ML19332B606
| ML19332B606 | |
| Person / Time | |
|---|---|
| Site: | Calvert Cliffs  |
| Issue date: | 10/31/1989 |
| From: | ABB COMBUSTION ENGINEERING NUCLEAR FUEL (FORMERLY |
| To: | |
| Shared Package | |
| ML19332B604 | List: |
| References | |
| B-MPS-89-1026, NUDOCS 8911060062 | |
| Download: ML19332B606 (29) | |
Text
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Final Report on i.
Low Temperature Overpressure Protection Pressure-Te:1perature Limits i
for Ba imore Gas and Electric Company Calvert Cliffs Unit 1
+
October 1989 Prepared by COMBUSTION ENGINEERING, INC.
Plant Structures Group 1000 Prospect Hill Road Windsor, CT 06095 0500 L
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TABLE OF CONTENTS Seetion lille f191
1.0 INTRODUCTION
11
2.0 BACKGROUND
2-1 r
3.0 DESCRIPTION
OF METHOD 3-1 3.1 ASSUMPTIONS OF ANALYSIS 3-3 3.2 HEAT TRANSFER ANALYSIS 35 3.3 FRACTURE MECHANICS INFLUENCE COEFFICIENTS 38 3.4 FRACTURE MECHANICS EVALUATION 3 10 4.0 RECOMMENDED APPLICATION 41 5.0 RESULTS S1 l
6.0 REFERENCES
61 L
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LIST OF TABLES r.
I Table No.
Tillt Etat l'
l.0 BGLE Calvert Cliffs Unit 1 Cooldown LTOP 5-2 P-1 Limit Data 2.0 BG&E Calvert Cliffs Unit 1 Heatup LTOP 53 P-T Limit Data t
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816. doc (8978)bh-4
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b LIST OF FIGURES 1
Fioure No.
1111g Eggt i
1.0 BGLE Calvert Cliffs Unit 1 54 LTOP P-T Limits - Cooldown
)
)
2.0 BG&E Calvert Cliffs Unit 1 5-5 1
3.0 BG&E Calvert Cliffs Unit 1 56 LTOP P-T Limits Heatup l'
4.0 BG&E Calvert Cliffs Unit 1 5-7 LTOP P T Limits - Heatup t
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816. doc (8978)bh 5 Section 1 INTRODUCTION This report documents the development of low temperature overpressurization transient Pressure-Temperature (P-T) limits for use in determination of Low Temperature Overpressure Protection (LTOP) setpoints. The purpose of establishing the LTOP Pressure-Temperature limits is to provide Baltimore Gas
& Electric with the ability to safely increase the Reactor Coolant Systen (RCS) operating window. As identified in the Reactor Vessel Service Life Evaluation for the Baltimore Gas & Electric Company Calvert Cliffs Units 1 &
2,(I) the operating window for Calvert Cliffs Unit 1 is anticipated to close early in plant life given the implementation of Regulatory Guide 1.99 Revision 02(2) as the accepted method for predicting irradiation embrittlement of low alloy steel.
This report provides a data base of LTOP Pressure Temperature limits for use in establishing LTOP requirements.
Pressure-Temperature limits are provided for plant heatup and cooldown transients, ranging from isothermal to a 100'F/hr rate of temperature change, in 10*F/hr increments.
These limits have been developed using Linear Elastic Fracture Mechanics (LEFM) and state of-the art techniques based upon finite element methods.
I 0
1-1
816, doc (8978)bh 6 Section
2.0 BACKGROUND
The application of the guidance for LTOP found in Standard Review Plan 5.2.2(3) coupled with the Pressure-Temperature limits for normal operation has resulted in operational difficulties. These difficulties have manifested
)
themselves as severely reduced or completely closed operating space, reduced plant heatup and cooldown rates, and as increased administrative restrictions.
These difficulties have been exacerbated by a greater shift in the normal operation pressure temperature limits toward higher minimum temperatures due to the material damage prediction methodology provided by Regulatory Guide 1.99 Revision 02. To partially offset the negative impact of Regulatory Guide 1.99 Revision 02, smaller heatup and cooldown rates, and improved Reactor Coolant Pump (RCP) operating curves have been utilized.
Although these refinements have been made, the RCS operating window can still become prohibitive to effective plant operation.
To address the operating window difficulties, Baltimore Gas & Electric, along with the Combustion Engineering Owners Group (CEOG) and CE, undertook a task to develop an analytical basis for re-opening the RCS operating window. An analytical solution was chosen since it permitted the use of updated and technically superior methods, i.e., detailed finite element linear elastic fracture mechanics and heat transfer methods versus the ASME Code Section 111 Appendix G approximations.
The details and technical justifications for the analytical solution are provided in CEOG report CEN 381-P, Reference 4.
Salient points of Reference 4 are provided below and in Section 3.0, I
Description of Method.
The approach documented within utilizes an alternate pressure-temperature limit for use only in LTOP analyses. The alternate pressure temperature.
limit, known as the low temperature transient pressure-temperature limit, provides the basis for determination of the LTOP relief valve setpoint and enable temperatures.
The low temperature transient pressure-temperature limits are based upon the detailed LEFM procedure described in Section 3.0.
2-1
$16. doc (8978)bh 7 j
Use of the low temperature transient pressure-temperature limit provides an appropriately conservative limit with respect to precluding crack initiation.
l These limits would not be exceeded during a low temperature overpressurization
]
~ transient since the LTOP system prevents exceeding the low temperature overpressurization transient pre.ssure-temperature' limits.
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816edoc(8978)bho8 i
Section 3 l
1 DESCRIPTION OF METHOD j
j
3.0 DESCRIPTION
OF METHOD 1
The approach described within is an analytical solution to re open the RCS operating window. An analytical solution was chosen since it permits the use of updated and technically superior metnods, 1.e., detailed Linear Elastic Fracture Mechanics procedures versus ASME Section !!!, Appendix G approximations (S), and Regulatory Guide 1.99 Revision 02 shift predictions versus Regulatory Guide 1.99 l
Revision 01, without imposing significant operational difficulty.
In addition, the analytical approach provides a superior fracture mechanies justification for LTOP system enable temperatures and the LTOP system pressure setpoints.
The approach utilizes an alternate pressure-temperature limit for i
use oniv in LTOP analyses.
The alternate pressure-temperature limit, known as the low temperature transient pressure-temperature limit, would provide the basis for the LTOP relief valve setpoint and enable temperature. The low temperature transient pressure-temperature limit is based upon detailed LEFM finite element analyses.
Use of the low temperature transient pressure-temperature limit provides an appropriately conservative limit with respect to precluding crack initiation and which would be protected by the LTOP system, it is emphasized that this limit is only for setting LTOP system relief valve setpoints and enable I
temperatures and represents the maximum permissible pressure for any low temperature transient event. However, the low temperature transient P-T limit is not used as an operational limit.
It is a L
limit used for analytical purposes for establishing relief valve setpoints to provide conservative protection of the reactor pressure l
vessel from overpressurization during a potential low temperature transient.
At all times during normal operation steady-state f
conditions and during heatup and cooldown evolutions, the RCS would still be required to operate below the limits prescribed by 10 CFR Part 50, Appendix G(6) 3-1
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816. doc (8978)bh 9 3
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The development of a pressure-temperature limit for low temperature transients is consistent with ASME Code philosophy. Appendix C to Section XI of the ASME CodeU) established a low temperature overpressure transient screening criteria, which recognized the need for an alternate limit for evaluation of low temperature transient events. The LTOP P-T limit method used the screening criteria as a precedent and as a basis to build upon. However, as described in Section 4, the LEFM method for low temperature transients is more appropriate than the ASME Section XI screening criteria due to the explicit treatment of clad and residual stresses over a complete range of flaw sizes. Appropriately, the low temperature transient-P T limit conservatively ensures preclusion of crack. initiation by 3
accounting for all loadings over a wide range of flaw sizes.
Therefore, these limits are also appropriately used as the post-event screening criteria for low temperature overpressure transients as opposed to the ASME Section XI screening criteria for low temperature transients.
This approach provides consistency between the pressure temperature limit which is protected during potential transients by the LTOP relief valves, and the screening method for post-event evaluation.
L Application of the definition of the L13/ enable temperature provided in Revision 02 to Standard Review Plan 5.2.2 does not sufficiently open the RCS operating window at Calvert Cliffs Unit 1.
However, application of the low temperature overpressurization j
transient P-T limit methodology resolves the concern of sufficient operating space and permits re-opening of the RCS operating window.
In addition, the approach permits operation with improved RCS subcooling, is likely to improve RCP seal performance due to operation further from the minimum seal pressure limit, is likely to reduce inadvertent actuation of the LTOP relief valves by providing t-sufficient operating margin to the setpoint, and provides an improved limit for post-event evaluations.
3-2
[
816. doc (8978)bh10 3.1 ASSUMPTIONS OF ANALYSIS 3.1.1 Calvert Cliffs Unit 1 Vessel Geometry The dimensions of the Calvert Cliffs Unit i vessel are given below.
Inside Radius of Cladding 86.9 Inches Inside Radius of Base Metal 87.2 Inches Outside Radius 95.8 Inches C1ad. Thickness 0.3125 Inches Base Metal Thickness 8.625 Inches a
Radius to Thickness Ratio 10.1 3.1.2 Lqadina Conditions The Calvert Cliffs Unit I vessel geometry was analyzed for general loading conditions including internal pressure and thermal gradients. The internal pressure case considered uniform pressure on the inside wall of the vessel acting on the cladding surface.
For this case, pressure on the crack face was also assumed to be present.
The thermal transients analyzed consist of an isothermal event, and a set of cooling and heating transients.
These events range from 10 degrees F per hour to 100 degrees F per hour. _ The transients were assumed to have a constant linear heatup or cooldown rate throughout the event. These rates cover the typical LTOP transients which are significantly below a 100'F/hr rate of temperature change.
l i
l Apart from the internal pressure on the vessel inside surface and the thermal gradient ' loadings, the loadings that are often ignored i
in the ASME Code calculations were also explicitly included in the study.
These loadings include' crack face pressure for the inside i
crack, weld residual stresses in the base metal welds, and the cladding induced stresses.
1 33
816. doc (8978)bh ll
]
l
,g In the crack face pressure case, the same values as the vessel i
' internal. pressure are assumed to act on the crack face from the vessel inside surface to the crack tip.
This is a realistic loading i
that is present since the crack face is pressurized and will cause additional stress intensity at the crack tip. For the vessel b
geometry under consideration-. crack face pressure will increase the stress intensity factor (SIF) by approximately'10 percent and is i
considered significant at low pressure and low temperature. The l
y crack tip SIF is explicitly computed for all the crack sizes and added at the time of the allowable pressure computations..The weld residual stresses that are induced at the time of fabrication of vessels from plates are also often neglected. These stresses also cause additional crack tip loadings that are significant especially for small cracks. The residual stresses are generally tensile on the inside and outside surfaces of the vessel and are compressive in the midwall. The stress profile _ is a self-equilibrating type and i
closely approximates a cosinusoidal shape over the wall thickness.
The crack tip Sif was computed for all of the assumed crack sizes for unit loadings and then was included in the evaluation with the appropriate. peak residual stress of 8 ksi. The peak residual stress c
magnitude assumed was based on previous measurements by C-E on Reactor Pressure Vessel geometries.
The stainless steel cladding on the inside surface of the vessel also causes significant stresses near the vessel inside surface due l
to large temperature variations from operating (550*F) to shutdown (100'F) temperatures. These stresses are caused by the difference in thermal expansion coefficients of the stainless steel cladding
[
and the carbon steel base material. The differential thermal l
expansion stresses are often called cladding residual stresses since they are permanent in nature once the vessel is cooled down to the shutdown temperature.
The cladding stresses are also often neglected in the traditional approaches due to the complexity they pose in the analysis.
In order to understand their effects, the cladding was explicitly 3-4
.I
^
' 816, doc (8978)bh-12 1
i modeled in-the finite element method and the crack tip SIF computed.
)
The unit thermal load cases that account for the cladding include the unit temperature drop and linear gradient in cladding material, j
These SIF's are then added to the base metal SIF's through a polynomial approximation to the temperature profile.
I i
3.1.3 Flaw Sizes This analysis exemined the effect of different hypothetical flaw sizes ranging from small sizes such as through clad.to a depth of 10% through the wall.
Large flaws may not always result in the limiting cases especially when weld residual stresses and the
^
cladding induced stresses are present.
It is the most limiting flaw case that is of importance in computing the pressure temperature limits and asses *,ing the margins available between the actual limits t
and those' set for operation of the plant.
Consequently, a range of flaw sizes were analyzed and most limiting cases provide the basis for the composite P T limits.
The flaw sizes analyzed in this study are semi-elliptical surface flaws oriented in the axial direction. The aspect ratio for the flaws analyzed is one to six. The aspect ratio provides the ratio between the depth and length of the analyzed flaw. The flaw sizes analyzed in the study include flaw depths on the order of the clad thickness up to ten percent through the vessel thickness from the inside surface.
Flaw sizes on the outside radius of the vessel had assumed depths of 10% through the vessel wall.
l The solutions for all the crack cases were enveloped, unless l
mentioned otherwise, to obtain the most limiting cases.
3.2 HEAT TRANSFER ANALYSIS i
L Prior to fracture mechanics analysis of the vessel geometry, a series of one-dimensional radial thermal analyses of the vessel were 3-5 L
816. doc (8978)bh-13 performed to obtain the temperature distribution through the wall as a function of transient time. The temperature history was computed using an axisymmetric one dimensional radial heat conduction -
convection type finite element analysis. A special purpose computer code was developed forethe vessel geometry consisting of three-noded isoparametric elements. The computer program essentially solves the heat conduction equation, 1 E(kr O)=pC N r or or p ar Subjected to the boundary condition on the inner surface of q
h(i T) c 3
i Temperature in the vessel wall, Degrees F where:
T(r,t)
Thermal conductivity, BTU /sec-inch 'F k
Specific Heat, BTV/lb *F C
=
p 3
Density,lb/ inch
=
p Radial distance through the wall, inches r
Time, seconds t
2 Heat flux rate, BTU /sec inch q
=
Surface heat transfer film coefficient, b
=
2 BTV/sec-inch,.7 Coolant temperature, 'F T
=
c Vessel Surface metal temperature, 'F g
and T
s l
Using the shape functions N(r) for temperatures in the isoparametric formulation for the finite element, T(r,t) = [N(r)) (T (t))
n i
l with T (t) as the nodal temperatures and T(r,t) a:: the temperature L
n field, the above equation can be reduced to, l
l 36
816. doc (8978)bh.14 l
i
[K)(Tn)+[C](h) - (Q (t))
I where [K) and [C] represent the thermal conductivity and capacitance matrices, and (Q) represents the external heat flux.
The heat transfer code, specifically written for the LTOP P-T limit analysis, solves the thermal transient problem for the vessel with heat ccnvection on the inside surface and insulation boundary corditions on the outside surface. The mesh spacing can be varied especially when the cladding is considered. The heat transfer analyses always consider the cladding in the model. The first two cladding element nodes are also included in the fracture mechanics analysis.
The computer code assumes the thermal properties K, p, and C of the p
cladding and base material and the film coefficient to be constants with temperature.
However, over the temperature range of interest, 50 550'F, these properties do not vary significantly and use of the -
average values was considered to be acceptable engineering practice.
The heat transfer material properties assumed in the analysis are as follows:
l Cladding Type 304 Vessel Base Material Stainless SA533 Grade B Class 1 Steel
[
?
Thermal conductivity 24 10 (BTV/hr-ft'F)
Specific heat 0.12
'0.12 (BTV/lb 'F) 489 489 Density 3)
(lb/ft 2
Surface heat transfer film coefficient = 1000 BTU /hr ft 'F l
3-7
e 816. doc (8978)bhol5 The analysis boundary condition on the vessel inside surface has an applied film coefficient representing the heat transfer effects i
between the tulk coolant and the metal surface. The coolant temperature was assumed to be a function of time and input in the l
fctm of a table in the analysis.
The analysis results from the heat transfer study consist of the temperature history at all the nodal points in the vessel wall cylinder at the time intervals requested. These are catalogued in a file and are read into the fracture mechanics analysis.
3.3 FRACTURE MECHANICS INFLUENCE COEFFICIENTS Several finite elenient models of the Calvert Cliffs Unit I vessel geometry with different crack depth were developed to compile the influence coefficients. A set of models were used that includec cladding as a separate layer of elements on the inside surface of the vessel. The models consisted of several 8 noded isoparametric plane strain finite elements with number of elements ranging from 68 to 89 and nodes from 259 to 324.
At the crack tip, special quarterpoint crack tip elements were employed to simulate the singularity present in the stress field due to the presence of the
{
crack. The cracktip elements provide excellent accuracy in calculating the stress intensity factors due to various loadings on the vessel.
y The technical approach used in developing the influence functions in
/
the LTOP study is somewhat different from those of the traditional approaches.
Typically, the traditional approaches compute the cracktip stress intensity factors either from handbook formulae or from the influence functions based on the stress profile in the uncracked structures.
In the LTOP study, the influence function approach is carried one step further in the analysis, that is, to influence coefficients based on the temperatures instead of stresses.
This alternate procedure, while providing the same 3-8
~
816; doc (8978)bh-16 l
accuracy as the former one, saves the stress analysis step for each
- )
L time increment in the heatup/cooldown transient. This facilitates an order of magnitude reduction in the computational process and also helps understand the complexities involved in studying the effects of cit:dding.
For example,. it was found that the average
[
difference.in the cladding ard base metal temperature is of significance when compared to the gradient within the cladding l
itself.
The temperature profile across the vessel wall was assumed to be of i
the form,
+ $c
.O T
+
T(r) - T4 = aTuniform + aTcladding c
+ (f 1)3
( f
- 1) T e(
- 1) T T
b b
b b
b the metal temperature at radius r, where T(r)
=
the initial uniform metal temperature, T
=
4 uniform drop in the vessel temperature, AT
=
uniform.
cladding
,- average uniform drop in cladding AT compared to base metal',
average temperature gradient in T
=
l
- cladding, linear, quadratic and cubic gradient Tb, Tb, T b
=
components of temperature in base
- metal, 1
distance along the wall measured from l
x
=
the center of cladding thickness, distance along the wall measured from y
=
I the inside surface of base metal wall, cladding and base metal wall and t,t
=
e b
thickness
+
3-9 l
l
I.
-816 doc (8978)bh 17 i
The influence coefficients were computed for the unit values of uni forr.,- Tcladding, T'e, T'b, T"b, and Tb.
These were combined using a polynomial fit to the temperature profile to calculate the i
total cracktip SIF's. A sample case was verified for accuracy using the traditional approach and the new approach and were found to be 4
in excellent agreement with each other, For other mechanical loadings, unit values of these load
.l distributions were used to compute the SIF and then appropriately j
combined at the time of final evaluation.
The values chosen for j
internal pressure and crack face pressure were 1000 psi each.
For residual stresses, a cosine wave distribution with a peak tensile
~
magnitude of 1000 psi on the inside and outside surfaces and a 1000 l
psi compressive stress at the midwall in the base metal were assumed. All the SIF's were cataloged and used appropriately at the time of transient evaluation.
3.4 FRACTURE MECHANICS EVALVATION 3.4,1 Fracture Touchness Once all the influence functions for the cracktip SIF's were 4
developed and the heat transfer transients analyzed for the heatup and cooldown cases, the fracture mechanics evaluation was performed.
For each time step selected in each transient, a polynomial fit was calculated for the temperature profile through the wall. Then the crack tip SIF was compiled for the thermal case.
The allowable pressure was then evaluated using,
_.g '
Kith + E!res + (SF) KIpr IKlall where K,)) was taken as KIc, at crack initiation for LTOP events.
g The fracture toughness value assumed is given by, 3-10
~
816.dbc(8978)bh-18 i
MI+N KIc = 33.2 + 2.806 e
, Ksi (inch)I where T' - T - RTNDT (deg.F), and T is the metal temperature at the-crack tip and RT is the cracktip adjusted reference temperature.
NDT The crack sizes in the evaluation included those cracks with a depth on the order of the clad thickness. When a conservative surface fracture toughness of the material is taken into account (approximately seventy percent of the actual surface toughness), the through clad crack is no longer a flaw size of concern.
Recent work j
sponsored by the USNRC has also shown that the heat affected zone metal appears to have superior transition temperature toughness compared to the base metal.8 This work-further supports the increased. fracture toughness properties at the surface of the material modeled in the fracture mechanics analyses.
The evaluation utilized a 264.58'F adjusted reference temperature at the inside surface of the reactor vessel corresponding to 16 EFPY.
The Regulatory Guide 1.99 Revision' 02 shift prediction and attenuation formula were used in these analyses to predict the adjusted reference temperatures at the various crack depths through the vessel wall.
i 3.4.2 Safety Factors The safety factors assumed in the ASME Section III Appendix G method are 2.0 for pressure loads and 1.0 for thermal loads for a single crack depth of 1/4 through the wall thickness. This case is represented by the following expression:
2K 2.0 Kipr + Kith la in the above case, base metal weld residual, crack face pressure, and cladding induced stresses are ignored and presumed to ve i
3 11
1816edoc(8978)bh-l'9 I
compensated by the safety margin on the pressure loading and by the use of the crack arrest toughness, K),, rather than the crack The present analysis, while accounting j
initiation toughness, KIc.
for all mechanical and thermal loadings, also uses different crack I
sizes including those on the order of clad thickness in depth and assumes K as the relevant parameter. This case is represented by i
ic the following expression:
('SF) K IEle ipr
- Kith + Elres includes where K includes crack face pressure effects and Kith Ipr cladding stress effects. A safety factor of 1.2 was utilized in the development of the Calvert Cliffs Unit 1 LTOP pressure-temperature limits.
Since all the loading cases are accurately accounted for and the f
was more critical crack sizes (small ones) were included, Kyc considered the appropriate fracture toughness parameter since it is the initiation that is being addressed.
The safety factor (SF)
'i indicates the real margin available under these composite loading conditions.
Even though this factor is somewhat lower than that used in the Appendix G curves, it is considered appropriate since all loading conditions and a range of flaw sizes were considered in the analysis and it is in the same range of that suggested by the I
ASME Section XI Appendix E screening criteria.
1 i
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3 12 I
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816. doc (8978)bh-20 Section 4.0 RECOMMENDED APPLICATION This report provides low temperature overpressurization transient P T limits for use in establishing LTOP requirements.
Combustion Engineering recommends using the enclosed P-T limits for analytical purposes only and not for establishing limits for normal heatup and cooldown operations. The limits for normal heatup and cooldown operations are those developed in accordance with 10 CFR 50 Appendix G as supplemented by the recommended procedures of Appendix G to Section 111 of the ASME Boiler and Pressure Vessel Code.
1 Combustion Engineering recommends that the LTOP system relief valve setpint E
be determined su that the maximum transient pressures for the limiting mais and energy addition transients do not exceed the allowable pressure limits at the reactor vessel for the chosen cooldown and heatep rates indicated in Figures 1 and 2.
In addition, Combustion Engineering recommends that the LTP enable temperature be established using Figures 1 and 2.
The LTOP enable temperature should be the temperature corresponding to the primary system safety valve pressure setpoint for the chosen cooldown and heatup limit indicated in Figures 1 and 2.
For temperatures greater than the LTOP enable l
temperature, the primary safety valves provide protection against exceeding i
the chosen P T limit used as the basis for low Temperature Overpressure Protection.
W Also, Combustion Engineering recommends that continuity be maintained between the(LTOP requirements and the Appendix G P T limits.
For example, if the LTOP analysis indicates that 10'F/br, 30'F/hr and 50*F/hr heatup rates are permitted over certain temperature ranges, then during normal operation heatup, the operator must maintain system pressure within the allowable heatup limits prescribed by Appendix G for 10'F/hr, 30*F/hr, and 50*F/hr over the same specified temperature ranges.
Since the LTOP relief valve setpoint.could be set above some of the Appendix G limits at low temperatures, administrative controls are required to prevent operation above the Appendix G limits during normal operation heatup and cooldown evolutions.
41
g 816.d2c(8978)bho21 py It is important to note that the low temperature overpressurization transient
- pressure-temperature limits of this report are for analytical purposes for
' determining the LTOP requirements associated with the mass and energy addition -
transients. The normal operation pressure temperature limits prescribed by Appendix G are required to be adhered to during normal plant heatup and cooldown evolutions.
E S
e i
i 4-2 i
816. doc (8978)bh-22 Section 5.0
[
f G E TJ Low Temperature Overpressure Protection Pressure Temperature limits have been developed for Calvert Cliffs Unit 1.
These limits have been calculated for heatup and cooldown rates ranging from isothermal to 100'F/hr in 10'F/hr increments.
The results of these analyses are presented in tabular form for cooldown and heatup transients in Table 1 and Table 2, respectively. The results are also presented in graphical form.
Figures 1 and 2 provide the LTOP P-T limits for cooldown, while Figures 3 and 4 provide the LTOP P-T limits for heatup.
The limits presented within this report provide a data base for the establishment of LTOP relief valve setpoints and LTOP administrative controls.
Application of these limits to LTOP, when compared to ASME Code Section !!!
Appendix G P-T limits for normal operation, should provide for significantly improved Reactor Coolant System operating space.
t The recommended application of the LTOP P-T limits is to provide a LTOP relief valve setpoint such that the chosen LTOP P-T limits will not be exceeded during the limiting mass and energy addition transients of interest.
In addition, it is recommended tnat the LTOP enable temperature be established at the temperature associated with the primary safety valve pressure setpoint on the' chosen LTOP P-T limit curve.
l 1
I 1
I l-5-1 L
. ~.
- ~...--.-..
I l
)
1 TABLE 1 i
i BG&E CALVERT CLIFFS UNIT 1 COOLDOWN LTOP P-T LIMIT DATA i
i Ito 8 C00 LING TRAusttW15 10P CSAct 8.Fe1.2
!$0there."*a ne a " "*' ' " a=" a a "**n n o u n CCWL ! be n n u..... u n n. u n o...........
. f *ktet L* loo L*iOF/hr L 20F/hr L 30F/hr L 40F/hr L 50F/hr L*60F/hr L*70F/hr L*40F/hr L 90F/hr L*100F/hr 210
.5M
.435
.346
.257 167
.078
.000 4
200
.M0 451
.362
.273
.183
.ON
.006 190
.557
.48
.378
.290
.201
.113
.023 180
.5 74 485
.397
.308
.220
.131
..M3
.000 i
170
.593
.504 416
.328
.240-
.152
.063
.000 160
.613
.525
.437
.349
.262 174
.087
.000 150
.635
.548
.40
.373
.286
.198
.112
.025 140
.659
.572 485
.398
.312
.28
.140
.054
.000 130
.685
.599
.513 428
.342
.257
.171
.087
.003 120
.715
.630
.545 460
.375
.291
.207
.123
.040 '
.000 110
-.749
.664
.581
.498 413
.331
.248
.1M
.0M
.004 100
.788 704
.622
.M0 458
.376
.295
.215 135
.057
.000 90
.832
.750
.M9
.590
.509 429
.350
.273
.195 118
.044 i
- fl0
.883
.403
.724
.648
.M8 492
.415
.340
.265
.193
.121 70
.M3
.866
.799
.715
.639
.565
.491 419
.344
.279
.212 60 1.013
.939
.866
.795
.723
.651
.582.
.513 447
.382
.320 50 1.097 1.027
.957
.590
.822
.755
.689
.L"
.565
.505 448 40 1.197 1.130 1.065 1.003
.939
.878
.818
.t
.7te
.653
.601 30 1.315 1.254 1.1M 1.139 1.081 1.026
.973
. VJ
.876
.429
.788 20 1.457 1.403 1.350 1.303 1.251 1.204 1.159 1.118 1.0?9 1.M4 1.013 i-L 10 1.628 1.582 1.537 1.499 1.457 1.414 1.384 1.3%
1.327 1.303 1.284 l
0 1.834 1.798 1.763 1.737 1.704 1.678 1.656 t.638 1.624 1.613 1.610 10 2.083 2.058 2.0M 2.024 2.005 1.992 1.985 1.983 1.987 1.9M 2.002 20 2.344 2.3 75 2.M7 2.3 73 2.368 2.373 2.3M 2.403 2.424 2.454 2.484 l
30 2.750 2.757 2.768 2.796 2.811 2.836 2.869 2.911 2.958 3.005 3.070 4
40 3.125 3.125 3.125 3.125 3.125 3.125 3.125 3.125 3.125 3.125 3.125 50 60 70 80 90 100 110
.g^*
l 120 130 s
140
^M 150 160 170 160 l
l 5-2 l
i
?
TABLE 2 BG&E CALVERT CLIFFS UNIT 1 i
180 8 NSAfib8 78Amtituft.10P CRACE*8.Fe1.2 i
l oothers ; - - - - - - - - - - - - - - - - NEAf l ut " - - - - - - - - '
f*RTST L*lso L*10F/hr L 20F/hr L*30F/hr L 40F/hr L $0F/hr 160F/hr L*70F/hr L 80F/hr L 90FW L 100F/hr 210
.524 200
.540
- 190
.557
- 180
.5 74
.614
.614
.614
- 170
.593
.670
.706
.718
.7M
.735
.734
.734
. 738
.758
.734
- 160
.613-
.698
.38
.792
.414
.831
.M1
.847
.851
.853
.856 150
.635
.722
.795
.44
.882
.908
.927
.MO
.950
.M 7
.M3 140
.659
.746
.826
.888
.936
.973 1.000 1.019, 1.032 1.N3 _
1.050 130
.685
.772
.855
.926
.983 1.028 1.058 1.078 1.093 1.895 1.003 120
.715
.801
.485
.961 1.024
- 1. 0 73 1.103 1.124 1.1M 1.899 1.074 110
.749
.833
.918
.M7 1.065 1.112 1.147 1.176 1.152 1.110 1.077 100
.788
.871
.954 1.034 1.106 1.152 1.190 1.218 1.193 1.134 1.004 90
.832
.913
.996 1.076 1.147 1.193 1.234 1.244 1.241 1.44 1.128 80
.843
.963 1.043 1.123 1.191 1.238 1.278 1.273 1.268 1.268 1.179 e
70
.M3 1.021 1.099 1.178 1.242 1.290 1.313 1.307 1.300 1.2M 1.248 60 1.013 1.088 1.1M 1.241 1.301 1.348 1.357 1.344 1.340 1.333 1.324 50 1.097 1.164 1.241 1.316 1.370 1.416 1.410 1.398 1.388 1.378 1.371 40 1.197 1.2M 1.333 1.403 1.453 1.492 1.475 1,41 1.448 1.436 1.426 30 1.315 1.378 1.441 1.500 1.550 1.575 1.554 1.537 1.5.J 1.505 1.492 20 1.457 1.513 1.571 1.633 1.M7 1.677 1.652 1.629 - 1,608 1.589 1.573 10 1.624 1.677 1.728 1.782 1.000 1.001 1.770 1.742
- 1. 716 1.693 1.671 0
1.834 1.873
-1.916 1.959 1.978 1.953 1.915 1.880 1.848 1.419 1.792 10 2.003 2.112 2.143 2.175 2.182 2.139 2.092 2.069 2.000 1.971 1.M0 20 2.384 2.399 2.414 2.436 2.428 2.365 2.300 2.254 2.206 2.141 2.120 30 2.750 2.748 2.750 2.752 2.719 2.M3 2.572 2.507 2.445 2.300 2.338 40 3.125 3.125 3.125 3.125 3.076 2.981 2.8%
2.815
- 2. 739 2.667 2.603 I
50 3.125 3.097 3.812 2.931 i
M 70 80 90 100 4
110 120 130 l'
140 l
150 160 1 70 180 5-3
FIGURE 1 i
BG&E CALVERT CLIFFS UNIT 1 LTOP P-T LIMITS - CCOLDOWN l
BG & E LTOP STUDY l
ISO & COOLING TRfNSIENTS-10P CRACK-SF-12, Thi.ck Clod L31251 25 e
I -L-Iso
-l I
-L-2Wk i
- L-1 T/hr i
-L-fiF&
-L-5F/hr 2
i
- L-10.-(F._/h_r_..
b I
I I i
Q 15
-l 4
j g
1 I
W t
3 13 w
i E
1
--l--
o-
- ~/,/
5
/
8
/
i t
s/
o/
/
\\
-200
-15 0
-10 0
-50 0
50 100 15 0 25)
T-SRDCITdog F)
Isormio/so/eOnoo d.e n
e FIGURE 2 BG8E CALVERT CLIFFS UNIT 1 LT0? P-T LIMITS - C00LDOWN BG & E LTOP STUDY Ic9 & COOLING TRfUSIDTIS-10P CRACK-ST-1.2, TH,ck Clod L31251 25 7---
- L-Iso 1
L-KFAr
-L-3FAr l
-L-STAR 4
2 r-
- L-7T hr i
- L-9Fh _.
i T
I
'3' l
Q 15 l
l d
l I
l
/
l m
/
/
'd i
/
/
yi i
/,,,/ s
/_
.s
//p-1
/
/
//
I 1
1-1 0
-200
-15 0
~-100
' -50 0
50 1T 150 200 T-SRTICITdeg F1 Is0/10/30/50/70/90,degar a
a
FIGURE 3 BG&E CALVERT CLIFFS UNIT 1 LTOP P-T LIMITS - HEATUP BG & E LTOP STUDY-ISO A HERTING TICSIENTS-10P CRACK-ST-12, -Thi.ck Clod L3125')
i
[
p I
- L-Iso
.if
/'/
l
- L-20FAr i
e/
-L-4(FAr i
i i
'/
-L-6(For t
l
'/
- - L-6(FAr -
2 f-
'i
-L-10lFAr I
'/
t-i i
l t
l I
T
/
t
/
I 13
}
<j.
=
1 1
I.
I i
i I
I 5'
j 8
t i
i i
t
}
s
+
+
~ _ _ -..
-200
-150
-10 0
-50 0
50 100 15 0 200 T-SRTM71Tdeg F) 4 ISO /20/40/60/EK)/1[B.@
FIGURE 4 BG&E-CALVERT CLIFFS UNIT 1 LTOP P-T LIMITS - HEATUP BG & E LTOP STUDY ISO A HERTING TRANSIDITS-10P CRPCK--SF-1.2, Thi,ck Clod (.31251 25 2
1 i. -L-Iso 1
i/
' -L-MF/hr i
e'
- L-3(F/hr i
./
-L-5[F/hr I
^
/
- - L-75/br-2
/
L-L-fMFAr I.
m i
/
b
~
r_,
15 t
b
/
I
?
l l
g.# /
w 2
m i
i i
i li t
0
-l-------
-I----
I d
-200
-150
-100
-50 0
50 100 15 0 200 T-SRTNDTideg n IS0n0/30/ son 0/90 daytr
,(816 doc (8978)bh235 60 '.
Section 6.0
.6 REFERENCES 1.
Reactor Vessel Service Life Evaluation for Baltimore Gas & Electric-Ccmpeif Calvert Cliffs Units 1 & 2, prepared by_ Combustion _ Engineering, Inc.,: Janusry 1988.
3; s
2.
U. S. Nuclear Regulatory Commission, Regulatory Guide 1.99, Revision 02,
,;)
" Radiation Embrittlement of Reactor Vessel _ Materials," May 1988.
[
+
3.
U. S. Nuclear Regulatory Commission Standard Review Plan, 5.2.2, 50-
" Overpressure Protection," Revision 02, November 1988, 4.
CEN-381 P, " Low Temperature Overpressurization Transient Pressure-Temperature Limit for Determination of Low Temperature Overpressu e Protection Setpoints,' prepared by Combustion Engineering,
-1 Inc., Dece:1ber 1988.
5.
ASME Boiler and Pressure Vessel Code Section III Appendix G, " Protection Against Non-ductile' Failure," 1988 Edition.
g 6.
Code of Federal Regulations,10CFR Part 50 Appendix G, " Fracture it Toughness Requirements for Light-Water Nuclear Power Reactors,"
January 1989.
..e.
^'*
7.
ASME Boiler and Pressure Vessel Code Section XI Appendix E, " Evaluation of Unanticipated Operating Events," Winter Addenda, 1986.
8.
NUREG/CR 5207, " Fracture Evaluation of Surface Cracks Embedded in Reactor Vessel Cladding," September 1988.
e6 l
6-1
. -