ML20133N153
| ML20133N153 | |
| Person / Time | |
|---|---|
| Site: | South Texas  |
| Issue date: | 10/22/1985 |
| From: | Wisenburg M HOUSTON LIGHTING & POWER CO. |
| To: | Knighton G Office of Nuclear Reactor Regulation |
| References | |
| CON-#485-934 OL, NUDOCS 8510280411 | |
| Download: ML20133N153 (57) | |
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I fhl k NE Mf lionston 1.ighting & Power I!O. Box 1700 llouston. Texas 77001 (713) 228-9211
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October 22, 1985 ST-HL-AE-1432 File No.: C9.17 Mr. George W. Knighton, Chief Licensing Branch No. 3 Division of Licensing U. S. Nuclear Regulatory Commission Washington, DC 20555 South Texas Project Units 1 and 2 Docket Nos. STN 50-498, STN 50-499 Responses to DSER/FSAR Items Update of Table 4.3-1. Construction Materials
Dear Mr. Knighton:
The attachments enclosed provide STP's response to Draft Safety Evaluation Report (DSER) or Final Safety Analysis Report (FSAR) items.
The item numbers listed below correspond to those assigned on STP's internal list of items for completion which includes open and confirmatory DSER items, STP FSAR open items and open NRC questions.
This list was given to your Mr. N. Prasad Kadambi on October 8, 1985 by our Mr. M. E.
Powell.
The attachments include mark-ups of FSAR pages which will be incorporated in a future FSAR amendment unless otherwise noted below.
The items which are attached to this letter are:
Attachment Item No.*
Subject 1
F 4.3-1 Update of Table 4.3-1, Construction Materials.
Note: This attachment also includes a general update of Chapter 4.
- Legend Opf D - DSER Open Item C - DSER Confirmatory Item F - FSAR Open Item Q - FSAR Question Response Item y
L1/DSER/ao 0510280411 851022 PDR ADOCK 05000498 E
PDR J
Ilouston Lighting & Power Company ST-HL-AE-1432 File No.: 09.17 Page 2 If you should have any questions concerning this matter, please contact Mr. Powell at (713) 993-1328.
Very t y yours, g
M. R. W senburg Manager, Nuclear Mcensing JSP/bl Attachments: See above I.
s Ll/DSER/ao i
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P-ST-HL-AE-1432 File No.: C9.17 Page 3 cc:
Hugh L. Thompson, Jr., Director Brian E. Berwick, Esquire Division of Licensing Assistant Attorney General for Office of Nuclear Reactor Regulation the State of Texas U.S. Nuclear Regulatory Commission P.O. Box 12548, Capitol Station Washington, DC 20555 Austin, TX 78711 Robert D. Martin Lanny A. Sinkin Regional Administrator, Region IV 3022 Porter Street, N.V. #304 Nuclear Regulatory Commission Washington, DC 20008 611 Ryan Plaza Drive, Suite 1000 Arlington, TX 76011 Oreste R. Pirfo, Esquire Hearing Attorney N. Prasad Kadambi, Project Manager Office of the Executive Legal Director U.S. Nuclear Regulatory Commission U.S. Nuclear Regulatory Commission 7920 Norfolk Avenue Washington, DC 20555 Bethesda, MD 20814 Charles Bechhoefer, Esquire Claude E. Johnson Chairman, Atomic Safety &
Senior Resident Inspector /STP Licensing Board c/o U.S. Nuclear Regulatory U.S. Nuclear Regulatory Commission Commission Washington, DC 20555 P.O. Box 910 Bay City, TX 77414 Dr. James C.
Lamb, III 313 Woodhaven Road M.D. Schwarz, Jr., Esquire Chapel Hill, NC 27514 Baker & Botts One Shell Plaza Judge Frederick J. Shon Houston, TX 77002 Atomic Safety and Licensing Board U.S. Nuclear Regulatory Commission J.R. Newman, Esquire Washington, DC 20555 Newman & Holtzinger, P.C.
1615 L Street, N.W.
Mr. Ray Goldstein, Esquire Washington, DC 20036 1001 Vaughn Building 807 Brazos Director, Office of Inspection Austin, TX 78701 and Enforcement U.S. Nuclear Regulatory Commission Citizens for Equitable Utilities, Inc.
Washington, DC 20555 c/o Ms. Peggy Buchorn Route 1, Box 1684 E.R. Brooks /R.L. Range Brazoria, TX 77422 Central Power & Light Company P.O. Box 2121 Docketing & Service Section Corpus Christi, TX 78403 Office of the Secretary U.S. Nuclear Regulatory Commission H.L. Peterson/G. Pokorny Washington, DC 20555 City of Austin (3 Copies)
P.O. Box 1088 Austin, TX 78767 Advisory Committee on Reactor Safeguards U.S. Nuclear Regulatory Commission J.B. Poston/A. vonRosenberg 1717 H Street City Public Service Board Washington, DC 20555 P.O. Box 1771 San Antonio, TX 78296 Revised 9/25/85 l
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l TABLE 4.1-1 (Continued) j REACTOR DESIGN COMPARISION TABLE W. B. McGuire South Texas Project CORE MECHANICAL DESIGN PARAMETERS UNITS 1 & 2 UNIT 3 1 & 2 26.
Design RCC Canless RCC Canless 17 x 17 17 x 17 t
27.
Number of Fuel Assemblies 193 193 28.
UO2 Rods per Assembly 264 264 29.
Rod Pitch, in.
0.496 0.496 1
30.
Overall Dimensions, in.
8.426 x 8.426 8.426 x 8.426 31.
Fuel Weight (as UO ), Ib 222,739 259,Mo s 2
32.
Zirca d y Weight, Ib 50,913 54,840 4' 0co%,%fj
, l 18 33.
Number of Crids per Assembly 8 - Type R 10 - Type R i
34.
Loadies Technique 3 region 3 region cn i.*
non-uniform non-uniform 4
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ANALYTICAL TECHNIQUES IN CORE DESIGN S$ction Analysis Technique Computer Code Referenced Nuclear Design (Continued)
Group constants for control HAMER-AIM 4.3.3.2 rods with self-shielding 2.
X-Y Power Distributions, 2-D, 2-Group Diffusion TURTLE 4.3.3.3 3
Fuel Depletion, Critical Theory
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Boron Concentrations, x-y 2-D m 3-D pik;m pug 4333 4
Xenon Distributions, Theent - hstet Modtd Reactivity Coefficients s"lE 3.
Axial Power Distributions, 1-D, 2-Group Diffusion
, PANDA 4.3.3.3
. pp p]
Control Rod Worths, and Theory o'
g&g Axial Xenon Distribution g-4.
Fuel Rod Power Integral Transport Theory LASER 4.3.3.1 Effective Resonance Monte Carlo Weighting REPAD j
Temperature Function i
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2.
4.2 FUEL SYSTEM DESIGN The plant conditions for design are divided into four categories in accord-ance with,their anticipated frequency of occurrence and risk to the public:
Condition Normal Operation; Condition II - Incidents of Moderate Frequency; Condition III - Infrequent Incidents; Condition IV - Limiting Faults. The bases and description of plant operation and events involving each Condition are given in the Accident Analysis Chapter 15.
The reactor is designed so that its components meet the following performance and safety criteria:
1.
The mechanical design of the reactor core components and their physical arrangement, together with corrective actions of the reactor control, protection and emergency cooling systems (when applicable) assure that:
Fuel damage
- is not expected during Condition I and Condition II a.
events. It is not poerible, however, to preclude a very small number of rod failures. These are within the capability of the plant cles p system and are consistent with plant design bases.
b.
The reactor can be brought to a safe state following a Condition III event with only a small fraction of fuel rods damaged
- although sufficient fuel damage might occur to preclude immediate resumption of operation.
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c.
The reactor can be brought to a safe state and the core can be kept
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suberitical with acceptable heat transfer geometry following transients arising from Condition IV events.
2.
The fuel assemblies are designed to withstand, without exceeding the criteria of Section 4.2.1.5, loads induced during shipping, handling and core loading.
3.
The fuel assemblies are designed to accept control rod insertions in order to provide the required reactivity control for power operations and reactivity shutdown conditions.
4.
All fuel assemblies have provisions for the insertion of incore instru-mentation necessary for plant operation.
5.
The reactor internals in conjunction with the fuel assemblies and in-core control components direct reactor coolant through the core. This achieves acceptable flow distribution and restricts bypass flow so that the heat transfer performance requirements can be met for all modes of i
operation.
e eFuel damage as used here is defined as penetration of the fission
,h product barrier (i.e., the fuel rod cladding).
i 4.2-1 Amendment 45
ATTACHMENT /
I ST HL AE / Y31.
PAGE '/ OF 5 y STP FSAR the clad has some capability for accommodating plastic strain, the yield stress has been accepted as a conservative design basis.
from tat uni <raAla.ked cerdtico.
The claMic hue simh cturusb. frardent 2)
Clad Tensile Strain a
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The(strain is less than one percen This limit is consistent with proven practice.
c.
Vibration and Fatigue 1)
Strain Fatigue The cumulative strain fatigue cycles are less than the design strain fatigue life. This basis is consistent with proven prac-tice.
f 2;2)
Vibration Potential fretting wear due to vibration is prevented assuring that the stress-strain limits are not exceeded during design life.
Fretting of the clad surface can occur due to flow-induced vibra-tion between the fuel rods and fuel assembly grid springs. Vibra-tion and fretting forces vary during the fuel life due to clad diameter creep-down combined with grid spring relaxation.
C d.
Chemical Properties of the Cladding - This is discussed in Refer-ence 4.2-2.
4.2.1.2 Fuel Material.
y a.
Thermal-Physical Properties Fuel Pellet Temperatures - The center temperature of the hottest pellet is to be below the melting temperature of the UO (melting point of 5080*F [Ref. 4.2-3] unitradiated and decreasing by 58'F per 10.000 mwd /NIU). While a limited amount of center melting can be tolerated, the design conservatively precludes center melting.
A calculated fuel centerline temperature of 4700*F has been se-1ected as an overpower limit to assure no fuel melting. This pro-vides sufficient margin for uncertainties as described in Subsection 4.4.2.9.
The normal design density of the fuel is 95 percent of theoreti-cal. Additional information on fuel properties is given in Ref-erence 4.2-2.
b.
Fuel Densification and Fission Product Swelling The design bases and models used for fuel densification and swell-ing are provided in References 4.2-4 and 4.2-5.
c.
Chemical Properties w
Amendment 30
ATTACHMENI f ST HL AE 3 Y.72-STP FSAR PAGE5 OF5e/
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These limits are applied to the design and evaluation of the top and bottom nozzles, guide thimbles, grids, and the thimble joints.
The design bases for evaluating the structural integrity of the fuel assen-blies are:
Non-operational - 4 g axial and 6 g lateral loading with dimensional h9 a.
stability.
b.
For the normal operating and upset conditions, the fuel assembly 1
component structural design criteria are established for the two primsry material categories, namely austenitic steels and Zircaloy.
The stress categories and strength theory presented in the ASME B&PV Code,Section III, are used as a general guide. The maximum shear theory (Tresca criterion) for combined stresses is used to determine the stress intensities for the austenitic steel compon-ents. The stress intensity is defined as the numerically largest difference between the various principal stresses in a three dimen-sional field. The allowable stress intensity value for austenitic i
steels, such as nickel-chromium-iron alloys, is given by the lowest of the following:
1)
One-third of the specified minimum tensile strength or 2/3 of the specified minimum yielded strength at room temperature l(
2)
One-third of the tensile strength or 90 percent of the yield strength at temperature but not to exceed 2/3 of the specified minimum yield strength at room temperature.
The stress limits for the austenitic steel components are given l
below. All stress nomenclature is per the ASME B&PV Code,Section III.
Stress Intensity Limits Categories Limit General Primary Membrane Stress Intensity Sm Local Primary Membrane Stress Intensity 1.5 Sm Primary Membrane plus Bending Stress Intensity 1.5 Sm Total Primary plus Secondary Stress Intensity 3.0 Sm The Zircaloy structural components which consist of guide thimble and fuel tubes are in turn subdivided into two categories because of material differenpas and functional requirements. The fuel tube d
design criteria AYeovered separately in Section 4.2.1.1. The maxi- #
mun shear theory is used to evaluate the guide thimble destgn. For conservative purposes, the Zircaloy unirradiated properties are used to define the stress limits.
4.2-5 Amendment 49
ATTACHMENT /
ST HL-AE- /V32-PAGE & OF64/
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1.
Absorber Rods The material properties and compatibilities are given in Refs. 4.2-2 30 and 4.2-7.
The design bases include a stress intensity limit, Se, of 2/3 of the 0.2 percent offset yield stress for the 304 stainless steel AJad tubing during the 15 year minimum RCCA design life. The design bash of 4---
the absorber material is that it does not exceed its minimum melting 30 point of 3913*F (Ref. 4.2-7).
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2.
Burnable Poison Rods
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The burnable poison rod clad is designed as a Class 1 Component under i
Article NB-3000 of the ASME B&PV Code,Section III, 1973 for Conditions i
I and 11.
For abnormal loads during Conditions III and IV code stresses I
are not considered limiting. Failures of the burnable poison rods dur-ing these conditions must not interfere with reactor shutdown or cooling of the fuel rods.
!J The burnable poison absorber material is non-structural. The structural elements of the burnable poison rod are designed to maintain the absorber geometry even if the absorber material is fractured. The rods are designed '
i so that the absorber material is below its softening temperature (1492*F*
7 for reference 12.5 withove horon rods).
In addition, the structural ele-s ments are designed to prevent excessive slumping.
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3.
Neutron Source Rods The neutron source rods are designed to withstand the following:
I a.
The external pressure equal to the Reactor Coolant System (RCS) operating pressure with appropriate allowance for overpressure transients and, b.
An internal pressure equal to the pressure generated by released gases over the source rod life.
4.
Thimble Plug Assembly The thimble plug assembly is needed to restrict bypass flow through those thimbles not occupied by absorber, source or burnable poison rods.
The thimble plug assemblies satisfy the following:
I a.
Accommodate the differential thermal expansion between the fuel j
assembly and the core internals, b.
Maintain positive contact with the fuel assembly and the core j
internals.
1 i
- Borosilicate glass is accepted for use in burnable poison :ods if the softening temperature is 1510 1 18'F.
The softening temperature is defined i
in ASTM C 338.
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l and seal welded at the ends to encapsulate the fuel. A schematic of the fuel rod is shown in Figure 4.2-3.
The fuel pellets are right circular cylinders consisting of slightly enriched uranium dioxide powder which has been com-i cacted by c:;1d pressing and then sintered to the required density. The ends of each pellet are dished slightly to allow areater axial expansion at t j
center of the pellets aM but o. SmaU cMrder at tM cuhr eq[inder surfic
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To avoid overstressing of the clad or seal velds, void volume and clearances i
are provided within the rods to accommodate fission gases released from the fuel, differential thermal expansion between the clad and the fuel, and fuel Jensity changes during irradiation.
Shifting of the fuel within the clad j
dt. ring handling or shipping prior to core loading is prevented by a stainless l
steel helical spring which bears on top of the fuel. At assembly the pellets are stacked in the clad to the required fuel height, the spring is then in-i serted into the top end of the fuel tube and the end plugs pressed into the f
ends of the tube and welded. All fuel rods are internally pressurized with helium during the welding process in order to minimize compressive clad
. stresses and prevent clad flattening due to coolant operating pressures.
t The fuel rods are presently being designed and pre-pressurized so that: 1) the internal gas pressure mechanical design limit given in Subsection 4.2.1.3 (B) is not exceeded and, 2) the cladding stress-strain limits (Subsection 4.2.1.1) are not exceeded for Condition I and II events, and 3) clad flat-
[
tening will not occur during the fuel core life.
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4.2.2.2 Fuel Assembly Structure. The fuel assembly structure consists of a bottom nozzle, top nozzle, guide thimbles and grids, as shown in Figure 4.2-2.
l 4.2.2.2.1 Bottom Nozzle: The bottom nozzle serves as a bottom struc-l tural element of the fuel assembly and directs the coolant flow distribution to the assembly. The square nozzle is fabricated from Type 304 stainless j
steel and consists of a perforated plate and four angle Ir:gs with bearing plates as shown in Figure 4.2-2.
The legs form a plenum for the inlet cool-ant flow to the fuel assembly. The plate also prevents accidental downward ejection of the fuel rods from the fuel assembly. The bottom nozzle is fas-l tened to the fuel assembly guide tubes by locked screws which penetrate j30 through the nozzle and mate with a threaded plug in each guide tube.
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Coolant flow through the fuel assembly is directed from the plenum in the i
bottom nozzle upward through the penetrations in the plate to the channels between the fuel rods. The penetrations in the plate are positioned between the rows of the fuel rods.
Axial loads (holddown) imposed on the fuel assembly and the weight of the j
fuel assembly are transmitted through the bottom nozzle to the lower core i
support structure.
Indexing and positioning of the fuel assembly is control-l 1ed by alignment holes in two diagonally opposite bearing plates which mate i
with locating pins in;the lower core support. Any lateral loads on the fuel assembly are transmitted to the lower core support through the locating pins.
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PAGE # OF 6Y STP FSAR The clad'in the roPa'ssemblies is slightly cold worked Type 304 stainless steel. All other structural materials are Types 304 or 308 stainless steel except for the springs which are Inconel-718. The borosilicate glass tube provides sufficient boron content to meet the criteria discussed in Section 4.3.1.
4.2.2.3.3 Neutron Source Assembiv: The purpose of a neutron source assembly is to provide a base neutron level to ensure that the detectors are operational and responding to core multiplication neutrons.
Since there is very little neutron activity during loading, refueling, shutdown, and ap-proach to criticality, a neutron source is placed in the reactor to provide a positive neutron count of at least 2 counts per second on the source range detectors attributable to core neutrons. The detectors, called source range detectors, are used primarily when the core is suberitical and during special suberitical modes of operations.
The source assembly also permits detection of changes in the core multiplica-tion factor during core loading refueling, and approach to criticality. This can be done since the multiplication factor is related to an inverse func-tion of the detector count rate. Therefore a change in the multiplication factor can be detected during addition of fuel assemblies while loading the core, a change in control rod pocitions, and changes in boron concentration.
Both primary and secondary neutron source rods are used. The primary source rod, containing a radioactive material, spontaneously emits neutrons during initial core loading and reactor startup. After the primary source rod decays beyond the desired neutron flux level, neutrons are then supplied by the secondary source rod. The secondary source rod contains a stable mat-erial, which must be activated by neutron bombardment during reactor opera-tion. The activation results in the subsequent release of neutrons. This becomes a source of neutrons during periods of low neutron flux, such as during refueling and subsequent startups.
The reactor core employs four source assemblies:
two primary source assem-blies and two secondary source assemblies.
Each primary source assembly contains one primary source rod and a number of burnable poison rods.
Each secondary source assembly contains a symmetrical grouping of four secendary source rods'Md-may-centain a nuber-of--burnable-peisen-redsr--LocatimMl 18 f4Hed with a :=rce-or-burnable-poison-rod-contaiM thimble plug! The source assemblies are shown in Figures 4.2-13 and 4.2-14.
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Neutron source assemblics are employed at opposite sides of the core. The i
assemblies are inserted into the rod cluster control guide thimbles in fuel assemblies at selected unrodded locations.
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As shown in Figure / 4.2-13. cad 4.2 -1d thebsource assemblies-contains a hold-down assembly identical to that of the burnable poison assembi fThe secondary sourceJassembly shown in Figure 4.2-14 contains a spider assembly. The spider assembly is in the forn of a' central hub with radial vanes centaining cylindrical fingers from which the secondary source rods and thinble plugs are suspended, J i
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ATTACHMENT /
ST.HL AE W3b-PAGE 9 OF5y STP FSAR The primary and secondary source rods utilize the same cladding material as the absorber rods. The secondary source rods contain antimony-beryllium pellets stacked to a height of approximately 88 in.
The primary source rods contain capsules of californium (plutonium-beryllium possible alternate) source material and alumina spacer pellets to position the source material within the cladding.JThe r'ods7n each issembly~are permanently faste (hetopendtoaholddownassembly.)
The other structural members are constructed cf Type 304 stainless steel except for the springs. The springs exposed to the reactor coolant are Inconel 718.
W-Thimble \\PlugAssembly:[@Inordertolimitbypassflowthrough 4.2.2.3.4 the rod cluster control guide thimble ~ in fuel assemblies which do not con-assembliesarefittedwiththimbleplugassembliesatthoselo[catibus.,@-
tain either control roder, source rod, or burnable poison rode the fuel M
The thimble plug assemblies as shown in Figure 4.2-15 corisist of a flat base plate with short rods suspended from the bottom surface and a spring pack assembly. The twenty-four short rods, called thimble plugs, project into the upper ends of the guide thimbles to reduce the bypass flow.
Each thimble plug is permanently attached to the base plate by a nut which is lock-welded to the threaded end of the plug. Similar short rods are also used on the source assemblies and burnable poison assemblies to plug the ends of all
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vacant fuel assembly guide thimbles. At TustallatTon in core, the thimb1 M
- plug assemblies interface with both the upper core plate and with the fuel assembly top nozzles by resting on the adaptor plate. The spring pack is compressed by the upper core plate when the upper internals assembly is lowered into place.
All components in the thimble plug assembly, except for the springs, are con-structed from Type 304 stainless steel. The springs are Inconel 718.
4.2.3 Design Evaluation QL
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The fuel assemblies,4.2, the mechanical design bases of,.2.1, and other ip-the p uel rods are designed to sati an safety criteria of4 e
terfacingnuclearandthermal-hydraulicdesignbasesspecifiedinSectioni4.3 and 4.4.
Effects of Accident Conditions II, III, IV or anticipated transi-ents without trip (ATWT) on fuel integrity are presented in Chapter 15 or supporting topical reports.
The initial step in fuel rod design evaluation for a region of fuel is to determine the limiting rod (s). Limiting rods are defined as those rod (s) whose predicted performance provides the minimum margin to each of the design criteria.
For a number of design criteria the limiting rod is the lead burn-up rod of a fuel region. In other instances it may be the maximum power or the minimum burnup rod. For the most part, no single rod will be limiting with respect to all design criteria.
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UNITS 1 & 2 Figure 4.213.
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reactivity compensation. The core is also designed to have an overall negative moderator temperature coefficient of reactivity so that average coolant temperature or void content provides another, slower compensatory effect. Nominal power operation is permitted only in a range of overall negative moderator temperature coefficient.
The negative moderator ten-perature coefficient can be achieved through use of fixed burnable poison and/or control rods by limiting the reactivity held down by soluble boron.
Burnable poison content (quantity and distribution) is not stated as a design basis other than as it relates to accomplishment of a non-positive moderator temperature coef ficient at power operating conditions discussed above.
4.3.1.3 control of Power Distribution.
Basis The n'uclear 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 operating conditions including an allowance of 2 percent for calori-metric error and not including power spike factor due to densification.
2.
Under abnormal conditions including the maximum overpower condition,
(
the fuel peak power will not cause melting as defined in Subsection 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 DNBR shall not be less thanf4r28;"as discussed in Section 4.4.1) under g
Condition I and 11 events including the maximum overpower condition.
l.30 4.
Fuel management will be suc 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 GDC10.
Discussion Calculation of extreme power shapes which hffect fuel design limits is performed with proven methods and verified frequently with measurementa from operating reactors.
The conditions under which limiting power shapes are assumed to occur are chosen conservatively with regard to any permis-sible operating state.
Even though there is good agreement between measured peak power calcu-l lations and measurements, a nuclear uncertainty margin (Subsection l
4.3.2.2-7) is applied to calculated peak local power. Such a margin is provided both for the analysis for normal operating states and for antici-(
pated transients.
4.3-3
TIT ' HMENT
/
]
C ST HL AE-W3 >
Limits for alarms, reactor trip, etc. will be given in the Technical ecifications. Descriptions of the systems provided are given in Section 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. he reactivity coefficients reflect the changes in the neutron multiplication due to varyirg 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 Chapter 15.
ne reactivity coefficients are calculated on a corewise basis by radial and axial diffusion theory methods. The effect of radial and axial power distribution on core average reactivity coefficients is implicit in those calculations and is not significant under normal operating conditions. For example, a skewed xenon distribution which results in changing axial offset by 5 percent changes the moderator and Doppler temperature coef ficients by less than 0.01 pcm/*F and 0.03 pcm/*F respectively. Anartgficialyskewedxenondistributionwhichresultsin changing the radial by 3 percent changes the moderator and Doppler temperature coeffici s by less than 0.03 pcm/*F and 0.001 pcm/*F respectively. The spatial effects are accentuated in some transient conditions; for example, in postulated rupture of the main steamline break and rupture of RCCA mechanism housing described in Sections 15.1.5 and 15.4.8, and are included in these analyses.
The analytical methods and calculational models used'in calculating the reactivity coefficients are given in Section 4.3.3.
Rese models have been confirmeo 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.
~
The reactivity requirements at EOL of a typical cycle for a 168 in and a 144 in 17 x 17 four loop core are listed on a comparable basis in Table 9 e <J 4.3-4.
The Doppler defect is slightly less for the 168 in core _due-to the y 3ch lower average linear power density (5.20 vs. 5.44 Kw/ft). The moderator defect is higher due to the slightly more negative moderator temperature 9. L M. $
coefficient at the higher temperature of the 168 in core.
ne redistribution requirement is greater for the longer core (1.20 percentA p vs. 0.85 percentap). More excess margin is available to the 168 in core
{
than the 12 ft core due to the use of 57 rather than 53 control rods in this example. Both cores operate in the same range of expected reactivity l
ters as shown in Table 4.3-5.
4.3-20 Amendment 27
/.lL ATTACHMENT /
ST-HL.AE 3 VJJ-STP FSAR PAGE 21 0F 66f 4
(-
4.3.2.3.1 Fuel Temperature (Doppler) Coefficient: The fuel tempera-l ture (Doppler) coefficient is defined au the change in reactivity per degree change in effective fuel temperature and is primarily a measure of the Doppler broadening of uranium-238 and plutonium-240 resonance absorp-tion peaks. Doppler broadening of other isotopes such as uranium-236, neptunium-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 l
corresponding reduction in reactivity.
r The fuel temperature coefficient is calculated by perfo X-YcalculationsusinganupdatedversionoftheTURTLE{gigggjo-group Code.
Moderator temperature is held constant and the power level is varied.
Spatial variation of fuel temperature is taken into account by calcula-ting the effective fuel temperature as a function of power density as discussed in Subsection 4.3.3.1.
The Doppler temperature coefficient is shown on Figure 4:3-27 as a func-tion of the effective fuel temperature (at BOL and EOL conditions). The effective fuel temperature is lower than the volume averaged fuel tempera-ture since the neutron flux distribution is non-uniform through the pellet i
and gives preferential weight to the surface temperature. The Doppler-only contribution to the power coefficient, defined later, is shown on Figure 4.3-28 as a function of relative core power. The integral of the differential curve on Figure 4.3-28 is the Doppler contribution to the power defect and is shown on Figure 4.3-29 as a function of relative power.
(
The Doppler coefficient becomes more negative as a function of life as the plutonium-240 content increases, thus increasing the plutonium resonance absorption, but overall beenmes less negative since the fuel temperature changes with burnup as described in Subsection 4.3.3.1.
The upper and lower limits of Doppler coefficient used 10 accident analyses j
are given in Chapter 15.
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 co.ef-ficients so obtained are moderator density, temperature, pressure and void coefficients.
Moderator Density and Temperature Coefficients (den:it4 The moderator temperature e icient is defined as the change in reac-tivity per degree change in the moderator temperature. Generally, the effect of the changes in moderator density as well as the temperature are considered together. A decrease in moderator density means less moderation which results in a negative moderator coefficient. An increase in coolant temperature, keeping the density constant, leads to a hardened neutron spectrum and results in an increase in resonance absorption in uranium-238, plutonium-240 and other isotopes. The hardened spectrum also causes a s
4.3-21
ATTACHMENY I STP FSAR ST-HL AE W3 A PAGE WF 9/
decrease in the fission to capture ratio in uranium-235 and plutonium-239.
I Both of these effects make the moderator coefficient more negative. Since water density changes more rapidly with temperature as temperature increases, the eoderator temperature coefficient become more negative with increasing temperature.
temfemkre The soluble boton used in the reactor as a means of reactivity control of. -
also has an effect on moderatoradeneity* coefficient since the soluble boron poison density as well as the water density is decreased when the b, Ifa b W
L_
coolant temperature rises. A-decreas (in the soluble poison concentra-tion introduces a positive component in the moderator coeffisient.
emperstwb-r Thus, if the concentration of soluble poison is large enough, the net value of the coefficient may be positive. With the burnable poison rods present, however, the initial hot boron concentration is sufficiently low that the moderator temperature coefficient is negative at operating tem-
~~
The effect of control rods is to make the moderator (ccief-peratures.
ficient more negative by reducing the required soluble baron concentration 3
and by increa' sing the " leakage" of the core.
%pe rskrg With burnup, the moderator (coef ficient becomes ac,re negative primarily as a result of boric acid dilution but also to a significant extent from the effects of the buildup of plutonium and fission products.
@ pro 4--
The moderato H coefficient is calculated for the various plant conditions discussed above by performing two-group X-Y calculations, varying the moderator temperature by about 5'F about each of the mean temperatures. f The moderatorfcoeMicient is shown as a function of core temperature and QfMM boron concentration for the unrodded and rodded core on Figures 4.3-30 through 4.3-32.
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 4.3-33 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 Figure 4.3-3.
(icmperatue. (4mM The moderatorkcoef ficients presented here are calculated on a corevide basis, since they are used to describe the core behavior in normal and accident situations when the moderator temperature changes can be con-sidered to affect the entire core. Moderator temperature coefficient and moderator density coefficient are used interchangeably according to which is more appropriate as input for the codes used, Moderator Pressure Coefficient The moderator pressure coefficient relates the change in moderator den-sity, resulting from a reactor coolant pressure change, to the corres-ponding effect on neutron production. This coefficient is of much less A significance in comparison with the moderator temperature coefficient.
4.3-22
~__-,,___.,___,-___m.
o i
~
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change of 50 psi in pressure has approximately the same effect on reac-4 tivity as a half-degree change in moderator temperature. This coefficient can be detprained from the moderator temperature coefficient by relating change in pressure to the corresponding change in density. The moderator pressure coefficient is negative over a portion of the moderator tempera-ture range at BOL (-0.004 pcm/ psi, BOL) but is always positive at operating conditions and becomes more positive during life (+0.3 pcm/ psi EOL)[ glue e
principally to the change in boron concentration of the moderator Moderator Void Coefficient i,
The moderator void coefficient relates the change in neutron multipli-cation to the presence of voids in the moderator. In a PWR this coeffi-cient is not very significant because of the low void content in the 3
coolant. The core void content is less than one-half of one percent and is due to local. or statistical boiling. The void coefficient varies from 50 pcm/ percent void at BOL and at low temperatures to -250 pcm/ percent 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 tem-perature and fuel temperature change as the core power level changes is j
called the total power coefficient and'is expressed in terms of reactivity change per percent power change. The power coefficient at BOL and EOL conditions is given on Figure 4.3-34.
It becomes more negative with burnup reflecting the combined effect of moderator and fuel temperature coefficients with burnup. The power defect l
(integral reactivity effect) at BOL and EOL is given on Figure 4.3-35.
i 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 pre-sented there, the accuracy of the current analytical model is:
10.2 percent op for Doppler and power defect 12 pcm/*F for the moderator coefficient.
Experimental eva?.uation of the calculated coefficients will be capleted during the physics star:up tests described in Chapter 14.
4.3.2.3.5 Reactivity Coefficients Used in Transient Analysis: Table 4.3-2 gives the limi'.ing values as well as the best estimate values for the i
reactivity coefficients. The limiting values are used as design limits in the transf.ent analysis. The exact values of the coefficient used in the analysir. depend on whether the f.ransient of interest is examined at MBOL or EOL, whether the most negative or the most positive (least i
negative) coefficients are appropriate, and whether spatial nonuniformity i
must be considered in the analysis. Conservative values of coefficients, considering various aspects of analysis are used in the transient analysis.
1 This is described in Chapter 15.
l 4.3-23 i
1
ATTACHMENT i ST-HL-AE-1 YJ 3-PAGE J Y OF 59 STP FSAR 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 shutdown to the hot zero power (HZP) 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 tempera-ture is conservatively increased by 4*F to account for the control dead band and measurement errors.
m peu h rt emptabrd e
Since the moderator oe ficient is negative, there is a reactivity addi-tion with power reduction. The moderator (E efficient becomes more nega-tive as the fuel depletes because the boron concentration is reduced.
This effect is the major contributor to the increased requirement at EOL.
3 4.3.2.4.3 Redistribution: During full power operation the coolant density decreases with core height, and this, together with partial inser-tion 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 ccre, 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 calcu-lated with an allowance for effects of xenon distribution.
h0 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 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 of 10 percent Ap (hot) 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 the addition of soluble boron to the coolant and by burnable poiso The soluble boron concentration for ar--
~
ATTACHMENT /
.PAGE J6 OFC/
)
several core configurations, the unit boron worth, and burnable poison worth are given in Tables 4.3-1 and 4.3-2.
Since the excess rea4tivity for burnup is controlled by soluble boron and/or burnable poisonf it is e
not included in control rod requirements.
gg 4.3.2.4.7 Xenon and Samarium Poisoningk Changes in zenon and samarium concentrations in the corc occur at a sufficiently slow rate, even fol--
lowing rapid power level changes, that the resulting reactivity change is controlled by changing th soluble boron concentrag Q 2.4.8 pH ef f e~Etsl hanges in reactivity due to a change in 4
~
coolant pH, if any, are sufficiently small in magnitude and occur slowly enough to be controlled by the boron system. Further details are provided in Reference 4.3-13.
I A.lSECT 4.3.C.4.O h
4.3.2.4.9 Experimental Confirmation: Following a normal shutdown, the total core reactivity change during cooldown with a stuck rod has beeh measured on a 121 assembly, 10 ft high core and 12J assembly, 12 ft high core. In each case, the core was allowed to coo 16own until it reached 4--
criticality simulating the steamline break accident. For the 10 ft core, i
the total reactivity change associated with the cooldown is overpredicted by about 0.3 percent Ap with respect to the measured result. This repre-sents an error of about 5 percent in the total reactivity change and is about half the uncertainty allevance for this quantity. For the 12 ft core, the difference between the measured and predicted reactivity change was an even smaller 0.2 percent op. These measurements and others demon-strate theM(CapMoiIi bil of the methods described in Section 4.3.3 4.3.2.4.10 Control: Core reactivity is controlled by means of a chemical poison dissolved in the coolant, RCCA's, and burnable poison rods as described below.
4.3.2.4.11 Chemical Poison: Boron in solution as boric acid is used to control relatively slow reactivity changes associated with:
1.
The moderator temperature defect in going from cold shutdown at ambient temperature to the hot operating temperature at zero power, 2.
The transient xenon and samarium poisoning, such as that following power changes or changes in rod cluster control position, 3.
The excess reactivity required to compensate for the effects of fissile inventory depletion and buildup of long-life fission products.
1 4.
The burnable poison depletion.
The boron concentrations for various core conditions are presented in Table 4.3-2.
l l
4.3-26
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ATTACHMENT I
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ST HL AE l'/31 PAGE at, OF 6V
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pck reactivity requirements at E01,of a typical cycle for a 168 M and a 144 17 'it 17fourloop.corearelistedonaco:parablebagpy'inTable 4.3-4 The Doppler defect is slightly less for the 168 W core due to the lower average linear power density (5.20 vs. 5.44 Kw/ft).
The moderator defect is higher due to the slightly more negative mpfgrator temperature coefficient at the higher temperature of the 168 it' core.
The redistribution requirement is greater for the longer core (1.20 percent A p vs. 0.85 percento p ).
More excess margin is available to the 168 W ore, g than the 12 ft core due to the use of 57 rather than 53 control rods in this exa ple.
Both cores operate in the same range of expected reactivity paraceters as shown in Table 4.3-5.
t l
l l
i
F ATTACHMENT /
ST-HL-AE- / {3 ;L STP FSAR Only
(
4.3.2.4.12 Rod Cluster Control Assemblies: /'ull-length assemblies are employed etesrvaWin this reactor. The number of assemblies is shown in g
Table 4.3-1.
The RCCA's are used for shutdown snd 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 tempera-ture, or menon concentration (with rods not exceeding the allowable rod insertion limits),
4.
Reactivity ramp rates resulting ftom load changes..
The allowed control bank reactivity insertion is limited at full power to
@7 maintain shutdown capability. As the pouer level is reduced, control rod reactivity requirements are also reduced and more rod insertion is allowed. The contcol bank position is r.onitored 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
,, lmegg determining the maximum worth of an inserted RCCA ejection accident.
CM Further discussion will be provided in the Technical Specifications on rod l 27 insertion limits.
Power distribution, rod ejection and rod misalignment analyses are based on the arrangement of the shutdown and control groups of the RCCA's shown on Figure 4.3-36.
All shutdown RCCA's are withdrawn before withdrawal of the control banks is initiated.
In going from zero to 100 percent power, control banks A, B, C and D are withdrawn sequentially. The limits of rod positions and further discussion on the basis for rod insertion limits will 27 be provided in the Technical Specifications.
+
4.3.2.4.13 Reactor Coolant Temperature:
Reactor coolant (or moderator) temperature control has added flexibility in reactivity control of the Westinghouse PWR.
This feature takes advantage of the negative moderator temperature coefficient inherent in a PWR to:
1.
Maximize return to power capabilities 2.
Provide + 5 percent power load regulation capabilities without requiring control rod compensation 3.
Extend the time.in cycle life to which daily load follow operations can be accomplished s
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STP FSAR 6f mpe rAhtre Reactor coolant temperature control supplements the dilution capability of the plant by lowering the reactor coolant l temperature to' supply positive reactivity thfough the negative moderatoracoefficient of the reactor. After 4---
the transient is over, the system automatically recovers the reactor coolant temperature to the programmed value.
Moderator temperature control of reactivity, like soluble boron control, has the advantage of not significantly affecting the core power distribution.
However, unlike boron control, temperature control can be rapid enough to achieve reactor power change rates of 5 percent / minute.
4.3.2.4.14 Burnable Poison Rods: The burnable poison rods provide partial control of the excess reactivity available during the first fuel cycle. In doing so, these rods prevent the moderator temperature coefficient from being positive at normal operating conditions. They perform this function by reducing the requirerent for soluble poison in the moderator at the begidning of the first fuel cycle as described previously. For purposes of illustration a typical burnable poison rod pattern in the core together with the number of rods p,er assembly is shown on Figure 4.3-5, while the arrangements within an assembly are displayed on Figure 4.3-4.
The reactivity worth of these rods is shown in Table 4.3-1.
The boron in the rods is depleted with burnup but at a sufficiently slow rate so that the resulting critical concentration of soluble boron is such that the moderator temperature coefficient remains negative at all times for power operating conditions.
4.3.2.4.15 Peak Xenon Startup: Compensation for the peak xenon buildup is accomplished using the boron control system.
Startup from the peak xenon condition is accomplished with a combination of rod motion and boron dilution.
The boron dilutien may be made at any time, including during the shutdown period, provided the shutdown margin is maintained.
[5 4.3.2.4.16 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 in the Technical Specifications 1 30
/
and discussed in Subsections 4.3.2.4.12 and 4.3.2.4.13.
The power distri-bution is maintained within acceptable limits through the location of the rod l 30 bank. Reactivity changes due to the changing xenon concentration can be controlled by rod motion and/or changes in the soluble boron concentration.
Late in cycle life, extended load follow capability is obtained by augmenting the limited boron dilution capability at low soluble boron concentrations by temporary moderator temperature reductions.
Rapid power increases (5 percent / minute) from part power during load follow operation are accomplished with a combination of rod motion, moderator temperature reduction, and boron dilution. Compensation for the rapid power i
e AneroE
~
ATTACHMENT f ST-HL AE- /f3 2 PAGE M OF 54 STP FSAR c-A increase is accomplished initially by a combination of rod withdrawal and moderator, temperature reductioc. As the slower boron dilution takes affect after the' initial rapid power increase, the moderator temperature returns to the programmed value.
4.3.2.4.17 Burnup:
Control of the excess reactiv ty for burnup is accomplished using soluble boron and/or burnable poiso. The boron concen-
+-
tration must be limited during operating conditions to ensure tge
,e, rat o r 3,
temperature coefficient is negative.
Sufficient burnable poisont thstal-e led at the beginning of a cycle to give the desired cycle lifetime with-out exceeding the boron concentration limit. The practical minimum boron concentration is 10 ppm.
4.3.2.5 Control Rod Patterns and Reacti ity Worth. The RCCAs are l27 designated by function as the control group and the shutdown groups.
The terms " group" and " bank" are used sy onymously throughout this report we--
to describe a particular grouping of control assemblies. The rod cluster saspembly pattern is displayed on Figure 4.3-36.
The control banks are labeled A, B, C, and D and the shutdown banks are labeled SA, SB, etc.,
as applicable.
Each bank, although operated and controlled as a unit is comprised of two subgroups. The axial position of the RCCAs may be con-127 trolled manually or automatically. The RCCAs are all dropped into the core following actuation of reactor trip signals.
' (.
~
Two e iteria have been employed for selection of the control groups.
Firs the total reactivity worth must be adequate to meet the require-r ments specified in Table 4.3-3.
Second, in view of the fact that these rods may be particily inserted at power operation, the total power peaking factor should be low enough to ensure that the power capability require-ments 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 amou s.
The axial power shape would be more peaked following movement of a single group of rods worth three to four percent op than following movement of more banks each worth less; therefore, four l 18 banks (described as A, B, C, and D on Figure 4.3-36) each worth approxi-mately One percent AP have been selected. Typical control bank vorths are shown in Table 4.3-2.
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 pf with control rods above the insertion limit set by shutdown and other considerations which will be given in the Technical Specifications. Early l 27 in some cycles there may also be a withdrawal limit at low power to maintain a negative moderator temperature coefficient. As xenon and other fission products accumulate, this restriction is relaxed. However for the reference final core design described in this chapter, no such withdrawal limit is required.
Ejected rod worths are given in Section 15.4.8 for several different
(
conditions.
Allowable deviations due to misaligned control rods will be discussed in l 27 the Technical Specifications.
l 1
~
ATTACHMENT i ST-HL AE-IV3%
.PAGE 30 0F 5%/
Confirmatory critical experiments on burnable poisons are described in 4 --
Reference 4.3-3 4.3.3.3 Spatial Few-Group Diffusion Calculations.
Spatial few-group calculations consist primarily of two group dif fusion X-Y calculations
-.iising an' updated version of the TURTLE Code, two-grgup x y nodal calcula-D71 cNy 'and two-group axial e
gonusingbnupdaledversionoftheFLARE
/ calculatione_using an updated version of the PANDA Code.
(ths. I'AL ADcM E4 3-33] cMEd uiscrete X-Y calculations (1 mesh per cell) are carried out to determine critical boron concentrations and power distributions in the X-Y plane.
An axial average in the X-Y plane is obtained by synthesis from unrodded and rodded planes. Axial effects in unrodded depletion calculations are accounted for by the axial buckling, which varies with burnup and is determined by radial depletion calculations which are matched in reactivity to the analogous R-Z depletion calculation. The moderator coefficient is evaluated by varying the inlet temperature in the same X-Y calculations a
used for power distribution and reactivity predictions.
Validation of TURTLE reactivity calculations is associated with the validation of the group constants themselves, as discussed in Subsection 4.3.3.2.
Validation of the Doppler calculations is associated with the fuel temperature validation discussed in Subsection 4.3.3.1.
Validation of the moderator coefficient calculations is obtained by comparison with plant measurements at hot zero power conditions as shown in Table 4.3-13.
I VEC.%
Q 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 con-stants and the radial buckling used in the axial calculation are obtained o %e PANDA' radialTalEblat'io~nT~id which group ~constadts in annular' rings representing the various material regions in the X-Y plane are homogenizedbyflux-volumeweighting.)
W Validation of the spatial codes for calculating power distributions involves the use of incore and excore detectors and is discussed in Subsection 4.3.2.2.7.
Based on comparison with measured data it is estimated that the accuracy of current analytical methods is:
1 0.2 percent Ap for Doppler detect
+ 2 x 10-5/*F for moderator coefficient
-1 50 ppm for critical boron concentration with depletion 1 3 percent for power distributions 10.2 percent Ap for rod bank worth 1 4 pcm/ step for differential rod worth 10.5 pcm/ ppm for boron worth 10.1 percent Ap for moderator defect
)
4.3-40
ATTACHMENT I
- d. 3-M lA/3 EF'TI-fo V.*5 3.3 ST-HL AE-g3p PAGE a/
jgg g g pALA0btl is uuc9 ig 4wn c9imen siev a.l ad +ktse _eQimens; m,(
co\\ cu lati cws. FAL A Do Al ca n k us48 in.
sa E4 4 y a wa l st.s y
ea\\eviaktens, M LA=v aJ k &<4< r min e cri+ic l bovew cew cvub 4;ews em b I re-8 wo r f A s,
od reachivi +y coe ffici<wls.
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/ 4 fe f#'
icw LE mla 1 << - l/
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ATTACHMENT /
~
ST-HL AE / y3L PAGE 9 AOF @/
q STI' TS.A.It f
I 4.3-28 Suich, J. E. and Honeck, H. C., "The RAMMER System, Hetergeneous Analysis by Multigroup Hethods of Exponentials and Reactors,"
DP-1064, January, 1967 4.3-29 Flatt, H. P. and Buller, D. C., " AIM-5, A Multigroup, 0$e Dimen-sional Dif fusion Equation Code," NAA-SR-4694, March,1960.
- 4. 3-Moore, J.
S., " Nuclear Design of Westinghouse Pressurized Water Reactors with Burnable Poison Rods," WCAP-9000-L, Revision 1 (Proprietary), July,1969 and WCAP-7806, Dece=ber,1971.
- 4. 5 -33. Caml.s w, T.vn.,4t t 4 3 " fat g ooia - t. Mat sunjhouse detal j
CompAr Code," WC AP - 9VS r A (Propristory) om0 g WC AP 9VS6 A (Hn - Prepriefory ), hel,* r-i g 7 c), ad s
6eid i, serkm bu N-9 g
h
~~
Y 3 -30 g
Medvi k., R.J., " Sado t Cove.T.I Fue(
P<<Ermance
\\
Evaladim," u>c A P - 338f-SG, Park II, " Eva lua fim e [
nicss Spc4 rome.fr c awa Ractio cksmical Ana lyses i
e [ 1rva&:ak<k
% 4em. Pla teniu ~
Fu < l, " Jal;;,197o.
/.
4.~5 -3 t 6,ne r, R. c. 3.ee al., "P vo - vo _ Fuelc4 Cri Heo l z
2 Experiwks, " uxR P-17z<,-l, Jaly, twe7
~
4 J-49
~
~
ATTACHMENT /a-ST-HL AE /t/?
PAGE 4 4 OF<t/
(
TABLE 4.3-1 (Continued)
~~
REACTOR CORE DESCRIPTION (First Cycle)
Diameter of Guide Thimbles (upper part), in.
0.450 I.D.
0.482 0.D.
Diameter of Guide Thimbles (lower part), in.
0.397 1.D.
0.429 0.D.
Diameter of Instrument Guide Thimbles, in.
0.450 I.D'.
0.482 0.D.
Fuel Rods Number 50,952 Outside Dir. meter, in.
0.374 Diametral Gap, in.
0.0065 Clad Thickness, in.
0.0225 Clad Material Zircaloy-4 Fuel Pellets Material UO Sintered Density (percent of Theoretical) 95 Fuel Enrichments, wt %
Region 1 1.50 Region 2 2.20 Region 3 2.90 Diameter, in.
0.3225 Length, in.
0.530 [ -
Mass of UO per Foot of Fuel Rod, Ib/ft OM Rod Cluster Control Assemblies Neutron Absorber Hafnium Composition M
Diameter, in.-
0.341 4.3-45 Amendment 30
~
ATTACHMENT /
ST-HL AE- /V3>-
(
TABLE 4.3-1 (Continued) n REACTOR CORE DESCRIPTION (First Cycle) s Density, Ib/in 0.454 (min) 30 Cladding Material Type 304, Cold Worked Stainless Steel Clad Thickness, in.
0.0185 Number of Clusters 57 30 Number of Absorber Rods per Cluster 24 Burnable Poison Rods (First Core)
Number 946 l18 Material Borosilicate Class C. tad Outside Diameter, in.
0.381 Inner Tube 0.D.,
in.
0.1815 Clad Material Stainless Steel Inner Tube Material Stainless Steel
(
Boron Loading (w/o B Os in glass rod) 12.5 Weight of Boron-10 per foot of rod, Ib/ft
.000419 l30 Initial Reactivity Worth, %Ap 4.65 (HFP), 4.65 (HZP) 3.40 (cold)
Excess Reactivity Maximum Fuel Assembly k. (Cold, Clean, Unborated Water) 1.39 Maximum Core Reactivity (Cold. Zero Power, Beginning of Cycle) 1.22
(
I 4.3-46 Amendment 30
[
^
1 TABLE 4.3-2 NUCLEAR DESIGN PARAMETERS 1
(First Cycle)
Core Average Linear Power, kW,'f t, including densification effects 5.20 Total Heat Flux Hot Channel Factar, F 2.50 q
Nuclear Enthalpy Rise Hot Channel Jactor, F 1.52 aH
+
Reactivity Coefficients Design Limits Best Estimate l
7 Doppler-only Power, Upper Curve
-19.4 to -12.6
-b u
Y Coefficients, pen /*F++
C-125 5
0 (See Figure 15.0-5),
Lower Curve
-10.2 to -6.7 4TAto -9 l3 g
i Doppler Temperature Coefficient
-2.9 to -1.1
-2.5 to M _g'g g
I pcm/*FM Moderator Temperature Coefficient.
0 to -40
-6.
to -30.0 pcm/*F++
Boron Coefficient, pcm/ ppm ++
-16 to -7
-14. to -9
> b3 y
i Rodded Moderator Density Coefficient, pcm/ga/cc++
<0.43 x 105
<0.34 x 105 i
w2 r MN i
! (
@b S
j
'j
+ Uncertainties are given in Section 4.3.3.3 y
W I
N E
b.
j 4
I p
A
'1%BLE 3.3-2 (Continued)
NUCLEAR DESICN PARAMETERS a
s (Pirst Cycle)
.i Radial Factor (BOL to EOL)
,p Unrodded to 1.28 D bank 1.50 to 1.45 D+C 1.60 to 1.45 D+C+B 1.80 to 1.55 I'l u
Boron Concentrations. BOL. ppm y
s.
vs Zero Power, k,gg = 0.99, Cold, Rod Cluster y,
Control Assemblies out, clean 1080 Zero Power, k,gg = 0.99 Hot, Rod Cluster Control Assemblies Out, clean 1030 Design Basis Refueling Boron Concentration 2500 j$g Zero Power, k,gg <0.95, Cold, Rod Cluster M
Control Assemblies In, clean 910 dfg.
Zero Power, k,gg = 1.00, Hot, Rod Cluster 31kk control Assemblies Out, clean 930 Full Power, No Xenon, k,gg = 1.0, Hot, Rod Cluster Control Assemblies Out 835 Zero Power, K
=.99, Cold, Rod eff Cluster Control Assemblies in Less 730 Most Reactive Rod Stuck in Full Out Position s
ATTACHMENT 1 ST HL AE /W)L/
(
TABLE 4.3-4 COMPARISON OF REACTIVITY REQUIREMENTS End of Life (Equilibrium Cycle) 3800 MWt 3411 MWt 168 inch fuel 144 inch fuel 1.
Control Requirements a.
Fuel Temperature (Doppler), % Ap 2.95 2.94
+ Moderator Temperature, % Ap
+ Void, % Ap
+ Rod Insertion Allowance, % Ap b.
Redistribution, % Ap 1.20 0.85 2.
Total Control, % Ap 4.15 3.79 3.
Estimated Rod Cluster Control Assembly Worth 130 a.
Number of Control Rod Clusters 57 53 b.
Worth of all assemblies, % Ap 8.50 7.30 l30 c.
Worth of all but one Assembly 6.90 6.20 (highest worth), % Ap 4.
Estimated Rod Cluster Control Assembly
)
credit with 10 percent adjustment to i
accommodate uncertainties 6.20 5.58 (3c - 10 percent), % Ap i
5.
Shutdown Margin Available 2.05 (4-2), % Ap M *I Ibl I
1.79 l
[a] The design basis minimum shutdown is 1.75% Ap
[b] The design basis minimum shutdown is 1.60% Ap m
~
3 PAGE h OF 64/
f TABLE 4.3-6
~-
BENCFMARK CRITICAL EXPERIMENTS Description of Number of LEOPARD k Using df Experiments
- Experiments Experimental Bucklings UO 2
~
Al clad 14 1.0012 SS clad 19
'O.9963
)
Berated H O 7
0.9989 2
Subtotal 40 0.9985 U-Metal Al clad 41 0.9995 Unclad 20 0.9990 Subtotal 61 O.9993 Total 101 0.9990
- Reported in Reference [4.3 4.3-55
~
ATTACHMENT /
/Y3A ST HL Ag 0Frit/
PAGE W TABLE 4.3-10
(
SAXIUN CORE II ISOTOPICS ROD MY, AXIAL ZONE 6 LEOPARD Atom Ratio Measured
- 20 Precision (%)
Calculation
-5
-5 U-234/U 4.65 x 10 29 4.60 x 10
-3
-3 U-235/U 5.74 x 10
+0.9 5.73 x 10
-4
-4 U-236/U 3.55 x 10 15.6 3.74 x 10 U-238/U 0.99386 10.01 0.99385
~3 Pu-238/Pu 1.32 x 10 12.3 1.222 x 10
~
Pu-239/Pu 0.73971 10.03 0.74497 Pu-240/Pu 0.19302 10.2 0.19102
-2
-2 Pu-241/Pu 6.014 x 10
+0.3 5.74 x 10
-3
-3 Pu-242/Pu 5.81 x 10
+0.9 5.38 x 10
(
~
-2 Pu/U**
5.938 x 10 0.7 5.970 x 10
~
-4
~0
~
Np-237/U-238 1.14 x 10 9
0.86 x 10
-2
-2 Am-241/Pu-239 1.23 x 10 15 1.08 x 10 b 242/Pu-239 1.05 x 10 '
110 1.11 x 10-4
~
Cm-244/Pu-239 1.09 x 10 120 0.98 x 10 '
-4
~
t, 30 l
- Reported in Reference [4,3-l
- Weight ratio 4.3-59
ATTACHf..ENT /
STp FSAR ST.HL E l'/3A PAGE a_OF 6 'l
{
TABLE 4.3-12
-COMPARISON OF MEASURED AND CALCUIATED ROD WORTH 2-Loop Plant, 121 Assemblies, 10 foot core Measured (per)
Calculated (pcic)
Group B 1885 1893 Group A 1530 1649 Shutdown Group 3050 2917 ESADA-Critical *, 0.69" Pitch.
2 w/o PuC, 8% Pu 2
9 Control Rods 6.21" rod eeparation 2250 2250 2.07" rod separation 4220 4160
(
1.38" rod separation 4010 2010 i
l
- Reported in Reference [4.3-4.3-61
ATTACHMENT ST HL AE- / V PAGE LJ/ OF 19996 1
, W{EPLACE u)/ TN
(
A)Eu)
FK,u2E LA 2000 NOTE: HOT FULL POWER RODS OUT 1600 -
Ee
~
WITHOUT BURN ABLE W
POISONS S
8
$ 800 BURNUP DIFFERENCE g
1500 MWD /MTU 5
L 5;g __ WITH BURNABLE E
POISONS u
O O
2000 4000 6000 8000 10000 12000 14000 CORE AVERAGE BURNUP (MWD /MTU)
Amendment 26 SOUTH TEXAS PROJECT
- (
UNITS 1 & 2 Boron Concentration versus First Cycle Burn-up With and Without Burnable Poison Rods Figure 4.3 3
~
ATTACHMENT I 19986 i ST.HL.AE /Y3>
PAGE t/ LOF4//
2000 NOTE: HOT FULL POWER RODS OUT 1600 E
C, z
WITHOUT BURNABLE b
POISONS z
8
$ 800 BURNUP DIFFERENCE
{
8 900 MWD /MTU c
t
. WITH BURNABLE 400 -
E POISONS u
e i
i i
o 0
2000 4000 6000 8000 10000 12000 14000 CORE AVERAGE BURNUP (MWD /MTU)
Amendment 26 SOUTH TEXAS PROJECT UNITS 1 & 2 Boron Concentration versus First Cycle Burn up With and Without Burnable Poison Rods
~
Figure 4.3 3
' ATTACHMENT I ST-HL AE-if 3Adi PAGE 4hO -
g EPLAC E LDITH n
12,373. o J) ElO F160LE
(
R P
N M
L K
J H
~G F
E D
C 8
A 180o 1
4 4
4
/
$l
/9 2
9 24 "3
/
3 3
24 4
3 24 3
8 4
24 12 8
8
/
12 24
[
8 12 9
5 9
12 8
8 6
4 4
8 8
'8 8
4 4
j 7
24 8
8 J2 8
8 24 8
90 4
8 8
12, 12 8
8 4
270 9
24 8
8
,/
12 8
8 24
(
10 4
4 8
/
8 8
8 4
4 11 9
12 8
8 8
12 9
12 24 12' 8
8 12 24 "3
13 3
24
/
4 4
24 3
8 9
14 9
24 15 4
4 4
0 NUM8ER INDICATES NUMBER OF 8URNABLE POISON ROOS 946 8P ROOS S - INDICATES SOURCE ROD 12.5W/080 23
/
/
/
/
/
Amendment 18. 5/1/81
/
SOUTH TEXAS PROJECT UNITS 1 & 2
{
Figure 4.3-5.
Burnable Poison Loading Pattern
~. _ _ _. _ _ _
-~
ATTACHMENT /
hfE O g t/
12.373 to R
P
.N N
L K.
J H
G F
E D
C 8
A 180c I
4 4
4
$3 2
9 24 9
3 3
24 4
8 4
24 3
4 24 12 8
8 12 24 5
,3 12 8
45 8
8 12 9
6 4
4 8
8 8
8 4
4 7
24
.! 8 8
12 8
8 24 8
90 4
'8 8
12 12 8
8 4
270 9
24,
8 8
12 8
8 24 10 4
}'
4 8
8 8
8 4
4 il 9
12 8
8 45 8
12 9
12 24 12 8
g 12 24 13 3
' 24 4
8 4
24 3
ll 14 9
24 9
15 4
4 4
0:
NUMBER INDICATES N'.NSER OF BURNASLE POISON RODS 946 8P RODS S - INDICATES SOURCE ROD 12.5 W/0 B 0 2
" C"d'?nW. b/1/N
^
SOUTH TEXAS PROJECT UNITS 1 & 2 l
Figure 4.3-5.
~
Burnable Poison Loading Pattern
~
ATTACHMENT /
ST HL.AE 14.3x YAGE l/40F 5 tl 2.0 POWER :.5
- SI 0:.55 Ao
-2 6.03 i
POWER :.5 D
1.5 -
POWER 1.0 D =.I Ao : +8.20 A0 = -6.92 i
a:
Y
@ 1.0 -
a I.
-(
t I
0.5 I
I I
I I
I I
I I
O O
0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0 PERCENT OF ACTIVE CORE HEIGHT FROM BOTTOM LEGEND:
Amendment 1, 7/19/78 D
= FRACTIONAL INSERTION OF SOUTH TEXAS PROJECT A0 = AXtAL 0FFSET
{
Typical Axial Power Shapes Occuring
- (
at Middle of Life Figure 4.3-15.
1 s
ST-HL AE-14
>]
ATTACHMEN STP FSAR PAGE WOF 4 I
9"g,y = 9DNB, EU (4.4-5) i F
cnd 9DNB EU is the uniform DNB heat flux as predicted by the W-3 DNB correlation, Reference 4.4-10 all flow cell walle are heated.
F is the flux shape factor to account for nonuniform axial heat flux dis-tributions, Reference'4.4-10, with the "C" term modified as in Reference 4.4-3.
Fs is the modified spacer factor defined by Equation (4.4-1) in Subsection 4.4.2.2.1 and using an axial gridsspacing coefficient, Ks = 0.059, and a thermal diffusion coefficient (TDC) of 0.059, based on the 22 in. grid spacing data prev:,gusly described. Since the actual grid spacing is 19.8 in., the modified spact' factor is conservative since the DNB performance was found to improve and TDC 18 6 increases se axial grid spacing is decreased, References 4.4-8 and 4.4-12.
The TDC value for 20 in. grid spacing (approximately the same spacing as this design) is 0.061.
?
9,e is the actus1 local heat flux.
3 Tha DNB heat flux ratio as applied to this design when a cold wall is present ist DNBR = 9bNB,N CW x Fj (4.4-6) 9 oc 1
t whare:
9"NB,N,CW = 9 NB,EU Dh
- N (h.4-7)
F whare:
DNB,EU Dh is the uniform DNB heat flux as predicted by the W-3 cold wall DNB i
9 correlatien, Reference 4.4-3, when not all flow cell walls are heated (thimble cold wall cell).
CWFI4*4-33 = 1.0-Ru [13.76-1.372e *78*-4.732 ( G ) 0.0535 (4.4-8) l f
10
-0.0619 ( P 3 0.14 - 8.509Dh.107j 0
1000 s
cnd Ru =.1 - De/Dh IS dafined by Equation (4.4-1) in Subsection 4.4.2.2.1 is the same sa used for typical cell.
Valu2s of minimum DNB provided in Table 4.4-1 and 4.4-2 are the limiting valu 4---
obtained by applying the above two definitions of DNBR to the appropriate cell (typical cell with all walls' heated, or a thimble cold wall cell with a partial h20ted wall condition).
Tha procedures used in the evaluation of DNB margin for this application show th:t the calculated minimum DNBR for the peak rod or rods in the core will be chova 1.30'during Class I and II incidents, even when all the engineering hot l18 4.4-8 Amendment 18,5/1/81 l..
~ ;<.
STP FSAR ATTACHME 1 1 ST HL AE-I l18 1.
Pellet diameter, density and enrichment
'(
Design values employed in the THINC analysis related to the above fabrica-tion variations are based on applicable limiting tolerances such that these design values are met for 95 percent of the limiting channels at a 95 per-cent conf'ldence level. Measured manufacturing data on Westinghouse 17 x 17 fuel show the tolerances used in this evaluation are conservative. The 118 effect of variations in pellet diameter, enrichment and density is employed in the THINC analysis as a direct multiplier on the hot channel enthalpy l18 rise.
2.
Inlet Flow Ma1 distribution The consideration of inlet flow ma1 distribution in core thermal performances l
is discussed in Section 4.4.4.2.2.
A design basis of 5 percent reduction in coolant flow to the hot assembly is used in the THINC-IV analysis.
3.
Flow Redistribution The flow redistribution accounts for the reduction in flow in the hot channel resulting from the high flow resistance in the channel due to the local or bulk boiling. The effect of the non-uniform power distribution is l
inherently considered in the THINC analysis for every operating condition i
which is evaluated.
4.
Flow Mixing The subchannel mixing model incorporated in the THINC Code and used in
(
~
reactor design is based on experimental data (4.4-17) discussed in Section 4.4.4.5.1.
The mixing vanes incorporated in the spacer grid design induce additional flow mixing between the various flow channels in a fuel assembly as well as between adjacent assemblies. This mixing reduces the enthalpy rise in the hot channel resulting from 3ocal power peaking or unfavorable mechanical tolerances.
4.4.2.2.5 Effects of Rod Bow on DNBR: The phenomenon of fuel rod bowing, as described in Reference 4.4-84, must be accounted for in the DNBR safety analysis of Condition I and Condition II events for each plant application. Applicable generic credits for margin resulting from retained conservatism in the evaluation of DNBR and/or margin obtained from measured y
plant operating parameters (such as or core flow), which are less limiting than those required by the p ne safety analysis, can be used t 43 offset the effect of rod bow.
a i
l The safety analysis for South Texas cores maintained sufficient margin (3.3
- lti id'"tiii'd ""
l percent) to acconanodate full and low flow DNbR p!"/I scaling factor (I = fuel l
Reference 4.4-85 with the incorporation of the L j
rod bending moment of inertia, L = span length) to account for 17X17 XL span lengths. A design limit DNBR of 1.30 vs. 1.28, a grid spacing coefficient (K ) of.059 vs.
066, and a thermal diffusion coefficient (TDC) of.059 vs.
r
.0$1areexamplesofconservatismutilizedinthesafetyanalysis.
- f The maximum rod bow penalties accounted for in the design safety analysis
(
are bssed on an assembly average burnup of 33,000 mwd /NTU. At burnups af only (used for mcd$ied SfM2f bhor, s
4.4-11 Amendment 43
ATTACHMENT l
~
ST-HL-AE /41 A PAGE uf OF 64 STP FSAR i
greater than 33,000 mwd /MTU, credit is taken for the effect of F 43 burndown,duetothedecreaseinfissionableisotopesandthebubdupof
)-
fission product inventory, and no additional rod bow penalty is required.
4.4.2.3 Einear Heat Generation Rate. The core average and maximum Linear Powers are given in Table 4.4-1.
The method of determining the maximum Linear Powers is given in Section 4.3.2.2.
4.4.2.4 Void Fraction Distribution. The calculated core average and the hot subchannel maximum and average void fractions are presented in Table 4.4-3 for operation at full power with design hot channel factors. The void fraction distribution in the core at various radial and axial locations is presented in Reference (4.4-18). The void models used in the THINC-IV computer code are described in Section 4.4.2.7.3.
Normalized core flow and enthalpy rise distributions are shown on Figures 4.4-5 through 4.4-7.'
4.4.2.5 Core Coolant Flov Distribution. Assembly average coolant mass velocity and enthalpy at variots radial and axial core locations are given below. Coolant enthalpy rise and flow distributions are shown for the 1/3 core height elevation on Figure 4.4-5, and 2/3 core height elevation on Figure 4.4-6 and et the core exit. on Figure 4.4-7.
These distributions are for the full power conditions as given in Table 4.4-1 and for the radial power density distribution shown cn Figure 4.3-7.
The THINC Code analysis 3
for this case utilized a uniform core inlet enthalpy and inlet flow I
distribution. No orificing is employed in the reactor design.
4.4.2.6 Core Pressure Drops and Hydraulic Loads.
4.4.2.6.1 Core Pressure Drops: The analytical model and experimental data used to calculate the pressure drops shown in Table 4.4-1 are described in Section 4.4.2.7.
The core pressure drop includes the fuel assembly, lower core plate, and upper core plate pressure drops. The full power operation pressure drop values shown in Table 4.4-1 are the unrecoverable pressure drops across the vessel, including the inlet and outlet nozzles, and across the core. These pressure drops are based on the best estimat flow for actual plant operating conditions as described in Section 5.1.lhg h Sedh also defines and describes the thermal design flow (minimum flow) which is the basis for reactor core thermal performance and the mechanical design l
flow (maximum flow) which is used in the mechanical design of the reactor i
vessel internals and fuel assemblies. Since the best estimate flow is that l
flow which is most likely to exist in an operating plant, the calculated core pressure drops in Table 4.4-1 are based on this best estimate flow
]
rather than the thermal design flow.
i Uncertainties associated with the core pressure drop values are discussed in 1
Section 4.4.2.9.2.
4.4.2.6.2 Hydraulic Loads: The fuel assembly hold down springs, l18 Figure 4.2-2, are designed to keep the fuel assemblies in contact with the I
lower core plate under all Condition I and 11 events with the exception of the turbine overspeed transient associated with a loss of external load.
The hold down springs are designed to tolerate the possibility of an over deflection associated with fuel assembly liftoff for this case and provide contact between the fuel assembly and the lower core plate following this i
Amendment 43
ATTACHMENT ST HL AE-/(I3>/
Line lengths and sizes for the Safety Injection System (SIS) are deter-a mined so as to guarantee a total system resistance which will provide,
)
as a minimum, the fluid delivery rates asstaned in the safety analyses described in Chapter 15.
7.
The parameters for components of the RCS are presented in Section 5.4 component and subsystem design.
l 8.
He steady state pressure drops and temperature distributions through i
the RCS are presented in Table 5.1-1.
l 4.4.3.2 Operating Restrictions on Ptanps. The minimum net positive cuction head (NPSH) and minimum seal injection flow rate must be estab-lished before operating the reactor coolant pumps. With the minimum 6 gps (,.
labyrinth seal injection flow rate established,/the operator will have to verify that the system pressure satisfies NPSH' requirements.
4.4.3.3 Power-Flow Operating Map (BWR). Not applicable to STP.
~.
. 4.4.3.4 Temperature-Power Operating Map.
The relationship between RCS temperature and power is shown on Figure 4.4-21.
De effects of reduced core flow due to inoperative pumps is discussed in Sections 5.4.1, 15.2.5, and 15.3.4.
Natural circulation capability of the system is shown in Table 15.2-2.
l 4.4.3.5 Load Following Characteristics. D e RCS is designed on the i
basis of steady state operation at full power heat load.
H e reactor cool-ent ptsnps utilize cons. ant speed drives as described in Section 5.4 and the reactor power is controlled to maintain average coolant temperature at a valse which is a linear function of load, as described in Section 7.7.
i 4.4.3.6 Dermal and Hydraulic Characteristics Summary Table. Th j
thermal and hydraulic characteristics are given in Tables 4.3-1 4.4-1[e diid 4.4h 4.4.4 Evaluation j
4.4.4.1 Critical Heat Flux.
The critical heat flux correlation utilized in the core thermal ana. lysis is discussed in Section 4.4.2.
4.4.4.2 Core Hydraulics.
'f-4.4.4.2.1 Flow Paths Considered in Core Pressure Drop and Thennal Design: The following flow paths or core bypass flow are considered:
)
1.
Flow through the spray nozzles into the upper head for head cooling purposes.
2.
Flow entering into the rod cluster control guide thimbles to cool the control rods.
4.4-22 l
)
ATTACHMENT /
SIP FSAR ST HL AE W PAGE <fo OF54 ThewayinWichTfHisusedintheDSBcalculationisimportant.
The location of minimum DNBR depends on the axial profile and the value of DNBR depends on the enthalpy rise to that point. Basically, the maximun value of the rod integral is used to identify the most likely rod for mini-mum DNBR. An axial power profile is obtained dich den nonnalized to the design value of FfH recreates the axial heat flux along the limiting rod.
The surrounding rods are assumed to have the same axial profile with rod average powers Wich are typical distributions found in hot assemblies.
In this manner,vorst case axial profiles can be combined with worst case radial distributions for reference DNB calculations.
It shouldbenotedagainthatF{gisanintegralandisusedassuchin DNB calculations. Local heat fluxes are obtained by using hot channel and adjacent channel explicit power shapes Wich take into account variations in horizontal power shapes throughout the core. The sensitivity of the THINC-IV analysis to radial power shapes is discussed in Reference 4.4-18.
For operation at a fraction P of full power, the design FjH used is given by:
F$H=1.5211+0.3(1-P))
(4.4-19)
The permitted relaxation of F{g is included in the DNB protection setpoints and allows radial power shape changes with rod insertion to the insertion limits 14.4-66), thus allowing greater flexibility in the nuclear design.
f:
4.4.4.3.2 Axial Heat Flux Distributions: As discossed in Subsection 4.3.2.2, the axial heat flux distribution can vary as a result of rod motion, power change, or due to spatial menon transients Wich may occur in the axial direction. Consequently it is necessary to measure the axial power imbalance by means of the excore nuclear detectors (as discussed in Subsection 4.3.2.2.7) and protect the co.re from excessive axial power imbalance 4 The reactor trip system provides automatic reduction of the trip setpcint in the overtemperature AT channels on excessive axial power imbalance; that is, den an extremely large axial offset corresponds to an axial shape Wich could lead to a DNBR W ich is less than that calculated for the reference DNB design axial shape.
i ITne n rmal re rence DNB desi axial shape is either a chopped cosine shap a pe to av rage v lue of 1. 5 or a kewed- -the op a pe, d endin upo whi is no e cons rvative each plicat on.
e ref rence NB axi sh Ref / c e-us a in e ablish ng core. a 1121. (i.,e.
e3, npera ure 4 protee on s d accid nts p--L $5. g ed cosine s
-c pe vi a pea s tpoint and C ndition e valu EtHght gre m
--t +
xcepti/ns t this a fA> <
o aver coas ng dos free
, a sidgle oppedfull loss o flow dth pupp(
bN lengt contr rod, a stati 11y mi align d full engt contro rod whic are i itiated om no al ful power operat n.
ince th re at k y -/
few axia power s apes W ch giv DNBR' less an is she as c par d to e refer ce DNB esign xial hape, greater axial over i balane can be allowed and, thus, increases plant operating flexibility.
4.4-25 Amendment A
e o
ATTACHME ST.HL.AE /VS PAGE 6/ OF TV INSERT 4.4-1 The nomal reference DNB design axial shape is either a chopped cosine shape with a peak to average value of 1.S5 or a skewed-to-the-to'p shape, depending upon which is more conservative in each application. The reference DNB axial shapeusedinestablishingcoreDNBlimits(i.e.,overtemperatureaTprotec-tion system setpoints) and Condition II accidents for the South Texas plants is a chopped cosine shape with a peak to average value of 1.61. With respect to minimum DNBR, this axial shape bounds all of the shapes which could occur during power operation including overpower conditions as generated for the nucleardesign(refertoSection4.3.2.2.6). Accidents which are initiated from nonnal full power operation including loss of flow with pump (s) coasting down freely, a single dropped control rod, and a statically misaligned control rod are analyzed with a 1.55 chopped cosine axial shape because this shape bounds all of the possible shapes which could occur at nomal full power operating conditions.
.l e
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. ~.,,. - - -
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ATTACHMENT /
PAGE % CF S Qs (loil Ck*fff*I C' S,.MJ t
To determane the penalty to be taken in protect n set points for extreme values of flux dif fere'nce, Aia, reference shape is supplemented by other axial shapes skewed to the bottom and top of the core.
The course of those accidents in which DNB is a concern is analyzed in Chapter 15 assuming that the protection set points have been set on the basis of these shapes.
In many cases the axial power distribution in the hot channel changes throughout the course of the accident due to rod motion, coolant temperature and power level changes.
The initial conditions for the accidents for which DNB protection is required are assumed to be those permissible within the constant axial offset control strategy for the load maneuvers described in Reference 4.4-67.
In the case of the loss of flow accident the het chacral heat flux profile is very similar to the power density profile in normal operation preceding the accident.
It is therefore possible to illustrate the calculated minimum DNB ratio for conditions representative of the loss of flow accident as a function of the flux difference initially in the core.
A plot of this type is provided on Figure 4.4-10 for first core initial conditions.
As noted on this figure, all power shapes were evaluated with a, full power radial peaking factor (
) of 1.52.
The radial contribution to the hot rod power sha is conservative both for the initial condition and for the condition at the tir.e of einirui DNBE during the loss of flow transient.
Also shown is the sinimus DKBR calculated for the reference, power shape at the same conditions.
U& $4]c L O'" " W **"O j
4.4.4.4
_ Core Thermal Response.
A general summary of the steady-state thermal-hydraulic design parameters including thermal output, flow rates, etc., is provided in Table 4.4-1.
As stated in Section 4.4-1, the design bases of the application are to prevent DNB and to prevent fuel melting for Condition 1 and 11 events.
The protective systems described in Chapter 7 are designed to meet these bases.
The response of,the core to Condition 11 transients is given in Chapter 15.
4.4.4.5' Analytical Technicues.
.4.4.4.5.1 Core Analysis:
The objective of reactor core thermal design is to determine the maximum heat removal capability in all flow subchannels and to show that the core safety licits (as will be presented in the Technical Specifications) are not exceeded while compounding engineering and nuclear effects.
Tne thermal design considen le:a1 variations in dimensions, power generation, flow redistributien, and mixing.
THINC-lv is a realistic three-dimensional matrix e.odel which has been developed to account for hydraulic and nuclear effects en the enthalpy rise in the core.
(References 4.4-18 and 4.4-49) The behavior of the hot assembly is determined by. superimposing the power distribution among the assemblies upon the inlet flow distribution while allowing for flow mixing and flow distribution between assemblies.
The average flov and enthalpy in the hottest assembly is obtained from the core-wide, asseib.ly by assembly analysis.
The local variations in power, fuel red and pellet fabrication, and mixing within the hottest ae.embly are then superimpesed on the 4verage conditions of the hottest assembly in order to determine the conlitions in the hot channel.
4.4-26
/.4ndrint 27
ATTACHMENT I ST-HL AE / W>
STP FSAR PAGE 5 30F5 y 4.4-71.
Weismann, J., Wenzel, A. R., Tong, L.S., Fitzsinunens, D., Thorne,
W. and Batch, J., " Experimental Determination of the Departure i
from Nucleate Boiling in Large Rod Bundles at high Pressures,"
Chem. Eng. Prog. Symp. Ser. 64, No. 82,114-125 (1968).
4.4-72.
Boure, J. A., Bergies, A. E. and Tong, L.S., " Review of Two-Phase Flow Instability," Nucl. Engr. Design 25 (1973) p. 155-192.
4.4-73.
Lahey, R. T., and Moody, F.
J., "The Thermal Hydraulics of e Boiling Water Reactor," American Nuclear Society, 1977.
4.4-74.
Saha P.,
Ishii, M., and Zuber N., "An Experimental Investigation of the Thermally Induced Flow Oscillations in Two-Phase Systems,"
. of Heat Transfer, Nov. 1976, pp. 616-622, 4.4-75.
er6Fi FSAR, Docket #50-395.
6 u
4.4-76.
Byron /Braidwood FSAR,.)ocket #50-456.
~.
4.4-77., Comanche Peak FSAR, Docket #50-445.
4.4-78.
Kakac, S., Veziroglu, T. N., Akyuziu, K., Berkol, O., " Sustained and Tranient Boiling Flov Instabilities in a Cross-Connected Four-Parallel-Channel Upflow System," Proc. of 5th International Heat Transfer Conference, Tokyo, Sept. 3-7, 1974.
4.4-79.
Kao, H.
S., Morgan, C.
D.,
and Parker, W.
B., " Prediction of Flow
~
Oscillation in Reactor Core Channel," Trans. ANS, Vol. 16, 1973, pp. 212-213.
4.4-10.
(This reference has beeen deleted.)
4 l
4.4-&1.
- Ohtsubo, A., and Uruwashi, S., " Stagnant Fluid due to Local Flow Blockage," J. Nucl. Sci. Technol., h, No. 7, 433-434, (1972).
4.4-82.
- Basamer, P., Kirsch, D. and Schultheiss, G.
F., " Investigation of the Flow Pattern in the Recirculation Zone Downstream of Local Coolant Blockages in Pin Bundles," Atomwirtschaft, 17, No. 8, 416-417, (1972).
(In German).
4.4-83.
Burke, T.
M., Meyer, C. E. and Shefcheck J., " Analysis of Data from the Zion (Unit 1) MIINC Verification Test," WCAP-8453 (Proprietary) December, 1974 and WCAP-8454, December, 1974.
4.4-84.
- Skaritka, J., (Ed.), " Fuel Rod Bow Evaluation " WCAP-8691, Rev.1 (Proprietary) July 1979.
4.4-85.-
" Partial Response to Request Number 1 for Additional Information on WCAP-8691, Revision 1" Letter, E. P. Rahe, Jr., (Westinghouse) to 43 J. R. Miller (NRC), Nf,-EPR-2515, dated October 9, 1981; " Remaining Response to Request Number 1 for Additional Information on WCAP-8691, Revision 1" Letter, E. P. Rahe, Jr., (Westinghouse) to J. R. Miller (NRC) NS-EPR-2572, dated March 16, 1982.
4.4-41 Amendment 43
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