. The completed steel wall liner was braced internally and locally with temporary bracing to prevent distortion during concrete placement. On completion of the concrete exterior walls to Elevation 100 feet (PS datum), a rubber waterproof membrane was attached to the exterior concrete surface with adhesives and a 1/8-inch protection board installed to protect the membrane. The space between the excavation and the reactor containment structure was then backfilled and compacted in 8-inch layers to 95 percent of the maximum dry density in accordance with procedures. The liner was completed, finishing with the construction of the 1/2-inch thick steel dome, with all welds constructed and tested in accordance with procedures previously outlined. The steel dome liner was supported during erection from a steel tower which was erected at Elevation 130 feet. Sections of the 3.8-62 SGS-UFSAR Revision 6 February 15, 1987 steel dome liner were lifted into place, braced with wind girders, and welded into their final locations. 3.8.1.6.8.8 Penetrations In general, a penetration consists of a sleeve embedded in the concrete wall and welded to the containment liner. The weld to the liner is shrouded by a continuous channel which is test pressurized to demonstrate the integrity of the penetration-to-liner weld joint. The pipe, electrical conductor, duct, or equipment access hatch passes through the embedded sleeve and the end.of the resulting annulus is closed off, either by welded end plates, bolted flanges, or a combination of these. Provision has been made for differential expansion and misalignment between each pipe and sleeve. No piping loads are imposed on the liner. Pressurizing connections are provided to demonstrate the integrity of the penetration assemblies. There are three large openings that significantly perturb the reinforcing pattern. One is the equipment hatch with an 18-foot diameter outer barrel; the others are two personnel hatches with 9 foot-9 inch diameter outer barrels. As a rule of thumb, only openings having diameters greater than 2 1/2 times the wall thickness are considered to require detailed analysis. The design and analysis for these three large openings are described in Section 3.8.1.6.8.9. The main wall reinforcing, consisting of vertical and horizontal reinforcing bars, is bent around all the openlngs. Contlnui ty of shell reinforcement is therefore maintained. For large openings, in addl.tion to these bars, circular reinforcing bars have been provided to take care of axial thrust and principal moments around the opening. Radial stirrups have been provided to take care of the torsion and shear. This combination of reinforcing bars takes care of all primary and secondary stresses. 3.8-63 SGS-UFSAR Revision 6 February 15, 198?

3.8.1.6.8.9 Equipment and Personnel Access Hatches Equipment and personnel access hatches, except as noted for the outage equipment hatch, OEH, are fabricated from A516, Grade 60 steel normalized to A300 requirements. All personnel locks and the portion of the equipment access hatch extending inside the containment structure beyond the concrete shell are designed in accordance with ASME Boiler and Pressure Vessel Code,Section III, Class B. waived. The Code was used as a guide; therefore the N Stamp requirement is The equipment and personnel hatch details are as shown on VTDs 301051 and 301075, respectively. The OEH is designed in accordance with ASME VIII. The hatch barrel is embedded in the containment wall and welded to the liner. Provision is made to test pressurize the space between the double gaskets of the door flanges and the weld seam channels at the liner joint, hatch flanges, and dished door. The personnel hatches will be double door, mechanically latched, welded steel assemblies. A quick-acting type equalizing valve connects the personnel hatch with the interior of the containment vessel for the purpose of equalizing pressure in the two systems when entering or leaving the containment. The personnel hatch doors are interlocked to prevent both 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 door operational status. Provision is made to permit bypassing the door interlocking system to allow doors to be left open during plant cold shutdown. Each door lock hinge is designed to be capable of independent three-dimensional adjustment to assist proper seating. An Emergency Lighting and Communication System powered from an external emergency power supply are provided in the lock interior. Emergency access to either the inner door, from the containment interior; or to the outer door, from outside, is possible by the use of special door unlatching tools. Plan and elevation drawings of the personnel air lock with all electrical and piping penetrations identified are provided on Figure 3. 8-37 and VTDs 30107 5 and 301059. 3.8-64 SGS-UFSAR Revision 27 November 25, 2013 Pressure and monitoring taps are provided to pressure test the double gaskets on each door to a "between the seals test pressure" of 10 psig. When testing for seal leakage during periods of when the air locks are frequently opened. Test connections are also provided to permit pressurization of the entire air lock. Leakage rate testing of the airlocks is described in Section 6.2.1. Tie-downs are used to prevent the inner door from becoming unseated during pressure tests. The yoked ends of the tie-downs are pin connected to the horizontal stiffeners at the door, and the threaded ends of the tie-downs are slipped through the holes of the tie-down beams and secured with nuts. The instruction manual for the personnel air locks requires tightening the nuts to draw the door flange to approximately 1/16 inch from the bulkhead flange before pressurization. There is no monitoring device to read the force exerted on the door. This mechanism cannot be operated from within the air lock. Around these hatch openings, thickened concrete edge beams are provided to take care of the high membrane and bending stresses in the opening areas. Loading combinations 1, 2, and 3 as listed in section 3.8.1.3 were used in the large opening analysis. Tornado stresses were found to be far less critical than stresses resulting from accident pressure or earthquake. Manual design utilized the elastic center method for ring beam analyses. The method is described and illustrated in "Statically Indeterminate Structures" by L. c. Maugh (4). The ring beams are designed to resist biaxial bending moments, axial tension, torsion, and biaxial shear. conrad Associates* finite element analysis (3) for the three large openings is based on the procedure formulated for linearly elastic, thin shells of arbitrary geometry which uses a fully compatible plate bending element to obtain the bending stiffness, and a constant strain triangle element to determine the membrane 3.8-65 SGS-UFSAR Revision 16 January 31, 1998 stiffness. The results of these analyses are shown on Figures 3.8-39 through 3.8-48. From the more refined computer analysis, modifications have been made in the preliminary penetration details to strengthen the reinforcement arrangement in areas where higher stresses have resulted from the independent check. Details for personnel and equipment hatch reinforcements are as shown on Plant Drawing 201131. 3.8.1.6.8.10 Piping Penetrations High integrity piping penetrations are provided for all piping passing through the containment. Figures 3. 8-49 and 3. 8-50 show typical cold and hot pipe penetrations respectively. The pipe is centered in the embedded sleeve which is welded to the containment liner. Seal plates are welded to the pipe at both ends of the sleeve. In some instances several pipes pass through the same embedded sleeve to minimize the number of penetrations required. In such cases, each pipe is welded to the inside seal plate and to the expansion bellows which is, in turn, welded to the outside seal plate. Large single pipe containment penetrations were installed with expansion test bellows, attaching the process piping to the penetration sleeves, which allowed for Appendix J type "B" pressure testing of the compartment formed between the process piping and the embedded sleeve, via a test connection on the bellows. Containment piping penetrations designed for Salem are not required to be type "B" tested for 10CFR50 Appendix J (Ref. Safety Evaluation S-C-R700-MSE-0253 Rev. 0). The type "B" test is applicable to piping penetrations that utilize expansion bellows as the leakage limiting boundary. The piping penetrations at Salem rely on partial/full penetration seal welds inside containment as the leakage limiting boundary, which are leak rate tested as part of the Appendix J type "A" containment Integrated Leak Rate Test (ILRT). 3.8-66 SGS-UFSAR Revision 27 November 25, 2013 Therefore, for containment penetrations, leak rate testing of separate penetrations (type .i)" testing) has been replaced by the containment integrated leak rate test (type "A" testing) as allowed by lOCFRSO Appendix "J". As a result, the abandoned expansion test bellows on the (10) service water lines penetrating the containment have been eliminateo as part of the service water system piping upgrade program. Figure 3.8-SOA shows a typical service water piping containment penetration detail. In the case of piping carrying hot fluid, the pipe is insulated and cooling is provided to limit the concrete temperature adjacent to the embedded sleeve to lSO*F. For the larger hot pipe penetrations, strong anchoring is necessary and is provided as shown. The anchors engage a large segment of the wall to adequately resist thrusts. Should a piping failure occur within the containment, the additional loading imposed upon the penetration is transmitted through the anchor to the containment structure. Therefore no permanent deformation of the pen.,tration will be realized. Moment eliminators are installed outside of the containment structure. Hangers and limit stops assist in supporting and reducing any moment loading of a free-hanging pipe. 3.8-66a Revision 13 June 12, 1994 THIS PAGE INTENTIONALLY LEFT BLANK 3.8-66b SGS-UFSAR Revision 13 June 12, 1994 A multiple guide arrangement is used to limit the high frequency vibrations which could be imposed on the piping penetrations and surrounding liner regions by vibrating pipes which pass through the containment structure. Further, the penetration sleeve and containment liner are anchored in the concrete so that the liner does not participate in vibration. The design of the equipment, pipes, support structures, and penetrations is in accordance with the stress limitations of Section III, ASHE Nuclear Vessel Code, and the ANSI B31.1, Code for Pressure Piping. A thermally induced loading is deformation at the liner-sleeve stress was calculated in both considered to produce uniform interface. The average liner the horizontal and vertical directions and the maximum moment and maximum hoop stress in the sleeve was computed. The thickness of the s] eeve was chosen so that the stresses do not exceed the allowable stresses. The material used for penetrations is carbon steel and conforms with the requirements of the ASHE Nuclear Vessel Code. As required by the Nuclear Vessel Code, the penetration material meets the necessary Charpy V-Notch impact values at a temperature of 20°F. The penetrations are not exposed to the elements and, therefore, are not subject to 0°F outside temperature. Penetration expansion bellows are suitably protected against field damage; such protection remains a part of the permanent installation. The originally specified material for all penetration sleeves was ASTM A-155-KC 70, Class I. Due to design modifications, A106 Grade B, A106 Grade C, Al55-KC 60, and A333 Grade I material has been utilized for some sleeves. In such cases, seamless pipe was used where possible. Where it was required, a heavier wall thickness than that specified for the origjnal Al55-KC 70 Class I was used, so that in all these cases, sleeves of equal or greater strength than the original design have been used. 3.8-67 SGS-UFSAR Revision 6 February 15, 1987 3.8.1.6.8.11 Electric Penetrations Power, control, fiber optic, and shielded conductors are assembled in canisters which have been inserted in and welded to nozzles in the field. Figure 3.8-51 shows typical electrical penetrations. A prototype of each type of penetration has been factory tested at 271F and 62 psig in a steam chamber. Tests prove the ability of prototypes to function properly, electrically and mechanically, before, during, and after subjection to these conditions. Each penetration is factory tested before shipment to verify that the leakage rate does not exceed 1 x 10-6 cc/sec at one atmosphere differential when tested with dry helium.

There are 56 electrical penetrations per unit.

The penetration sleeves to accommodate the electrical penetration asse mblies are Schedule 80 carbon steel, except where otherwise noted.

3.8.1.7 Testing and Inservice Inspection Shop leak testing procedures have been conducted in accordance with Appendix A of ANS 7.60, "Proposed Standard for Leak Rate Testing of Containment Structures for Nuclear Reactor." A proof test is applied to each penetration which pressurizes the necessary areas to 54 psig. This pressure is maintained a sufficient time to allow soap bubble tests of all welds and mating surfaces.

It is retested to a pressure of 47 psig for leakage rate.

All penetrations, the personnel locks, and the equipment hatches are designed

with double seals which are pressure tested at 54 psig.

Individual penetration internals and closures and sleeve weld channels are leak

tested at 47 psig after installation.

3.8-68 SGS-UFSAR Revision 29 January 30, 2017

Information and inservice testing and additional information on pre-service testing is contained in Section 6.2.1. 3.8.2 Steel Containment System Since the containment is a reinforced concrete structure lined with only a thin steel plate leak-tight barrier, this section does not apply. For equipment hatch, personnel access hatches, and penetrations, see Section 3. 8. 1. 3.8.3 Internal Structures 3.8.3.1 General Description For arrangement of the walls, elevated slabs, and components of the internal structures see Figure 3.8-1 and Plant Drawing 208900. see Sections 3.8.1.1 and 3.8.3.4. 3.8.3.2 Design Codes For further description, The internal structures have been designed under the following codes: 1. Building Code Requirements for Reinforced Concrete, ACI 318-63. 2. AISC Manual of Steel Construction, 6th Edition or later edition, as applicable. 3.8-69 SGS-UFSAR Revision 27 November 25, 2013 3.8.3.3 Loads and Loading Combinations The following design load criteria were used in the design of all compartment structures. (a) C -1.0 D +/- 0.05 D + 1.5 P + 1.0 T (b) C -1.0 D +/- 0.05 D + 1.25 P + 1.0 T + 1.25 E (c) C -1.0 D +/- 0.05 D + 1.0 P + 1.0 T + 1.0 E' + J (d) C-1.0 D +/- 0.05 D + 1.0 T + 1.0 E' + J' where: c D p T E E' J the capacity of the structure the dead load the pressure differential of 15 psi the thermal effect the OBE the DBE the greater of 950 kips jet force or pipe rupture load transmitted by restraints J' the greater of 1500 kips jet force or pipe rupture load transmitted by restraints For additional information see Internal Pressure Analysis in Section 3.8.1.4. The load combinations utilized in the design of the internal structures were either equivalent to or more conservative than those outlined in the SRP. The following tabulations provide a comparison combinations utilized with the SRP criteria. of 3.8-70 SGS-UFSAR Revision 13 June 12, 1994 load (2) SRP Salem (2b) SRP Salem (3) SRP Salem (4) SRP Salem (5) SRP Salem (6) SRP Salem INTERNAL CONCRETE STRUCTURES 1. 4D + 1. 7E 1. 9E Not Critical 0.75 (1.4D + 1.71 + 1.9E = 1.7 T + 1.7R) 0 0 Less Critical Than (5) D + L + T + R + E' 0 0 Less Critical Than (6) D + L + T + R + l.SP a a D + L + T + R + l.SP a a a D + L + T + R + I. 25P a a + 1. 25E a + (Yr + Yj + Ym) D + L + T + R + 1.25P + (Yr + Yj + Ym) + 1.25E a a D + L + T + R + P + (Yr + Yj + Ym) + E' a a a D + L + T + R + P + (Yr + Yj + Ym) + E' a a See FSAR Section 3. 8. 1. 3 for identification of symbols. Although the R and Y forces are not listed in the overall structural analysis load combination formulae, the local effects under piping load, jet load, and missile impingement were taken into account. 3.8.3.4 Design and Analysis Provisions The reactor vessel is surrounded and supported by a massive circular wall known as the primary shield, which also supports the central portion of the refueling canal walls above it and the operating deck which rests on the top of the refueling canal walls. 3. 8-71 SGS-UFSAR Revision 6 February 15, 1987 I The reactor vessel cavity wall, known as the "Primary Shield," was designed as a thick cylinder subjected to internal pressure, dead load, design and operating basis earthquakes, jet forces, and guillotine breaks of the reactor coolant loop piping entering and leaving the reactor and entering the steam generator. Cognizance was taken of the discontinuity moments and shears at the base of the cylindrical wall. Figures 3. 8-52 through 3. 8-55 show the plan and cross sections of the vessel cavity with pressure and forces indicated,* and a sketch showing thick wall stress analysis. These pressures and forces were used as conservative desi.gn criteria. The values resulting from the final analysis are well below these magnitudes. The reactor pressure vessel cavity extends vertically from Elevation 54 feet to Elevation 104 feet, Above Elevation 104 feet, the area becomes part of the refueling canal. Below Elevation 81 feet, the cavity walls are inherently safe from any disaster as they are embedded in the massive concrete foundation of the building. As of the original design of the primary shield, in the event of a failure of a reactor pressure vessel nozzle, pressure relief is provided by B blowout plugs, approximately 10 square feet in area each, venting to Elevation 104 feet. Additionally, there is always a 2-inch air gap between the vessel flange and the wall at Elevation 104 feet, except during refueling. A pressure of 900 psi was used for the design of the nozzle penetrations through the shield wall to withstand a longitudinal break in the nozzle extensions. The reactor cavity is designed for a pressure of 175 psi acting upon the walls following a break in the vessel nozzle. A guillotine break of the nozzle was also considered. The reaction in this case is taken by the reactor supports, which are integral with the reactor cavity walls. Subsequent to the original design, leak-before-break (LBB) was approved for the primary loop piping {see Section 3. 6. 4). LBB allows the elimination of the dynamic effects of pipe rupture, including sub-compartment pressuriZcation. Thus, the dynamic pressure loads resulting from a break in the primary piping can be eliminated from the design of the primary shield, also eliminating the need for the pressure relief function of the blowout plugs. In order to reduce personnel radiation exposure and critical path time during refueling outages, the blowout plug covers that were installed and removed during every refueling are now left in place auring normal operation. The existing analysis of the primary shield, which includes breaks in the primary piping with pressure relief through the blowout plugs, is conservatively retained. 3.8-72 SGS-UFSAR Revision 23 October 17, 2007 * * *

  • *
  • During seismic loading the entire internal structure could be subjected to torsional stresses resulting from a 5-percent accidental eccentricity. For information concerning the extent to which the internal structure and its walls are capable of withstanding the stresses resulting from the accidental eccentricity, see Section 3.7. The following structures are removable; provisions have been made to prevent them from becoming missiles. Hatch covers at Elevation 130 feet They are 4-foot thick massive concrete blocks bolted down to the slab to withstand any uplift pressure. Missile shield over Reactor Vessel The missile shield over the reactor vessel consists of a 181-inch diameter, 2-inch thick steel plate that is permanently attached to the Integrated Head Assembly (IHA). It is secured to prevent it from becoming a missile. See UFSAR Section 5.5.14.1 for a description of the IHA . Lead blocks at Elevation 100 feet in fuel transfer canal area and at Elevation 89 feet-6 inches to 95 feet-10 inches transfer chamber They are for radiation protection for the gap between containment exterior wall and interior structure. Blocks are held in place by angle frames and steel plates to hold them during earthquake motion. Hatch cover over fuel transfer chamber at Elevation 100 feet It is a steel enclosure filled with poured lead. It is laterally 3.8-73 SGS-UFSAR Revision 23 October 17, 2007 restrained, subject to no jet force and could never become a missile. 3.8.4 Other Category I Structures 3.8.4.1 Summary Description The Category I structures other than the containment structures are listed in Section 3. 2. They are Auxiliary Building, Fuel Handling Buildings, Service Water Intake Structure, Class I water tank foundations, and the Class I equipment supports. The orientation of the principal structures is shown on Plant Drawing 201012. The general arrangements of Containment Buildings, Auxiliary Building, and Fuel Handling Buildings are shown in Section 5. 3.8.4.2 Design Codes Category I structures were designed under the following codes: 1. Building Code Requirements for Reinforced Concrete, ACI 318-63. 2. AISC Manual of Steel Construction, 6th Edition or later edition, as applicable. 3.8.4.3 Loads and Loading Combinations Loads and load combinations for Category I structures under SSE or tornado conditions are similar to those of the containment except for LOCA pressure and temperature loadings. and OBE conditions. The working stress design method was used for operating (No load factor was used for OBE loadings.) For combination of containment loadings and load symbols, see Section 3.8.1.3. 3.8-74 SGS-UFSAR Revision 27 November 25, 2013 The tornado and tornado generated missile analyses for category l structures are provided in Section 3.8.1.4. The load combination caseD specified tornado loads combined with operating loads. The tornado load (Wt) includes the static forces produced by the 360 mph maximum wind velocity, a 3 psi negative pressure, and the structural response to the missile impact. The stresses on any structural member produced by the effective pressure transformed from the tornado wind, the impact of the missile, and also differential pressure were superimposed to obtain the most critical total stress, provided the induced stress from these three components are in the same direction. When one of the components induced an opposite stress, thereby reducing the total stress in the member, it was neglected. In other words, all six loading combinations listed in the SRP have been considered with factors of one instead of 0.5 for Wp in combinations iv and vi and also have taken into account stress directions as stated above. category I water storage tanks are not designed to withstand tornado-induced missiles. Sufficient backup water sources are available to assure a safe shutdown of the reactor and the ability to maintain the reactor in a safe condition. Certain Category I components, such as emergency diesel generator combustion air intakes and exhausts, etc., are located outside the category I buildings and are not designed to withstand tornado-generated missiles. Additionally, certain openings in the roof slabs and external walls of the category I buildings are not designed to withstand tornado-generated missiles. For these components and openings, a probabilistic evaluation of the tornado missile strikes was completed for Unit 1 and Unit 2. The TORMIS computer code that was used is based on the methodology given in EPRI Report NP-2005-CCM, "Tornado Missile Simulation and Design Methodology -Computer Code Manual," August 1981. This code calculates the probabilities of tornado missile strike on specified targets. The full spectrum of missiles described in NUREG-0800, Section 3. 5 .1. 4, which envelopes the Salem missiles described above, was used for these probability assessments. The cumulative probability of missile strikes on all the unprotected components and openings for each of the two Salem units was determined to be less that 10-6, which is acceptable per the guidance provided in Regulatory Guide 1.117 and SRP Section 2.2.3. Venting of structures under tornado generated differential pressure was not adopted as a design criterion for category I structures. These structures are capable of sustaining the differential pressure generated by a tornado. Metal siding on the Turbine Generator Building (non-Category I), however, was designed to be blown out to relieve tornado generated differential pressure. 3.8-75 SGS-UFSAR Revision 16 January 31, 1998 category I structures are not additionally loaded or affected by the adjacent non-category I structures since the non-Category I structures (adjacent to Category I structures) are heavily braced to withstand tornado wind forces such that they will not collapse on the Category I structures. Five heavy bays of steel bracing system have been installed in the turbine-generator area structure to withstand the static horizontal shear force under earthquake condition. With this bracing system, the Turbine Building will not collapse on the adjacent Category I structures under OBI or DBB loads. The load combinations utilized in the design of Category I structures were either equivalent to or more conaervative than those outlined in the SRP. The working stress design method waa used for the category I concrete structures under operating and OBB loadings. The Salem design is in conformance with load combination a(i)2a in SRP 3.8.4. The strength design method was used for the Category I structures under SSE/tornado and DBA loadings. The Salem design is in conformance with load combination (6) and (8) in SRP 3.8.4. Both load combinations represent the moat severe cases. Hydrostatic loadings from the hurricane condition were applied to the structures to check their stability. The procedures used by our consultant (Dames and Moore) for transferring the static and dynamic floor effects to load were as delineated in the u. s. Army coastal Engineering Research center Technical Report No. 4. Total head, including wave effects, was considered to investigate the lateral and overturning effects. The buoyancy effect of groundwater was included in the assessment of the sliding and overturning potential of all Category I structures. The buoyancy effect will reduce the dead weight and thus reduce the factors of safety against sliding and overturning. To include the buoyancy effect in assessing the sliding and overturning potential is the more conservative and correct approach. 3.8-76 SGS-UFSAR Revision 16 January 31, 1998 --

The safety against sliding, overturning, and flotation for all Category I structures under all loading combinations are within the limits set by the SRP 3.8.5. Masonry Walls For the loading criteria for non-structural masonry walls see Section 3.8.4.5.1. 3.8.4.4 The and Procedures structures have for normal been on ACI 318-63 OBE and "Ultimate "Working Design," for normal loads DBE or tornado. In the under OBE the allowable stresses are one-third above the normal code working stresses. Wind stresses are found to be less critical than those for an OBE. Load factors of have been used in the ultimate design under DBE or tornado loading. The stress of reinforcing steel under ultimate has been under 0. 9 Fy. The reduction factor "0" as described in Section 3.8.1.4.1 for concrete stress is for all Category I structures. A coefficient "k" of 0.85 for 3500 psi concrete has been used in addition to "0" for equivalent rectangular concrete stress distribution. During the design phase of the re-racking which was implemented in 1994, the Spent Fuel Pool was reanalyzed for increased fuel storage capacity. The following is the list of items incorporated in the analysis: 1) The "ultimate strength" design method based on NUREG-0800, 2) Standard Review Plan 3.8.4, Rev. 1, 1981 was used. The plant design spectra, given in Salem UFSAR, events were broadened, per provisions of Reg. used. for DBE and OBE Guide 1.122 and 3) The response spectrum method was used to determine the self-excitation loading on the pool structure. 4) The pool structure was modeled in three dimensions via a 3-D 5) SGS-UFSAR finite element model. The thermal across the slab and the pool walls was computed using finite element method. Thus, the effect of interaction between the ambient, pool water, and is fully incorporated in the 3.8-77 Revision 25 October 26, 2010

6) The pressure on the lower portion of the wall during a seismic event undergoes a cyclic pulsation due to the hydrodynamic coupling between racks and the pool walls. This loading was quantified using Whole Pool Multi-Rack analysis. This loading was included in the analysis. 7) Analyses have been performed that evaluate the spent fuel pool structure (reinforced concrete as well as the stainless steel liner and its anchors) for a boiling pool condition. The acceptance criteria used for these evaluations was ACI 359 (ASME Code Section III Division 2, Reference 7). The spent fuel pool structure, including the liner and its anchors, was shown to meet all requirements of this code for a boiling pool condition. Analyses were also performed for the spent fuel pool liner and its anchors for a maximum normal pool temperature of 150°F. The liner and its anchors were shown to meet the acceptance criteria of ACI 359. Thermal cycling of the liner and its anchors was evaluated and shown to not be a concern. Steel members inside the Category I structures are designed in accordance with the AISC Manual of Steel Construction (Sixth Edition or later edition, as applicable). Seismic design criteria are described in Section 3. 7. Tornado and tornado generated missile design is described in Sections 3.3.2 and 3.5.2. Four independent seismic analyses, similar to those for the Containment Building, have been performed for the 1) Auxiliary Building, 2) Fuel Handling Building, 3) Service Water Intake Structure, and 4) Outer Penetration Building. Conservative results have been utilized for the building design. The time history computer analysis calculations are kept on file with PSE&G. Specifically, the information for the Auxiliary Building and the Fuel Handling Building has been provided in the Reference 3 report, while the information for the Service Water Intake Structure and Outer Penetration Building has been provided in the Reference 5 report. The loading combinations used for Category I (seismic) steel structures other than containment are as follows: 3.8-78 SGS-UFSAR Revision 17 october 16, 1998 Working stress design 1. D + L + I + H Allowable stresses in accordance with AISC Manual of Steel Construction 2. D T 1 + I + H + E Allowable stresses are one-third above the normal allowable stresses. Ultimate 1. D + L + I + H + E' 2. D + L + I + H + T The stress in the ultimate strength design has been kept under 0.9 Fy 3.8-78a SGS-UFSAR Revision 25 October 26, 2010 THIS PAGE INTENTIONALLY BLANK 3.8-78b I SGS-UFSAR Revision 17 October 16, 1998 where: D "' Dead load L = Live load I "" Impact load where moving load is present H = Thermal load E = Operating Basis Earthquake E' "" Design Basis Earthquake T = Tornado loading Protection against tornado wind loads and tornado missiles is discussed in Sections 3.3.2 and 3.5.2. Protection against turbine disc rupture and missile generation is described in Section 3.5.4. Masonry Walls For analysis of non-structural masonry walla see section 3.8.4.5.1. 3.8.4.4.1 HVAC Duct and Support Methodology Original Seismic Design Methodology The original seismic design assessed selected members (i.e., those with potential nonductile failure modes) for compliance with the stress limits. In addition, limiting support dimensions and tolerances, configurations, spans and member sizes were provided which ensured that the duct systems and supports' frequencies are rigid. The effective sheet metal structural properties are based on the 1969 version of the American Iron and Steel Institute (AISI) derivation of effective width for thin sections with stiffened compression elements subject to local buckling. HVAC Duct Seismic Adequacy Verification Methodology An alternate approach has been developed to provide consistent criteria and methodology for functional and seismic qualification of HVAC duct systems and supports. This methodology utilizes an alternative approach based on review of industry standard codes and practices, past earthquake performance data and shake table test results. All known credible failure modes for HVAC duct systems and supports, when subjected to earthquake loadings, were specifically documented. Then, engineering efforts were focused onto these credible failure modes and specific guidance for elimination of the failure mode or criteria for maintaining 3.8-79 SGS-UFSAR Revision 16 January 31, 1998 a suitable margin of safety against the failure mode were developed. The methodology anaures that the resulting margins of aafaty for aeiamic loadings and documentation requirement& are conaiatent with those of USI A-46 program requirement& par the SQUG GIP. All applicable seismic desiqn basis data and criteria from the SQOG GIP are adopted, and supplemented by teat results and stress requirements, as follows: 1. The GIP criteria for fans, air handlers and dampers that are directly applicable to seismic evaluation of BVAC duct systems are used. 2. The GIP criteria for expansion anchors and welded anchorage& are ulled directly. The higher margine of aafety for new inatallationa aa specified in the GIP are also used directly. 3. The SQUG GIP includes detailed guideline& and criteria for evaluation of raceway system supports, including limited analytical reviews with equivalent static load factors. The BVAC duct methodology adopts a similar approach, revising the load factors to account for differences in damping between raceways (as evidenced by dynamic testa) and HVAC ducts (limiting damping to St of critical for Design Basis Earthquakes as stated in the UFSAR for bolted steel structures). 4. The SQUG GIP includes detailed caveat& and inclusion rules based on earthquake experience and dynamic teat programs for each equipment class and raceways. Similar caveats and inclusion rules, based on review of the same earthquake experience data base and other industry HVAC duct system teat programs, are included in the methodology. s. The GIP criteria for spatial seismic interaction evaluations are used directly. The acceptable stress limits for Design Basis Earthquake loading are baaed on the industry standard working level streas allowablea, increased by standard factors used for nuclear seismic designs. For pressure loading, stress limits for the ducting sheet metal and duct stiffeners are in accordance with the requirements of industry standard codes (SMACNA). Implementation of this methodology demonstrates that the existing installations are adequate for seismic and pressure loadings. These implementation results exemplify that the new methodology baa margins of safety consistent with or exceeding the original design basis. 3.8-79a SGS-UFSAR Revision 16 January 31, 1998 3.8.4.5 Materials, Quality Control, and Special Construction Techniques 3.8.4.5.1 Masonry Walls There is no masonry block construction in the Containment and Fuel Handling Building. In the Auxiliary Building and penetration area, removable block walls are reinforced with steel bars and also anchored to the slab. These provisions are used to prevent the wall from collapsing under earthquake forcesi however, they are not considered as major shear walls to carry the lateral forces for the building. 3.8-79b SGS-UFSAR Revision 16 January 31, 1998 All of the masonry walls that have been installed within (and between) Category I structures and adjacent to Category I tanks have been re-evaluated for seismic loadings and are found to be within the following two groups: 1. Those walls whose collapse would endanger or affect in any way the safety of any Category I structure 2. Those walls whose collapse would not affect the safety of any Category I structure A structural analysis has been performed on each of the walls whose collapse would affect any of the Category I structures to determine the shear and bending stresses to assess their margin of safety. For the re-evaluation and analysis, the masonry wall field testing and inspection, the design for the corrective action, the structural steel reinforcing, and the drawings, see "Report on Re-evaluation of Masonry Walls," dated November 28, 1980, which was submitted to the NRC on December 10, 1980. Corrective actions for those masonry walls which do not meet the NRC criterion (1.33 times allowable ACI shear or tensile stress for mortar when the wall is subjected to out of plane bending during an SSE) are detailed in a PSE&G letter (Liden to Varga) dated December 8, 1982. 3.8.5 Foundations 3.8.5.1 Description All major Category I structure foundations, except the tank foundations, are built on top of lean concrete fill, which in turn bears on the Vincetown formation at approximately Elevation 30 feet. The actual elevation of top Vincetown was established by visual inspection and additional borings after the excavation was completed. The Service Water Intake Structure was constructed within a 3.8-80 SGS-UFSAR Revision 17 October 16, 1998 -

steel sheet cofferdam. The material within the cofferdam has been removed down to the Vincetown formation and all the less compact materials on top of the formation were removed and replaced with tremie concrete to the bottom of the mat. The Category I water storage tank foundation is a 3-foot concrete mat serving as the top of a pipe trench. The trench foundation rests on compacted backfill brought up from the top of the lean concrete fill at Elevation 79 feet. The profile of the principal plant structures, cofferdams, and subsurface formations are shown on Plant Drawing 201012. plant structures is given in Table 3.8-11. A summary of foundations for Seismic separation joints for building foundation mats adjacent to each other are provided to allow independent motion of each building under earthquake conditions. 3.8.5.2 Applicable Codes, Standards, and Specifications All foundations were designed according to all of the applicable sections of the same codes, standards, and specifications as the buildings and structures which they support. These are listed in Sections 3.8.1.2, 3.8.3.2, and 3. 8. 4. 2. 3.8.5.3 Loads and Load Combinations All foundations were designed according to all of the applicable loads and load combinations as the buildings and structures which they support. listed in Sections 3.8.1.3, 3.8.3.3, and 3.8.4.3. These are 3.8.5.4 Design and Analysis Procedures The containment base mat is analyzed as a rigid circular plate subjected to loadings from the axisymmetric exterior cylinder 3.8-81 SGS-UFSAR Revision 27 November 25, 2013 wall, crane wall, interior walls, and equipment acting around an equivalent circle. The soil pressure is found in a conventional manner without the benefit of its elastic formation. Our manual analysis was based on the ACI Paper, Title No. 63-63, "Analysis of Circular and Annular Slab for Chimney Foundation," by Kuang-Han Chu and omar F. Afandi. A finite element program was used to check the rebar under five loading combinations. Since the mat is covered by 2 to 5-feet thick of concrete slab and also the lower 34 feet of cylinder liner is insulated, the thermal effect on the mat has been neglected. The Service Water Intake Structure and Category I water tank foundation seismic design has been based on the manual dynamic model analyses using the average response spectra as the ground motion input. Other category I structure foundation mats were designed for all loading combinations as described in Section 3.8.4.3. 3.8.6 References for Section 3.8 1. Timoshenko and Woinowsky-Kriegar, "Theory of Plates and Shells," McGraw-Hill, 1959. 2. Timoshenko and Goodier, "Theory of Elasticity," McGraw-Hill, 1951. 3. "Structural Analysis of Containment Vessel -Salem Nuclear Generating Station," Conrad Associates, van Nuys, california, 1970. 4. Maugh, L. c., "Statically Indeterminate Structures," John Wiley and Sons, New York, New York, 1946. 5. VTD 32023 7-01, "Design Basis Response Analysis of the Salem Nuclear Generating Station Structures," EQE Final Report, January, 1995. 6. 1995 ASME Boiler & Pressure Vessel Code,Section III, Division 2, Code for Concrete Reactor Pressure Vessels and Containments. 7. ACI 359-95 (ASME Boiler and Pressure Vessel Code section III Division 2), "Code for Concrete Reactor Vessels and Containments." 3.8-82 SGS-UFSAR Revision 17 October 16, 1998

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