ML17223A515
| ML17223A515 | |
| Person / Time | |
|---|---|
| Site: | Saint Lucie  |
| Issue date: | 02/07/1990 |
| From: | FLORIDA POWER & LIGHT CO. |
| To: | |
| Shared Package | |
| ML17223A512 | List: |
| References | |
| NUDOCS 9002150297 | |
| Download: ML17223A515 (107) | |
Text
ATTACHMENT 4 PRESSURE-TEMPERATURE LIMITS And LOW TEMPERATURE OVERPRESSURE PROTECTION For ST ~ LUCIE UNIT 2 For 15 EFFECTIVE FULL POWER YEARS 9002150297 900207 PDR ADOCK 05000389 P'DC
h "P,41 Ig4 t
619(86L4) ch-2 TABLE OF CONTENTS SECTION
1.0 INTRODUCTION
TITLE PAGE 2.0 2.1 2.2 2.3 2;4 2.5 2.6 2.7 2.8 3.0 3.1 3.2 3.3 3.4 3.5 4.0 PRESSURE -
TEMPERATURE LIMITS ADJUSTED REFERENCE TEMPERATURE PROJECTIONS GENERAL APPROACH FOR CALCULATING PRESSURE-TEMPERATURE LIMITS THERMAL ANALYSIS METHODOLOGY COOLDOWN LIMIT ANALYSIS HEATUP LIMIT ANALYSIS HYDROSTATIC TEST AND CORE CRITICAL LIMIT ANALYSIS LOWEST SERVICE TEMPERATURE, MINIMUM BOLTUP TEMPERATURE, AND MINIMUM PRESSURE LIMITS DATA LOW TEMPERATURE OVERPRESSURE PROTECTION GENERAL METHOD AND ASSUMPTIONS PRESSURE TRANSIENT ANALYSES 3.3. 1 Ener Addition Transients 3.3.2 Mass Addition Transients 3.3.3 Controllin Pressures LIMITING CONDITIONS FOR OPERATION
SUMMARY
OF PROPOSED CHANGES REFERENCES 15 17 18 21 22 24 25 25 26 28 28 30 31 32 32 33 Page 2
I,
619(86L4) ch-3 LIST OF TABLES NO.
TITLE St. Lucie Unit 2 Reactor Vessel Beltline Materials St. Lucie Unit 2 Controlling Materials and Their Adjusted Reference Temperatures St. Lucie Unit 2 Cooldown and Heatup Pressure-Temperature Limit Data, 10'F/hr to 50'F/hr and Isothermal St. Lucie Unit 2 Cooldown and Heatup Pressure-Temperature Limit Data, 60'F/hr to 100'F/hr St. Lucie Unit 2 Hydrostatic Test Pressure-Temperature Limit Data Maximum Transient Pressures Summary of Controlling Pressures LTOP Requirements, 15 EFPY PAGE 35 36
'7 38 39 40 41 42 Page 3
(l
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619(86L4) ch-4 LIST OF FIGURES NO.
TITLE St. Lucie Unit 2 P-T Limits 15 EFPY Cooldown St. Lucie Unit 2 P-T Limits 15 EFPY Cooldown St. Lucie Unit 2 P-T Limits '15 EFPY Cooldown St. Lucie Unit 2 P-T Limits 15 EFPY Cooldown St. Lucie Unit 2 P-T Limits 15 EFPY Cooldown St. Lucie Unit 2 P-T Limits 15 EFPY Heatup PAGE 43 44 45 46 47 48 St. Lucie Unit 2 P-T Limits 15 EFPY St. Lucie Unit 2 P-T Limits 15 EFPY Heatup Heatup 49 50 10 12 13 St. Lucie Unit 2 P-T Limits 15 EFPY Heatup St. Lucie Unit 2 P-T Limits 15 EFPY Hydrostatic and Core Critical Operation St. Lucie-2 RCP Start Transient, W/SDCRV, P set
= 350 psia St. Lucie-2 RCP Start Transient, Nominal PORV Setpoint 470 psia, Primary Temp.
= 290'F St. Lucie Unit 2, 15 EFPY Maximum Allowable Cooldown Rate 51 52 53 54 55 Page 4
ft I
619(86L4) ch-5 INTRODUCTION The following sections describe the basis for development of reactor vessel beltline pressure-temperature limitations and Iow temperature overpressure protection requirements for the St. Lucie Unit 2 Nuclear Generating Station.
These limits are calculated to meet the regulations of 10 CFR Part 50 Appendix A, Design Criterion 14 and (1)
Design Criterion 31.
These design criteria required that the reactor coolant pressure boundary be designed, fabricated,
- erected, and tested in order to have an extremely low probability of abnormal leakage, of rapid failure, and of gross rupture.
The criteria also require that the reactor coolant pressure boundary be designed with sufficient margin to assure that when stressed under operating, maintenance, and testing the boundary behaves in.a non-brittle manner and the probability of rapidly propagating fracture is minimized.
The pressure-temperature limits are developed using the requirements of 10 CFR 50 Appendix G
This appendix describes the requirements for developing the pressure-temperature limits and provides the general basis for these limitations.
The margins of safety against fracture provided by the pressure-temperature limits using the requirements of 10 CFR Part 50 Appendix G are equivalent to those recommended in the ASNE Boiler and Pressure Vessel Code Section III, Appendix G, "Protection Against Nonductile Failure."
The general guidance provided in those procedures has been utilized to develop the St. Lucie Unit 2 pressure-temperature limits with the requisite margins of safety for the heatup and cooldown conditions.
The Reactor Pressure Vessel beltline pressure-temperature limits are based upon the irradiation damage prediction methods of Regulatory Guide 1.99 Revision 02 This methodology has been used to calculate the limiting material Adjusted Reference Temperatures for St. Lucie Unit 2 and have utilized fluence values for 15 Effective Full Power Years (EFPY) assuming an 24 month fuel cycle beginning after 6.01 EFPY of.operation.
Page 5
lf<
619(86L4) ch-6 This report provides reactor vessel beltline pressure-temperature limits in accordance with 10 CFR 50 Appendix G for 15 EFPY.
The events analyzed are the isothermal, 10 through 100'F/hr cooldown conditions and the 10 through 100'F/hr heatup conditions.
These conditions were analyzed to provide a data base of reactor vessel P-T limits for use in establishing Low Temperature Overpressure Protection requirements.
Low Temperature Over pressure Protection (LTOP) requirements are established based upon the guidance provided in USNRC Standard Review Plan (SRP) 5.2.2.
Using this guidance the limiting (5) transient pressures have been determined for mass and energy addition-transients to establish the appropriate LTOP setpoints, heatup and cooldown rates, and administrative requirements.
Based upon the P-T limit analyses and LTOP requi rements provided within this report, no limiting vessel operability issues are anticipated to exist.
2.0 PRESSURE - TEMPERATURE LIMITS 2.1 ADJUSTED REFERENCE TEMPERATURE PROJECTIONS In order to develop pressure-temperature limits over the design life of the reactor vessel, adjusted reference temperatures (ART) for the controlling beltline material need to be determined.
The adjusted reference temperatures of reactor vessel beltline materials for St.
Lucie Unit 2 have been calculated at the I/4t and 3/4t locations after 15 EFPY operation.
By comparing ART data for each material, the controlling material for St. Lucie Unit 2 has been determined.
Page 6
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- CII, K~I
'4
619(86L4) ch-7 The adjusted reference temperatures (ART) have been calculated using the procedures in Regulatory Position 1.1 of Regulatory Guide 1.99 Revision 02 The calculative procedure for the ART values for each material in the beltline is given by the following expression:
ART = Initial RTNpT + <RTNpT + Margin Initial RTNpT is the reference temperature for the unirradiated material.
4RTNpT is the mean value of the adjustment in the reference temperature caused by irradiation and is given by the following expression:
(Cf) f(0.28 - 0. 10 log f)
NPT CF is the chemistry factor for the beltline materials which is a
function of residual element content, i.e., weight percent copper and nickel.
Regulatory Guide 1.99 Revision 02 provides values for the chemistry factors for wel'ds and for base metal plates and forgings.
The term f is the neutron fluence at any depth in the vessel.
The neutron fluence at any depth is given by the following expression:
'(
-0.24x) surf The term f f is the calculated value of the neutron fluence (10 n/cm, E>
1MeV) at the inner wetted surface of the vessel at the location of the postulated defect (1/4t or 3/4t), and x is the depth into the vessel wall from the inner wetted surface in inches.
Margin is the quantity that is added to obtain a conservative upper bound value of ART.
The margin term is given by the following expression:
Margin
~ 2/
+
oI cr Page 7
(*
I ~
619(86L4) ch-8 The terms a>
and a
represent the standard deviation for initial RT T and the standard deviation of the mean value for reference NDT temperature shift.
The following information provides the basis for the calculated ART values for St. Lucie Unit 2:
1.
Material data were obtained from Reference 6, including copper content, nickel content and initial reference temperature
( RTNDT )
These data are summarized in Tabl e 1 for St.
Lucie Unit 2.
2.
Peak neutron fluence for the Unit 2 beltline region was determined to be 1.826 x 10 n/cm (E>1 MeV) at 15 EFPY based 19 2
on an 24 month fuel cycle (Reference 7).
3.
Shell course minimum reference thickness is 8.625 in. for both the lower and intermediate shell of both units (References 8,
9, and 10).
4.
Calculations were based on the procedures in Regulatory Position 1.1 of NRC Regulatory Guide 1.99, Rev.
2 (Reference 4).
Uncertainty in initial RTNDT was taken as O'F for measured values and 17'F for welds without measured values (Reference 11).
Adjusted reference temperatures for all beltline materials at the I/4t and 3/4t locations after 15 EFPY were calculated using Regulatory Guide 1.99 Revision 02 and the results of the calculation are listed in Table 1 for St. Lucie Unit 2.
The controlling materials are shown in Table 2; the term "controlling" means having the highest ART for a given time and position within the vessel wall.
The highest, or limiting, ARTs are then used to develop the pressure-temperature limits for the corresponding time period.
I Page 8
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1 Is%
'I
619(86L4) ch-9 In the case of St. Lucie Unit 2, the intermediate shell plate M-605-1 is controlling at the I/4t and 3/4t locations after 15 EFPY based on the predicted ART values of 140'F and 119'F respectively.
According to Position 1.1 of Regulatory Guide 1.99, Revision 2 (4)
'he uncertainty in the value, of initial RT T is to be estimated from the precision of test method when a "measured" value of initial RTNDT i s ava i 1 able.
RTNDT is derived in accordance with NB2300 of the ASME Boiler and Pressure Vessel Code,Section III. It involves both a series of drop weight (ASTM E208) and Charpy impact (ASTM E23) tests on the material.
The RTNDT resulting from this two test method evaluation is conservatively biased.
The elements of this conservatism include:
I)
Choice for RTNDT is the higher of NDT or TCV -60'F.
The drop-weight test is performed to obtain NDT and a full Charpy impact curve is developed to obtain TCV for a given material.
The combination of the two test methods gives protection against the possibility of errors in conducting either test and, with the full Charpy curve, demonstrates that: the material is typical of reactor pressure vessel steel.
Choice of the more conservative of the two (i.e., the higher of NDTT or TCV-60'F) assures that tests at temperatures above the reference temperature will yield increasing values of toughness, and verifies the temperature dependence of the fracture toughness implicit in the KIR curve (ASME Code,Section III, Appendix G).
2)
Selection of the most adverse Charpy results for TCV.
In accordance with NB2300, a temperature, TCV, is established at which three Charpy specimens exhibit at least 35 mils lateral expansion and not less than 50 ft-lb absorbed energy.
The three specimens will typically exhibit a range of lateral expansion and absorbed energy consistent with the variables Page 9
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f1!
l IJ7e
>S(87V>) dd-'-0 inherent in the test:
Specimen temperature, testing equipment,
- operator, and test specimen (e.g.,
dimensional tolerance and mat rial homogeneity).
All of these variables are controlled using process and procedural controls, calibration and operator training, and they
=
are conservatively bounded by using the lowest measurement of the three specimens.
Furthermore, two related criteria are used, lateral expansion and absorbed
- energy, where consistency between the two measurements provides further assurance that they are realistic and the material will exhibit tl-e intended strength, ductility and toughness implicit in the KIR curve.
3)
Inherent conservatism in the protocol used in performing the drop-weight test.
The drop-weight test procedure was carefully designed to assure attainment of explicit values of deflection and stress concentration, eliminating a specific need to account for below nominal test conditions and thereby guaranteeing a conservative direction of these uncertainty components.
In addition, the test protocol calls for decreasing temperature until the first failure is encountered, followed by increasing the test temperature 10'F above the point where the last failure is encountered.
This in fact assures that one has biased the resulting estimate toward a low failure probability region of the temperature versus failure rate function diagrammed below.
The effect of this protocol is to conservatively accommodate the integrated uncertainty ccmpone
.s.
- Sc'entail
=:-:-'-::"" -- ='QjjjQR4zQ:,Effects
~ ~
~
r'.
~
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Page 10
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619(86L4) ch-11 Given the three elements of conservatism described
- above, values of initial RTNpT obtained in accordance with NB2300 will result in a conservative measure of the reference temperature.
The conservative bias of the NB2300 methodology and the drop-weight test protocol essentially eliminate the uncertainty which might result from the precision of an individual drop-weight or Charpy impact test.
Therefore, when measured values of RT T are available, the estimate of uncertainty in initial RTNpT is taken as zero.
2.2 GENERAL APPROACH FOR CALCULATING PRESSURE-TEMPERATURE LIMITS The analytical procedure for developing reactor vessel pressure-temperature limits utilizes the methods of Linear Elastic Fracture Mechanics (LEFM) found in the ASME Boiler and Pressure Vessel Code Section III, Appendix G (Reference
- 3) in accordance with the requirements of 10 CFR Part 50 Appendix G (Reference 2).
For these
- analyses, the Mode I (opening mode) stress intensity factors are used for the solution basis.
The general method utilizes Linear Elastic Fracture Mechanics procedures.
Linear Elastic Fracture Mechanics relates the size of a flaw with the allowable loading which precludes crack initiation.
This relation is based upon a
mathematical stress analysis of the beltline material fracture toughness properties as prescribed in Appendix G to Section III of the ASME Code.
The reactor vessel beltline region is analyzed assuming a
semi-elliptical surface flaw oriented in the axial direction with a depth of one quarter of the reactor vessel beltline thickness and an aspect ratio of one to six.
This postulated flaw is analyzed at both the inside diameter location (referred to as the 1/4t location) and the outside diameter location (referred to as the 3/4t location) to assure the most limiting condition is achieved.
The above flaw geometry and orientation is the maximum postulated defect size (reference flaw) described in Appendix G to Section III of the ASME Code.
Page 11
Nt+
619(86L4) ch-12 At each of the postulated flaw locations the Node I stress intensity factor, KI, produced by each of the specified loadings is calculated and the summation of the KI values is compared to a reference stress intensity, KIR, which is the critical value of KI for the material and temperature involved.
The result of this method is a relation of pressure versus temperature for each reactor vessel operating limits which preclude brittle fracture.
KIR is obtained from a reference fracture toughness curve for low alloy reactor pressure vessel steels as defined in Appendix G to Section III of the ASME Code.
This governing curve is defined by the following expression:
KIR = 26.78
+ 1.223 e
L.0145(T-ART + 160)]
- where, IR reference stress intensity factor, Ksi din T
temperature at the postulated crack tip, 'F ART adjusted reference nil ductility temperature at the postulated crack tip, 'F For any instant during the postulated heatup or cooldown, KIR is calculated at the metal temperature at the tip of the flaw, and the value of adjusted reference temperature at that flaw location.
Also for any instant during the heatup or cooldown the temperature gradients across the reactor vessel wall are calculated (see Section 2.3) and the corresponding thermal stress intensity factor, KIT, is determined.
Through the use of superposition, the thermal stress intensity is subtracted from the available KIR to determine the allowable pressure stress intensity factor and consequently the allowable pressure.
Page 12
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~
F sr p'q iP.
', a/
619(86L4)ch-13 In accordance with the ASME Code Section III Appendix G
requirements, the general equations for determining the allowable pressure for any assumed rate of temperature change during Service Level A and 8 operation are:
2KIM KIT KI R
- 1. 5K
+
K K
( Inserv ice Hydrostatic Test)
- where, K M
=
Allowable pressure stress intensity factor, Ksi/in IM K T
=
Thermal stress intensity factor, Ksi/in IT KIR
=
Reference stress intensity, Ksi/in The pressure-temperature limits provided in this report account for the temperature differential between the reactor vessel base metal and the reactor coolant bulk fluid temperature.
Correction for elevation and RCS flow induced pressure differences between the reactor vessel beltline and pressurizer, are included in the development of the pressure-temperature limits.
Consequently, the P-T limits are provided on coordinates of pressurizer pressure versus indicated RCS temperature.
The pressure correction factors calculated by Combustion Engineering are based upon the differential pressure due to the elevation difference between the reactor vessel wall adjacent to the bottom of the active core region, and the pressurizer pressure instrument nozzle.
This term of the pressure correction factor is equal to 26.6 psi.
The pressure correction factors are also based upon flow induced pressure drops across the reactor core through the hot leg pipe up to the surge line nozzle.
This term of the pressure correction factor has two values which are dependent upon the Reactor Coolant Pump (RCP) combination utilized Page 13
619(86L4) ch-14 during operation.
At temperatures of T
< 200'F, the flow induced c
pressure drop is based upon the RCS flow rates resulting from two operating RCPs and is equal to 36.1 psi.
At temperatures of T 200'F, the flow induced pressure drop is based upon the RCS flow rates resulting from three operating RCPs and is equal to 46.7 psi.
Consequently, two pressure correction factors are utilized in correcting the reactor vessel beltline region pressure to pressurizer pressure depending upon the cold leg'emperature.
The following pressure correction factors have been utilized:
PRESSURE CORRECTION FACTOR (PSI)
> 200'F
< 200'F 74 psi 63 psi Sy explicitly accounting for the temperature differential between the flow tip base metal temperature and the reactor coo1ant bulk fluid temperature, and the pressure differential between the beltline region of the reactor vessel and the pressurizer pressure measurement
- nozzle, the P-T limits are correctly represented on coordinates of pressurizer pressure and cold leg temperature.
Instrument uncertainties have not been included in the pressure-temperature limits.
These uncertainties which are on the order of 20 to 42 psi, and 10'F are insignificant when compared to the margin terms included in the ASIDE Section III Appendix G
methods.
Specifically, the pressure stress is multiplied by a factor of two resulting in conservative stress intensity values.
For example, a P-T limit which shows an allowable internal pressure of 400 psi actually is based upon a stress associated with an internal pressure of 800 psi.
In addition, the use of a lower bound allowable stress intensity, KIR which is shifted in accordance with Regulatory Guide 1.99 Revision 02 methods (i.e.,
2a margin on mean predicted shift) ensures a conservative measure of allowable stress Page 14
t
619(86L4) ch-15 intensity as a function of temperature, in the P-T limit computations.
Sased upon telephone conversations between FPSL and the NRC staff, the instrument uncertainties (errors) are insignificant relative to the conservatisms of stress intensity factors.
2.3 THERMAL ANALYSIS METHODOLOGY The Mode I thermal stress intensity factor is obtained through a
detailed thermal analysis of the reactor vessel beltline wall using a computer code.
In this code a one dimensional finite element radial conduction-convection heat transfer analysis is performed.
The vessel wall is divided into 10 elements and an accurate distribution of temperature as a function of radial location and transient time is calculated.
The code utilizes convective boundary condition on the inside wall of the vessel.
Variation of material properties through the vessel wall are permitted allowing for the change in material thermal properties between the cladding and the base metal.
In general, the temperature distribution through the reactor vessel wall is governed by a partial differential equation, aT a2T 1
aT pC~=K ~
+
subject to the following boundary conditions at the inside and outside wall surface locations:
At r = ri
-K~
~
h (T-T) aTr c
Atr r0 aTr 0
Page 15
t J 4
i Vp>
619(86L4) ch-16
- where, P
C K
T t
h T
re r density, lb/ft specific heat, btu/lb-'F thermal conductivity, btu/hr-ft-'F vessel wall temperature,
'F radius, ft time, hr convective heat transfer coefficient, btu/hr-ft -'F RCS coolant temperature,
'F inside and outside radii of vessel wall, ft The above is solved numerically using a finite element model to determine wall temperature as a function of radius, time, and thermal rate.
Thermal stress intensity factors are determined by the calculated temperature difference through the beltline wall using thermal influence coefficients specifically generated for this purpose.
The influence coefficients depend upon geometrical parameters associated with the maximum postulated defect, and the geometry of the reactor vessel beltline region (i.e., r /r,, a/c, a/t), along with the assumed unit loading.
The thermal stress intensity factors are determined by the temperature difference and temperature profile through the beltline wall using thermal influence coefficients and superposition.
ASIDE III Appendix G recognizes the limitations of the method it provides for calculating KIT because of the assumed temperature profile.
Since a detailed heat transfer analysis results in varying temperature profiles (and consequently varying thermal stresses),
an alternate method for calculating KIT was employed as required by Article G-2214.3 of Reference 3.
The alternate method employed used a polynomial fit of the temperature profile and superposition using influence coefficients to calculate KIT.
The influence coefficients were calculated using a 2-dimensional finite element model of the reactor vessel.
The influence coefficients were corrected for 3 dimensional effects using ASIDE Section XI Appe'ndix A procedures (Reference 13).
Page 16
i7 f.j
619(86L4) ch-17 2.4 COOLOOWN LIMIT ANALYSIS During cooldown, membrane and thermal bending stresses act together in tension at the reactor vessel inside wall.
This results in the pressure stress intensity factor, KIM, and the thermal stress intensity factor, KIT, acting in unison creating a high stress intensity.
At the reactor vessel outside wall the tensile pressure stress and the compressive thermal stress act in opposition resulting in a lower total stress than at the inside wall location.
Also neutron embrittlement, the shift in RTNOT and the associated reduction in fracture toughness are less severe at the outside wall compared to the inside wall location.
Consequently, the inside flaw location is more limiting and is analyzed for the cooldown event.
Utilizing the material metal temperature and adjusted reference temperature at the 1/4t location, the reference stress intensity is determined.
From the method provided in Section 2.3, the through wall temperature gradient is. calculated for the assumed cooldown rate to determine the thermal stress intensity factor.
In general, the thermal stress intensity factors are found using the temperature difference through the wall as a function of transient time as described in Section 2.3.
They are then subtracted from the available KIR value to find the allowable pressure stress intensity factor and consequently the allowable pressure.
The cooldown pressure-temperature curves are thus generated by calculating the allowable pressure on the reference flaw at the 1/4t location based upon KIR K
I Page 17
a f.p
(~
A'4'
619 (86L4) ch-18
- where, KI>
=
Allowable pressure stress intensity as a function of coolant temperature, Ksi/in KIR
=
Reference stress intensity as a function of coolant temperature, Ksi~in KIT
=
Thermal stress intensi ty as a function of coolant temperature, Ksi/in To develop a composite pressure-temperature limit for the cooldown
- event, the isothermal pressure-temperature limit must be calculated.
The isothermal pressure-temperature limit is then compared to the pressure-temperature limit associated with a cooling rate and the more restrictive allowable pressure-temperature limit is chosen resulting in a composite limit curve for the reactor vessel beltline.
Tables 3 and 4 provide the results for the isothermal, 10'F/hr through 100'F/hr cooldown pressure-temperature limits.
These tables provide the allowable pressure versus reactor coolant temperature for the various cooldown conditions.
The allowable pressure, is in units of Ksi while the temperature is in units, of 'F.
Figures 1, 2, 3,
4 and 5 provide a graphical presentation of the cooldown pressure-temperature limits found in Tables 3 and 4.
It is permissible to linearly interpolate between the cooldown pressure-temperature limits.
2.5 HEATUP LIMIT ANALYSIS During a heatup transient, the thermal bending stress is compressive at the reactor vessel inside wall and is tensile at the reactor 3
vessel outside wall.
Page 18
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619{86L4)ch-19 Internal pressure creates a tensile stress at the inside waII as well as the outside wall locations.
Consequently, the outside wall location has the larger total stress when compared to the inside wall.
However, neutron embrittlement (the shift in material RTNpT and the associated reduction in fracture toughness) is greater at the inside location than the outside.
Therefore, both the inside and outside flaw locations must be analyzed to assure that the most limiting condition is achieved.
As described in the cooldown case, the reference stress intensity factor is calculated at the metal temperature at the tip of the flaw and the adjusted reference temperature at the flaw location.
For heatup the reference stress intensity is calculated for both the 1/4t and 3/4t locations.
Using the finite element method described in Section 2.3, the temperature profile through the wall and the metal temperatures at the tip of the flaw are calculated for the transient history.
This information is used to calculate the thermal stress intensity factor at the 1/4t and 3/4t locations using the calculated wall gradient and thermal influence coefficients.
The allowable pressure stress intensity is then determined by superposition of the thermal stress intensity factor with the available reference stress intensity at the flaw tip.
The allowable pressure is then derived from the calculated allowable pressure stress intensity factor.
It is interesting to note that a sign change occurs in the thermal stress through the reactor vessel beltline wall.
Assuming a
reference flaw at the 1/4t location the thermal stress tends to alleviate the pressure stress indicating the isothermal steady state condition would represent the limiting P-T limit.
However, the isothermal condition may not always provide the limiting pressure-temperature limit for the 1/4t location during a heatup transient.
This is due to the correction of the base metal temperature to the Reactor Coolant System (RCS) fluid temperature at the inside wall by accounting for clad and film temperature differentials.
Page 19
'4 id 0
t" 1'h C ~
619(86L4) ch-20 For a given heatup rate (non-isothermal),
the differential tempera-ture through the clad and film increases as a function of thermal rate resulting in a higher RCS fluid temperature at the inside wall than the isothermal condition for the same flaw tip temperature and pressure.
Therefore to ensure the accurate representation of the 1/4t pressure-temperature limit during heatup, both the isothermal and heatup rate dependent pressure-temperature limits are calculated to ensure the limiting condition was achieved.
These limits account for clad and film differential temperatures and for the gradual buildup of wall differential temperatures with time, as do the cooldown limits.
At the 3/4t location the pressure stress and thermal stresses are tensile resulting in the maximum stress at that location.
Pressure-temperature limits were calculated for the 3/4t location accounting for clad and film differential temperature and the buildup of wall temperature gradients with time using the method described in Section 2.3.
The allowable pressure was derived based upon a flaw at the 3/4t location by superposition of the thermal stress intensity with the available reference stress intensity for the metal temperature and adjusted reference temperature at that position.
To develop composite pressure-temperature limits for the heatup transient, the isothermal, 1/4t heatup, and 3/4t heatup pressure-temperature limits are compared for a given thermal rate.
Then the most restrictive pressure-temperature limits are combined over the complete temperature interval resulting in a composite limit curve for the reactor vessel beltline for the heatup event.
Tables 3 and 4 provide the results for the 10'F/hr through 100'F/hr heatup pressure-temperature limits. These tables provide the allowable pressure versus reactor coolant temperature for the various heatup conditions.
The allowable pressure is in units of ksi while the temperature is in units of 'F.
Figures 6 through 9, Page 20
5f>
5p(
mt tl e
619(86L4) ch-21 provide a graphical presentation of the heatup pressure-temperature limits found in Tables 3 and 4. It is permissible to linearly interpolate between the heatup pressure-temperature limits.
2.6 HYOROSTATIC TEST AND CORE CRITICAL LIMIT ANALYSIS Both 10 CFR Part 50 Appendix G and the ASME Code Appendix G require the development of pressure-temperature limits which are applicable to inservice hydrostatic tests.
For hydrostatic tests performed subsequent to loading fuel into the reactor vessel, the minimum test temperature is determined by evaluating KI, the mode I stress intensity factors.
The evaluation of KI is performed in the same manner as that for normal operation heatup and cooldown conditions except the factor of safety applied to the pressure stress intensity factor is 1.5 versus 2.0.
From this evaluation, a
pressure-temperature limit which is applicable to inservice hydrostatic tests is established.
The minimum temperature for the inservice hydrostatic test pressure can be determined by entering the curve at the test pressure (1. 1 times normal operating pressure) and locating the corresponding temperature.
The inservice hydrostatic test limit is provided for 15 EFPY and is referenced on the core critical P-T limit figure.
Appendix G to 10 CFR Part 50, specifies pressure-temperature limits for core critical operation to provide additional margin during actual power operation.
The pressure-temperature limit for core critical operation is based upon two criteria.
These criteria are that the reactor vessel must be at a temperature equal to or greater than the minimum temperature required for the inser vice hydrostatic
- test, and be at least 40'F higher than the minimum pressure-temperature curve for normal operation heatup or cooldown.
The core critical limit has been developed based upon the 50'F/hr heatup P-T limit and the minimum temperature required for the inservice hydrotest for 15 EFPY.
The core critical limits are Page 21
V~
p a
<Wr
619 (86L4) ch-22 referenced on the inservice hydrostatic test P-T limit figure.
The core critical and inservice hydrostatic test P-T limits are provided in Figure 10.
Note, that the core critical limits established above are solely based upon fracture mechanics considerations, and do not consider core reactivity safety analyses which can control the temperature at which the core can be brought critical.
2.7 LOWEST SERVICE TEMPERATURE, MINIMUM BOLTUP TEMPERATURE, AND MINIMUM PRESSURE LIMITS In addition to the computation of the reactor vessel beltline P-T limits, additional limits have been provided fot reference.
These additional limits are the Lowest Service Temperature, Minimum Boltup Temperature, and Minimum Pressure Limits.
These limits are described below.
The Lowest Service Temperature is the minimum allowable temperature at pressures above 20% of the pre-operational system hydrostatic test pressure (625 psia).
This temperature is defined as equal to the most limiting RTNDT for the balance of Reactor Coolant System (RCS) components plus 100'F, per Article NB 2332 of Section III of the ASME Boiler and Pressure Vessel Code.
The maximum RTNDT for the balance of the RCS components is associated with the Reactor Coolant System piping and has a value of 60'F.
Therefore, the Lowest Service Temperature is equal to 160'F.
The minimum p'ressure limit is the break, point between the minimum boltup temperature and the Lowest Service Temperature.
Defined by the ASME Boiler and Pressure Vessel Code as 20% of the pre-operational Page 22
619(86L4) ch-23 hydrostatic test pressure, the minimum pressure is as follows when pressure correction factors for elevation and flow are taken into account:
562 psia 551 psia T
< 200'F c
Tc
> 200'F The minimum boltup temperature is the minimum allowable temperature at pressures below the 20% of the pre-operational system hydrostatic test pressure.
The minimum is defined as the initial RTN0T for the material of the higher stressed region of the reactor vessel plus any effects for irradiation per Article G-2222 of Section III of the ASME Boiler and Pressure Vessel Code.
The maximum initial RTN0T associated with the stressed region of the reactor vessel flange is 50'F.
The minimum boltup temperature is therefore 50'F.
- However, for conservatism a minimum boltup temperature of 80'F is utilized.
Page 23
f't<
"C
619(86L4)ch-24 2;8 DATA Reactor Vessel Data Reference Design Pressure Design Temperature Operating Pressure Beltline Thickness Inside Radius Outside Radius Cladding Thickness 2500 psia 650'F 2250 psia 8.625 in 86.813 in 95.66 in
.2187 in 14 14 14 14 14 14 14 Material-SA 302 Grade B
Reference Thermal Conductivity Youngs Modulus Coefficient of Thermal Expansion Specific Heat Density 23.8 BTU/hr-ft-'F 15 28 x 10 psi 15 7.77 x 10 in/in/'F 15
.12 BTU/lb-'F
.283 lb/in Stainless Steel Claddin Thermal Conductivity 10 BTU/hr-ft-'F Ad usted Reference Tem erature Values 15 EFPY 1/4t 3/4t 140'F 119'F Film coefficient on inside surface 1000 BTU/hr-ft -'F Pressure Correction Factors For Elevation and Flow RCS temperature
< 200'F RCS temperature
> 200'F dp
= 63 psia dp
~ 74 psia Page 24
S g4, fY,
619(86L4) ch-25 3.0 LOW TEMPERATURE OVERPRESSURE PROTECTION 3.1 GENERAL The primary objective of low temperature overpressure protection (LTOP) systems is to preclude violation of applicable Technical Specification P-T limits during startup and shutdown conditions.
These P-T limits are usually applicable to a finite time period such as one cycle, 5 EFPY, etc.
and are based upon the ir radiation damage prediction by the end of the period.
Accordingly, each time new P-T limits are to become effective, the LTOP system needs to be re-analyzed and modified, if necessary, to continue its function.
A typical LTOP system includes pressure religving devices and a
number of administrative and operational controls.
At St.
Lucie Unit 2, the current LTOP system makes use of two shutdown cooling relief valves (SDCRV) at the low RCS temperatures and two PORVs, in the remaining part of the LTOP temperature range, i.e.,
up to the LTOP enable temperature.
The SDCRVs (Tag Nos.
V3666 and V3667) are spring operated rel.ief valves with a lift setting of 350 psia each.
The PORVs (Tag Nos.
V1474 and V1475) are pilot operated relief valves with a nominal opening setpoint of 470 psia.
These relief valves, in combination with certain other limiting conditions for operation contained in Technical Specifications, comprise the St. Lucie Unit 2 LTOP System.
Since the new P-T limits described in this report cover the
'operating period ending at 15 EFPY (vs. the current 6 EFPY) the existing LTOP system was re-analyzed to determine the modifications required to be implemented in order for the system to provide adequate LTOP up to 15 EFPY.
The following sections document the method, assumptions and results of the new analysis.
Page 25
fg 0
<<pe I
et e,
619(86L4) ch-26 3.2 METHOD AND ASSUMPTIONS The approach taken was to maintain the existing SDCRV and PORV setpoints as indicated in Reference (7).
Accordingly, the existing SDCRV lift setting of 350 psia and the nominal PORV setpoint of 470 psia were assumed in the subject analysis.
The assumptions utilized in the existing pressure transient analyses (Reference
- 16) which support the St. Lucie Unit 2 LTOP'system in its present configuration were re-evaluated and modified as indicated below.
1.
RCS pressure just prior to PORV opening was conservatively assumed to be greater than the nominal PORV setpoint, due to relative pressure instrument uncertainty between the pressure indication and the PORV actuation channels.
This 3a uncertainty was determined by a statistical combination of indi,vidual errors in the non-common components.
The individual errors were calculated based upon a 30-month calibration interval which is representative of a 24-month fuel cycle.
This new relative uncertainty of 23.4 psi was added to the nominal PORV setpoint of 470 psia to arrive at a
conservative opening pressure of 493.4 psia.
This is a new assumption that maximizes transient pressures at the PORV opening.
2.
PORV opening time was assumed to equal 1.0 second which enveloped the opening times observed in applicable tests with a Garrett PORV.
Based on an evaluation of test data it was assumed that this total opening time consisted of solenoid delay time of 0. 12 seconds and stroke time of 0.88 seconds.
To account for ramp opening during stroke time, a
delay in PORV opening equal to a sum of the Page 26
l>>
I~I
~
619(86L4) ch-27 solenoid delay time (0.12 seconds) and one half of the stroke time (0.44 seconds) was assumed in the transient analyses.
This delay was assumed to be followed by instantaneous opening.
This is a
new assumption.
The existing analyses assumed instantaneous PORV opening, without any delay.
3.
The impact of PORV opening time was taken into account by adding transient-specific pressure accumulation during 4t
= 0.56 seconds (0. 12 seconds pIus 0.44 seconds) to the opening pressure of 493.4 psia to arrive at the maximum opening pressures.
Pressure accumulation was assumed to be a function of an applicable ramp rate.
(This is a new assumption.)
Based on the existing analyses and modified assumptions, new maximum transient pressures in the same design basis transients were determined as appropriate.
Out of these,'the most limiting pressures in given temperature ranges were selected as "controlling" the limiting temperatures for LTOP.
Finally, by applying these controlling pressures to the P-T limit curves for 15 EFPY, limiting conditions for operation were identified.
The procedure described above is similar to that utilized in the existing analyses, with one exception.
Previously, LTOP enable temperatures during heatup and cooldown have been determined at the intersections between a horizontal. line corresponding to the safety valve setpoint (2500 psia) and the most limiting P-T limit curves for heatup and cooldown, respectively.
Note that the enable temperature generally identifies the upper temperature limit below which the LTOP system has to be operable.
Page 27
t5 lW '
sl ~
619(86L4) ch-28 J
In this analysis, the LTOP enable temperatures were determined in accordance with a definition contained in the latest revision of Standard Review Plan 5.2.2 (Reference 5).
According to SRP 5.2.2, the LTOP enable temperature is "the water temperature corresponding to a metal temperature of at least RTNOT + 90 F at the beltline location
( 1/4t or 3/4t) that is controlling in the Appendix G limit calculations".
As a result, the enable temperatures for 15 EFPY determined by using the above guideline (247'F during heatup, and 230'F during cooldown) are less restrictive than the existing enable temperatures of 313'F during heatup and 304'F during cooldown, that are applicable to 6 EFPY.
3.3 PRESSURE TRANSIENT ANALYSES The modified assumptions presented in Section 3.2 affect only pressure transients, mitigated by a PORV.
Accordingly, the exiting analyses utilizing a SOCRV for transient mitigation remained unchanged (see Reference 16) and the maximum transient pressures calculated for the three design basis transients apply to this analysis.
These pressures are provided in Table 6 in the "SOCRV" column.
The RCP start transient mitigated by a SDCRV is illustrated in Figure 11.
As a result, new pressure transient analyses were performed only for the design basis transients mitigated by a PORV.
These transients are addressed below.
3.3.1 Ener Addition Transients The RCS pressure transient due to a reactor coolant pump (RCP) start when the secondary steam generator inventory is at a higher temperature than the primary (RCS) inventory is the design basis energy addition transient, as indicated in Reference 16.
Page 28
I 4 r 4
l v
f
619(86L4) ch-29 Accordingly, a
new analysis of this transient was performed using the modified assumptions, per Section 3.2 and the same computer model as was utilized previously.
The analysis was performed for the nominal PORV setpoint of 470 psia with instrument uncertainty and PORV opening time taken i nto account.
The assumed RCS fluid temperature was 290'F, i.e.,
equal to that in the existing analysis.
The analysis assumed that the pressure transient was taking place in the pressurizer.
The effect of the PORV inlet piping on the analysis results was taken into account by assuming PORV flow rates to be calculated at the PORV inlet pressure, rather than at the pressurizer pressure.
This assumption conservatively reduced PORV discharge, thus maximizing the transient pressures.
The following major assumptions were used in the analysis of the RCP start transient; in addition to the assumptions mentioned above and in Section 3.2:
,Water-solid conditions in the pressurizer.
One-PORV mitigation.
Additional energy inputs equal to 1'4 decay heat, full installed pressurizer heater capacity, and operating RCP motor rating.
Initial saturated conditions in the pressurizer.
PORV discharge as a function of fluid subcooling at the PORV inlet.
Letdown flow paths isolat'ed.
No heat absorption or metal expansion in the reactor coolant pressure boundary.
Page 29
4 tp
619 (86L4) ch-30 This pressure transient is graphically illustrated in Figure 12 and the resulting maximum transient pressure of 513 psia is provided in Table 6.
3.3.2 Mass Addition Transients The RCS pressure transient due to an inadvertent safety injection actuation is the design basis mass addition transient.
The most severe mass addition transient results from an actuation of two r
HPSI pumps with a simultaneous operation of all three charging pumps, with letdown isolated.
This transient,
- however, can only occur at RCS temperature above 200'F, based on Technical Specifications, LCO 3.5.3.
As a result, at RCS temperature below 200'F, the most limiting mass addition transient is due to one HPSI and three charging pumps input.
A review of the existing analyses of mass addition transients indicated that the equilibrium pressures in these transients apply to the subject analysis.
Note that an equilibrium pressure is the pressurizer pressure at which mass input into the RCS equals relief valve discharge.
Accordingly, these equilibrium pressures were utilized in the subject analysis to arrive at the maximum pressures in the two design basis mass addition transients mitigated by a PORV.
These equilibrium pressures are as follows:
Transient Equilibrium Pressure PORV Miti ation 2 HPSI
+ 3 Charging Pumps 1 HPSI + 3 Charging Pumps 535 psia 375 psia Page 30
C".
619(86L4) ch-31 The maximum transient pressures were determined based on the modified assumptions, Section 3.2 and the equi librium pressures above.
The approach was to compare the equilibrium pressures with the maximum pressures at the PORV opening.
The maximum transient pressures were assumed to equal the highest of the two, as illustrated below.
The equilibrium pressure in the two HPSI and three charging pumps transient is 535 psia (see above).
At the nominal PORV setpoint of 470 psia, a conservative value of opening pressure is 493.4 psia, Section 3.2, Item 1.
When the effect of the PORV opening time of 1.0 second is taken into account, the maximum pressure at the PORV opening becomes approximately 529 psia.
Since the latter is below the equilibrium pressure of 535 psia, the maximum transient pressure is assumed to equal 535 psia (see Table 6).
Conversely, for the one HPSI and three charging pumps transient, the equilibrium pressure is 375 psia vs. the maximum pressure at the PORV opening of 512 psia.
Accordingly, the latter was assumed as the maximum transient pressure
=-in this transient.
The final results of the mass addition transient analysis are provided in Table 6.
3.3.3 Controllin Pressures The pressure transient analysis results contained in Table 6
were evaluated to identify the controlling pressures and applicable temperature ranges.
The controlling pressures are the maximum transient pressures of all applicable transients in a particular temperature region.
Page 31
I W> I A
6],9(86L4)ch-32 Table 7 contains a summary of controllirg pressures that were utilized in the determination of limiting conditions for operation that result from LTOP requirements.
These limiting conditions for operation are provided in,Section 3.4.
3.4 LIMITING CONDITIONS FOR OPERATION The temperature requirements for aligning the SDCRVs and PORVs for LTOP, and the limitations on heatup and cooldown rates are provided in Table 8.
It should be noted that during heatup, the LTOP function can be transferred from the SDCRVs to the PORVs at any temperature above that required for SDCRV alignment (e.g.,
165'F in Table 8).
During cooldown, however, the SDCRVs must take over the LTOP function upon reaching the indicated temperature such as 163'F in Table 8.
The existing LTOP requirements related to the limitations on RCP starts and HPSI pump alignment to the RCS remain unchanged.
These are also included in Table 8.
Note the new applicable temperatures for the RCP start limitation in Table 8.
3.5
SUMMARY
OF PROPOSED CHANGES The proposed LTOP system is designed in accordance with the requirements set forth in the NRC Branch Technical Position RSB 5-2, Reference 5.
The proposed system is adequate to prevent violation of Appendix G
P-T limits during the operating period ending at 15 EFPY. 'n order to implement the proposed LTOP system the following is required:
1.
Modification of appropriate Technical Specifications and 2.
Modification of appropriate plant 'operating, procedures.
Page 32
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619(86L4) ch-33 The implementation of the proposed LTOP system will not result in a reduction in the margin of safety presently afforded by Technical Specifications.
4.0 REFERENCES
1.
Code of Federal Regulations, 10 CFR Part 50, Appendix A, "General Design Criteria for Nuclear Power Plants",
January 1988.
2.
Code of Federal Regulations, 10 CFR Part 50, Appendix G
"Fracture Toughness Requirements",
January 1988.
3.
ASME Boiler and Pressure Vessel Code Section III, Appendix G, "Protection Against Nonductile Failure",
1986 Edition.
4.
Regulatory Guide 1.99, "Radiation Embrittlement of Reactor Vessel Materials", U.S. Nuclear Regulatory Commission, Revision 2, May 1988.
5.
U. S. Nuclear Regulatory Commission Standard Review Plan (SRP) 5.2.2, Overpressure Protection, Revision 2, November 1988.
6.
Florida Power and Light Company letter L-86-25, C. 0.
Woody to F. J. Miraglia, dated 1-28-86.
7.
Florida Power and Light Company St. Lucie 1 and 2, Input Parameters for Revised P-T Limits and LTOP Requirements, JPN-PSL-SEMJ-89-007, Revision 1, March 1989.
8.
C-E Drawing, E-233-496, General Arrangement Elevation, 172 in I.D.
PWR.
9.
C-E Drawing, E-71172-124-001, Intermediate Shell.
Page 33
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Vg
~ I
619(86L4) ch-34 10.
C-E Drawing, E-71172-142-001, Lower Shell.
11.
Evaluation of Pressurized Thermal Shock Effects Oue to, Small Break LOCA with Loss of Feedwater for the C-E NSSS, CEN-189, December 1981.
12.
Input Parameters for Revised P-T Limits and LTOP Requirements, JPN-PSL-SEMJ-89-007, Revision 1 Attachment B,
March 1989.
13.
ASME Boiler and Pressure Vessel Code Section XI, Appendix A, "Analysis of Flaw Indications",
1986 Edition.
14.
Instruction Manual, Reactor Vessel Assembly, St. Lucie Unit No. 2, Florida Power and Light Company, C-E Book No.
- 71172, December 1977.
15.
ASME Boiler and Pressure Vessel Code Section III, Appendix I, "Design Stress Intensity Values, Allowable
- Stresses, Material Properties, and Design Fatigue Curves",
1986 Edition.
16.
Florida Power and Light Company Letter L-87-482, C. 0.
Moody to the NRC, dated 11-27-87, Attachment 4.
Page 34
F,
't 41 f 0
L4)ch-35 TABLE 1 ST.
LUCIE UNIT 2 REACTOR VESSEL BELTLINE MATERIALS LOCATION ID NO.
Intermediate Shell Plate Intermediate Shel 1 Long Seam Welds M-4116-3 M-605-1 M-605-2 M-605-3 101-124 A,B,C Lower Shell Plate M-4116-1 M-4116-2 44
.07
.60 20 74 92 74 31
.07
.13
.04
.60
.61
.62
.61
.07 20 30 10
-50 INITIAL CF CU,X NI,X RTNDTF 37
.06
.57 20
- 1. 826
- 1. 826 1.826
- 1. 826
- 1. 826
- 1. 826 1.826 92 74 99 84 99 84 140 119 138 lll-110 89 12
-4 PEAK ART NEUTRON FLUENCE 15 EFPY, 'F 10 N/M, E>lMev 1/4 T 3/4 T Lower Shell Long Seam Welds 101-142 A,B,C 38
.05
.10
-5.0 1.826 26 6
Intermediate/Lower Girth Seam Weld 101-171 41
,.07
.08
-70 1.826 14
-10 (a)
CF = Chemistry Factor determined by Regulatory Guide 1.99 Rev.2 (b)
Peak fluence on the vessel inside surface at 0 azimuth.
(Cl
TABLE 2 ST.
LUCIE UNIT 2 CONTROLLING MATERIALS AND THEIR ADJUSTED REFERENCE TEMPERATURES REACTOR VESSEL St. Lucie Unit 2 LOCATION Intermediate Shell Plate MATERIAL IO NO.
M-605-1 INITIA)
~DT~
~14T 3/4T
<<w t
I 1
t) qP<
a,~
4 2
eQ Pa J 0
TABLE 3 St. Lucie Unit 2 Cooldown and Heatuo Pressure-Temperature Limit Data, JO F/hr to 50oF/hr and Isothermal RCS lEHP OEG F
40 50 60 70 80 100 110 120 130 140 150 160 170 180 190 200 210 220 230 240 250 260 270 280 290 300 310 320 330 340 I SO THERNAL 0.447 0.457 0, 467 0.477 0.487 0.507 0.527 0.547 0.567 0.597 0.627 0.667 0.707 0.7S7 0.817 0.'876 0.956 1.046 1.146 1.266 1.406 1.566 1.756 1.966 2.21d 2.506 2.836 2.92d 10 F/
HQJR 0.407 O. 417 0.427 0.437 0.457 0.467 0.487 0.517 0.537 0.567 0.607.
0.647 0.687 0.747 0.807 0.866 0.946 1.0C6 1.1C6
'1.266 1.406 1 ~ 566 1.7$6 1.966 2.21d 2.506 2.836 2.926 20 F/
HOUR 0.377 0.387 0.407 0.417 0.437 0.457 0.487 0.507 0.547 0.577 0.627 0.667 0.727 0.797 0.856 0.946 1.046 1.146 1.266 1.406 1.566 1.7$ d 1.966 2.21d, 2.506 2.836 2.92d 30 F/
HOUR 0.337 0.347 0.357 0.367 0.387 0.407 0.427 0.457 O.C87 0.517 0.557 0.597 0.657 0.717 0.787 0.856 0.946 1.046 1.146 1.266 1.COS 1.S66 1.7$ 6 1.966 2.216
- 2. S06 2.836 2.926 CCOLOOQI P-ALLQJABLE (KBl) 40 F/
HOUR 0.297 0.307 0.317 0.337 0.357 0.377 0.397 0.427 0.457 0.487 0.537 0.577 0.637 0.707 0.777 0.856 0.9SS 1.046 1.146
'1.266 1.C06 1.566 1.756 1.966 2.21d 2.506 2.836 2.926 50 F/
HQJR 0 ~ 257 0.267 0.287 0.297 0.317 0.337 0.367 0.397 0.427 0.467 0.507 0.567 0.627 0.687 0.767 0.8$ 6 0.9SS 1.046 1.1C6 1.266 1.406 1.566 1.7$ 6 1.966 2.216 2.506 2.836 2.926 10 F/
HQJR 0.467 0.477 0.487 0.507 0.527 0.547 0.567 0.597 0.627 0.667 0.707 0.757 0.817 0.87d 0.9SS 1.046 1.1C6 1.256 1 ~ 386 1.536 1.716 1.91d 2.146 2.416 2.7M 2.926 20 F/
HQJR 0.467 0.477 0.487 0.497 0.517 0.537 0.567 0.597 0.627 0.667 0.707 0.757 0.817 0.876 0.95d 1.046 1.146 1.246 1.376 1.516 1.676 1.866 2.086 2.346 2.636 2.926 30 F/
HOUR 0.467 0.467 0.457 0.467 0.477 0'97 0.517 0.547 0.577 0.617 0.657 0.717 0.777 0.836 0.916 1.006 1.116 1.2CS 1.356 1.496 1.6C6 1.826 2.036 2.276 2.546 2.866 Z.926 HEATUP P.ALLQJABLE (KBl) 00
~ ~ ~
40 F/
HOUR 0.467 0.457 O.C47 0.447 0.447 0.457 0.477 0.497 0.527 0.557 0.597 0.647 0.697 0.746 0.816 0.906 1.006 1.11d 1.246 1.396 1.566 1.77S 1.986 2.206 2.466 2.766 2.926 50 F/
HOUR O.C67 0.457 0.437 0.427 0.427 0.437 O.CC7 O.C57 0.477 0.507 0.537 0.577 0.627 0.676 0.73S 0.816 0.896 0.996 1.116 1.246 1.406 1.586
'1.796 Z.OM 2.306 2.636 2.926 Page 37
I Pt
~1 A
,Pg tp j(f 1
TABLE 4 St. Lucie Unit 2 Cooldown and Heatup Pressure-Temperature Limit Data, 60oF/hr to 100 F/hr RCS TEHP OEG F
40 50 60 70 80 90 100 110 120 130 140 150 160 170 180 190 200 210 220 230 2CO 250 260 270 280 290 300 310 320 330 340 350 360 370 380 ISO THERHAL 0.447 0.457 0 ~ 467 O.C77 o.'48r 0.507 0.527 0.547 0.567 0.597 0.627 0.667 0.707 0.7S7 0.817 0.876 0.956 1.046
- 1. 1Cb 1 ~ 266 1.C06 1.566 1.7S6 1.966 2.21d 2.506 2.836 2.92d 60 F/
HOUR 0.227
- 0. 237 0.247 0.267 0.287 0.307 0.337 0.367 0.407 0.447 0.'487 0.547 0.607 0.687 o'.rer 0.856 0.956 1.046
- 1. 146 1.266 1.406 1.566 1.756 1.966 2.21d 2.506 2.836 2.92d 70 F/
HOUR 0.197 0.217 0.237 0.257 0.277 0.307 0.337 0.377 0.427 0.477 0.527 0.597 0.677 0.767 0.856 0.956 1.046 1 ~ 146 1 ~ 266 1.406 1.566 1.7Sd 1.966 2.216 2.506 2.836 2.926 75 F/
HOUR 0.167 0.187 0.197 0.217
- 0. 237 0.267 0.297 0.327 0.367 0.417 0.467 0.527 0.597 o'.err 0.767 0.866 0.956 1.046 1.046
- 1. 2&b 1.406 1.566 1 ~ 756
, 1.966 2.216 2.506 2.836 2.926 80 F/
HOUR 0.147
- 0. 167 0.187 0.207 0.227 0.247 0.277 0.317 0.357 0.407 0.457 0.517 0.587 0.667 0.767 0.866 0.956 1.046 1.146 1.266 1.406 1.566 1.756 1.966 2.216 2.506 2.836 2.92d COOLOOMH P-AL!.OMASLE (KS I )
90 F/
HOUR 0.117 0.137 0.147 0.167 0.197 0.227 0.257 0.287 0.337 0.387 0.437 0.507 0.577 0.667 0.767 0.876 0.956 1.0C6 1.146 1.266 1.C06 1.566 1.756 1.966 2.2'16
. 2.S06 2.836 2.92d 100 F/
HOUR 0.087 0.097 0.117 0.137
- 0. 167 0.197 0.227 0.267 0.317 0.367 0.427 O.C97 0.577 0.667
- 0. 777 0.876 0.956 1.046 1.146 1.266 1.406 1.566 1.7S6 1.966 2.216 2.506 2.836 2.926
&0 F/
HOUR 0.467 0.457 0.437 0.417 0.407 0.407 0.417 0.427 0.447 0.467 0.497 0.527 0.567 0.606 0.656 0.726 0'06 0.896 0.996.
1.11d 1.2S6 1.416 1.606 1.816 2.066 2.356 2.696 2.926 70 F/
HOUR 0.467 0.457 0.427 0.407 0.397 0.397 0.397 0.407 0.417 0.427 0.457 0.477
- 0. 517 0.546 0.596 0.656 0.716 0.796 0.896 0.996 1.126 1.2&6 1.436 1.626 1.8C6 2.106 2.406 2.7C6 2.926 80 F/
HOUR 0.467 0.457 0.427 0.407 0.387 0.377 0.377 0.377 0'87 0.397 0.417 0.437
- 0. 467 O.C96 0.536 0.586 0.646 0.716 0.796 0.886 1.006 1.126 1.276 1.446 1.646 1.876 2.146 2.456 2.816 2.926 90 F/
HOUR 0.457 0.427 0.407 0.387 0.367 0.367 0.367 0.367 0.377 0.387 0.407 0.427 0.446 O.C86 0.526 0.576 0.636 0.716 0.796 0.896 1.006 1.146 1.296 1.476 1.67d 1.916
- 2. 196 2.516 2.886 2.92d HEATUP P-ALLOMASLE (KSI) 100 F/
HOUR O.C&7 0.457 0.427 0.397 0.377 0.367 0.357 0.347 0.347 0.347 0.357 0.377 0.397 o.coe 0.436 0.476 0 ~ 526 0.576 0.636 0.716 0.806 0.90&
1.026 1.156 1.316 1.506 1.716 1.9&6 2'56 2.576 2.92&
Page 38
TABLE 5 St. Lucie Unit 2 Hydrostatic Test Pressure-Temperature Limit Data RCS TEHP HYORO 080 F
STATIC 40 50 0.617 60 0.627 70 0.637 80 0.657 90 0.677 100 0.697 110 0.717 120 0.747 130 0.777 140 0.817 150 0.857 160 0.907 170 0.967 180 1.037 190 1.117 200 1.196 210 1.296 220 1.416 230 1.556 240 1.716 250 1.906 260 2.116 270 2.366 280,2.d56 290 2.92d Page 39
f+I
619(86L4) ch-35 TABLE 6 MAXIMUM TRANSIENT PRESSURES Transient PORV RELIEF VALVE SDCRV RCP Start 513* psia 343 psia 2 HPSI+3 Charging Pumps 1 HPSI+3 Charging Pumps 535 psia 512" psia 355 psia 345 psia Note:
- New Va)ue Page 40
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0
619(86L4) ch-36 TABLE 7
SUMMARY
OF CONTROLLING PRESSURES RELIEF VALVE RCS Temperature SOCRV PORV
< 200'F 345 psia 513 psia
> 200'F 355 psia*
535 psia Note:
- This controlling pressure was not utilized since SDCRV alignment for LTOP is below 200'F.
Page 41
E
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~ tJ
619(86L4) ch-38 TABLE 8 LTOP REQUIREMENTS, 15 EFPY Both SDCRVs are required to be aligned to the RCS as follows:
During heatup, at T
< 165'F During cooldown, at T
< 163'F Both PORVs are required to be aligned to the RCS as follows:
During heatup, at 165'F
< T
< 247'F During cool down, at 163'F
< T
< 230'F Heatup rates shall be limited to a maximum of:
50'F/hr, at all temperatures Cooldown rates shall be limited to a maximum of:
50'F/hr, Applicability:
During heatup, at T
< 247'F During cooldown, at T
< 230'F One HPSI pump shall be rendered inoperable prior to entering Mode 5
during cooldown.
(This is an existing limitation).
30'F/hr, at T
< 103'F at 103'F
< T
< 126'F 75'F/hr, at 126'F T
137'F 100'F/hr, at T
> 137'F A RCP shall not be started with two idle loops, unless the secondary water temperature of each steam generator is less than 40'F above each of the RCS cold leg temperatures.
(This is an existing limitation).
Page 42
0 e 's g<a.
U:
~
,tw'4 '
2500 FIGURE 1 ST. LUCIE UNIT 2 P-T LIMITS 16'FPY COOLDOWN 2000 lO 0
W IC Po 1500 LOWEST SERVICE TEMP 160 0F a
1000 Cl I~
u MIN PRESSURE 10 ~F/HR MINIMUMBOLTUP TEMPERATURE 80 4F 100 200 300 INDICATED RCS TEMPERATURE Tc, 4F 400 Page 43
-r
2500 FIGURE 2 ST. LUCIE UNIT 2 P-T LIMITS 15 EFPY COOLDOWN 2000 g
1500 i
LOWEST SERVICE TEMP 160 4F aa 1000 Q
W I
O MIN PRESSURE MINIMUMBOLTUP TEMPERATURE 80 oF INDICATED RCS TEMPERATURE Tct 4F page 4a
'\\
0 C
0
1000 900 800 LOWEST SERVICE TfMP 160 0F FIGURE 3 ST. LUCIE UNIT 2 PT LIMITS 15 EFPY'OOLOOWN M
700 tO Pu 800 500 CC Pg 400 O
o 300 I~
Qz 209 MIN PRESSURE 4,o
~+
o aO MINIMUMBOLTUP TEMPERATURE 80 4F 80 100 150 200 250 INOICATEO RCS TEMPERATURE Tc, 4F 300 I
Page 45
gJ
1000 900 LOWEST SERVICE TEMP 160 4F.
FIGURE 4 ST. LUCIE UNIT 2 P-T LIMITS 15 EFPY COOLDOWN V) 0 700 M
(0 600 PL Lll 500 M
Vl Lll CL 400 O
Ul I
300 CIX 200 qO
~i++
gO MINIMUMBOLTUP TEMPERATURE 80 4F 60 100 150 200 250 INDICATED RCS TEMPERATURE Tc, 4F Page 46
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A~
1000 900 800 LOWEST SERVICE TEMP 160 0F FIGURE 5 ST. LUCIE UNIT 2 PT LIMITS 15 EFPY COOLDOWN V)L 700
. IL M
~
coo lC CL Ill~
5O0 M
V7 Pc 400 PL CI W
I o
300 CIX 200 MIN PRESSURE MINIMUMBOLTUP TEMPERATURE 80 4F 60 100 '
150 200 250 INDICATED RCS TEMPERATURE Tc~ 4F 300 Page 47
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2500 2000 FIGURE B ST. LUCIE UNIT 2 P-T LIMITS 15 EFPY HEATUP FLANGE LIMIT 50 4F/HR ISO 10 30 50 70 90 0
tLlK
~
g 1500 LOWEST SERVICE TEMP 160 0F 4
1000 A
I~
O MIN PRESSURE MINIMUMBOLTUP TEMPERATURE 80 4F 100 200 300 INDICATED RCS TEMPERATURE Tc, 4F 400 Page 48
2500 FIGURE 7 ST. LUCIE UNIT 2
. P-T LIMITS 15 EFPY HEATUP ISO 20 40 2000 80 g
1500 LOWEST SERVICE TEMP 1SO 0F a.
1000 Cl I
500 MIN PRESSURE yO MINIMUMBOLTUP TEMPERATURE 80 oF 100 200 300'NDICATED RCS TEMPERATURE Tc, oF Page 49
ir D
O~
l
~l 'I V
1000 900
&00 FIGURE 8 ST. LUCIE UNIT 2 P-T LIMITS 15 EFPY HEATUP LOWEST SfRVICE TEMP 160 OF Ch FLANGE LIMIT 50 oF/HR Ul W
(D Vl 600 PL Ill CC N
500 M
LLl L
400 4
Lll I
o 300 d
MIN PRESSURE 200 MINIMUM8OLTUP TEMPERATURE 80 oF 60 100 150 200 250 INDICATED RCS TEMPERATURE Tc~ oF 300 Page 50
r 1
1000 900 800 FIGURE 9 ST. LUCIE UNIT 2 P-T LIMITS 15 EFPY HEATUP LOWEST SERVICE TEMP 160 0F PL 700 600 500 e
400 4
I~
300 4z 200'IN PRESSURE gO qO MINIMUMBOLTUP TEMPERATURE 80 4f 60 100 150 200 250 INDICATED RCS TEMPERATURE Tc, 4F Page 51
1
2500 2000 FIGURE 10 ST. LUCIE UNIT 2 P T LIMITS 15 EFPY HYDROSTATIC AND CORE CRITICAL OPERATION INSERVICE HYDROSTATIC TEST C
M Q
V MINIMUM SERYICE TEMP 60 OF 1000 Ch I~
O MIN PRESSURE CORE CRITICAL OPERATION (50 oF/HR) 500 MINIMUMBOLTUP TEMPERATURE 80 oF 0
INDICATED RCS TEMPERATURE Tc, oF
\\
400 Page 52
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FIGURE 11 ST. LUCIE.2 RCP START TRANSIENT W/SOCRV, P.SET ~ 360 PSIA dt.p ~ 404F PNIAX 36<>
10 ELAPSED TIME. SECONDS 20 Page 53
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,~
0
FIGURE 12 ST. LUCIE UNIT 2 RCP START TRANSIENT NOMINALPORV SETPOINT = 470 PSIA 600 ht~.P = 4(PF 550 (1) 513 PSIA 500 450 I
V 400 Vl W
CC 0
350 3000'0 15 ELAPSED TIME, SECONDS 20 NOTE: (1) THE MAXIMUMTRANSIENT PRESSURE EQUALS THE PRESSURIZER PRESSURE AT THE PORV OPENING THAT INCLUDES:
- NOMINALPORV SETPOINT (470 PSIA)
- INSTRUMENT UNCERTAINTY(23A PSI)
- PRESSURE ACCUMULATIONDUE TO 1.0 SEC.
Page 54
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~C 0
0
FIGURE 13 FIGURE 3.4-4 ST. LUCIE-2 P/T LIMITS, 15 EFPY MAXIMUMALLOWABLECOOLDOWN RATES 100 80 gf 00 l~
4O O0V 20 RAT+ F/HR
.30 50 75 100 TE
. LIMI,
<105 05-130
"=130-140
>140 80
- 100, 120 140 180 200 To - INDICATED REACTOR COOLANT TEMPERATURE, F
NOTE: A MAXIMUMCOOLDOWN RATE OF 100 F/HR IS ALLOWED AT ANY TEMPERATURE ABOVE 14{IF Page'5 ST. LUCK NET 2 3'32
.l I