LR-N17-0034, Salem Generating Station, Units 1 & 2, Revision 29 to Updated Final Safety Analysis Report, Chapter 4, Sections 4.3.1 to 4.3-61
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4.3 NUCLEAR DESIGN 4.3.1 Design Bases This section describes the design bases and functional requirements used in the nuclear design of the Fuel and Reactivity Control System and relates these design bases to the General Design Criteria (GDC) in lOCFRSO Appendix A. Where appropriate, supplemental criteria such as the Final Acceptance Criteria for Emergency Core Cooling Systems are addressed. Before discussing the nuclear design bases, it is appropriate to briefly review the four major categories ascribed to conditions of plant operation. The full spectrum of plant conditions is divided into four categories, in accordance with the anticipated frequency of occurrence and risk to the public: 1. Condition I -Normal Operation, 2. Condition II -Incidents of Moderate Frequency, 3. condition III -Infrequent Faults, 4. Condition IV -Limiting Faults. In general, the Condition I occurrences are accommodated with margin between any plant parameter and the value of that parameter which would require either automatic or manual protective action. condition II incidents are accommodated with, at most, a shutdown of the reactor with the plant capable of returning to operation after corrective action. Fuel damage* is not expected during Condition I and Condition II events. It is not possible, however,
- Fuel damage as used here is defined as penetration of the fission product barrier (i.e., the fuel rod clad). 4.3-1 SGS-UFSAR Revision 6 February 15, 1987 to preclude a very small number of rod failures. These are within the capability of the plant cleanup system and are consistent with the plant design basis. Condition III incidents shall not cause more than a small fraction of the fuel elements in the reactor to be damaged, although sufficient fuel element damage might occur to preclude immediate resumption of operation. The release of radioactive material due to Condition III incidents should not be sufficient to interrupt or restrict public use of these areas beyond the exclusion radius. Furthermore, a Condition III incident shall not, by itself, generate a Condition IV fault or result in a consequential loss of function of the Reactor Coolant System (RCS) or reactor containment barriers. Condition IV occurrences are faults that are not expected to occur but are defined as limiting faults which must be designed against. Condition IV faults shall not cause a release of radioactive material that results in an undue risk to public health and safety. The core design power distribution limits related to fuel integrity are met for Condition I occurrences throuqh conservative design and maintained by the action of the Control system. The requirements for Condition Il occurrences are met by providing an adequate protection system which monitors reactor parameters. The Control and Protection Systems are described in section 7, and the consequences of Condition II, III, and IV occurrences are given in Section 15. 4.3.1.1 Fuel Burnup Basis The fuel rod design basis is described in section 4.2. The nuclear design basis is to install sufficient reactivity in the fuel to attain the average region discharge burnup 4.3-2 SGS-UFSAR Revision 17 October 16, 1998 values given in reference 28. The above, along with the design basis in Section 4.3.1.3, Control of Power Distribution, satisfies GOC-10. Discussion Fuel burnup is a measure of fuel depletion which represents the integrated energy output of the fuel (MWD/MTU} and is a convenient means for quantifying fuel exposure criteria. The core design lifetime or design discharge burnup is achieved by installing sufficient initial excess reactivity in each fuel region and by following a fuel replacement program (such as that described in section 4.3.2) that meets all safety-related criteria in each cycle of operation. Initial excess reactivity installed in the fuel, although not a design basis, must be sufficient to maintain core criticality at full power operating conditions throughout cycle life with equilibrium xenon, samarium, and other fission products present. The end-of-design cycle life is defined to occur when the chemical shim concentration is essentially zero with control rods present to the degree necessary for operational requirements. In terms of chemical shim boron concentration this represents approximately 10 ppm with no control rod insertion. A limitation on initial installed excess reactivity is not required other than as is quantified in terms of other design bases such as core negative reactivity feedback and shutdown margin discussed below. 4.3.1.2 Negative Reactivity Feedbacks (Reactivity Coefficient) The fuel temperature coefficient will be negative and the moderator temperature coefficient of reactivity will be non-positive for power operating conditions, thereby providing 4.3-3 SGS-UFSAR Revision 17 October 16, 1998 negative reactivity feedback characteristics. The design basis meets GDC-11. compensation for a rapid increase in is there are two or effects. These are the resonance absorption effects (Doppler) associated with changing fuel temperature and the spectrum effect resulting from changing moderator density. These basic physics characteristics are often identified by reactivity coefficients. The use of slightly enriched uranium ensures that the Doppler coefficient of reactivity is negative. This coefficient provides the most designed to have an overall reactivity compensation. The core is also coefficient of moderator so that average coolant or void content slower effect. Nominal power is only in a range of overall non-positive moderator coefficient. The non-positive moderator temperature coefficient can be achieved through use of fixed burnable absorber, integral fuel burnable absorber (IFBA) and/or control rods by limiting the reactivity held down by soluble boron. Burnable absorber content (quantity and distribution) is not stated as a design basis other than as it relates to accomplishment of a non-positive moderator coefficient at power conditions discussed above. 4.3.1.3 Control of Power Distribution Basis The nuclear design basis is that, with at least a 95 percent confidence level: 1. The fuel will not be operated at greater than 13.3 kW/ft under normal conditions SGS-UFSAR an allowance 4.3-4 Revision 25 October 26, 2010 of 0.6 percent for calorimetric error and including densification effects. 2. Under abnormal conditions including the maximum overpressure condition, the fuel peak power will not cause melting as defined in Section 4. 4. 1. 2. 3. The fuel will not operate with a power distribution that violates the departure from nucleate boiling (DNB) design basis (i.e., the DNB ratio (DNBR) shall not be less than the safety limit, as discussed in Section 4.4.1.1) under Condition I and II events including the maximum I overpower condition. 4. Fuel management will be such as to produce rod powers and burnups consistent with the assumptions in the fuel rod mechanical integrity analysis of Section 4.2. The above basis meets GDC-10. Discussion Calculation of extreme power shapes which affect fuel design limits is performed with proven methods and verified frequently with measurements from operating reactors. The conditions under which limiting power shapes are assumed to occur are chosen conservatively with regard to any permissible operating state. To ensure that the axial profile meets with the linear heat rate limit and the DNB limit, ex-core detector signals are used to provide a top to bottom flux difference, AI, which is input, through F{AI), into both the overpower AT and overtemperature AT trip points. Even though there is good agreement between measured peak power calculations and measurements, a nuclear uncertainty margin is applied to calculated peak local power. Such a margin is provided 4.3-5 SGS-UFSAR Revision 20 May 6, 2003 both for the analysis of normal operating states and for anticipated transients. 4.3.1.4 Maximum Controlled Reactivity Insertion Rate The maximum reactivity insertion rate due to withdrawal of rod cluster control assemblies or by boron dilution is limited. This limit, expressed as a maximum reactivity change rate (75 pcmjsec)* is set such that peak heat generation rate and DNBR do not exceed the maximum allowable at overpower conditions. satisfies GDC-25. This The maximum reactivity worth of control rods and the maximum rates of reactivity insertion employing control rods are limited so as to preclude rupture of the coolant pressure boundary or disruption of the core internals to a degree which would impair core cooling capacity due to a rod withdrawal or ejection accident (See Section 15). Following any condition IV event (rod ejection, steamline break, etc.) the reactor can be brought to the shutdown condition and the core will maintain acceptable heat transfer geometry. This satisfies GDC-2B. Discussion Reactivity addition associated with an accidental withdrawal of a control bank (or banks) is limited by the maximum rod speed (or travel rate) and by the worth of the bank(s). For this reactor the maximum control rod speed is 45 inches per minute and the maximum rate of reactivity change considering two control banks moving is less than 75 pcm per second.
- l pcm = 105 4p (See footnote Table 4.3-2) 4.3-6 SGS-UFSAR Revision 6 February 151 1987 4.3.1.5 Minimum shutdown margin as in the Technical is at any power operating condition in the hot shutdown condition and in the cold shutdown condition. In all analyses involving reactor trip, the single, highest worth Rod Cluster Control Assembly (RCCA) is postulated to remain um::ripped in its full-out position (stuck rod criterion). This satisfies GDC-26. Two independent and soluble boron in control systems are provided, namely control rods the coolant. The Control Rod System can compensate for the reactivity effects of the fuel and water temperature changes accompanying power level changes over the range from full-load to no-load, Control Rod System provides the minimum shutdown margin In addition, the under Condition I events and is capable of making the core subcritical rapidly enough to prevent exceeding fuel damage limits assuming that the worth control rod is stuck out upon 'I'he Boron can for all xenon burnout and will maintain the reactor in cold shutdown. Thus, backup and emergency shutdown provisions are provided by a mechanical and a chemical shim control system which satisfies GDC-26. When fuel assemblies are in the pressure vessel and the vessel head is not in kef£ will be maintained at or below 0.95 with control rods and soluble boron. Further, the fuel will be SGS-UFSAR 4.3-7 Revision 25 October 26, 2010 maintained sufficiently subcritical that removal1of all RCCAs will not result in criticality. ANS Standard NIB. 2 specifies a keff not to exceed 0. 95 in spent fuel storage racks and transfer equipment flooded with pure water and a keff not to exceed 0. 98 in normally dry new fuel storage racks assuming optimum moderation. No criterion is given for the refueling operation; however a 5 percent margin, which is consistent with spent fuel storage and transfer and 3 percent below the new fuel storage, is adequate for the controlled and continuously monitored operations involved. 4.3.1.6 Stability Basis The core will be inherently stable to power oscillations of the fundamental mode. This satisfies GDC-12. Discussion Oscillations of the total power output of the core, from whatever cause, are readily detected by the loop temperature sensors and by the nuclear instrumentation. The core is protected by these systems and a reactor trip would occur if power increased unacceptably, preserving the design margins to fuel design limits. The stability of the Turbine/Steam Generator/Core Systems and the Reactor Control System is such that total core power oscillations are not normally possible. The redundancy of the protection circuits ensures an extremely low probability of exceeding design power levels. Basis Spatial power oscillations within the core, with a constant core power output, should they occur, can be reliably and readily detected and suppressed. 4.3-8 SGS-UFSAR Revision 6 February 15, 1987 Discussion The core is designed so that diametral oscillations due to spatial xenon effects are self-damping and no operator action or control action is required to suppress them. The stability of diametral oscillations is so great that this excitation is highly improbable. Convergent azimuthal oscillations can be excited by prohibited motion of individual control rods. Such oscillations are readily observable and alarmed, using the ex-core long ion chambers. Indications are also continuously available from in-core thermocouples and loop temperature measurements. The BEACON core monitoring system (Reference 34) uses nearly continuous measurements of the excore detector ion chambers, core thermocouples, control rod indications, and loop temperature measurements for updating of the BEACON core model using the 3D-ANC (Advanced Nodal Code) neutronics solution (Reference 36) to provide nearly continuous power distribution monitoring. Moveable in-core detectors can be activated to provide more detailed and accurate information for calibration of the thermocouple data and measured spatial power distribution data used by BEACON. If BEACON were to become inoperable, the in-core flux mapping process would be used in place of BEACON to monitor core power distribution. In all presently proposed cores these horizontal plane oscillations are self-damping by virtue of reactivity feedback effects designed into the core. However, axial xenon spatial power oscillations may occur late in core life. The control bank and ex-core detectors are provided for control and monitoring of axial power distributions. Assurance that fuel design limits are not exceeded is provided by reactor overpower and overtemperature trip functions which use the measured axial power imbalance as an input. 4.3.1.7 Anticipated Transients Without Trip The effects of anticipated transients with failure to trip are not considered in the design bases of the plant. Analysis has shown that the likelihood of such a hypothetical event is negligibly small. Furthermore, analysis of the consequences of a hypothetical failure to trip following anticipated transients has shown that no significant core damage would result and system peak pressures would be limited to acceptable values and no failure of the RCS would result (1). 4.3-9 SGS-UFSAR Revision 27 November 25, 2013 4.3.2 Description The reactor cores consist of a specified number of fuel rods which are held in bundles by spacer grids and top and bottom The fuel rods are f . J . l TM l ' d ' l constructed o.: .. oy or Zlr o cy n.ca tubes 002 fuel pellets. 'l.'he bundleR, which approximates a known as fuel as.<;;emblies, circular cylinder. are arranged in a pattern E:ach fuel assembly contains a 17 :x: n rod array composed of 264 fuel rods, 24 rod cluster control {RCC) th:l.mbles and an in-cor.:*e instrumentation thimble. Figure 4. 2-1 shows a cross sectional view of a 17 x 17 fuel assembly and the related RCC locations. Further details of the fuel assembly are given in Section 4.2.1. For initial core loading, the fuel rods within a given assembly have the same uranium enrichrnen*t in both the .r.adial and cndal J<'uel af.;scmblies of three different enrichments are used in the initial core loadings to establish a favorable radial power cU.stribution. regions consisting of two lower enrichments are so as to form a checkerboard pattern in the central portion of the core. The third region is arranged around the periphery of the core and contains the highest enrjchment. The enrichments for the first cores are shown in Table 4.3-1. A reJ.oad pattern is a low leakage loading pattern or a low-low leakage A low loading pattern has either burned (depleted) or fresh (.feed) assemblies arranged on the core periphery. A low-low leakage pat tern has only burned (depleted) assemblies on the core periphery. The reload cores a:re normally designed to be able to operate approximately eighteen months between refuelings. The typical loading patterns for both Salem Units are shown in Figures 4.5-1 and 4.5-3. 4.3-10 SGS-UE'SAR Revision 23 October 17, 2007 *
- The core average enrichment is determined by the amount of fissionable material required to provide the desired cycle energy requirements. The physics of the burnout process is such that operation of the reactor the amount of fuel available due to the of neutrons by the U-235 atoms and their fission. The rate of U-235 is proportional to the power level at which the reactor is In addition, the fission process results in the formation of fission products, some of which absorb neutrons. These effects, depletion and the buildup of fission products, are partially offset by the buildup of plutonium shown on Figure 4.3-1 for the 17 x 17 fuel assembly, which occurs due to non-fission absorption of neutrons in U-238. Therefore, at the beginning of any cycle a reactivity reserve to the depletion of the fissionable fuel and the buildup of fission reactor. This excess over the life must be "built" into the is controlled by boron dissolved 1n the coolant and burnable absorbers. '::'he concentration of boric acid in the primary coolant is varied to provide control and concentration to of compensate the soluble for long-term reactivity requirements. neutron absorber is varied c:o compensate The for reactivity xenon and changes due to fuel burnup, fission product poisoning including samarium, burnable absorber depletion, and the cold-to-operating moderator temperature Using its normal makeup path, the Chemical and Volume Control (CVCS) is of at a rate of 30 when the reactor coolant boron concentration is 1000 ppm and approximately 35 pcm/min when the reactor coolant boron concentration is 100 pprn. The peak burnout rate for xenon is 25 pcm/min. Rapid transient reactivity requirements and safety shutdown requirements are with control rods. As the boron concentration is increased, the moderator coefficient becomes less The use of a soluble absorber alone would result in a moderator coefficient at (BOL). Therefore, burnable absorbers are used to reduce the soluble boron concentration to ensure that the moderator temperature coefficient is for 9ower operating conditions. During the absorber content in the burnable absorbers is depleted thus adding positive reactivity to offset some of the negative reactivity from fuel depletion and fission product build'Jp. The 4. 3-11 SGS-UFSAR Revision 25 October 26, 2010 I depletion rate of the burnable absorbers is not critical since chemical shim is always available and flexible enough to cover any possible deviations in the expected burnable absorber depletion rate. Figure 4.3-2 is a graph of a typical core depletion with burnable absorbers. As stated previously, the purpose of burnable absorbers (integral and discrete) is to provide enough boron loading in the core to decrease core soluble boron concentrations to the point that a negative moderator temperature coefficient is maintained at all hot operating conditions. IFBA boron loadings in the range from 1. 57 to 2. 35 mg/inch per rod is typical. Two different discrete burnable absorbers (borosilicate glass-PYREX) and Wet Annular Burnable Absorbers (WABA) may be used. Both discrete absorbers have similar burnout characteristics and provide similar benefits to core design development which include reduction of the number and loading of the IFBA, which is a fuel performance (rod internal pressure, corrosion, and DNB propagation) benefit. Typical stack lengths for IFBA and WABA range from 108 to 132 inches. PYREX rods are 139 inches in length. Design data for IFBA, WABA, an PYREX can be seen in Table 4. 3-1. Reference 32 of this section provides a more detailed discussion of WABA. In addition to reactivity control, the burnable absorbers, both discrete {WABA and/or PYREX) and integral (IFBA), are strategically located to provide a favorable radial power distribution. Reload loading patterns utilize various burnable absorber types and distributions. These are determined on a cycle-specific basis. The typical burnable absorber loading patterns are shown in Figures 4.5-2 and 4.5-4. Tables 4.3-1 through 4.3-3 contain a summary of the reactor core design parameters for a typical fuel cycle, including reactivity coefficients, delayed neutron fraction, and neutron lifetimes. Some of these parameters change on a reload basis. The current values or allowable range of values may be found in the appropriate Reload Safety Evaluation (RSE)/Safety Assessment (SA) or Nuclear Design Report (NDR)/Curvebook (CB)/Plant Operations Package (POP), see Section 4.5. The RSE/SA is a cycle-specific document which evaluates the fuel assembly design and core loading pattern configuration for plant safety. The NDR/CB/POP are cycle-specific documents which include such information as power distributions, reactivity coefficients, boron and control rod worths under a variety of plant operating conditions. There is sufficient information in the RSE/SA and NDR/CB/POP to permit an independent calculation of the nuclear performance characteristics of the core. 4.3-12 SGS-UFSAR Revision 23 October 17, 2007 * *
- 4.3.2.2 Power Distribution The accuracy of power distribution calculations has been confirmed through approximately one thousand flux maps during some 20 years of operation under conditions very similar to those expected for the plant described herein. Details of this confirmation are given in Reference 2 and in Section 4.3.2.2.7. 4.3.2.2.1 Definitions Power distributions are quantified in terms of hot channel factors. These factors are a measure of the peak pellet power within the reactor core and the total energy produced in a coolant channel and are expressed in terms of quantities related to the nuclear or thermal design, namely: Power density is the thermal power produced per unit volume of the core (kW/liter). Linear oower density is the thermal power produced per unit length of active fuel (kW/ft). Since fuel assembly geometry is standardized, this is the unit of power density most commonly used. For all practical purposes it differs from kW/liter by a constant factor which includes geometry effects and the fraction of the total thermal power which is generated in the fuel rod. Average linear power density is the total thermal power produced in the fuel rods divided by the total active fuel length of all rods in the core. Local heat flux is the heat flux at the surface of the cladding (Btu-ft-2-hr-1). For nominal rod parameters this differs from linear power density by a constant factor. Rod power or rod integral power is the length integrated linear power density in one rod (kW). 4.3-13 , SGS-UFSAR Revision 17 October 16, 1998 Average rod power is the total thermal power produced in the fuel rods divided by the number of fuel rods (assuming all rods have equal length). The hot channel factors used in the discussion of power distributions in this '- section are defined as follows: FQ, Heat Flux Hot Channel Factor, is defined as the maximum local heat flux on the surface of a fuel rod divided by the average fuel rod heat flux, allowing for manufacturing tolerances on fuel pellets and rods. FNQ, Nuclear Heat Flux Hot Channel Factor, is defined as the maximum local fuel rod linear power density divided by the average fuel rod linear power density, assuming nominal fuel pellet and rod parameters. E F Q, Engineering Heat Flux Hot Channel Factor. is the allowance on heat flux required for manufacturing tolerances. The engineering factor allows for local variations in enrichment, pellet density and diameter, surface area of the fuel rod, and eccentricity of the gap between pellet and clad. Combined statistically the net effect is a factor of 1.03 to be applied to fuel rod surface heat flux. N F 4H, Nuclear Enthalpy Rise Hot Channel Factor. is defined as the ratio of the integral of linear power along the rod with the highest integrated power to the average rod power. Manufacturing tolerances, hot channel power distribution and surrounding channel power distributions are treated explicitly in the calculation of the departure from nucleate boiling ratio described in Section 4.4. 4. 3-14 SGS-UFSAR Revision 6 February 15, 1987
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- It is convenient for the purposes of discussion to define sub factors of F 0; however, design limits are set in terms of the total peaking factor . F0 Total peaking factor or heat flux hot-channel factor Maximum kW/ft Average kW/ft without densification effects where: F N and FE are defined above. Q Q the measurement uncertainty; 1.05 if a full core flux map was performed with the movable detectors. ( u ) 1.0+-Q-100.0 if BEACON (PDMS) was used for power distribution monitoring; where U0 is defined by Equation 5-19 of Reference 34 for the use of BEACON. Note: All peaking factor uncertainties for power distribution measurements are located in the Core Operating Limits Report for each specific cycle. N F XY ratio of peak power density to average power density in the horizontal plane of peak local power . 4.3-15 SGS-UFSAR Revision 19 November 19, 2001 P(Z) = ratio of the power per unit core height in the horizontal plane at height Z to the average value of power per unit core height. 4.3.2.2.2 Radial Power Distribution The power shape in horizontal sections of the core at full power is a function of the fuel and burnable absorber loading patterns and the presence or absence of a single bank of control rods. Thus, at any time in the cycle a horizontal section of the core can be characterized as unrodded or with group D control rods. These two situations combined with burnup effects determine the radial power shapes which can exist in the core at full power. The effect on radial power shapes of power level, xenon, samarium, and moderator density effects are considered also but these are quite small. The effect of non-uniform flow distribution is negligible. While radial power distributions in various planes of the core are often illustrated, the core radial enthalpy rise distribution as determined by the integral of power up each channel is of greater interest. Figures 4.3-3 through 4.3-7 show 4.3-16 . SGS-UFSAR Revision 17 October 16, 1998 representative low leakage radial power distributions for one eighth of the core for representative operating conditions. These conditions are ( 1) Hot Full Power (HFP) at BOL -unrodded -no xenon, (2) HFP at BOL -unrodded -equilibrium xenon, {3) HFP at BOt -Bank D in -equilibrium xenon, (4) HFP at Middle-of-Life (MOL) -unrodded -equilibrium xenon, and (5) HFP at EOL -unrodded -equilibrium xenon. The radial power distribution, and hence radial enthalpy rise distribution, changes on a cycle-specific basis. The appropriate NOR should be referenced. see section 4.5 for the current cycles* predicted radial enthalpy distribution. Since the position of the hot channel varies from time to time a single reference radial design power distribution is selected for DNB calculations. This reference power distribution is chosen conservatively to concentrate power in one area of the core, minimizing the benefits of flow redistribution. Assembly powers are normalized to core average power. 4.3.2.2.3 Assembly Power Distribution since the detailed power distribution surrounding the hot channel varies based on within-assembly design features and time in life, a conservatively flat assembly power distribution is assumed in the DNB analysis, described in Section 4.4, with the rod of maximum integrated power artificially raised to the design value of F:a* The design value of is determined during the RSE. The current cycles* is given in Section 4.5. Care is taken in the nuclear design of all fuel cycles and all operating conditions to ensure that a flatter assembly power distribution does not occur with limiting values to 4.3.2.2.4 Axial Power Distribution The shape of the power profile in the axial or vertical direction is largely under the control of the operator through either the manual operation of the control rods or automatic motion of control rods responding to manual operation of the eves. Nuclear 4.3-17 SGS-UFSAR Revision 17 October 16, 1998 effects which cause variations in the axial power shape include moderator density, Doppler effect on resonance absorption, spatial xenon, burnable absorbers, and burnup. Automatically controlled variations in total power output and control rod motion are also important in determining the axial power shape at any time. Signals are available to the operator from the ex-core ion chambers which are long ion chambers outside the reactor vessel running parallel to the axis of the core. Separate signals are taken from the top and bottom halves of the chambers. The difference between top and bottom signals from each of four pairs of detectors is displayed on the control panel and called the flux difference, 4!. Calculations of core average peaking factor for many plants and measurements from operating plants under many operating situations are associated with either 41 or axial offset in such a way that an upper bound can be placed on the peaking factor. For these correlations axial offset is defined as: axial offset and and are the top and bottom detector readings. Representative axial power shapes from Reference 3 for BOL, MOL, and EOL conditions are shown on Figures 4.3-8 through 4.3-10. These figures cover a wide range of axial offset including values not permitted at full power. 4.3.2.2.5 Deleted SGS-UFSAR 4.3-18 Revision 17 October 16, 1998 THIS PAGE INTENTIONALLY BLANK 4.3-19 SGS-UFSAR Revision 17 October 16, 1998 4.3.2.2.6 Limiting Power Distribution According to the ANS classifications of plant conditions (see section 15), condition I occurrences are those which are expected frequently or regularly in the course of power operation, maintenance, or maneuvering of the plant. As such, Condition I occurrences are accommodated with margin between any plant parameter and the value of that parameter which would require either automatic or manual protective action. In as much as Condition I occurrences occur frequently or regularly, they must be considered from the point of view of affecting the consequences of fault conditions (Conditions II, III, and IV). In this regard, analysis of each fault condition described is generally based on a conservative set of initial conditions corresponding to the most adverse set of conditions which can occur during Condition I operation. The list of steady state and shutdown conditions, permissible deviations (such as one coolant loop out of service) and 4.3-20 SGS-UFSAR Revision 17 October 16, 1998 operational transients is given in Section 15. Implicit in the definition of normal operation is proper and timely action by the reactor operator. That is, the operator follows recommended operating procedures for maintaining appropriate power distributions and takes any necessary remedial action when alerted to do so by the plant instrumentation. Thus, as stated above, the worst or limiting power distribution which can occur during normal operation is to be considered as the starting point for analysis of Condition II, III, and IV events. Improper procedural actions or errors by the operator are assumed in the design as occurrences of moderate frequency (Condition II). Some of the consequences which might result are listed in Section 15.1. Therefore, the limiting power shapes which result from such Condition II events are those power shapes which deviate from the normal operating condition at the recommended axial offset band, e.g., due to lack of proper action by the operator during a xenon transient following a change in power level brought about by control rod motion. Power shapes which fall in this category are used for determination of the reactor protection system setpoints so as to maintain margin to overpower or DNB limits. The means for maintaining power distributions within the required hot channel factor limits are described in the core Operating Limit Report ( COLR). A complete discussion of power distribution control in Westinghouse Pressurized Water Reactors (PWRs) is included in Reference 5. Detailed information on the design constraints on local power density in a Westinghouse PWR, on the defined operating procedures and on the measures taken to preclude exceeding design limits is presented in the Westinghouse topical report on peaking factors (Reference 6). The following paragraphs summarize these reports and describe the calculations used to establish the upper bound on peaking factors. The calculations used to establish the upper bound on peaking factors, FQ and include all of the nuclear effects which 4.3-21 SGS-UFSAR Revision 17 October 16, 1998 influence the radial and/or axial power distributions throughout core life for various modes of operation including load follow, reduced power operation, and axial xenon transients. Radial power distributions are calculated for the full power condition, and fuel and moderator temperature feedback effects are included for the average enthalpy plane of the reactor. The steady state nuclear design calculations are done for normal flow with the same mass flow in each channel and flow redistribution effects neglected. The effect of flow redistribution is calculated explicitly where it is important in the DNB analysis of accidents. The effect of xenon on radial power distribution is small (compare Figures 4.3-3 and 4.3-4, respectively) but is included as part of the normal design process. Radial power distributions are relatively fixed and easily bounded with upper limits. The core average axial profile, however, can experience significant changes which can occur rapidly as a result of rod motion and load changes and more slowly due to xenon distribution. For the study of points of closest approach to axial power distribution limits, several thousand cases are examined. Since the properties of the nuclear design dictate what axial shapes can occur, boundaries on the limits of interest can be set in terms of the parameters which are readily observed on the plant. Specifically, the nuclear design parameters which are significant to the axial power distribution analysis are: 1. Core power level 2. Core height 3. Coolant temperature and flow 4. Coolant temperature program as a function of reactor power 4.3-22 SGS-UFSAR Revision 17 October 16, 1998
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- 5. Fuel cycle lifetimes 6. Rod bank worths 7. Rod bank overlaps 8. Burnable absorber length and placement Normal operation of the plant assumes compliance with the following conditions: 1. Control rods in a single move together with no individual rod insertion differing by more than 12 steps above 85% RTP or more than 18 steps at or below 85% RTP from the bank demand position. Reference 35 documents Salem specific analysis performed to allow up to an 18 step rod misalignment at or below 85% RTP conditions. 2. Control banks are sequenced with overlapping banks. 3. The control bank insertion limits are not violated. 4, Axial power distribution procedures, which are given in terms of flux difference control and control bank position, are observed . The axial power distribution procedures referred to above are part of the required operating procedures which are followed in normal operation. They require control of the axial offset {flux difference divided by fractional power) at all power levels within a permissible operating band of a target value corresponding to the full power value. In the first cycle, the target value changes from about -10 percent to 0 percent linearly through the life of the cycle. Target values in a reload cycle vary based on previous cycle length and number of fresh assemblies. These cycle-specific target values are available in the appropriate NDR (see Section 4.5). This minimizes xenon transient effects on the axial power distribution since the procedures essentially keep the xenon distribution in phase with the power distribution. Calculations are performed for normal operation of the reactor including load following maneuvers. Beginning, middle, and end-of-cycle conditions are included in the calculations. Different histories of operation are assumed prior to calculating the effect of load follow transients on the axial power 4.3-23 SGS-UFSAR Revision 23 October 17, 2007 I I distribution. These different histories assume base loaded operation and extensive load following. The calculated points have been synthesized from axial calculations combined with radial factors appropriate for rodded and unrodded planes. The calculated values have been increased by a factor of 1.05 for conservatism and a factor of 1.03 for the engineering factor, Figure 4. 3-11 represents these results as an upper bound envelope on local power density versus elevation in the core. It should be emphasized that this envelope is a conservative representation of the bounding values of local power density. Expected values are considerably smaller and, in fact, less conservative bounding values may be justified with additional analysis or surveillance requirements. Finally, as previously discussed, this upper bound envelope is based on procedures of load follow which require the operator to operate within an allowed deviation from a target equilibrium value of axial flux difference, observing certain D bank insertion limits. These procedures are detailed in Technical Specifications and are predicated only upon ex-core surveillance supplemented by the normal monthly full core map requirement and by computer based alarms on deviation and time of deviation from the allowed flux difference band. Allowing for fuel densification effects, the average kW/ft for both Units 1 and 2 is 5.52. From Figure 4.3-11, the conservative upper bound value of normalized local power density, including allowances for densification effects, is 2. 40 corresponding to a peak local power density of 13.3 kW/ft at 100.6 percent power for Units 1 and 2. 4.3-24 SGS-UFSAR Revision 20 May 6, 2003
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- To determine Reactor Protection System setpoints, with respect to power distributions, three categories of events are considered, namely rod control equipment malfunctions, operator errors of commission, and operator errors of omission. The first category comprises uncontrolled rod withdrawal (with rods moving in the normal bank sequence) for full length rod banks. Also included are motions of the full length rod banks below their insertion limits, which could be caused, for example, by uncontrolled dilution or primary coolant cooldown. Power distributions were calculated throughout these occurrences assuming short-term corrective action, that is no transient xenon effects were considered to result from the malfunction. The event was assumed to occur from typical normal operating situations which did include normal xenon transients. It was further assumed in determining the power distributions that total power level would be limited by reactor trip to below 118 percent. Since the study is to determine protection limits with respect to power and axial offset, no credit was taken for trip setpoint reduction due to flux difference. Results are given on Figure 4.3-12 in units of kW/ft. The peak power density which can occur in such events, assuming reactor trip at or below 118 percent, is less than that required for centerline melt including uncertainties and densification effects. The second category, also appearing in Figure 4.3-12, assumes that the operator mis-positions the full length rod bank in violation of the insertion limits and creates short-term conditions not included in normal operating conditions. The third category assumes that the operator fails to take action to correct a flux difference violation. Representative results shown on Figure 4.3-13 are F Q multiplied by 100.6 percent power including an allowance for calorimetric error. The figure shows that provided the assumed error in operation does not continue for a period which is long compared to the xenon time constant, the maximum 4.3-25 SGS-UFSAR Revision 19 November 19, 2001 local power does not exceed 22.4 kW/ft including the above factors. However, the COLR restrict AI at 100. 6 percent power such that the peak linear power density is less than 18 kW/ft. These events are considered Condition II events. It should be noted that a reactor overpower accident is not assumed to occur coincident with an independent operator error. Analyses of possible operating power shapes show that the appropriate hot channel factors FQ and for peak local power density and for DNB analysis at full power are the values given in the COLR. The current unit and cycle's COLR reference is given in Section 4.5. Typical values are given in Table 4.3-2. FQ can be increased with decreasing power as shown in the Technical Specifications. Increasing with decreasing power is permitted by the DNB protection setpoints and allows radial power shape changes with rod insertion to the insertion limits as described in Section 4.4.3.2. It has been detsrmined that provided the above Conditions I through IV are observed, the Technical Specification limits are met. When a situation is possible in normal operation which could result in local power densities in excess of those assumed as the pre-condition for a subsequent hypothetical accident, but which would not itself cause fuel failure, administrative controls and alarms are provided for returning the core to a safe condition. These alarms are described in detail in Sections 7 and 16. 4.3.2.2.7 Experimental Verification of Power Distribution Analysis This subject is discussed in depth in Reference 2. A summary of this report is given here. In a measurement of peak local power density, F Q' with the moveable detector system described in Section 7.6, the following uncertainties have to be considered: 4.3-26 SGS-UFSAR Revision 19 November 19, 2001 * * *
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- 1. Reproducibility of the measured signal 2. Errors in the calculated relationship between detector current and local flux 3. Errors in the calculated relationship between detector flux and peak rod power some distance from the measurement thimble. The appropriate allowance for 1 above has been quantified by repetitive measurements made with several inter-calibrated detectors by using the common thimble features of the In-core Detector System. This system allows more than one detector to access any thimble. Errors in category 2 above are quantified to the extent possible, by using the fluxes measured at one thimble location to predict fluxes at another location which is also measured. Local power distribution predictions are verified in critical experiments on arrays of rods with simulated guide thimbles, control rods, burnable poisons, etc. These critical experiments provide quantification of error of types 2 and 3 above. Reference 2 describes critical experiments performed at the Westinghouse Reactor Evaluation Center and measurements taken on two Westinghouse plants with In-core Detector Systems of the same type as used in the plant described herein. The report concludes that the uncertainty associated with the peak nuclear heat flux factor, F Q' is 4. 58 percent at the 95 percent confidence level with only 5 percent of the measurements greater than the inferred value. This is the equivalent of a 2cr limit on a normal distribution and is the uncertainty to be associated with a full core flux map with moveable detectors reduced with a reasonable set of input data incorporating the influence of burnup on the radial power distribution. The uncertainty is usually rounded up to 5 percent. For use of the PDMS (BEACON) system, the measurement uncertainty {00) is defined in Equation 5-19 of Reference 34. On a cycle specific basis, peaking factor measurement uncertainties can be found in the COLR. The BEACON measured power distribution must be calibrated to flux map measurements during the initial ascension in power above 25 % power and at 180 EFPD intervals from the initial calibration flux map . 4.3-27 SGS-UFSAR Revision 19 November 19, 2001 I In comparing measured power distributions (or detector currents) against the calculations for the same situation, it is not possible to subtract out the detector reproducibility. Thus a comparison between measured and predicted power distributions has to include some measurement error. Such a comparison is given on Figure 4.3-14 for one of the maps used in Reference 2. Since the first publication of the report, hundreds of maps have been taken on these and other reactors. The results confirm the adequacy of the 5 percent uncertainty allowance on the calculated F0. A similar analysis for the uncertainty in FAH (rod integral power) measurements using the in-core movable detector system results in an allowance of 3. 60 percent at the equivalent of a 20' confidence level. For historical reasons, an 8 percent uncertainty factor is allowed in the nuclear design basis; that is, the predicted rod integrals at full power must not exceed the Technical Specification I COLR FN AH limit less 8 percent. This 8 percent may be reduced in final design to 4 percent to allow a wider range of acceptable axial power distributions in the DNB analysis and still meet the design bases of Section 4.3.1.3. For use of the PDMS (BEACON) system, the measurement uncertainty (UAHl is defined in Equation 5-19 of Reference 34. On a cycle specific basis, peaking factor measurement uncertainties can be found in the COLR. The BEACON measured power distribution must be calibrated to flux map measurements during the initial ascension in power above 25 % power and at 180 EFPD intervals from the initial calibration flux map. A measurement in the second cycle of a 121 assembly, 12 foot core is compared with a simplified one dimensional core average axial calculation on Figure 4. 3-15. This calculation does not give explicit representation to the fuel grids. The accumulated data on power distributions in actual operation is basically of three types: 1. Much of the data is obtained in steady state operation at constant power in the normal operating configuration. 2. Data with unusual values of axial offset have been obtained in the past as part of the multi-point ex-core detector calibrqtion exercise which is performed monthly. 3. Special tests have been performed in load follow and other transient xenon conditions which have yielded useful information on power distributions. 4.3-28 SGS-UFSAR Revision 19 November 19, 2001 * * *
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- These data are presented in detail in Reference 6. Figure 4.3-16 contains a summary of measured values of FQ as a function of axial offset for five plants from that report. 4.3.2.2.8 Testing A very extensive series of physics tests is performed on first cores. These tests are described in Section 14. Since not all limiting situations can be created at BOL, the main purpose of the tests is to provide a check on the calculational methods used in the predictions for the conditions of the test. Low power physics tests are performed at the beginning of each reload cycle to confirm the validity of the cycle-specific design models that are used in the calculations supporting the Reload Safety Evaluation. The physics tests measure the critical boron concentration, isothermal temperature coefficient and control rod worth and compare the results to predictions calculated by the design models. Additionally, moveable in-core detectors are used to measure nuclear and safety related parameters during the initial power ascension and periodically throughout the cycle operation. The measurements are used to ensure that the measured parameters are within the limiting values contained in the Technical Specifications, as well as to compare the corre.sponding parameters to the design models for validation . 4.3.2.2.9 Monitoring Instrumentation The adequacy of instrument numbers, spatial deployment, required correlations between readings and peaking factors, calibration and errors are described in References 2, 5, and 6. The relevant conclusions are summarized here in Section 4.3.2.2.7. Provided the limitations given in Section 4.3.2.2.6 on rod insertion and flux difference are observed, the Ex-core Detector System provides adequate monitoring of power distributions. The addition of BEACON {PDMS) provides nearly continuous on-line monitoring of the core power distribution through thermocouple measurements, rod position indications, loop temperature measurements, and excore detector readings which are fed from the plant computer. These readings are calibrated to in-core flux map data during the initial power ascension above 25 % power conditions and at 180 EFPD intervals thereafter. Further details of specific limits on the observed rod positions and flux difference are given in the COLR (See Section 4.5). Limits for alarms, reactor trip, etc. are given in the Technical Specifications. Descriptions of the systems provided are given in Section 7.7 . 4.3-29 SGS-UFSAR Revision 19 November 19, 2001 4.3.2.3 Reactivity Coefficients The kinetic characteristics of the reactor core determine the response of the core to changing plant conditions or to operator adjustments made during normal operation, as well as the core response during abnormal or accidental transients. These kinetic characteristics are quantified in reactivity coefficients. The reactivity coefficients reflect the changes in the neutron multiplication due to varying plant conditions such as power, moderator, or fuel temperatures, or less significantly due to a change in pressure or void conditions. Since reactivity coefficients change during the life of the core, ranges of coefficients are employed in transient analysis to determine the response of the plant throughout life. The results of such simulations and the reactivity coefficients used are presented in Section 15. The analytical methods and calculational models used in calculating the reactivity coefficients are given in Section 4. 3. 3. These models have been confirmed through extensive testing of more than thirty cores similar to the plant described herein; results of these tests are discussed in Section 4.3.3. Quantitative information for calculated reactivity coefficients, including fuel Doppler coefficient, moderator coefficients (density, temperature, pressure, void) and power coefficient is given in the following sections. 4.3.2.3.1 Fuel Temperature (Doppler) Coefficient The fuel temperature (Doppler) coefficient is defined as the change in reactivity per degree change in effective fuel temperature and is primarily a measure of the Doppler broadening of U-238 and Pu-240 resonance absorption peaks. Doppler broadening of other isotopes such as 0-236, Np-237 etc. are also considered but their contributions to the Doppler effect is small. An increase in fuel temperature increases the effective resonance absorption cross sections of the fuel and produces a corresponding reduction in reactivity. 4.3-30 SGS-OSFAR Revision 6 February 15, 1987 * *
- The fuel temperature coefficient is calculated by performing two-group three-dimensional calculations using ANC (Reference 31). The fuel temperature coefficient is calculated by subtracting the MTC from the ITC from HZP to HFP. A typical Doppler temperature coefficient is shown on Figure 4. 3-17 as a function of the effective fuel temperature (at SOL and EOL conditions). The effective fuel temperature is lower than the volume averaged fuel temperature since the neutron flux distribution is non-uniform through the pellet and gives preferential weight to the surface temperature. A typical Doppler-only contribution to the power coefficient, defined later, is shown on Figure 4.3-18 as a function of relative core power. The integral of the differential curve on Figure 4.3-18 is the Doppler contribution to the power defect and is shown on Figure 4.3-19 as a function of relative power. The Doppler-only power coefficient and defects were calculated using three-dimensional ANC (Reference 31). The Doppler coefficient becomes more negative as a function of life as the Pu-240 content increases, thus increasing the Pu-240 resonance absorption but less negative as the fuel temperature changes with burnup as described in Section 4.3.3.1. The upper and lower limits of Doppler coefficient used in accident analyses are given in Section 15. The Doppler-only coefficient and defect change slightly on a cycle-specific basis. The appropriate NDR should be referenced for the current cycle's information (see Section 4.5). 4.3.2.3.2 Moderator Coefficients The moderator coefficient is a measure of the change in reactivity due to a change in specific coolant parameters such as density, temperature, pressure, or void. The coefficients so obtained are moderator density, temperature, pressure, and void coefficients. Moderator Density and Temperature Coefficients The moderator temperature (density) coefficient is defined as the change in reactivity per unit change in the moderator temperature (density}. Generally, the effect of the changes in moderator 4.3-31 SGS-UFSAR Revision 17 October 16, 1998 density as well as the temperature are considered together. An increase in moderator density results in more moderation and hence an increase in reactivity. Therefore, the moderator density coefficient is positive. As temperature increases, density decreases (for a constant pressure); hence the moderator temperature coefficient becomes more negative. An increase in coolant temperature, keeping the density constant, leads to a hardened neutron spectrum and results in an increase in resonance absorption in U-238, Pu-240 and other isotopes. The hardened spectrum also causes a decrease in the fission-to-capture ratio in U-235 and Pu-239. Both of these effects make the moderator temperature coefficient more negative. Since water density changes more rapidly with temperature as temperature increases, the moderator temperature (density) coefficient becomes more negative (positive) with increasing temperature. The soluble boron used in the reactor as a means of reactivity control also has an effect on moderator density coefficient since the soluble boron poison density as well as the water density is decreased when the coolant temperature rises. A decrease in the soluble poison concentration introduces a positive component in the moderator temperature coefficient. Thus, if the concentration of soluble poison is large enough, the net value of the coefficient may be positive. With the burnable absorbers present, however, the initial hot boron concentration is sufficiently low that the moderator temperature coefficient is negative at operating temperatures. The effect of control rods is to make the moderator coefficient more negative by reducing the required soluble box:,on concentration and by increasing the "leakage" of the core. With burnup, the moderator temperature coefficient becomes more negative primarily as a result of boric acid dilution but also to an extent from the effects of the buildup of plutonium and fission products. 4.3-32 SGS-UFSAR Revision 17 October 16, 1998 The moderator coefficient is calculated for the various plant conditions discussed above by performing two-group three-dimensional calculations, varying the moderator temperature (and density) by about +/- 5°F about each of the mean temperatures. The moderator temperature coefficient is shown as a function of core temperature and boron concentration for the unrodded and rodded core on Figures 4. 3-20 through 4. 3-22. The temperature range covered is from cold ( 68°F) to about 600°F. The contribution due to Doppler coefficient (because of change in moderator temperature) has been subtracted from these results. Figure 4. 3-23 shows the hot, full power moderator temperature coefficient plotted as a function of first cycle lifetime for the just critical boron concentration condition based on the design boron letdown condition. The moderator coefficients presented here are calculated on a corewise basis, since they are used to describe the core behavior in normal and accident situations when the moderator temperature changes can be considered to affect the whole core. Moderator temperature coefficients change on a cycle-specific basis. The appropriate NOR should be referenced for the current cycle' s*information (see section 4.5). Moderator Pressure Coefficient The moderator pressure coefficient relates the change in moderator density, resulting from a reactor coolant pressure change, to the corresponding effect an neutron production. This coefficient is of much less significance in comparison with the moderator temperature coefficient. A change of 50 psi in pressure has approximately the same effect on reactivity as a half-degree change in moderator temperature. This coefficient can be determined from the moderator temperature coefficient by relating change in pressure to the corresponding change in density. The moderator pressure coefficient is negative aver a portion of the moderator temperature range at beginning-of-life (-0.004 pcm/psi, BOL) but is always positive at operating conditions and becomes more positive during life (+0.3 pcm/psi, EOL). 4.3-33 SGS-UFSAR Revision 17 October 16, 1998 Moderator Void Coefficient The moderator void coefficient relates the change in neutron multiplication to the presence of voids in the moderator. In a PWR this coefficient is not very significant because of the low void content in the coolant. The core void content is less than one-half of one percent and is due to local or statistical boiling. The void coefficient at BOL varies from 50 pam/percent void at low temperatures to -250 pcmjpercent void at EOL and at operating temperatures. The negative void coefficient at operating temperature becomes more negative with fuel burnup. 4.3.2.3.3 Power Coefficient The combined effect of moderator temperature and fuel temperature change as the core power level changes is called the total power coefficient and is expressed in terms of reactivity change per percent power change. A typical plot of the power coefficient at BOL and EOL conditions is given on Figure 4.3-24. It becomes more negative with burnup reflecting the combined effect of moderator and fuel temperature coefficients with burnup. A typical plot of the power defect (integral reactivity effect) at BOLand EOL is given on Figure 4.3-25. The power coefficient and defect change on a cycle-specific basis. The appropriate NOR should be referenced for the current cycle (see section 4.5). 4.3.2.3.4 Comparison of Calculated and Experimental Reactivity Coefficients section 4.3.3 describes the comparison of calculated and experimental reactivity coefficients in detail. Based on the data presented there, the accuracy of the current analytical model is: +/- 0.2 percent 4p for Doppler and power defect i 2 pcm/°F fer the moderator coefficient Experimental evaluation of the calculated coefficients are done during the physics startup tests described in Section 14. 4.3-34 SGS-UFSAR Revision 17 October 16, 1998 -
4.3.2.3.5 Reactivity Coefficients used in Transient Analysis Table 4. 3-2 gives the representative ranges for the reactivity coefficients used in transient analysis. The exact values of the coefficient used in the analysis depend on whether the transient of interest is examined at the BOL or EOL, whether most negative or the most positive (least negative) coefficients are appropriate, and whether spatial nonuniformity must be considered in the analysis. Conservative values of coefficients, considering various aspects of analysis, are used in the transient analysis. This is described in section 15. The values listed in Table 4.3-2 and illustrated on Figures 4.3-17 through 4.3-25 apply to a typical reload cycle. The coefficients appropriate for use in subsequent cycles depend on the core
- s operating history, the number and enrichment of fresh fuel assemblies, the loading pattern of burned and fresh fuel, the number and location of any burnable absorber rods, etc. The need for a reevaluation of any accident in a subsequent cycle is contingent upon whether or not the coefficients for that cycle fall within the identified range used in the analysis presented in Section 15. Control rod requirements for typical Unit 1 and Unit 2 reload cores are given in Table 4.3-3. 4.3.2.4 Control Requirements To ensure the shutdown margin stated in the Technical Specifications under conditions where a cooldown to ambient temperature is required, concentrated soluble boron is added to the coolant. Typical boron concentrations for several core conditions are listed in Table 4.3-2. For all core conditions including refueling, the boron concentration is well below the solubility limit. The RCCAs are employed to bring the reactor to the hot 4.3-35 SGS-UFSAR Revision 17 october 16, 1998 shutdown condition. 'l'he minimum required shutdown margin is given in the Technical Specifications. The ability to accomplish the shutdown for hot conditions is demonstrated in Table 4.3-3 by comparing the difference between the RCCA's reactivity available with an allowance for the worst stuck rod with that required for control and protection purposes. The shutdown margin includes an allowance of 10 percent for analytic uncertainties (see Section 4.3.2.4.9). The largest reactivity control requirement appears at the EOL when the moderator temperature coefficient reaches its peak negative value as reflected in the larger power defect. The control rods are required to provide sufficient reactivity to account for the power defect from full power to zero power and to provide the required shutdown margin. The reactivity addition resulting from power reduction consists of contributions from Doppler, moderator temperature, flux: redistribution, and reduction in void content as discussed below. 4.3.2.4.1 Doppler The Doppler effect arises from the broadening of U-238 and Pu-240 resonance peaks with an increase in effective pellet temperature. This effect is most noticeable over the range of zero power to full power due to the large pellet temperature increase with power generation. 4.3.2.4.2 Variable Average Moderator Temperature When the core is shut down to the hot, zero power condition, the average moderator temperature changes from the equilibrium full load value determined by the steam generator and turbine characteristics (steam pressure, heat transfer, tube fouling, etc) to the equilibrium no load value, which is based on the steam generator shell side design pressure. The design change in 4.3-36 SGS-UFSAR Revision 17 October 16, 1998 --
temperature is conservatively increased by 4°F to account for the control dead band and measurement errors. Since the moderator coefficient is negative, there is a reactivity addition with power reduction. The moderator coefficient becomes more negative as the fuel depletes because the boron concentration is reduced. This effect is the major contributor to the increased requirement at EOL. 4.3.2.4.3 Redistribution During full power operation the coolant density decreases with core height, and this, together with partial insertion of control rods, results in less fuel depletion near the top of the core. Under steady state conditions, the relative power distribution will be slightly asymmetric towards the bottom of the core. On the other hand, at hot zero power conditions, the coolant density is uniform up the core, and there is no flattening due to Doppler. The result will be a flux distribution which at zero power can be skewed toward the top of the core. The reactivity insertion due to the skewed distribution is calculated with an allowance for the most adverse effects of xenon distribution. 4.3.2.4.4 Void Content A small void content in the core is due to nucleate boiling at full power. The void collapse coincident with power reduction makes a small reactivity contribution. 4.3.2.4.5 Rod Insertion Allowance At full power, the control bank is operated within a prescribed band of travel to compensate for small periodic changes in boron concentration, changes in temperature, and very small changes in the xenon concentration not compensated for by a change in boron 4.3-37 SGS-UFSAR Revision 6 February 15, 1987 concentration. When the control bank reaches either limit of this band, a change in boron concentration is required to compensate for additional reactivity changes. Since the insertion limit is set by a rod travel limit, a conservatively high calculation of the inserted worth is made which exceeds the normally inserted reactivity. 4.3.2.4.6 Burnup Excess reactivity is installed at the beginning of each cycle to provide sufficient reactivity to compensate for fuel depletion and fission products throughout the cycle. This reactivity is controlled by the addition of soluble boron to the coolant and by burnable absorbers. The soluble boron concentration for several core configurations, and unit boron worths are given in Table 4.3-2. Since the excess reactivity for burnup is controlled by soluble boron and/or burnable absorbers, it is not included in control rod requirements. 4.3.2.4.7 Xenon and Samarium Poisoning Changes in xenon and samarium concentrations in the core occur at a sufficiently slow rate, even following rapid power level changes, that the resulting reactivity change is controlled by changing the soluble boron concentration. 4.3.2.4.8 pH Effects Changes in reactivity due to a change in coolant pH, if any, are sufficiently small in magnitude and occur slowly enough to be controlled by the Boron System. Further details are available in Reference 8. 4.3-38 SGS-UFSAR Revision 17 October 16, 1998 4.3.2.4.9 Experimental Confirmation a normal shutdown, the total core with a stuck rod has been measured on a 121 cool down 10-foot core, and a 121 12-foot core. In each case, the core was allowed to cool down until it reached criticality simulating the steamline break accident. For the 10-foot core, the total reactivity change associated with the cooldown was over-predicted by about 0. 3 percen: L'1p with respect to the measured result. This represents an error of about 5 percent in the total change and For the 12-foot is about half the allowance :':or this core, the difference between the measured and was an even smaller 0.2 percent L'1p. These measurements and others demonstrate the ability of the methods described in Section 4. 3. 3 to accurately predict the total shutdown reactivity of the core. 4.3.2.5 Core is controlled by means of a chemical dissolved in the coolant, RCCAs, and burnable absorber rods as described below. 4.3.2.5.1 Chemical Poison Boron in solution as boric acid is used to control relatively slow reactivity changes associated with: 2. SGS-Uf'SAR The moderator temperature defect in going from cold shutdown at ambient temperature to the hot The transient xenon and samarium power or in RCC 4.3-39 temperature at zero power such as that following Revision 25 October 26, 201C
- 3. The excess reactivity required to compensate for the effects of fissile inventory depletion and buildup of long-life fission products 4. The burnable absorber depletion Typical boron concentrations for various core conditions are presented in Table 4.3-2. 4.3.2.5.2 Rod Cluster Control Assemblies Fifty-three RCCAs are employed. These are used for shutdown and for control purposes to offset fast reactivity changes associated with: 1. The required shutdown margin in the hot zero power, stuck rod condition 2. The reactivity compensation as a result of an increase in power above hot zero power (power defect including Doppler, and moderator reactivity changes) 3. Unprogrammed fluctuations in boron concentration, coolant temperature, or xenon concentration (with rods not exceeding the allowable rod insertion limits) 4. Reactivity ramp rates resulting from load changes The allowed control bank reactivity insertion is limited at full power to maintain shutdown capability. As the power level is reduced, control rod reactivity requirements are also reduced and more rod insertion is allowed. The control bank position is monitored and the operator is notified by an alarm if the limit is approached. The determination of the insertion limit uses conservative xenon distributions and axial power shapes. In addition, the RCCA withdrawal pattern determined from these analyses is used in determining power distribution factors and in 4.3-40 ' SGS-UFSAR Revision 17 October 16, 1998 -
determining the maximum further discussion refer limits. worth of -:::o the an inserted RCCl\ ejection Technical Specifications on accident. For rod insertion Power rod ection, and rod are based on the arrangement of the shutdown and control groups of the RCCAs shown on 4. 3-26A and B. All shutdown RCCAs are withdrat-vn before li'lithdrawal of the control banks is ::_ni tiated. In going from zero to 100 percent power, control banks A, B, C, and D are v.Jithdrawn sequentially. The limits of rod posi ticns are provided in the COLR (see Section 4. 5)
- Further discussion on the basis of rod insertion limits are provided in the Technical Specifications. 4.3.2.5.3 Burnable Absorbers The burnable absorbers provide control of the excess reactivity available during the fuel cycle. In doing so, the temperature coefficient is prevented from being positive at normal operating conditions. They perform this function by reducing the requirement for soluble poison in the moderator at the beginning of the fuel cycle as described previously. The number of burnable absorber rods per is shown in Section 4. 5 for a The boron in the burnable absorbers is with but at a slow rate so that the critical cor:centration of soluble boron is such that the moderator temperature coefficient remains non-at all times for power operating conditions. 4.3.2.5.4 Peak Xenon Startup Compensation for the peak xenon buildup is accomplished using the Boron Control System. Startup frorr the peak xenon condition is accomplished with a combination of rod motion and boron dilution. 4.3-41 SGS-UFSAR Revision 25 October 26, 2010 The boron dilution may be made at any time, even during the shutdown period, provided the shutdown margin is maintained. 4.3.2.5.5 Load Follow Control and Xenon Control During load follow maneuvers, power changes are accomplished using control rod motion and dilution or boration by the Boron System as required. Control rod motion is limited by the control rod insertion limits as provided in the COLR (see Section 4.$) and discussed previously in this section. Reactivity changes due to the changing xenon concentration can be controlled by rod motion and/or changes in soluble boron concentration. 4.3.2.5.6 Burnup Control of the excess reactivity for burnup is accomplished using soluble boron and/or burnup absorbers. The boron concentration must be limited during operating conditions to ensure the moderator temperature coefficient is non-positive. Sufficient burnable absorber is installed at the beginning of a cycle to give the desired cycle lifetime without exceeding the boron concentration value which would yield a positive MTC per Technical Specifications. The practical minimum boron concentration is 10 ppm. 4.3.2.6 Control Rod Patterns and Reactivity Worth The RCCAs are designated by function as the control groups and the shutdown groups. The terms "group" and "bank" are used synonymously throughout this report to describe a particular grouping of control assemblies. The RCCA pattern is displayed on Figures 4.3-26A and B which is not expected to change during the 1 ife of the plant. The control banks are labeled A, B, C, and D and the shutdown banks are labeled SA, SB, SC, and SD. Each bank, although operated and controlled as a unit, is comprised of two subgroups. The axial position of the RCCAs may be controlled manually or automatically. The RCCAs are all dropped into the core following actuation of reactor trip signals. 4.3-42 SGS-UFSAR Revision 17 October 16, 1998 -
Two criteria have been employed for selection of the control groups. First, the total reactivity worth must be adequate to meet the requirements specified in Table 4.3-3. Second, in view of the fact that these rods may be partially inserted at power operation, the total power peaking factor should be low enough to ensure that the power capability requirements are met. Analyses indicate that the first requirement can be met either by a single group or by two or more banks whose total worth equals at least the required amount. The axial power shape would be more peaked following movement of a single group of rods worth 3 to 4 percent Ap; therefore, four banks (described as A, B, C, and D on Figures 4. 3-26A and B) each worth approximately 1 percent Ap have been selected. The position of control banks for criticality under any reactor condition is determined by the concentration of boron in the coolant. On an approach to criticality, boron is adjusted to ensure that criticality will be achieved with control rods above the insertion limit set by shutdown and other considerations (See the Technical Specifications). Early in the cycle there may also be a withdrawal limit at low power to maintain a negative moderator temperature coefficient. Usual practice is to adjust boron to ensure that the rod position lies within the so-called maneuvering band, that is such that an escalation from zero power to full power does not require further adjustment of boron concentration. Ejected rod worths are given in Section 15.3.1.6 for several different conditions. Experimental confirmation of these worths can be found by reference to startup test reports (9). Allowable deviations due to misaligned control rods are discussed in the Technical Specifications. 4.3-43 SGS-UFSAR Revision 17 October 16, 1998 A representative calculation for two banks of control rods withdrawn simultaneously (rod withdrawal accident) is given on Figure 4.3-27. Calculation of control rod reactivity worth versus time following reactor trip involves both control rod velocity and differential reactivity worth. The rod position versus time of travel after rod release assumed is given on Figure 4.3-28 for Vantage+ fuel. The drop time to the dashpot increases from 2.2 to 2.7 seconds for Vantage 5H, Vantage+, and RE'A, with the other times increasing proportionately. For nuclear design purposes, the reactivity worth versus rod position is calculated. by a series of steady state calculations at yarious control rod positions assuming all rods out of the core as the initial position in order to minimize the initial reactivity insertion rate. Also, to be conservative, the rod of highest worth is assumed stuck out of the core and the flux distribution (and thus reactivity importance) is assumed to be skewed to the bottom of the core. The result of these calculations is shown on Figure 4.3-29. The shutdown groups provide additional negative reactivity to assure an adequate shutdown margin. Shutdown margin is defined as the amount by which the core would be subcritical at hot shutdown if all RCCAs are tripped, but assuming that the highest worth assembly remains fully withdrawn and no changes in xenon or bo:ron take place. The loss of control rod wort;h due to the material irradiation is negligible since only D bank rods may be in the core under normal operating conditions. The values given in Table 4.3-3 show that the available reactivity in withdrawn RCCAs provides the design basis minimum shutdown margin allowing for the highest worth cluster to be at its fully withdrawn position. An allowance for uncertainty in the calculated worth of N-1 rods is made before determination of the shutdown margin. 4.3-44 SGS-UFSAR Revision 18 April 26, 2000 * *
- 4.3.2.7 Criticality of Fuel Assemblies Criticality of fuel assemblies outside of the reactor is precluded by adequate design of fuel transfer and fuel storage facilities and by administrative control procedures. This section identifies those criteria important to criticality analyses. New fuel is generally stored in fuel facilities with no water present but which are designed so as to prevent accidental criticality even if unborated water is present. In the analysis for the storage facilities, the fuel assemblies are assumed to be in their most reactive condition, namely fresh or undepleted and with no control* rods or removable neutron absorbers present. Assemblies cannot be closer than the by the except in cases such as in fuel shipping containers where are carried out to establish the of the design. The mechanical integrity of the fuel assembly is assumed and no credit is taken for neutron absorption properties of the storage facility unless specifically included in the design. For full flooding with unborated water, the fuel assembly spacing of the facility provides essentially full nuclear isolation, and for the array is no greater than keff for the single most reactive fuel The criterion for full flooding is 0.95. For the analysis of new (dry) fuel storage racks, an additional criterion of keff 0.98 is confirmed for optimum (low density) moderation conditions. These fresh fuel rack (J]l include allowances for uncertainties, biases, and manufacturing tolerances and assure 95 percent probability I 95 percent confidence level that the keff critically design criteria are met. The fuel assembly (17 x 17 fuel rods) of standard design and 3.5 w/o enriched uranium without a control rod or burnable absorbers, fully flooded and reflected with cold clean water, has a keff of about 0. 85. Two such fuel assemblies spaced 1 inch apart with parallel axes 9.5 inches apart have a of about 0.99. Three such fuel assemblies 1 inch with axes would be supercritical. An infinite number of dry fuel assemblies of this design would have a keff < 0.80. SGS-UFSAR 4.3-45 Revision 26 May 21, 2012 4.3.2.8 Stability 4.3.2.8.1 Introduction The stability of the PWR cores against spatial oscillations and the control of such transients are discussed extensively (References 5, 10, 11, 12) . A surrunary of these reports is in the discussion and the bases are given in Section 4.3.1.6. In a large reactor core, xenon-induced oscillations can take place with no corresponding change in the total power of the core. The oscillation may be caused by a power shift in the core which occurs rapidly by comparison with the xenon-iodine time constants. Such a power shift occurs in the axial direction when a plant load change is made by control rod motion and results in a change in the moderator density and fuel distributions. Such a power shift could occur in the diametral control action. of the core as a result of abnormal Due to the negative power coefficient of reactivity, PWR cores are inherently stable to oscillations in total power. Protection against total power stabilities is provided by the Control and Protection System as described in Chapter 7. Hence, the discussion on the core stability will be limited here to xenon-induced oscillations. 4.3.2.8.2 Stability Index Power distributions, either in the axial direction or in the X-Y plane, can undergo oscillations due to perturbations introduced in the equilibrium distributions without changing the total core power. The overtones in the current PWRs, and the stability of the core against xenon-induced oscillations can be determined in terms of the eigenvalues of the first flux overtones. either in the axial direction or in the X-Y of the first flux harmonic as 4.3-46 SGS-UFSAR the I Ct Revision 17 October 16, 1998 e
- b + ic, (Reference 10} then b is defined as the stability index and T as the oscillation period of the first harmonic. The time-dependence of the first harmonic o4l in the power distribution can now be represented as bt ae cos ct, (Reference 12) where A and a are constants. approximately by: The stability index can also be obtained where A , A 1 are the successive peak amplitudes of the oscillation and T is the n n+ time period between the successive peaks. 4.3.2.8.3 Prediction of the core Stability The stability of the core described herein (i.e., with 17 x 17 fuel assemblies) against xenon-induced spatial oscillations is expected to be equal to or better than that of earlier designs. The prediction is based on a comparison of the parameters which are significant in determining the stability of the core against the xenon-induced oscillations, namely ( 1) the overall core size is unchanged and spatial power distributions will be similar, coefficient is expected to be similar, and ( 2) the moderator temperature (3) the Doppler coefficient of reactivity is expected to be similar at full power. Analysis of both the axial and X-Y xenon transient tests, discussed in Section 4.3.2.8.5, shows that the calculational model is adequate for the prediction of core stability. 4.3-47 SGS-UFSAR Revision 17 October 16, 1998 4.3.2.8.4 Stability Measurements Axial Measurements TWo axial xenon transient tests conducted in a PWR with a core height of 12 feet and 121 fuel assemblies are reported in Reference 13, and will be briefly discussed here. The tests were performed at approximately 10 percent and 50 pe1.*cent of cycle 1" Both a free-running oscillation test and a controlled test were performed during the first test. The second test at mid-cycle consisted of a free-running oscillation test only. In each of the free-running oscillation tests, a perturbation was introduced to the equilibrium power distribution through an impulse motion of the control bank 0, and the subsequent oscillation was monitored to measure the stability index and the oscillation period. In the controlled test conducted early in the cycle, the part-length rods were used to follow the oscillations to maintain an axial offset (AO) within the prescribed limits. The AO of power was obtained from the ex-core ion chamber readings (which had been calibrated against the in-core flux maps) as a function of time for both free-running tests as shown on Figure 4.3-30. The total core power was maintained constant during these spatial xenon tests, and the stability index and the oscillation period were obtained from a least-square fit of the AO data in the form of Equation 2. The AO of power is the quantity that properly represents the axial stability in the sense that it essentially eliminates any contribution from even order harmonics including the fundamental mode. The conclusions of the tests are the following: 1. The core was stable against induced axial xenon transients both at the core burnups of 1550 MWO/MTU and 7700 MWO/MTU. The measured stability indices are -0.041 hr-l for the first test (Curve 1 of 4.3-48 SGS-UFSAR Revision 17 October 16, 1998 -
Figure 4.3-30) and -0.014 hr-1 for the second test (Curve 2 of Figure 4.3-30). The corresponding oscillation periods are 32.4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br /> and 27.2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br />. respectively. 2. The reactor core becomes less stable as fuel burnup progresses and the axial stability index was essentially zero at 12,000 MWD/MTO. Measurements in the X-Y Plane Two X-Y xenon oscillation tests were performed at a PWR plant with a core height of 12 feet and 157 fuel assemblies. The first test was conducted at a core average burnup of 1540 MWD /MTU and the second at a core average burnup of 12900 MWD/MTU. Both of the X-Y xenon tests show that the core was stable in the X-Y plane at both burnups. The second test shows that the core became more stable as the fuel burnup increased and all Westinghouse PWRs with 121 and 157 assemblies are expected to be stable throughout their burnup cycles. In each of the two X-Y tests, a perturbation was introduced to the equilibrium power distribution through an impulse motion of one RCCA located along the diagonal axis. Following the perturbation, the uncontrolled oscillation was monitored using the moveable detector and thermocouple system and the ex-core power range detectors. The quadrant tilt difference is the quantity that properly represents the diametral oscillation in the X-Y plane of the reactor core in that the differences of the quadrant average powers over two symmetrically opposite quadrants essentially eliminates the contribution to the oscillation from the azimuthal mode. The quadrant tilt difference data were fitted in the form of Equation 2 through a least-square method. A stability index of -0.076 hr -l with a period of 29.6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br /> was obtained from the thermocouple data shown on Figure 4.3-31. 4.3-49 SGS-UFSAR Revision 17 October 16, 1998 It was observed in the second X-Y xenon test that the PWR core with 157 fuel assemblies had become more stable due to an increased fuel depletion, and the stability index was not determined. 4.3.2.8.5 comparison of Calculations with Measurements The analysis of the axial xenon transient tests was performed in an axial slab geometry using a flux synthesis technique. The direct simulation of the AO data was carried out using the PANDA Code (Reference 14). The analysis of the x-Y xenon transient tests was performed in an X-Y geometry using a modified TURTLE code (Reference 7). Both the PANDA and TURTLE codes solve the two-group time-dependent neutron diffusion equation with time-dependent xenon and iodine concentrations. The fuel temperature and moderator density feedback is limited to a steady state model. All the X-Y calculations were performed in an average enthalpy plane. The basic nuclear cross sections used in this study were generated from a unit cell depletion program which was evolved from the codes LEOPARD (Reference 15) and CINDER (Reference 16). The detailed experimental data during the tests including the reactor power level, enthalpy rise, and the impulse motion of the control rod assembly, as well as the plant follow burnup data, were closely simulated in the study. The results of the stability calculation for the axial tests are compared with the experimental data in Table 4.3-4. The calculations show conservative results -1 for both of the axial tests with a margin of approximately 0.01 hr in the stability index. An analytical simulation of the first X-Y xenon oscillation test shows a -1 calculated stability index of -0.081 hr , in good agreement with the measured -1 value of -0.076 hr
- As indicated earlier, the second X-Y xenon test showed that the core had become more stable compared to the first test and no evaluation of the stability index was attempted. This increase in the core 4.3-50 SGS-UFSAR Revision 17 October 16, 1998 stability in the X-Y plane due to increased fuel burnup is due mainly to the increased magnitude of the negative moderator temperature coefficient. Previous studies of the physics of xenon oscillations, including three-dimensional analysis, (References 10, 11, 12). are reported in the series of topical reports A more detailed description of the experimental results and analysis of the axial and X-Y xenon transient tests is presented in Reference 13 and Section 1 of Reference 17. 4.3.2.8.6 Stability Control and Protection The Ex-core Detector System is utilized to provide indications of xenon-induced spatial oscillations. The readings from the ex-core detectors are available to the operator and also form part of the protection system. Axial Power Distribution For maintenance of proper axial power distributions, the operator is instructed to maintain an axial offset within a prescribed operating band, based on the ex-core detector readings. Should the axial offset be permitted to move far enough outside this band, the protection limit will be reached and the power will be automatically cut back. Twelve-foot PWR cores become less stable to axial xenon oscillations as fuel burnup progresses. However, free xenon oscillations are not allowed to occur except for special tests. The full length control rod banks present in all modern Westinghouse PWRs are sufficient to dampen and control any axial xenon oscillations present. Should the axial offset be inadvertently permitted to move far enough outside the control band due to an axial xenon oscillation, or any other reason, the protection limit on axial offset will be reached and the power will be automatically cut back. 4.3-51 SGS-UFSAR Revision 17 October 16, 1998 Radial Power Distribution The core described herein is calculated to be stable against X-Y xenon induced oscillations at all times in life. The X-Y stability of large PWRs has been further verified as part of the startup physics test program for PWR cores with 193 fuel assemblies. The measured x-Y stability of the cores with 157 and 193 assemblies was in good agreement with the calculated stability discussed in Sections 4. 3. 2. 8. 4 and 4. 3. 2. 8. 5. In the unlikely event that X-Y oscillations occur, backup actions are possible and would be implemented, if necessary, to increase the natural stability of the core. This is based on the fact that several actions could be taken to make the moderator temperature coefficient more negative, which will increase the stability of the core in the X-Y plane. Provisions for protection against nonsymmetric perturbations in the x-Y power distribution that could result from equipment malfunctions are made in the protection system delllign. This includes control rod drop, rod misalignment, and asymmetric loss-of-coolant flow. A more detailed discussion of the power distribution control in the PWR cores is presented in Reference 5. 4.3.2.9 Vessel Irradiation A brief review of the methods and analyses used in the determination of neutron and gamma ray flux attenuation between the core and the pressure vessel is given below. A more complete discussion is given in the pressure vessel irradiation and surveillance program. The materials that serve to attenuate neutrons originating in the core and gamma rays, from both the core and structural component consist of the core baffle, core barrel, the neutron pads, and 4.3-52 . SGS-UFSAR Revision 6 February 15, 1987
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- associated water annuli, all of which are within the region between the core and the pressure vessel . In general, few-group neutron diffusion theory codes are used to determine fission power density distributions within the active core, and the accuracy of these analyses is verified by in-core and/or BEACON power distribution measurements on operating reactors. Region and rodwise power sharing information from the core calculations is then used as source information in two-dimensional transport calculations which compute the flux distributions through the reactor. The neutron flux distribution and spectrum in the various structural components varies significantly from the core to the pressure vessel. Representative values of the neutron flux distribution and spectrum are presented in Table 4. 3-5. The values listed are based on time averaged equilibrium cycle reactor core parameters and power distributions, and thus, are suitable for long-term nvt projections and for correlation with radiation damage estimates. The irradiation surveillance program utilizes actual test samples to verify the accuracy of the calculated fluxes at the vessel . 4.3.3 Analytical Methods Calculations required in nuclear design consist of three distinct types, which are performed in sequence: 1. Determination of effective fuel temperatures 2. Generation of macroscopic few-group parameters 3. Space-dependent, few-group diffusion calculations These calculations are carried out by computer codes which can be executed individually; however, at Westinghouse most of the codes required have been linked to form an automated design sequence 4.3-53 SGS-UFSAR Revision 19 November 19, 2001 which minimizes design time, avoids errors in transcription of data, and standardizes the design methods. 4.3.3.1 Fuel Temperature <Doppler) Calculations Temperatures vary radially within the fuel rod depending on the heat generation rate in the pellet; the conductivity of the materials in the pellet, gap, and clad; and the temperature of the coolant. The fuel temperatures for use in most nuclear design Doppler calculations are obtained from a simplified version of the Westinghouse fuel rod design model described in Section 4.2.1.3.1, which considers the effect of radial variation of pellet conductivity, expansion-coefficient and heat generation rate, elastic deflection of the clad, and a gap conductance which depends on the initial fill gas, the hot open gap dimension, and the fraction of the pellet over which the gap is closed. The fraction of the gap assumed closed represents an empirical adjustment used to produce good agreement with observed reactivity data at BOL. Further gap closure occurs with burnup and accounts for the decrease in Doppler defect with burnup which has been observed in operating plants. For detailed calculations of the Doppler coefficient, a more sophisticated temperature model is used which accounts for the effects of fuel swelling, fission gas release, and plastic clad deformation. Radial power distributions in the pellet as a function of burnup are obtained from LASER {Reference 18) calculations. The effective U-238 temperature for resonance absorption is obtained from the radial tempe.t'ature distribution by applying a radially dependent weighting function. The weighting function was determined from REPAD (Reference 19) Monte Carlo calculations of resonance escape probabilities in several steady state and transient temperature distributions. In each case a flat pellet temperature was determined which produced the same resonance escape probability as the actual distribution. The weighting function was empirically determined from these results. The effective Pu-240 temperature for resonance absorption is determined by a convolution of the radial distribution of Pu-240 number densities from LASER burnup calculations and the radial weighting function. The resulting temperature is burnup dependent, but the difference between U-238 and Pu-240 temperatures, in terms of reactivity effects, is small. 4.3-54 SGS-UFSAR Revision 17 october 16, 1998 -
The effective pellet temperature for pellet dimensional change is that value which produces the same outer pellet radius in a virgin pellet as that obtained from the temperature model. The effective clad temperature for dimensional change is its average value. The temperature calculational model has been validated by plant Doppler defect data as shown in Table 4.3-6 and Doppler coefficient data as shown on Figure 4. 3-32. Stability index measurements also provide a sensitive measure of the Doppler coefficient near full power (see section 4.3.2.8). It can be seen that Doppler defect data is typically within 0.2 percent 4p of prediction. 4.3.3.2 Macroscopic Group Constants There are two lattice codes which have been used for the generation of macroscopic group constants needed in the spatial, few-group diffusion codes. One is a version of the LEOPARD and CINDER codes, which has historically been the source of the macroscopic group constants. The other is PHOENIX-P, which is used in present reload designs (Reference 30). Macroscopic few-group constants and analogous microscopic cross sections (needed for feedback and microscopic depletion calculations) were previously generated for fuel cells by a version of the LEOPARD (Reference 15) and C!NDER (Reference 16) codes, which are linked internally and provide burnup dependent cross sections. Normally a simplified approximation of the main fuel chains is used; however, where needed, a complete solution for all the significant isotopes in the fuel chains from Th-232 to cm-244 is available (Reference 20). Fast and thermal cross section library tapes contain microscopic cross sections taken for the most part from the ENDF/B {Reference 21) library, with a few exceptions where other data provide better agreement with critical experiments, isotopic measurements, and plant critical boron values. The effect on the unit fuel cell of non-lattice components in the fuel assembly is obtained by supplying an appropriate volume fraction of these materials in an extra region which is homogenized with the unit cell in the fast (MUFT) and thermal (SOFOCATE) flux calculations. In the thermal calculation, the fuel rod, clad, and moderator are homogenized by energy-dependent disadvantage factors derived from an analytical fit to integral transport theory results. Group constants for discrete burnable absorber cells, guide thimbles, instrument thimbles, and interassembly gaps are generated in a manner analogous to the fuel cell calculation. Reflector group constants are taken from infinite medium LEOPARD calculations. Baffle group constants are calculated from an average of core and radial reflector microscopic group constants for stainless steel. 4.3-55 SGS-UFSAR Revision 17 October 16, 1998 Group constants for control rods are calculated in a linked version of the HAMMER (Reference 22) and AIM (Reference 23) codes to provide an improved treatment of self shielding in the broad resonance structure of these isotopes at epithermal energies than is available in LEOPARD. The Doppler broadened cross of the control rod materials are represented as smooth cross sections in the 54-group LEOPARD fast group structure and in 30 thermal groups. The four-group constants in the rod cell and appropriate extra region coupled space-energy transport HAMMER calculation. calculation of the homogenized rod cell with extra region are generated in the A corresponding AIM is used to adjust the absorption cross sections of the rod cell to match the reaction rates in HAMMER. These transport-equivalent group constants are reduced to two-group constants for use in space-dependent diffusion calculations. In discrete X-Y calculations only one mesh interval per cell is used, and the rod group constants are further adjusted for use in this standard mesh by reaction rate matching the standard mesh unit assembly to a fine-mesh unit assembly calculation. Validation of the cross section method is based on analysis of critical experiments as shown in Table 4. 3-7, isotopic data as shown in Table 4. 3-8, plant critical boron (C8) values at HZP, BOL, as shown in Table 4.3-9 and at HFP as a function of burnup as shown on Figures 4.3-33 through 4.3-35. Control rod worth measurements are shown in Table 4.3-10. Confirmatory critical experiments on discrete burnable absorbers are described in Reference 24. PHOENIX-P has been approved by the USNRC as a lattice code for the generation of macroscopic and microscopic few group cross sections for PWR analysis {Reference 30). PHOENIX-P is a two-dimensional, multigroup, transport-based lattice code capable of providing all necessary data for PWR analyais. Since it is a dimensional lattice code, PHOENIX-P does not rely on predetermined spatial/spectral interaction assumptions for can provide a more accurate multigroup LEOPARD/CINDER. the heterogeneous fuel lattice and flux solution than versions of The solution for the detailed spatial flux and energy distribution is divided into two major steps in PHOENIX-P (Reference 30). First, a two-dimensional fine energy group nodal solution is obtained, coupling individual subcell regions (pellet, clad, and moderator) as well as surrounding pins, using a method based on Carlvik's collision probability approach and heterogeneous response fluxes which preserve the heterogeneity of the pin cells and their surroundings. The nodal solution provides an accurate and detailed local flux distribution, which is then used to homogenize the pin cells spatially to fewer groups. 4.3-56 SGS-UFSAR Revision 17 October 16, 1998 -*
Then, a standard S4 discrete ordinates calculation solves for the angular distribution, based on the group-collapsed and homogenized cross-sections from the first step. These S4 fluxes normalize the detailed spatial and energy nodal fluxes, which are then used to compute reaction distributions and to deplete the fuel and burnable absorbers. calculation evaluates the fundamental mode critical improved fast diffusion coefficient for the core codes. PHOENIX-P energy group energy groups from ENDF/B files. the of t:'1e and to model resonance parameters neutronics data necessary for modeling fuel, fission rates, power A standard Bl an of 42 or more group contains all and structural materials, coolant, and control and burnable absorber materials present in the PWRs. constants for burnable absorber cells, control rod cells, thimbles and instrumentation thimbles, or other non-fuel cells, can be obtained directly from PHOENIX-P without any adjustments such as those required in the cell or ID lattice codes. 4.3.3.3 Spatial Few-Group Diffusion Calculations Historically, spatial few-group diffusion calculations consisted primarily of two-group X-Y calculations using an updated version of the TURTLE code and two-group axial calculations using an updated version of the PANDA code. Discrete X-Y calculations ( 1 mesh per cell) were carried out to determine critical boron concentrations and pm,yer distributions in the X-Y p::Cane. An axial average in the X-Y plane was obtained by syn-thesis from unrodded and rodded planes. Axial effects in unrodded depletion calculations were accounted for by the axial buckling, which varies with burnup and is determined by radial depletion calculations depletion calculation. which are matched in reactivity to the analogous R-Z The moderator coefficient is evaluated by varying the inlet temperature in the same X-Y calculations used for power distribution and reactivity predictions. Validation of the calculations is associated with the validation of the group constants themselves, as discussed in Section 4.3.3.2. the calculations is associated with the fuel discussed in Section 4.3.3.1. Validation of the moderator calculations is obtained by with measurements power conditions as shown in Table 4.3-11. 4.3-57 Validation of validation coefficient at hot zero SGS-UFSAR Revision 25 October 26, 2010 I Axial calculations are used to determine differential control rod worth curves (reactivity versus rod insertion) and axial power shapes during steady state and transient xenon conditions (flyspeck curve). Group constants are obtained from the three-dimensional nodal model by flux-volume weighting on an axial slice-wise basis. Radial bucklings are determined by varying parameters in the buckling model while forcing the one-dimensional model to reproduce the axial characteristics (axial offset, mid-plane power) of the three-dimensional model. Recent few-group spatial calculations have input PHOENIX-P supplied two-group cross-sections to the Advanced Nodal Code (ANC) . ANC is a two-group, two or three-dimensional nodal code capable of determining typical nuclear design analyses, such as critical boron concentrations, average assembly and pin powers, control rod worths, reactivity coefficients, assembly and pin burnups and axial power Through the use of advanced nodal techniques, ANC is able to produce solutions similar to the fine mesh, finite difference diffusion theory codes such as TURTLE/TORTISE. ANC has been benchmarked against TORTISE (an improved version of TURTLE) for normal and off-normal conditions, such as ejected rod, stuck rod and dropped rod (Reference 31). The qualification of the PHOENIX-P/ANC methodology against measured data is given in Reference 30. Validation of the spatial codes for calculating power distributions involves the use of in-core and ex-core detectors and the BEACON core monitoring system (PDMS) and is discussed in Section 4.3.2.2.7. Based on comparison with measured data it is estimated that the accuracy of current analytical methods is: +/- 0.2 percent for Doppler defect +/- 2 x 10-5/°F for moderator coefficient +/- 50 ppm for critical boron concentration with depletion +/- 3 percent for power distributions +/- 0.2 percent for rod bank worth +/- 4 pcm/step for differential rod worth +/- 0.5 pcm/ppm for boron worth +/- 0.1 percent for moderator defect 4.3.4 References for Section 4.3 1. "Westinghouse Anticipated Transients Without Reactor Trip Analysis," WCAP-8330, August 1974. 2. Langford, F. L. and Nath, R. J., Jr., "Evaluation of Nuclear Hot Channel Factor Uncertainties," WCAP-7308-L, April 1969 (Westinghouse Proprietary) and WCAP-7810 (Non-Proprietary), December 1971. 4. 3-58 . SGS-UFSAR Revision 19 November 19, 2001 * * *
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- 3. McFarlane, A. F., "Core Power Capability in Westinghouse PWRs," WCAP-7267-L, October 1969 (Proprietary) and WCAP-7809 (Non-Proprietary), December 1971. 4. Hellman, J. M. (Ed.), "Fuel Densification Experimental Results and Model for Reactor Application," WCAP-8218-P-A (Proprietary) and WCAP-8219-A (Non-Proprietary), March 1975. 5. Moore, J. s., "Power Distribution Control of Westinghouse Pressurized Water Reactors," WCAP-7208, September 1968 (Proprietary) and WCAP-7811 (Non-Proprietary), December 1971. 6. McFarlane, A. F., "Power Peaking Factors," WCAP-7912-P-A, (Proprietary) and WCAP-7912-A, (Non-Proprietary), January 1975. 7. Altomare, S. and Barry, R. F., "The TURTLE 24.0 Diffusion Depletion Code," WCAP-7213-P-A (Proprietary) and WCAP-7758 (Non-Proprietary), February 1975. 8. Cermak, J. 0., et al, "Pressurized Water Reactor pH-Reactivity.Effect," Final Report, WCAP-3696-8 (EURAEC-2074), October 1968 . 9. Outzs, J. E., "Plant Startup Test Report, H. B. Robinson Unit No.2," WCAP-7844, January 1972. 10. Poncel, C. G. and Christie, A. M., "Xenon-Induced Spatial Instabilities in Large PWRs," WCAP-3680-20 (EURAEC-1974), March 1968. 11. Skogen, F. B. and McFarlane, A. F., "Control Procedures for Xenon-Induced X-Y Instabilities in Large PWRs," WCAP-3680-21, (EURAEC-2111), February 1969. 12. Skogen, F. B. and McFarlane, A. F., "Xenon-Induced Spacial Instabilities in Three-Dimensions," WCAP-3680-22 (EURAEC-2116), September 1969. 13. Lee, J. C., et al, "Axial Xenon Transient Tests at the Rochester Gas and Electric Reactor," WCAP-7964, June 1971. 14. Altomare, S. and Minton, G., "The PANDA Code," WCAP-7048-P-A (Proprietary) and WCAP-7757-A (Non-Proprietary), February 1975
- 4.3-59 SGS-UFSAR Revision 19 November 19, 2001 I I 15. Barry, R. F., "LEOPARD -A Spectrum Dependent Non-Spatial Depletion Code for the IBM-7094," WCAP-3269-26, September 1963. 16. England, T. R., "CINDER -A One-Point Depletion and Fission Product Program," WAPO-TM-334, August 1962. 17. Kubit, C. J., "Safety Related Research and Development for Westinghouse Pressurized Water Reactors, Program Summaries, Spring-Fall 1973," WCAP-8204, October 1973. 18. Poncelot, C. G., "LASER -A Depletion Program for Lattice Calculations Based on MUFT and THEMOS," WCAP-6073, April 1966. 19. Olhoeft, J. E., "The Doppler Effect Distribution in Reactor Fuel Elements," 1962. for a Non-Uniform Temperature Final Report, WCAP-2048, July 20. Nodvik, R. J., et al, "Supplementary Report pn Evaluation of Mass Spectrometric and Radiochemical Analysis of Yankee Core I Spent Fuel, Including Isotopes of Elements Thorium Through Curium," WCAP-6086, August 1969. 21. Drake, M. K. (Ed.), "Data Formats and Procedure for the ENDF Neutron Cross Section Library," BNL-50274, ENDF-102, Vol. 1, 1970. 22. Suich, J. E. and Honeck, H. C., "The HAMMER System, Heterogeneous Analysis by Multigroup Methods of Exponentials and Reactors," DP-1064, January 1967. 23. Flatt, H. P. and Baller, D. C., "AIM-5, A Multigroup, One Dimensional Diffusion Equation Code," NAA-SR-4694, March 1960. 24. Barry, R. F., "Nuclear Design of Westinghouse Pressurized Water Reactors with Burnable Poison Rods," WCAP-7806, December 1971. 4.3-60 SGS-UFSAR Revision 19 November 19, 2001 * * *
- 25. Strawbridge, L. E. and Barry, R. F., "Criticality Calculations for Uniform Water-Moderated Lattices," Nuclear Science and Engineering 23, 58, 1965. 26. Nodvik, R. J., "Saxton Core II Fuel Performance Evaluation," WCAP-3385-56, Part II, "Evaluation of Mass Spectrometric and Radiochemical Materials Analyses of Irradiated Saxton Plutonium Fuel," July 1970. 27. Leamer, R. D., et al, "PU02-U02 Fueled Critical Experiments," WCAP-3726-1, July 1967. 28. Davidson, S. L., Ed., et al., "Extended Burnup Evaluation of Westinghouse Fuel," WCAP-10125-P-A, Appendix B, December 1985. 29. Henderson, W. B., "Results of the Control Rod Worth Program," WCAP-9217, October 1977. 30. Nguyen, T. Q. et al., "Qualification of the PHOENIX-P/ANC Nuclear Design System for Pressurized Water Reactor Cores," WCAP-11596-P-A, June 1988. 31. Liu, Y. S., et al., "ANC: A Westinghouse Advanced Nodal Computer Code," WCAP-10966-A, September 1986. 32. Iorii, J. A. and Petrarca, D. J., "Westinghouse Wet Annular Burnable Absorber Evaluation Report", WCAP-10021-P-A, Revision 1, October 1983 33. Bradfute, J. L., et al, "Criticality Analysis of the Salem Units 1 and 2 Fresh Fuel racks", NFU-VTDWW-94-083-00, January 1994 34. C.L. Beard and T. Morita, "BEACON Core Monitoring and Operations Support System", WCAP-12472-P-A, August, 1994. 35. T. R. Wathey, "Conditional Extension of the Rod Misalignment Technical Specification for Salem Units 1 and 2," WCAR 14962/14963, August 1997. 36. T. Morita and W. H. Slagle, "BEACON Core Monitoring and Operations Support System (WCAP-12472-P-A) ," Addendum 1-A, January 2000. 4.3-61 SGS-UFSAR Revision 27 November 25, 2013