ML20244D601
| ML20244D601 | |
| Person / Time | |
|---|---|
| Site: | Trojan File:Portland General Electric icon.png |
| Issue date: | 04/14/1989 |
| From: | Cockfield D PORTLAND GENERAL ELECTRIC CO. |
| To: | NRC OFFICE OF INFORMATION RESOURCES MANAGEMENT (IRM) |
| References | |
| 2926W, NUDOCS 8904240025 | |
| Download: ML20244D601 (79) | |
Text
-
Portland General ElectricCOiway
~
David W. Cockfield Vice President, Nuclear April 14, 1989 Trojan Nuclear Plant Docket 50-344 License NPF-1 U.S. Nuclear Regulatory Commission Attn
- Document Control Desk Washington DC 20555
Dear Sirs:
Trojan Nuclear Plant Long-Term Pipe Support Design Verification Program In February 1989 the Nuclear Regulatory Commission (NRC) reviewed our Long-Term Pipe Support Design Verification Program at the offices of our Architect-Engineer. The NRC requested written responses to six questions following the review. Preliminary answers were developed during the audit and reviewed by the team prior to the exit meeting. Our formal responses to the six questions are provided as Attachments 1 through 6.
During the review, Portland Ceneral Electric Company (PGE) informed the NRC that the due Jates for several actions committed to you in our letter of January 20, 1989, " Design Verification Program," required extension.
The new dates were discussed with Mr. Roby Bevan and then with Mr. Terence Chan because of their relevance to the review. One of those dates must be further extended. Additionally, several more actions committed to you in the January 20th Ircter also require due date extension. Attachment 7 is a list of the items with extended due dates, showing the current status of each item and providing an explanation for the date extensions not previously discussed with the stat..
We would be pleased to discuss any questions or comments you may have regarding this correspondence.
Sincerely, 8904240025 890414
~
PDR ADOCK 05000344-P PNU 7
Attachments I
c:
Mr. John B. Martin Regional Administrator, Region V l
U.S. Nuclear Regulatory Commission l
Mr. William T. Dixon State of Oregon
. Department of Energy 90 Mr. R. C. Barr i
l g
NRC Resident Inspector Trojan Nuclear Plant l
121 S W Sanon SUeet Portuna Oregon 97204
.i
e Trojen Nuclosr Plant Document Control Dssk Docket 50-344 April 14, 1989 License NPF-1 Page 1 of 2 Question 1 What is the basis for using a ductility ratio of 20 for shear? How does it compare with the strain at shear ultimate? We would like to review your References 7 and 8 from Calculation DV/C-14.
Response to Question 1 The allowable shear ductility ratio of 20 was determined by review of Final Safety Analysis Report (FSAR) commitments, by comparison to the ductility ratio allowed for tension, and by comparison to the reserve ductility margin inherent in the tension ductility allowable.
Each of these three factors is discussed below.
Final Safety Analysis Report Commitments:
In accordance with Section 3.6.1.4 of Trojan's FSAR:
"A maximum ductility ratio of 20 is allowed.
While the FSAR does not explicitly state that this ductility allowable is to be used for both tension and shear conditions, the use of the same value, 20, for both conditions is consistent with the energy balance methodology document cited in the FSAR, Architect-Engineer (AE) Topical Report BC-TOP-9, Revision 1, " Design of Structures for Missile Impact." BC-TOP-9 recommends allowablo ductilities of 20 for both flexural tension and for shear.
Copies of the applicable pages of the FSAR and BC-TOP-9 are provided as Attachments 1A and 1B.
Allowable tension ductility:
Trojan's pipe whip design criteria allows a flexural tension ductility of 20.
The same value, 20, was used to evaluate shear ductility.
This is appropriate for two reasons.
It is consistent with the energy balance methodology as presented in BC-TOP-9 (see Attachment IB), and it provides approximately the same reserve margin as provided for tension (see discussion below).
Reserve ductility margins:
For tension, the ratio of ultimate strain (cult) to yield strain (cy) is approximately:
y = o /E = 36 ksi/29,000 ksi = 0.0012 in./in.
c y
Where oy = tensile yield stress E = modulus of elasticity
l c'
Trojtn Nuclsar Plant.
Docum nt Control D2sk
. Docket-50-344 April 14, 1989 License NPF-1<
Page 2 of 2 c lt = 0.30 in./in. (from Attachments.1C and ID),
u ultimate ductility = c lt 8/ y = 0.30/0.0012 e'250 u
. Ductility margin = 250/20 = 12.5 allowable ductility t For shear, the ratio of ultimate strain (v lt) to yield strain u
(v ) is approximately:
y vy = Ty/G'= 16.5 ksi/11,500 ksi = 0.0014 in./in.
l
.Where Ty = shear yield stress l
G = shear modulus
~
v lt = 0.30 in./in, (from Attachaient 1E) u ultimate ductility = v lt U/ y = 0.30/0.0014 e 215 u
Du'etility margin = 215/20 = 10.8
' allowable ductility t Thus, ductility margins in excess of 10 are provided in Trojan's pipe whip restraint (PWR) designs. The use of an allowable shear ductility ratio of 20 is a reasonable value.
The applicable pages of References 7 and 8 from Calculation 11760-DV/G-14 are provided as Attachments ID and 1E, respectively.
i
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4 LGD/ deb 2926W
ATTACHMENT 1A (4 pages) 3.6 CRITERIA FOR PROTECTION AGAINST DYNAMIC EFFECTS ASSOCIATED WITH THE POSTULATED RUPTURE OF PIPING The design and construction of the Trojan Nuclear Plant has included special ef forts to protect the public against the consequences of major mechanical accidents, including a design basis Loss-of-Coolant Accident (LOCA). Design measures have been taken to assure that the Containment structure and all essential equipment inside or outside of the Contain-dent
, including components of the reactor coolant pressure boundary and other safety-related components, are adequately protected against the ef fects of jet impingement and pipe whip resulting from postulated rupture of piping located either inside or outside Containment. The criteria utilized in the analysis of the different pipe rupture dynamic effects both inside and outside Containment ir summarized in Table 3.6-8.
This section discusses the analyses of pipe system breaks inside Contain-ment.
It defines the extent of allowable mechanical damage resulting from the pipe rupture, the various systems and equipment necessary to recover from these accidents and the mechanical provisions for preventing unacceptable extension of the consequences of a LOCA.
In particular, this section describes:
(1) The systems, or portions of systems, in which design basis piping breaks are postulated to occur, (2) The design basis piping break criteria with respect to location in the piping systems, and the pipe break sizes and orientations postulated to occur at these break locations,
[a] The Containment is defined as the Containment structure, liner and penetrations, the steam generator shell, the steam generator steam side instrumentation connections, and the steam, feedwster, blowdown and steam generator drain pipes within the Containment. Etsential equipment, as used herein, includes components and supports of the reactor coolant pressure boundary (see Section 5.1) and Engineered Safety Features (ESF) and their supports.
[b] Reactor coolant pressure boundary is defined in Section 5.1.
j' Amendment 3 3.6-1 (July 1985) 1
i ATTACHMENT 1A 3.6.1.3 Jet Impingement A jet impingement force will result from any of the pipe breaks postu-lated above. The jet impingement force, caused by the momentum change of fluid flowing through the break, is a function of the upstream fluid conditions, fluid enthalpy, source pressure, pipe flow restrictions, friction and dimensions.
Structural barriers, shields around safety-related components, and physical separation by Plant layout have been used in the Trojan design to limit the effects of impingement. Where i
necessary, the jet forces resulting from the pipe break have been com-1 J
puted using a jet model based on the geometry of the piping system and
{
4 l
the characteristics of the fluid forming the jet, as described above.
1 The jet is conservatively assumed to expand at a half-angle of 10 degrees.
The force of the jet at a distance removed from the jet is computed using the effective jet area and effective jet fluid properties at the target.
A homogeneous flow model is assumed for the jet.
The jet impingement forces inside Containment from postulated breaks are insuf ficient to damage structures or safety-ralated piping to preclude 5
a safe shutdown. The important ESF Electrical System consists of redun-dant elements designed to provide reliable power for all necessary equipment during even the most severe emergency situations, including jet impingement.
Electric isolation and physical separation of cables and equipment asJociated with redundant elements of the ESF ensure this reliability.
Section 8.0 contains a detailed description of electrical systems design.
3.6.1.4 Pipe Whip Pipe whip can result from the jet reaction force caused by the rupture f
of high-energy piping. This force is a time dependent function, asymp-totically acquiring a constant value if the blowdown rate is steady.
In practical situations, however, the blowdown rate varies with time along with changes in the nature of blowdown, resulting in a complicated force-time relationship.
In order to simplify the analysis, a conserva-tive assumption has been made that the jet reaction forcing function a
3.6-8 i
ATTACHMENT 1A at a postulated pipa ruptura builds up instantenasusly to its maxirum The situation where this assump--
value and remains conetant thereafter.
tion is not valid is for a cyclic or three step forcing function where corrective factore are applied using the time history approach.
Reference 5 contains a detailed presentation of jet reaction force computation on high-energy fluid system piping, other than for the reactor coolant loops.
Where necessary to prevent pipe whip, restraints are provided with the proper arrangement and spacing to prevent a plastic hinge mechanism (unrestrained rotation) from forming /as a result of the forces associ-ated with a pipe rupture.
The plastic hinge moment, g,isevaluatedfrom Ka I (3.6-1) g =y o
where I = { (R - R ), and R and R are pipe outer and inner radii.
g A plastic section coefficient (K) value of 2.5 has been selected for stainless steel piping based on tests. The yield stress, o, for Type 316 stainless steel is 19,000 psi for 600*F application and 29,000 The applied force in each case is calcu-psi for 120*F application.
I) lated in accordance with the methods outlined in BN-TOP-2 Whipping in bending of a broken stainless steel pipe section does not This has been demonstrated by cause the section to become a missile.
performing bending tests on large and small diameter, heavy-and thin-walled stainless steel pipe (0) 3 A-9
ATTACHMENT 1A
.4 The required pipe whip restraints were located in accordance with the criteria of Regulatory Guide 1.46.
Sections of high-energy piping which did not require restraint were first eliminated from consideration by examining pipe layout to determine where all safety components were protected by physical separation, barriers, or sufficient distance away from any postulated pipe whip about a plastic hinge formed at the restraint.
Finally, the remaining piping runs are restrained to pro-tect against the limited number of pipe breaks postulated as described above.
Once the restraint location has been identified, the restraint is designed using the jet thrust force, determined by the methods presented in Section 2.0 of BN-TOP-2, Revision 1.
This jet thrust force is applied to the restraint as an energy input due to the pipe moving through the gap between the pipe and the restraint. The restraint structure is designed using the energy balance method-described in Section 3.5 of BC-TOP-9, Revision 1 A maximum ductility ratio of 20 is allowed,
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and a dynamic increase factor of 1.2 is applied to the static yield strength. Justification for a dynamic increase factor of 1.2 for struct2:al steel in flexure and tension and 1.25 for concrete in com-pression is found in Chapter 6 of the Air Force Design Manual - Princi-I) ples and Practices for Design of Hardened Structures The Air Force Design Manual discusses the material found in References 9 through 16.
The following subsections deal with the specific arrangement and restraint details of the piping components inside Containment used to assure adequate protection.
3.6.2 PROTECTION CRITERIA FOR RUPTURE OF BRANCH LINES The piping connected to the primary reactor coolant loops (referred to as " branch lines" in this section) is arranged such that, in the event of a rupture of these lines, the Emergency Core Cooling System (ECCS) will be capable of providing the necessary core cooling. Continued operability of the system is assured by restraining the lines, physically separating the branch lines from the safety-related system lines or by Protecting the high-head safety injection lines, as required.
f 3.6-10
ATTACHMENT 1B',
'(2 pages).
t l
TOPICAL REPORT t
Revision 1 l
4 DESIGN OF STRUCTURES FOR MISSILE IMPACT
.i 4
PREPARED BY:
R. B. Linderman M. Fakhari J. V. Rotz
~
E. Thomas G. A. Tuveson G. C. K. Yeh APPROVED BY:
W. A. Brandes r4<sM-d J / /!,
', ~7 F+ 'r. *
' n..? ;.
L. G. Hinkleman 4
BECHTEL POWER CORPORATION lssue Oate: July 1973
. ATTACHMENT:1B Table 4 4 w
.i DUCTILITY. RATIOS
- l (From Reference 28)
Max. Value ofn-Reinforced Concrete
-Flexure
]
.i
' *fs10 Beams p
- Slabs
- f530
,p
' Compression 13 Walls & Columns where A,
s p is the ratio of tensile reinforcement g
=
I A'%
p' is the ratio of compressive reinforcement
=
Steel Elements Members proportioned to preclude lateral and local buckling Flexure, compression and shear 20 Steel columns Proportioned to preclude 1.3 elastic. buckling Members stressed in tension only 0. 5 [e j
y e
= ultimate strain u
e
= yield strain
9 ATTACHMENT IC
?
(2 pages)
STEEL STRUCTURES Design and Behavior CHARLES G. SALMON -
Professor of Civil Engineering The University of Wisconson Madison, Wisconson I
JOHN E. JOHNSON Professor of CivilEngineering The University of Wisconsin Madison, Wisconsin I
i I
INTEXT EDUCATIONAL PUBLISHERS IE Coilege Division of Intext 5
Scranton Son fronensco Toronto London
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4 ATTACHMENT 1C 32 Structural Steels-0.2*r offset -
. Minimum tensile strengin Heat treated Constructional alloy steets. As14 ouencned 100
. and tempered 4Hoy stel Meimum vieid strengtn.
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'i-80 -!'
~ds:
H'9nstrengtn. low-l a
allow Carbon steers.
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A440.A441 A572 !
[ 60 jl
,,, = s0 ksi A
. Carson E
_ g 40 steels; A36 tas i
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,.. m si Eult=.35 '
20 l
0.05 0.10 0.15 0.20 0.2s 0.30 0.35 strein. incha per inen Fig. 2.2.1. Typical stress. strain curves.
A36. Structural Steel (F, = 36 ksi). This is the primary structural steel for construction, replacing the previous primary steel known as A7.
First adopted in 1960 as a steel with more consistent properties than A7.
A36 had limitations imposed on carbon. manganese, sulfur, and phos.
phorus and on silicon in-plates exceeding. I% in, in thickness. The need for.better weldability for bridge construction led to a revision in 1962 to further restrict the carbon and manganese content.' Subsequently.
A36 steel was approved by the U.S. Bureau of Public Roads for welded bridges. Later revisions removed an earlier limitation of an 8 in, maxi-mum thickness (min. E, = 32 ksi for these plates thicker than 8 in.).
Where high strength to weight ratios are not important, and where bulk for rigidity is desired. A36 is usually the best choice. It is easily welded or bolted, and is available in the large variety of standard shapes, as well as in nearly any plate width and thickness.
A245. Flat Rolled Carbon Steel Sheets of Structural Quality (F, = 25 to 33 ksis. This steel has been used for light gage shapes, including steel joists.
. A283. Low and latermediate Tensile Strength Carbon Steel Plate of Struc-tural Quality (F, 6 33 ksil. For use in machine and equipment manu-i l
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Trojtn Nucler Picnt Dscumint Crntrol Dick Docket 50-344 April 14, 1989 i
License NPF-1 Page 1 of 2 Ouestion 2 2.1 STRUDL computer code models of pipe whip restraint (PWR) 11.2 assume pinned boundary conditions at each end. Therefore, there are no moment reactions at the anchorage. On the rock bolt evaluation (Page 10, Step 7), the interaction equation shows that loads are up to the limit of 1.0.
It appears that if the anchorage had been treated as a fixed end, this limit may have been exceeded. Justify the pinned end anchorage assumption.
4 d
1 2.2 Why is plastic section modulus used in the base plate stress check?
l I
Response to Question 2.1 M4/I r/o w/I4
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,S f(2"6 AGE) l 1/" GROUT l
A pin was used to represent the lack of fixity at the rock bolt base plate because this plate is only 6 inches wide. The lever arm, L (see I
sketch above), to resist moments is only about 2-1/2 inches. Thus, this l
connection has very little capability to resist moments.
Under low loading some moment capability will exist, but at the limiting condition (ultimate strength of this PWR), moments attracted by this connection would induce high compression stresses in the grout / concrete and stretching of the rock bolts. The rock bolts are 1 inch in diameter embedded 12 inches deep, so ductile bolt behavior is assured. The net effect would be that any moment would ultimately dissipate, and a pinned condition would result.
Thus, it is appropriate to model this connection as a pinned joint.
Response to Ouestion 2.2 The design of the PWR is such that the PWR's strength is limited by yielding of structural elements in the frame.
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- Trojan Nuc1sar Plant Document Control Dask-Docket 50-344 April 14, 1989 l<
License NPF-1 Page 2 of 2 l
The function of the base plate is to transfer loads from the frame to the anchor bolts..The ability of a base plate to transfer loads is. limited by its plastic section modulus. In the evaluation of the PWR base plates, the full plastic moments were not allowed to develop.
If 'the clastic strength of a base plate were exceeded, but the plastic strength was not exceeded, some yielding of the extreme fibers of the base plate would result.. Partial yielding would not impair the ability of the base plate to transfer loads.
Therefore, it is appropriate to use the plastic modulus When checking-the-adequacy of PWR base plates.
LCD/ml 2926W
Trojen Nucletr Plant Docum:nt Control Desk Docket 50-344 April 14, 1989 License NPF-1 Page 1 of 6 Question 3 The pipe whip restraint (PWR) analysis dealing with the factor of safety of 1.5 applied to the resistance capacity of the PWR.
This increased load was utilized to check the rock bolt anchorage design.
The margin of 1.5 may not be sufficient to check the anchorage when concrete capacity governs because the actual PWR resistance may be higher due to:
1.
Strain rate effects may be higher than the 20 percent (Dynamic Increase Factor (DIF) = 1.20).
2.
Using average yield rather than minimum static yield (=25 percent increase).
3.
The 0.9 factor should not be utilized.
4.
Assumptions made for boundary conditions in the model may be conservative for the frame analysis, but unconservative for anchorage.
5.
Strain rate effect was only used for tension, not shear.
6.
For supports (PWRs) with ductility (p) >~12, the stress-strain curve grows (not flat).
Response to Question 3 Parts 1, 2, 3, 5, and 6 of this question address the issue of whether the 1.5 factor of safety used in the PWR rock bolt designs is sufficient to envelope potential uncertainties in the analysis assumptions. Each of these items will be discussed individually, and then the effects of each of these items will be combined to show that the 1.5 allowance in the PWR designs is adequate.
Part 4 of this question regards the adequacy of modeling certain PWR anchorages containing rock bolts as pinned ends.
This is similar to Question 2.1, except that this question regards the adequacy of pinned end assumptions on a generic basis, rather than just for the specific PWR cited in Question 2.1.
3.1 The mechanical properties of some structural materials are affected by the rate at which straining takes place.
This offect could increase the resistance loads at which the PWRs develop plastic behavior.
This could also result in increased demand loads on the rock bolts.
In the analysis of Trojan's PWRs, a 1.20 strain rate dynamic increase factor was applied to the static yield strength of PWR frame members subject to axial or flexural loads.
This value is consistent with the Final Safety Analysis Report (FSAR) and is well
Trajcn Nucloce Pltnt Document Crntral Desk Docket 50-344 April 14, 1989 License NPF-1 Page 2 of 6 Response to Question 3 (continued) supported by published strain rate data, such as that contained in Attachments 3A, 3B, and 3C.
Strain rate dynamic strength increase factors were conservatively neglected when calculating rock bolt strengths.
Table 3.1, below, summarizes the strain rate data contained in Attachments 3A, 3B, and 3C.
Strain rates for typical Trojan PWRs are in the range of 0.12 to 0.16 in./in./sec. The corresponding maximum response times range ~from 0.005 to 0.010 second. Table 3.1 lists the increase factors for the entire range of strain rates addressed in the attachments, as well as the increase factors corresponding to the strain rates associated with Trojan's PWRs.
The balance of the table lists the increase factors explicitly included in the PWR designs and the potential conservatism or unconservatisms inherent in the methodology used at Trojan.
Strain rate dynamic increase factors for shear loads are addressed in the Response to Question 3.5, below.
TABLE L1 STRAIN RATE EFFECTS ON RfATIC TEktf LE YlELD STENtN OF IMLTERIALE RANGE OF OWERALL INCREASE INCREASE AWERAM AWERAGE NATERIAL AANGE OF AT STRAIN USED IN POTENTIAL POTENTIAL 15 CREASE RATES OF PW DEllas talCONSERVAflWI CDNSERVAflIM INTEREST STEEL SNAPES 1.00 - 1.60 1.20 - 1.30 1.20 1.25/1.20 e 1.0B (MILD STEEL)
ROCK BOLTS 1.20 (INTERMEDIATE 1.00 1.50 1.15 1.25 1.00 STEEL)
CONCRETE IN CsmPREssl0N 1.12 (OR O! AGONAL 1.00 1.60 1.10 1.1$
1.00
-~
TEN 510m ON SMEAR CONE) 3.2 The tensile yield stress used in the PWR analyses is 36.0 ksi, which is the minimum specified yield strength for A36 steel. Actual yield stresses for A36 steel typically range from 40 ksi to 44 ksi. Thus, the resistance loads at which the PWRs develop plastic behavior could increase.
This could also result in increased demand loads on the rock bolts.
Trojcn Nuclear Plent Documsnt Control Dask Docket 50-344 April 14, 1989
)
License NPF-1 l
Page 3 of 6
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Response to'Ouestion 3 (continued) D, excerpted from NUREG/CR-ll61, " Recommended Revisions j
to Nuclear Regulatory Commission Seismic Design Criteria," states i
that actual material strengths are normally higher than specified values.
The average strength increase cited for structural steel is
)
17 percent:
36.0 ksi
- 1.17 = 42.1 ksi.
In order to further quantify the actual yield strength increase for A36 steel, a limited review of Certified Material Test Reports (CMTRs) for steel used at Trojan was performed.
Since the PWRs l
contain mainly wide-flange shapes ranging from W4s to W12s, the
{
review focused on these sizes.
The first 14 members, from different heats, that were identified by scanning Trojan's CMTR files wero selected. Attachment 3E contains a few of the CMTRs selected by this sampling process.
The actual yield strengths of the 14 selected members, obtained from the CMTRs, were analyzed. The j
yield strengths ranged from 38.3 ksi to 48.3 ksi, with an average value of 42.2 ksi. Note that this average is consistent with results cited in Attachment 3D.
Based on the above research, the average potential unconservatism in the PWR analysis is estimated to be:
42.2 ksi / 36.0 ksi = 1.17 3.3 A 0.9 strength reduction factor was applied to the tensile yield strength of PWR frame members. Use of this factor leads to con-servative evaluation of the PWR frames. However, this factor could also lead to underestimating the resistance strength of PWR frames and calculation of unconservative demand loads on the rock bolts.
No unconservatism in Trojan's PWR rock bolt designs resulted. This is because similar reduction factors were applied to the calculated strengths of the rock bolts. A 0.9 factor was used when steel was the rock bolt's governing failure mode, and a 0.85 factor was used when concrete controlled the failure.
Thus, no not unconservatism (or conservatism) was induced by use of the 0.9 factor.
3.5 Strain rato dynamic increase factors for shear were not included in the PWR analyses. As described above for tensile conditions, the properties of some structural materials are affected by the rate at Which straining takes place. While the FSAR does not specify the use of a strain rate dynamic strength increase factor for shear, and the energy balance methodology document, Architect-Engineer (AE)
Topical Report BC-TOP-9, Revision 1, Design of Structures for Missile Impact," recommends a shear factor of 1.0, use of a larger shear strain rate factor would increase the resistance loads at which the PWRs develop plastic behavior. This would also result in increased demand loads on the rock bolts.
_____________a
Trajtn Nuclcir Plcnt Dscumint Crntrol Drk Docket 50-344 April 14, 1989 License NPF-1 Page 4 of 6 Response to Question 3 (continued)
In Trojan's PWR designs, no net unconservatism in the rock bolt designs resulted. This is because shear strain rate dynamic increase factors applicable to the PWR frames would be offset by similar factors which could have been applied to the calculated strength's of the rock bofts.
The shear strain rate data contained in Attachments 3B and 3C has been used to quantify the potential effects of this factor.
Table 3.2, below, summarizes the strr.in rate data contained in the attachments. Strain rates and maximum response times for typical Trojan PWRs are as stated in the Response to Question 3.1, above.
Note that, as explained in c..tachment 3B, the dynamic shear yield strength is proportional to the dynamic tensile yield strength, (ty) dynamic = 0.60 * (a ) dynamic. This is the same value y
used in the Trojan PWR analyses (Ty = oy (3 = 0.60 are directly o).
Thus, strain rate increases In ty y
shown in proportional to the strain rate increases in oy Table 3.1.
TABLE 3.2 STRATH RATE EFFE M ETAT 1C SMEAR STRENGTN OF MATERIALS I
RANGE OF OVERALL INCREASE INCREASE AVERAGE AWRAGE NATERIAL RANGE OF AT STRAIN USED IN POTENTIAL POTENTIAL l# Car $3E RATES OF MA DESIONS UNCONSERVATIEN CONSERVATISM INTEREST STEEL sNAPES 1.00 1.40 1.20 1.30 1.00 1.25
-~
(MILD STEEL)
ROCK BOLTS 1.20 (INTERMEDIATE 1.00 1.50 1.15 1.25 1.00 STEEL) 1 3.6 A36 steel, when subject to a?ik1 tress at ductilities greater than about 12, will experience stt-in aardening. The effects of this strain hardening were not included in the PWR analyses.
Consideration of this factor could increase the resistance loads at which the PWRs develop plastic behavior. This could also result in increased demand loads on the rock bolts.
Strain hardening of A36 steel was neglected because it has only a small effect on the PWR resistance loads. Most PWRs have " stopped" their pipes before reaching a ductility of 12, thus most PWR rock
')
i-Trojen Nuclsar Plant Dccum:nt Control Dask 4
Docket 50-344 April 14, 1989 License NPF-1 Page 5 of 6 Response to Question 3 (continued) bolts are not affected by this factor at all. For the other PWRs, most of their energy absorption has occurred before etrain hardening begins, thus the maximum potential unconservatism is small.
Trojan's PWR design criteria limits ductilities (p) to a maximum value of 20.
As can be seen by review of the A36 steel tensile
)
stress-strain curves in Attachments IC and ID, the curves are relatively flat in the p = 12 to 20 range. Potential unconservatisms can be quantified from these curves:
At y = 12:
Fy = 36.0 ksi At y = 30:
Fy = 41.0 ksi While an average unconservatism is less, the maximum potential unconservatism = 41.0/36.0 = 1.14.
Cumulative Effects:
In the Responses to Questions 3.1, 3.2, 3.3, 3.5, and 3.6, uncertainties in certain aspects of the PWR rock bolt anct.orage designs were quantified. The cumulative effects of all of these factors are as follows:
Potential Potential Unconservatism Conservatism Factor Factor 3.1: Tensile strain 1.08 1.12 rate effects 3.2:
Actual tensile yield strength 1.17 effects 3.3:
0.9 strength
~~
~~
reduction 3.5:
Shear strain 1.25 1.20 rate effects 3.6:
Strain hardening 7,34 effects Cumulative Effect 1.08*7fgj*1.17*1.25*7pg*1.14=143A
.4 Trojin Nuclear Plant Docum:nt Control D2sk Docket 50-344 April 14, 1989 License NPF-1 Page 6 of 6 Response to Question 3 (continued) j i
A factor of safety of 1.50 has been used in the PWR rock bolt designs.
This factor is larger than any of the individual potential uncertainties and even if all of the potential uncertainties identified above are assumed to effect a single PWR at the same time, the cumulative uncertainty is still less than the factor of safety explicitly included in the rock bolt designs:
1.50 > 1.34.
Therefore, the 1.50 factor of safety included in the rock bolt designs provides an adequate design margin.
3.4 In a few cases, PWR anchorages to concrete structures were modeled as pinned ends in the refined analysis calculations. When anchorages are modeled as fixed ends, moments develop which induce tensile loads in the anchor bolts.
Since pinned ends do not attract moments, inappropriate use of a pinned end assumption could lead to calculation of unconservative demand loads on rock bolts.
In the Response to Question 2.1, justification for modeling a rock bolt base plate on PWR 11.2 as a pinned end was provided. This question asks if any other PWR analyses modeled rock bolt base plates as pinned ends, and if so, to provide justification for using the pinned-end assumptions.
A review of all PWR refined analysis calculations was performed.
In only one additional case was a PWR base plate modeled as a pinned end.
This case occurred for PWR 41.6 under conditions similar to those at PWR 11.2.
PWR 41.6 has a very narrow base plate; only 5 inches wide. The rock bolts are located on the centerline of the base plate, thus the very short lever arm, about 2-1/4 inches, provides little capability for resisting moments. The 1 inch diameter rock bolts are embedded more than 8 inches deep, which is sufficient to develop the tensile strength of the rock bolt's steel shank, thus the bolts can stretch, ultimately dissipating any moment attracted by this connection.
Finally, the interaction ratio for PWR 41.6's rock bolts is low, 0.43, thus this connection has the capability of resisting additional loads. Therefore, the use of a pinned end assumption for PWR 41.6 is appropriate.
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ATTACHMENT 3B SWC AFS WC - T DR-6 2-15 8 (11 Pages)
TDR j
62-138 j
2 AIR FORCE DESIGN M ANUAL '
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PRINCIPLES AND PRACTICES FOR DESIGN OF HARDEN.ED STRUCTURES TECHNIC AL DOCUMENTARY REPORT NUMBER AFSWC-TDR-62-138 -
December 1961 i;
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Research Directorate AIR FORCE SPECIAL WEAPONS CENTER!
Air Force Systems Command Kir: land Air Force Base New Mexico Project Number 1080, Task Number 10801 4
(Prepared under Contract AF 19(601)-2390 by N.M. Newmark and J. D. Halttwanger, The Department of Civil Engineering,'
University of Illinois, Urbana, Illinois)
ATTACHMENT 3B CHAPTER 6.
DYNAMIC PROPERTIES OF MATERIALS
6.1 INTRODUCTION
Materlat properties under dynamic loads are of interest In two i
respects. First, the normal static stress-strain relationship may be altered permitting different deformations and energy absorption. Secondly, the dynamic loading may af fect the circumstances under which brittle failure can occur.
Such conditions as severe restraint, residual stresses, discon-tlnultles, flaws, and thickness of materials and Joints must be studied in their interrelation and influence on cracking tendency. Experimentation and study continue, with many questions as yet not answered. The purpose of this chapter is to discuss the problems and to summarize the present state of knowledge of dynamic material behavior.
6.2 METALS 6.2.1 General Olscussion. Metals can be grouped into two classes
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with respect to their behLvior under dynamic loading. In the first class are those metals with continuous unbroken stress-strain curves showing no sharp yielding zone. This group includes all metals with a basic crystal structure which is face-centered cubic, i.e., aluminum, copper, etc., and in addition those steels which are heat-treated or worked until they lose their definite yleid points. Metals of this group do not generally exhibit significant changes In their mechanical properties under dynamic loadings of the type encountered in blast resistant design. The second group Includes those metals which have a body-centered cubic crystalline lattice. This group Ireludes the standard structural alloy steels, etc. Metals of this group show a O
6-1 e
ATTACHMENT 3B marked variation in mechanical properties with c'hanges in rate of loading.
j In studying the influence of rapid loading on the behavior of these latter
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metals two types of investigations have been employed. In one, a constant strain-rate function is applied to the specimen and its rate is varied between tests (Refs. 6-1, 6-2, 6-3).
In the other a loading pulse with a sharp rise is applied to a specimen which has very little mass (Refs. 6-4, 6-5, 6-6).
Under the strain-rate testing, the stresses associated with the Initiation of yleiding are found to increase as the straln-rate Increases.
1 The magnitude of this increase is a direct function of the straln-rate. The few strain-rate tests that have been reported have been constant straln-rate tests.
The other data existing are those from the loading tests. In this case a load causing stresses in excess of those normally associated with I
yleiding is applied rapidly to a test specimen which has little mass. The specimen responds directly to the loading with strains increasing elastically with stress. During this time the loading test is essentla!!y equivalent to a constant strain-rate test. If the peak value of the stress is equal to or greater than the yield stress established by the strain-rate associated with the rate of loading, the material will yield with no delay. If the sttets is less than that value, the straining will stop for some finite time after the load has reached its peak before yleiding commences. There is thus a i
critical value of peak stress for each strain-rate to yleid relationship which determines if the time delay occurs. During this delay time the specimen supports a stress in excess of that conimonly associated with static yielding at a strain corresponding to elastic behavior (Refs. 6-5, 6-6).
This delay time decreases as the excess stress increases until zero delay 6-2 1
ATTACHMENT 3B j
tin 2 is re6ched*at a stress squal to the yleid stress value for the strain-I k
rate applied.
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In Fig. 6-1 the ratio of dynamic yield stress to the static yleid J
stress (values correspoedias to maximum strain-rate permissible under ASTM specification) is plotted against time to initiate yleiding, measured from zero time.
In applying these data to the design of structural systems it I
must be remembered that it is the response of the system which determines the dynamic effect felt by the material. The delayed yield behavior can only be found in essentially massless systems. In actual members possessing mass and susceptible to inertia loading the time-rate effect of delaying the yleid does not ordinarily manifest itself and can be neglected.
Strain-rates govern the dynamic material properties of most systems.
In an actual member, the strain-rate varies with both time and position in the member. For typical members or systems, the response generally carries the member to yleiding at a time when the strain-rate is near the maximum.
Although time to yielding can be computed, sufficient data are not available l
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to determine the ef'ects of this time variation of strain-rate. Thus average values are used, realizing that they are in general conservative.
In the response to dynamic loadings the Modulus of Elasticity of steel has been demonstrated not to change significantly.
6.2.2 Structural Steel. In Fig. 6-2 the dynamic stresses associated with the initiation of yielding are plotted for varying times to yleld. The static or base value of yield stress is taken as that corresponding to a time to yield of one second. At this rate the value of 37 ksi was selected based upon a study by Jackson and Moreland (Ref. 6-7) which shows that approximately
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ATTACHMENT 3B 90 percent of A-7 steel should have a yleid stress in excess of this value.
The approximate time to yleld is in the range 0.2-0.3 times the fundamental parlod of vibration of the mem6er being loaded. Using this relation it can 01 be seen from Fig. 6-2 that for structures with a period of approximately 100 esec or greater a dynamic yleid stress of 45,000-50,000 psl is reasonable, while for structures with a period of less than 100 esec e value greater than 50,000 psl is Indicated. Although this hedges somewhat on the longer porlods, it Is reasoned that the conservatism of using constant strain-rate data and the low static stress allow such a selection. A more refined yleid value might be selected after a detailed analysis of the original dasign; however, this does not seen Justified by the nature of the problem.
The results cited above are considered to apply equally well to
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i steels under the new ASTM Specification A-36 For sheer, numerous data and theoretical studies Indicate that the
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dynamic shear yield value is about 0.6 times the dynamic tension yield value, cnd failure in shear takes place at about 0.75 times the tensile strength.
6.2.3 Fich strength I.m Alloy steels. The results of laboratory tests indicate that the steels with higher static yield stresses do not echieve as high a percentags of Incresse In yloid stress under dynamic lood-Ings as do weaker steels. For the low carbon steels of this group which exhibit the flat yield zone, although little specific test data are available, the yield stress may be Increased slightly as the rate of loading increases.
A !!alted nu=her of straln-rate tests conducted by Jones and Moore (Ref. 6-2) show a flattening out of the dynamic yloid increase at higher strain rates.
6-4
ATTACHMENT 3B For this ress'on, until more complete data become available, increases above 5 percent are questionable. As with A-7 steel, the yield stress In shear is taken as 0.6 times the dynamic yloid stress in tension.
6.2.4 Reinforcing Steel. Behavior of Intermediate and structural grade reinforcing steals at the strain-rates associated with the response to blast loadings is analogous to the behavior of A-7 steel. Figure 6-2 gives the dynamic stresses associated with the Initiation of yleiding.versus the time to yield. For Intermediate gre 4. the static yleid value is taken as 45,000 pol. This value represents an everage value. For smaller bars the l
tendency is toward higher values and fa the larger bars, lower values
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(Ref. 6-23).
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From Fig. 6-2 the following dynamic yield values are Indicated:
Intermediate Grade. fg = 50,000 - 60,000 pel Structural Grade f4 = 40,000 - 50,000 psl Because of the variability of static properties of reinforcing bars, further increases in these values cannot be generally recomunended. For grades of reinforcing steels which do not exhibit definite yield zones, negligible I
dynamic Increases occur, as mentioned in Sect. 6.2.1.
For those grades of reinforcing which have higher initial yleid values than Intermediate grade and finite yield zones, dynamic Increases consistent with the approach used for low alloy steels can be used.
In regard to the dynamic properties of reinforcing bars, the question of the effect of welding is of Interest. There has been recently concluded at the University of Illinois a series of tests in which this question was studied (Ref. 6-24). Tests were made on No. 6 deformed bars 1
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ATTICHMENT 3B
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of two gradas: ASTM Designation A15 (Intermediate Grade) and ASTM Designation A-431 (a high strength steel). Sars of both grades were tested in the as-rolled condition and also after having been Joined by 60-degree, single-Vee butt walds with a 1/8 inch gap between the welded parts. The bars were tested under slowly applied loads as well as under rapidly app!!ed loads. For the rapid load tests, an inflntte duration pulse was used. The~ time required to reach failure in the high strength bar tests varied from 0.005 to 0.012 sec, and the time to yleid for the intermediate bars was about 0.003 to 0.004 sec.
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Interpreting the results of these tests, Ref. 6-24 states that
"...these tests have shown the high-strength reinforcing bars having the same physical and chemical properties of those used in this program can be welded to produce a bar having properties under slow and rapid loading not I
significantly different from those of the as-rolled unwelded bar. On the other hand, these results provide a warning that weld defects, perhaps even minor ones, can greatly reduce the ultimate strength and elongation under rapid loading".
Concerning the tests on the Intermediate grade bars, Ref. 6-24 states that the welded specimens developed yield points under rapid loading from 26 to 32 percent higher then that for the welded bars tested slowly, while the increase in yleid due to rapid loading of the as-rolled bars was usually about 33 percent.
Thus, it may be concluded that if the welds are properly made, butt-welded reinforcing bars behave in essentially the same manner as un-welded bars with both slow and rapid loading. However, weld defects can 6-6
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ATTACHMENT 3B I
tension, but by the nature of their structural function the impact loading i
does not have the same relation to the basic stresses as in the esse of bolts in shear. Thus it seems likely that high-strength bolts in shear would not be as susceptible to brittle failure as other connections. This has been confirmed by the A.R.E.A.'s conclusions that such joints are superl.ir to riveted joints and they are not affected adversely by extremely low temperatures.
Rivets and bolts undergoing impact loads must have consideration given to the materials used, the time-rate of loading anticipated and the lowest ambient temperature of the structural elements. In addition, con-sideration should be given to the type of loading normally carried prior to the anticipated dynan le loading. For example, a member normally subjected to reversals of load may suffer from fatigue and thus become more susceptible to brittle fracture. These considerations determine how critical will be the connection material properties, f abrication methods and worlonanship.
Information on significant properties of the materials and prescribed fabrication procedures are normally available from manufacturers, societies, and institutes concerned with the use of the material in question.
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6.3 CONCRETE Tests reported in Refs. 6-17 end 6-18 show that with increased rates of straining concrete properties vary as shown in Fig. 6-5.
In the referenced tests, two basic concrete strengths were tested. The '%eak" concrete had an f' of approximately 2500 pnt while the " strong" concrete 6-10
ATTACHMENT 3B had an (' of approximately 6000 psi. The curve of Figure 6-5 represents an average of the effect upon these two strengths. In the general region of interest, the increase in the ultimate strength ranges from 20 to 40 Flexural members are generally proportioned such that the reinforcing percent.
J steel' governs the resistance capacity. For such members concrete strength variations of the amounts given above have little or no ef fect upon the resistance capacity. Therefore, there is little necessity for using in-creased concrets flexural strength values for dynamic loads. There may, however, be good reason for using increased compressive strengths in columns and similar members.
It should also be noted that the dynamic increases in compressive strength discussed above should be considered as being applicable to the
. static strength at the time of loadino, not the so-called standard or 28-day i
static strength. Thus, from an attack point of view, it would be proper to take into account an increase in strength due to aging be forc applying the
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dynamic increase factor. For design, it is probably also reasonable to consider at least some increase due to aging; however, the amount of time that should be assumed to axist between the construction of a facility and the date of potential loading by blast is a subjewt of which is beyond the scope of this manual.
Shear properties of concrete should increase under dynamic conditions.
Little data exists on this subject. Because of the brittis nature of this mode of falture, no allowance should be made for any increases that might occur.
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ATTACHMENT 3D (2 pages)
Recommended Revisions to l
Nuclear ' Regulatory Commission Seismic Design Criteria D. W. Coats Project Manager gLawrence Livermore Laboratory Prepared for U.S. Nuclear Regulatory Commission l
April 1988
ATTACHMENT 3D DESIGN CRITERIA MARGIN The basic conservatism that results from the actual J
strength of material being normally higher than specified values is documentd6 in UCID 17965.
The average FOC that
{
results from this effect is 1.17 for steel and 1.27 for
<([
reinforced concrete.
The role of quality assurance pro-grams is maintaining this FOC is discussed.
It would not appear unreasonable to expect this FOC to diminish natural-ly as manufacturing facilities across the country and even around the world become more uniform and delivered more uniform products.
On the other hand there appear to be few advantages to artificially lowering this traditional and easily understood source of conservatism.
UCID 18100, in a nice piece of work, showed that elastic floor spectra may be expected to be generally higher than floor spectra generated from motion containing some plastic action.
In particular, peak responses were lower as was expected.
In some respects this conservatism is Design Criteria Margin and in some respects it is Calculational Margin.
In any event, if other sources of Calculational Margin were eliminated, it would be comforting to know that in an extra severe earthauake the plasticity dampens the floor spectra.
The work reported in " Elastic-Plastic Seismic Analysis of Power Plant Braced Frames" by Nelson and Murray is an j
exploration of another aspect of the plastic reserve strength in nuclear structures.
This study is particularly l
l l
[
189
ATTACHMENT 3E (3 pages)
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- Trojen Nucioar Plant Document Control Desk Docket 50-344 April 14, 1989 License NPF-1 Page 1 of 5 Question 4 Regarding rock anchor bolt allowables, the licensee has not demonstrated that the mode of failure is controlled by ductile behavior of steel. A factor of safety of 4 exists for the tension component of load only.
In the tension-shear ~ interaction formula, a safety factor of 1.33 is used.
To provide additional confidence, the licensee should provide a summary of additional explicit and inherent factors of safety.
Response to Question 4 The use of rock bolts anchoring pipe supports at Trojan has been studied extensively over the last 1-1/2 years.
These studies have developed a clearer understanding of how rock bolts perform and have resulted in a design criteria which is appropriate for their use in pipe support applications. The following discussions summarize the key elements of those studies, the provisions of the new design criteria, and the explicit and inherent factors of safety in Trojan's rock. bolt designs.
In addition to these discussions, an interaction diagram graphically depicting the factors of safety in the criteria and an updated version of Table SB* listing the interaction ratios and tension-only factors of safety for individual pipe supports are provided in Attachments 4A and 4B.
l Performance of Rock Bolts:
Rock bolts are not typical expansion PAchors. While they do employ l
an expansion mechanism at their base, they differ from typical expansion anchors (e.g., Phillips Red-Heads Hilti Kwik-Bolts, and Parabolts) in the following ways:
a.
Rock bolts are made from deformed rebar and are grouted in place after the expansion wedges are set.
This provides assurance that the rock bolts will not fail by a slip-out mechanism. Although no credit is taken in the design calculations, the grouting also provides means of resisting tension loads (through the grout bond) in addition to that provided by the expansion assembly.
I b.
Rock bolts use heavy-duty expansion hardware and are set at high torque values.
l
- Table 5B was contained in Portland General Electric Company's (PGE's) 4 August 18, 1987 submittal to the Nuclear Regulatory Commission (NRC) regarding Trojan's desi;n of pipe supports anchored with rock bolts.
L
k Trojan Nucisar Plant Document Control Dask H
Docket =50-344
' April 14, 1989
, License NPF-1 Page 2 of 5 Response to Question 4 (continued)
The heavy. expansion hardwaro allows the rock bolts to be set at high torque values,' thereby obtaining a very strong " bite" into the concrete, c.
Rock bolt installation is less sensitive to variances in installation workmanship.
The mechanics of the rock bolt expansion, assembly is different from that used by expansion anchors.
For rock bolts, the expansion assembly is constructed such that pulling of the bolt shaft completely through the expansion shell is. precluded. This makes-installation variances in items such as the diameter of the hole drilled in the concrete and the amount of pull-up occurring during the setting process less.significant.
Trojan's Rock. Bolt Design Criteria Trojan's criteria for. design and evaluation of rock. bolts used in pipe supports contains two independent parts.
Part 1:
The combination of shear and tension on a bolt is checked using a straight-line interaction equation.
Part 2:
If the tension capacity of a bolt is controlled by the strength of the concrete (i.e., a nonductile failure mode),
a tension-only load check is made to ensure that a factor of safety of at least 4 is maintained.
Since the shear strength of a rock buit is limited by the shear strength of the bolt's steel shank:
a.
If a rock bolt's tension strength is also controlled by its steel shank's strength, then a ductile failure mode is ensured, b.
If the rock bolt's tension strength is controlled by the concrete cone strength, then shear and tension are controlled by two i
different elements:
the bolt steel for shear and the concrete cone for tension.
In this case, shear-tension interaction i
effects will be less than that predicted by the straight-line j
interaction equation.
This two part criterion was developed and discussed with NRC personnel. PCE understood this criteria to be acceptable to the NRC, and therefore implemented the criteria by revising the design calculations for all pipe supports that use rock bolts.
Pipe supports that did not meet the two part criteria (a total of l
14 supports) were modified during the 1987 and 1988 refueling outages to meet the new criteria.
This work is now complete.
1
i Trojtn Nuclear Plent Docum:nt Control D3sk Docket 50-344 April 14, 1989 License NPF-1 Page 3 of 5 Response to Question 4 (continued)
Factors of Safety in Trojan's Rock Bolt Designs In Part 1 of Trojan's rock bolt design criteria, the shear-tension interaction equation factors of safety are applied to the rock bolt tension.and shear loads.
The major factors of safety for the tension loads are quantified below.
Some of these factors are explicitly included in the interaction equation, while others are inherent in the assumed material strengths. The data presented below are for the faulted Safe Shutdown Earthquake (SSE) load case as this case contains the governing (lowest) safety factors.
Explicit Factors of Safety (FOS)
If the rock bolt's tension strength is controlled by its steel shank's strength:
Tas =
0.9 v*Abolt
- 1.5 1.5 A
I dy bolt 1.11 Where ay = tensile yield strength Abolt = area of bolt shank Thus, the FOS = 1.11.
If the rock bolt's tension strength is controlled by the concrete cone strength:
0.85 (4/f'c) *Aconc
- 1.5 Tac
=
2.0 y, 7(4/f'c) A
=
conc Where f'c = compressive strength of concrete Aconc = area of concrete cone Thus, the FOS = 1.57.
1 l'
L
I Trojan Nucisar Plant Docum:nt Control Dask.
Docket 50-344' April.14, 1989 License NPF-1 Page 4 of 5 1
--Response to'Ouestion 4 (continued)
Inherent Factors of Safety
-If the rock bolt's tension strength is controlled by its steel shank's strength, for 1 inch diameter rock bolts:
Tyleid = 37 kips Tultimate = 50 kips 50/37 = 1.35 Thus, the FOS = 1.35.
If the rock bolt's tension strength is controlled by the concrete cone strength:
The 90-day concrete compressive strength was used to calculate the' pullout shear cone strength.
Concrete, Which is now 15 years old at Trojan, gains strength with age.
As shown in
' Attachments AC and 4D, a strength increase of about 15 percent is expected. Since the pullout strength of concrete is proportional to the square root of the concrete compressive strength, a factor of safety can be quantified as follows:
FOS = /1.15 = 1. 0 7 The pullout strength of the concrete cone is calculated at 4/f'c.
The actual strength is larger due to effects such as confinement of the concrete, presence of rebar, elliptical rather than 45 degree shear cones, and conservatism in the 4 factor. A lower bound on the actual pullout strength can be estimated by comparison of the calculated cone strength to the rock bolt failure test
- performed by PGE:
Based on PGE test:
tension capacity >51 kips Based on 4/f'e
- 1.15 : tension capacity = 35 kips FOS 2 51/35 = 1.46
- Although failure testing was performed on only one bolt, this has significance because (1) nine other proof load tests were performed which did not show any premature failures, and (2) the failure test resulted in failure of the bolt's steel shank, not concrete cone pullout. This indicates that the concrete pullout strength is higher than 51 ksi and a nonductile failure mode did not exist.
1 1_1______________
)
m _..
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LTrojcn,Nuclstr. Plant
.Documint Costrol D3sk; Docket 50-344 April.14, 1989 License NFir-1L Page 5 of 5 Summary.of Tension Factors of Safety in Trojan's--Rock Bolt Interaction
. Equation:
Explicit Inherent' Total FOS FOS FOS' Steel-controlled
,1.11.
1.35-
'1.50 failure Concrete-controlled 1.57
>1.07
- 1.46
- >2.45 failure
= 1.56 l
l l
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L'
. ATTACHMENT 4Ai (1 page)
Trojan 'Nuclearf Plant Rock 1 Bolts Used lIn Pipe Supports
' Design Criteria interaction Envelopes-Tult (steel controlled; concrete Le 2 7.6")
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MATEAIAL.$ ANO $PfCIPICATIONS 5
ATTACHMENT 4D i
DESIGN OF CONCRETE STRLCTERES c
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Trojen Nucicer Plant Docum nt Control Dask Docket 50-344 April 14, 1989 License NPF-1 i
Page 1 of 1 l
1 Question 5 Rock anchor bolt safety factors:
the factor of safety of 4 (for tension loads only) is not currently part of the Civil and Structural Design
)
Criteria 11760-C1.
This requirement should be incorporated into that
{
document.
Response to Question 5
)
k A requirement for maintaining a factor of safety of at least 4 for tension j
loads only will be added to Section 8.1 of the Civil and Structural Design Criteria, Standard 11760-C1.
This action was completed March 9, 1989 (see Attachment SA, Section 8.1.4).
Note specifically that no allowable stress increase factor, S, was used when checking for a minimum factor of safety of 4 against concrete pullout.
l LGD/ml 2926W
ATTACHMENT SA (4 pages) m CIVIL AND STRUCTURAL DESIGN CRITERIA FOR THE TROJAN NUCLEAR PLANT PORTLAND GENERAL ELECTRIC COMPANY PORTLAND, OREGON i
DESIGN CRITERIA DOCUMENTS i
COVER SHEET STANDARD 11760-C1
(
JOB NO:
11760 DISCIPUNE: crvzt/sraccrunAt 5 ' @saa:aW:.Sirar1%^ (4% 4 s.s A M %LA 6 '%/,2 uuds**ao ub.S d42., t& M T %'la n u <6 u e n e.o um% s_ssh.fJklW 8 %k idYs!JO L*d%T:l.?'" 44A aw $.-g R, Pacea0y y-mata navisCN Ds9CRIPTON eRista.
cmstus ATOR appgoyAL i
_'m._._m
._____.._____._..-_.-1_.._.
ATTACHMENT 5A 7.0' PIPE SUPPORTS
(
All pipe support design'.and evaluation of existing civil designed pipe.
supports'shall be performed in accordance with the latest revision of DC-11760-P-003, " Pipe Support Design' Criteria for Trojan Nuclear Plant."
8.0 ROCK BOLTS Evaluation of existing. anchorages which use Hollow-Core Groutable Rebar Rock Bolts (rock bolts), manufactured by Williams Form Engineering Corp., shall be performed in accordance with these provisions.
8.1 Components Subject to Recurring Loads The following criteria are applicable when the anchored components are subject to recurring, repetitive, or cyclic loads. This includes components such as pipe supports and equipment anchorages, which are subject to 1 cads from normal operation and OBE and SSE events.
8.1.1 Allowable Tension Loads:
Ta = allowable tension load
= lesser of Tas and Tac:
Tas = steel controlled tension allowable
= (0.9*Py*S) / 1.5 0.9 = strength reduction factor Py = tensile yield strength of the rock bolt's steel shank:
bolt dia.:
1" l-3/8" 2"
Py:
37k 74k 148k 1.5 = factor of safety for working level loads S
= allowable stress increase factor
= 1.00 for normal loads
= 1.25 for OBE loads
= 1.50 for SSE loads
(
TPF61/5-22 -
Standard 11760-C1 ATTACHMENT 5A 1
l k"
Tac = concret'e cofLtrolled tension allowable
=(0.85*4* fc'*Ac*S) / 2.0 h -( -
0.85
=- strength reduction factor fc' design (or actual) concrete compressive
=
strength 4* % '
concrete pullout strength
=
2.0.
factor of safety for working level loads
=
S allowable stress increase factor i
=
1.00 for normal loads
)
=
1.25 for OBE loads
=
1.50.for SSE loads
=
Ac
- projected area of shear cone
,s
=
W *(le)4 (raduced for spacing & edge 3,
=
distance limitations) le effective embedment length
=
(L - a).
=
L specified (or measured) embedment depth
=
1 to the bottom of the bolt distance from bottom of bolt to bottom a
=
of expansion shell
,{
1 bolt'dia.:
1" 1-3/8" 2"
I a:
0.5" 0.75" 1.0" 8.1 2 Allowable Shear Loads:
(.
Va
= allowable shear load (Vy*3) / 2.0
=
Vy
= shear yield strength of the rock bolt's steel shank = Py / ( I' Py
= bolt tensile yield strength, as. defined above 2.0 = factor of safety for working level loads S
= allowable stress increase factor
= 1.00 for normal loads
= 1.25 for OBE Loads
= 1.50 for SSE loads 8.1.3 Shear and Tension Interaction Straight line interaction shall be used to combine shear and tension loads:
(T / Ta) + (V / Va) < 1.0 T
= demand tension load V
= demand shear load Ta = allowable tension load Va = allowable shear load TPF61/5-23
-22
_u: _ - - _ _
].'.-
Standard 11760-C1
-ATTACHMENT 5A
~
8.1.4~
Factor-of Safety for Tension Loads In addition t'o the interaction' criteria stated above, a minimum factor of safety (FS) of 4 against concrete
~
pullout'shall be provided. Only the tension component of the load shall be used in this criteria:
Pc uit
>4
.FS
=
Pd
'Where Pc, uit 4 *.85
- k fe'*A
=
c largest tension demand load on the bolt (usually Pd
=
from'the SSE load cese) 8.2 Components Subject to One-Time Loads The following criteria are applicable when the: anchored components'are subject:to.one-time loads. This includes components.such as pipe whip restraints and guard pipes, which are subject to loads from extreme
. environmental.or design' basis accident conditions.
8.2.1 Allowable. Tension Loads:
Ta = allowable tension load
= lesser of Tas and Tac:
Tas = steel controlled tension allowable
=(0.9*Pult) / 1.5 0.9
= strength reduction factor..
1.5
= factor of safety Fult = ultimate tensile strength.cf the rock bolt's steel shank:
bolt dia.:
1" l-3/8" 2"
Pult:
50k 100k 200k
(
TPF61/5-24 Tro'jan Nuclear Plant' Documant Control Desk
- i..
[
Dockot" 50-344 April 14, 1989 License NPF-1 Attachment'6 Page 1 of 2 1
l' Question 6
-The refined analysis of pipe whip restraint (PWR) 5.3 did not address the differences in structural members. Original calculations used TS4x4x0.375 members in frame. Refined calculation shows W4x13 members.
This difference i
should have been reconciled.
1 Response to Question 6 The. difference in' member sizes was acknowledged in the' configuration screening process, documented in Calculation 11760-DV/G-7, -Sheet 24.
Because the substitution was allowed (per Note 7 on Drawing 6478-C-399, Sheet 1), and because the difference in section properties between the LTS4x4x3/8 and a W4x13 (with two 3/8-inch side plates) is minor, this substitution alone was not judged to be cause for initiating a refined analysis calculation. The section properties of_the different members, which are repeated below, are documented in the design calculations.
Note that the refined analysis of PWR 5.3 was initiated because a subsequent modification of PWR 5.3's design changed one of the W4x13 members to a 1-inch x 4-inch flat bar. Documentation of the adequacy of this substitution'was judged to be required, and hence a refined analysis of the flat bar was performed.
TS4 and W4 Section Property Comparison:
For a TS4x4x3/8:
from Calculation 6478 H-10, No. 5.3:
Mp = 352 in-kips Where Mp = plastic yield moment For a W4x13 with two 3/8-inch side plates:
from Calculation 11760-DV/G-10:
Mp = 332 in-kips Comparing the M values:
p 352/332 = 1.06 Thus, only a 6 percent difference in Mp resulted from the substitution. This small difference was judged not to be significant enough to require reanalysis of this PWR.
A second example of a TS4 to W4 substitution was also reviewed by the Nuclear Regulatory Commission (NRC) during the audit.
In this case, for PWR 19.1, a TS4x4x1/2 was replaced with a W4x13 (with two 3/8-inch side l
i
Trojan Nuclose Plent Document' Control Dask' 4
Docket 50-344
. April 14, 1989
. License NPF-1.
Page 2 of 2 Response to Question 6. continued
. plates). This. substitution was also judged to have only a minor effect.
The section' property = comparison for this case shows:
For a TS4x4x1/2:
from' Calculation 6478 H-10, No. 19.1:
Mp = 298 in-kips For a W4x13 with two 3/8-inch sido plates:
from Calculation 11760-DV/G-10:
Mp = 332 in-kips 298/332 = 0.89 Thus, only.a'll percent difference in Mp.
resulted from the substitution. This small difference was judged not to be significant enough to require reanalysis of this PWR, Note that the TS4 to W4 substitution was the only substitution allowed on a
" generic" basis.
LGD/ deb 2926W i
1
I
=
j Trojen Nuciter Plent Docum:nt Control D32k Docket 50-344 April 14, 1989 j
License NPF-1 j
Page 1 of 3 0
EXTENDED DUE DATES FOR COMMITTED ACTIONS IN LETTER DATED JANUARY 20. 1989 Original Extended Current Item Description Duo Date Due Date Status ACTIONS WITH PREVIOUSLY EXTENDED DUE DATES An evaluation of tornado missile effects on 02/15/89 04/10/89 C
the Intake Structure will be completed to ensure compliance with Final Safety Analysis Report (FSAR) commitments.
A pseudo-dynamic seismic analysis of the 02/24/89 04/10/89 C
Intake Structure will be completed, and an evaluation of structural-resistance will be made to ensure compliance with the governing codes and criteria.
Calculat2ons will be completed to document 02/24/89 04/10/89 C
the adequacy of the structural design for the bearing pressure at the support points on the separator walls as part of the pseudo-dynamic analysis discussed above.
An evaluation of Nuclear Steam Supply System 02/28/89 03/24/89 C
support anchorages will be completed using final loads of record.
Evaluation of the seismic design of the Main 03/03/89 03/17/89 C
Steam Support Structure (MSSS) will be completed to confirm compliance with FSAR commitments.
Evaluation of the MSSS for tornado missile 02/15/89 04/14/89 C
effects will be completed to confirm com-pliance with FSAR commitments.
Calculation M-6 will be revised to correctly 03/10/89 03/24/89 C
reflect the as-built structural concrete configuration of the Auxiliary Feedwater (AFW) pump room walls.
The AFW pump room structural design drawings 03/01/89 03/10/89 C
will be corrected to reflect the as-built conditions.
Legend:
C - Complete IP - In Progress L - __-___
o Trojen'Nuclser'PIEnt Documtnt Control Dask Docket ~50-344 April 14, 1989 License NPF-1 Page 2 of 3 original Extended Current Item Description Due Date Due Date Status ACTIONS WITH PREVIOUSLY EXTENDED DUE DATES (continued)
'An evaluation of tornado missile effects 02/03/89 03/21/89 C
on buried yard piping and duct banks will be completed to ensure compliance with FSAR commitments.
Civil Structural Design Criteria Cl will be 01/27/89 03/10/89 C
revised to include appropriate tornado load definition, design methods, and load combinations.
Final pressure build-up and jet force loads 03/03/89 06/30/90*
IP from High-Energy Line Break (HELB) of Main Steam and Feedwater lines will be confirmed, and evaluation of the MSSS for HELB effects will be completed to confirm compliance with FSAR commitments.
Analysis of the MSSS anchorage of the upper 03/03/89 06/30/90*
IP chord of the steel tower support trusses will be_ completed to document the structural adequacy of the connection.
Evaluation of the MSSS for tornado wind 02/15/89 06/30/90*
IP effects will be completed to confirm compliance with FSAR commitments.
Legend:
C - Complete IP - In Progress
- The due date was previously extended to 04/14/89. The commitment date i
is being further extended to facilitate the use of a new computer code for analyzing the effects of a HELB.
This new code will be used to predict compartment pressures and temperatures following the rapid introduction of steam into the MSSS.
Several Portland General Electric programs, including environmental qualification, will utilize these results. For reasons of standardization and economics, these results will be used for input into the HELB structural effects analysis which will be performed by the Architect-Engineer (A-E).
It is expected to take four months to obtain the input data for the HELB analysis, six months to perform the actual HELB analysis, and four months for the A-E to complete the structural ef fects analysis of the MSSS.
l
(
Ed' Trojen Nuclear Plant Document Control Dask Docket 50-344 April 14, 1989 License NPF-1 Page 3 of 3 Original Extended Current Item Description Due Date Due Date Status ADDITIONAL ACTIONS REQUILrUG EXTENDED DUE DATES **
~
'FSAR Section 5.4.12.3 will/be revised to refer 07/15/89 08/31/89 IP to Section 3.8.3 for allowable stress levels for the Reactor Coolant System supports.
FSAR Section 3.8.4.1.1.2 will be amended to 07/15/89 08/31/89 IP clarify that the analysis of the capability of the Fuel' Building'to withstand a dropped spent fuel-shipping cask will be in Sec-tion 9.1.4 or 9.1.5 when cask movement into the Fuel Building is permitted.
FSAR Section 2.5.4.1 will be amended and 07/15/89 08/31/89 IP Note 9 on Drawing C-101 will be revised to agree with the design in that not all Seismic Category I structures are. founded on rock.
FSAR Section 3.8.4.6 will be revised to 07/15/89 08/31/89 IP show.3000 psi strength concrete was used in the MSSS.
FSAR Section 3.8.4.1.5 will be revised to 07/15/89 08/31/89 IP correctly reflect the as-built framing of the AFW pump rooms.
Legend:
C - Complete IP - In Progress
- These commitment dates are extended to coincide with the next scheduled amendment to the FSAR.
2926W.0489