ML20092N185

Affidavit of Rc Iotti & Jc Finneran Re Clinching Down of U-bolts.Related Correspondence

ML20092N185
Person / Time
Site:	Comanche Peak
Issue date:	06/22/1984
From:	Finneran J, Iotti R TEXAS UTILITIES ELECTRIC CO. (TU ELECTRIC)
To:
Shared Package
ML20092N160	List: ML20092N157 ML20092N165 ML20092N185 ML20092N192
References
NUDOCS 8407020527
Download: ML20092N185 (300)

Text

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%: ,g' T-ATTACHMESp UNITED STATES OF AMERICA NUCLEAR REGULATORY COMMISSION S-.

'r BEFORE THE ATOMIC SAFETY AND LICENSING BOARD  !

In the Matter of )

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TEXAS UTILITIES ELECTRIC ) Docket Nos. 50-445 COMPANY, _et _al. ) 50-446 (Comanche Peak Steam Electric ) (Application for Station, Units 1 and 2) ) Operating License)

AFFIDAVIT OF ROBERT C. IOTTI AND JOHN C. FINNERAN, JR. REGARDING CINCHING DOWN OF U-BOLTS We , Robert C. Iotti and John C. Finneran, Jr. , being first duly sworn, hereby depose and state as follows:1 (Iotti) I am employed by Ebasco Services, Inc. as Chief Engineer of Applied Physics. In this position, I am responsible for directing analytical and design work in diverse technical areas, including analyses of the response of piping and support systems for dynamic events, including earthquakes. I have been engaged by TUECO to coordinate and oversee the technical activities performed to respond to the Board's Memorandum and 1 Except as otherwise indicated, each Affiant attests to all parts of this affidavit.

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• Order of December 28, 1983. A statement of my educational and professional qualifications is attached to Applicants' letter of May 16, 1984 to the Licensing Board.  !

l (Finneran) I am the Pipe Support Engineer for the Pipe Support Engineering Group at Comanche Peak Steam Electric  !

i Station. In this position, I oversee the. design work of all pipe [

I design organizations for Comanche Peak. I have previously l provided testimony in this proceeding. A statement of my professional and educational qualifications was received into evidence as Applicants' Exhibit 142B.

Q. What is the purpose of this Affidavit?

A. The purpose of this Affidavit is to respond to CASE's l concerns regarding the practice of cinching down U-bolts.

i' These concerns have been summarized by the Board and can be categorized into the following four areas: [

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1. The acceptability of cinching down U-bolts as a resolution of the potential instability of non-  :

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< rigid U-bolt supports (see the December 28, 1983 Memorandum and Order at 77 ~and 33, and the February 8, 1984 Memorandum and Order (Reconsideration) at 20).  !

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2. The use of SA-307 (or SA-36) steel in U-bolts which are cinched-down considering the ASME Code prohibition against the use of such material in friction type connections (see December 28, 1983  :

Memorandum and Order at 28 and 33, and February 8, 1984 Memorandum and Order (Reconsideration) at i 22-4). A fair restatement of this second concern -

is that there needs to be reasonable assurance that adequate clamping force can be produced and .

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maintained by the U-bolt connection to prevent rotation of the U-bolt cross piece assembly around the pipe.

3. The forces and stresses that are induced in the U-bolt itself (see December 28, 1983 Memorandum and Order at 33-41, and February 8, 1984 Memorandum and Order (Reconsideration) at 24-5).

4. The local (and global, if any) stresses induced in the pipe by the cinching down practice (see December 28, 1983 Memorandum and Order at l33-41, and February 8, 1984 Memorandum and Order (Reconsideration) at 25-6).

Q. In overview fashion, how have you responded to these concerns?

A. .To address each of these concerns, Applicants committed to provide the following (see Applicants' Plan to Respond to Memorandum and Order (Quality Assurance for Design) at 5-6 (itens 3, 4 and 5)):

"3. Provide evidence that the use of U-bolt cinching is appropriate to eliminate potential local instability without introducing adverse effects in the piping and the U-bolt itself."

"4. Provide evidence that there are no adverse long-term effects from U-bolts caused by heat-up and cooldown and related friction on the pipe."

"5. Provide evidence of the acceptability of stresses on pipes caused by thermal expansion in local areas around cinched U-bolts."

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s s The primary method used to obtain this information is testing. However, Applicants have also performed finite element analyses 2 of the tested configurations. The reason for also having finite element analyses performed for the tested configurations are two-fold:

1. Data obtained by tests are limited to locations where strain gauges are placed. A tool is needed to obtain information at other locations and to interpret the data obtained by test. This tool is the finite element model of the tested configuration, which is correlated to the test data at the locations where data are directly available.

2. A model is needed to predict (with good confidence) the behavior of configurations which are different from those being tested, i.e.,

different pipe size, U-bolt size, preload, etc.

The finite element model, once it is verified against the test results, can be used for this purpose. Moreover, it can also be used to verify the adequacy of other, more simplistic models, which are developed to assess U-bolt loads, pipe stresses, and stability questions.

Q. Before describing the test program, its results, and the results of the finite element analyses, are there any items of apparent misunderstanding that you wish to clarify?

A. Yes. We would like to clarify an apparent misunderstanding by the Board. On page 28 of its December 28, 1983 Memorandum and Order, the Board states:

2 These finite element analyses employ idealized but realistic models of the piping and U-bolt cross piece assemblies, subdivide the models into many " finite elements" and theoretically predict the states of stress and strains in each of the elements.

1 5-

"The fact that this material [SA-3073] was incorporated into the U-bolts is not surprising, since they were not initially designed.to be cinched down and to develop friction forces to hold the pipe."

The Board is apparently under the impression that no U-bolts at CPSES were initially intended to be cinched down.

This is not the case. A significant number of U-bolt supports at CPSES were always intended to be cinched down.

On only a relatively small number (less than 15) was the initial design changed such that U-bolts were cinched down because of potential pipe support instability. It should be noted that there are other U-bolt supports at CPSES which are not cinched down, e.g., U-bolts on rigid frames used as one- or two-way supports.

For the Board's information, Table 1 provides a partial list of the cinched-down U-bolts at CPSES. Considering normal, upset and emergency loads, the ten highest. loaded U-bolt supports for each pipe size are given in this table, except for the smaller and the larger pipe sizes which have less than 10 such supports.

t 3 Even though~the Board refers to SA-307 material, the i

designation of the U-bolt material is SA-36. Applicants recognize that the material is the same in any case, with A-307 being the designation employed for headed bolts.

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, , There is another item which we wish to clarify with the Board. This item deals directly with the second concern, i.e., the use of S A-36 U-bolts where SA-307 bolts (similar to SA-36) are prohibited by ASME Code Table XVII-2641.1-1 Note 1. We do not take issue with the fact that the material employed, SA-36, is known to relax 4 under applied thermo-mechanical loads. We do, however, take issue with the Board's apparent characterization of the U-bolt / cross piece connection as a friction connection, and therefore, with the Board's implied conclusion as to the applicability o# Note 1 of ASME Table XVII-2461.1-1 (quoted below) to this connection.

" Friction type connections loaded in shear are not permitted. The amount of clamping force developed

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l, by SA-307 is unpredictable and generally insufficient to prevent complete slippage."

4 Relaxation is here denoted as a characteristic of certain materials which when stressed to certain levels will not l maintain that :evel, but will " relax" to a lower stress level. For instance, a bar loaded to an initial stress of 40,000 psi and then held at constant strain and temperature may after a time period have a remaining stress of only 30,000 psi. This time dependent stress reduction is called stress relaxation. The total strain remains fixed, but a part of the elastic strain is replaced with inelastic strain. It should be noted that stress ralaxation stops after a material reaches a certain level of stress, e.g.,

for material such as SA-36 this level is approximately 1/2 of the yield stress.

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t This note is clearly intended for a friction type connection i in which the load, as transferred in shear, is reacted by  !

the friction between the two surfaces bolted together. In g

the case of the U-bolt (as shown in Figure 1), the load is ,

intended to be carried by the bolt in tension or by the j cross piece in compression, and not in shear. Further, slippage between the U-bolt nut and the surface of the cross  !

piece (either plate or built-up tube steel) as would occur l i

if F, exceeds 2f + uFp (See Figure 1) is perfectly to.lerable provided there is sufficient tension in the U-bolt to  :

maintain contact between the pipe and the U-bolt and the  !

pipe and the cross piece, which is needed to keep the U-bolt / cross piece assembly from rotating (slipping) about the pipe. Significantly, the U-bolt / cross piece connection is ,

not-a friction type connection, and is not intended to be l l loaded in shear. Obviously, it could be loaded in shear r

under U-bolt lateral or axia15 loads, but in this instance I~ it is inconsequential whether it acts as a friction or a bearing connection. Accordingly, we do not see the direct

-relevance of the ASME Note (quoted above) to the U-bolt

" clamp" configuration used at Comanche Peak. ,

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j 5- Lateral direction is defined as the direction aormal (perpendicular) to the pipe axis and parallel to the plane of the U-bolt. Axial direction is the direction of the pipe I e

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. Fb,Fb = Tension in U-bolt legs '

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We-disagree that the use of SA-36 threaded rods is in contradiction to the ASME Code, and hence that their use is a design error. As we have already stated, we do agree that the relaxation 6 characteristics of this material are a i concern only to the extent that a sufficient clamping force i

must be maintained to insure contact between pipe and U-bolt / cross piece. The test program specifically addresses  ;

this question along with others.

~Q. Describe the objectives of the testing program.

A. The test program had seven overall objectives:

1. To determine the relationship between applied torque and the tension developed in the U-bolt;

2. To determine the force required to initiate slip between the U-bolt or cross piece and the pipe as a function of preload (applied torque);

3. To determine the load (and stresses) induced in the pipe and the cross piece as a function of preload; i

4. To determine the additional loads (and stresses) induced in the pipe,.the cross piece and ,

! the U-bolt by differential thermal expansion of the pipe with respect to the U-bolt / cross piece, and internal pipe pressure;

5. To determine the additional load (and stresses) induced in the pipe, U-bolt and cross piece by application of mechanical loads, which I t

6 Relaxation has been previously defined (see note 4). Here, however, Applicants wish to inform the Board that stress relaxation does not continue but ceases after the material  ;

reaches a certain level of stress. For a material like SA-36, the level of stress at which stress relaxation ceases is i

in the range of one half of the yield stress.  ;

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i are applied in both directions of intended constraint. This is done to verify any preloading t effects in the U-bolt / cross piece assembly;

6. To determine method and extent of U-bolt i material stress relaxation (i.e., loss of preload) under applied mechanical load, thermal cycling and normal plant vibration, to verify whether for the expected torque levels (preload), material relaxation would result in sufficient loss of preload to impair the capability of the U-bolt assembly to function as intended; and

7. To determine whether the U-bolt assembly would be stable under a seismic event.

Q. In that the objectives of the tests rely on the preload (torque applied to the bolts) in the field, how did you determine the range of preloads to use?

A. Until now, the Board's sole information regarding the preload level that would exist in the field was based on the Brown and Root Design Change Notice No. 1, dated October 8, 1982, to Construction Procedure No. 35-1195-CPM 9.10 Rev. 8, which states:

"When U-bolts are specified on the design document as not having any clearances, the U-bolt shall be snug tight so that the U-bolt cannot be moved by hand . . . . Snug tight is defined as the tightness attained by a few impacts on an impact wrench or the full effort of a man using an ordinary spud wrench."

This process has been interpreted by Mr. Doyle as resulting in approximately 800 in-lbs. of torque (May 4, 1983 Surrebuttal Testimony of Mr. J. Doyle, at 11-12). Mr. Doyle L

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made no attempt to quantify whether this value of torque is applicable to all pipe sizes and all pipe schedules.

Applicants have done so.

To determine the range of torques which exists in the field, Applicants inspected the torque of a randomly selected representative sample of cinched down U-bolt supports. The results of this sampling are summarized in Table 2. This data was used to determine the range of torques to be applied to each of the test specimens. From the data,' Applicants established that for the four inch pipe, tests should be conducted with preload varying from 5 to 60 f t-lbs. Similarly, a torque range of 10 to 100 f t-lbs.

was established for the 10-inch pipe tests, and a 20 to 240 ft-lb. range was established for the 32-inch pipe tests. In all cases, the upper value equals or exceeds the torques measured in the field for that particular pipe size. The 240 ft-lbs. was selected for the 32-inch specimen, even though it is considerably higher than the maximum value measured in the field for similar diameter pipes. However, since the torques noted in the field were progressively higher on larger pipe sizes, this was viewed as a conservative approach and reasonable upper bound.

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.Q . Is it possible that there might be considerably higher

torques applied to U-bolts in the plant than those which you l i

have described and were used in the tests?

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A. We consider that this likelihood is very remote. For the l smaller lines, as is discussed later in this Affidavit, initially high torque values that'would stress the U-bolt l

above 1/2 of the yield stress would have decreased as the ,

material relaxed to a state where the stresses are about 1/2 {

of the yield.7 For the small lines (below 10 inches), then, j the upper value of the torque that should be present in the field is that which corresponds to a U-bolt load which is stressed to about 1/2 yield.

l The maximum torque achieved in the torquing process for l l  !

10-inch pipe would-place the U-bolt at about the 1/2 yield i

stress. Above the 10 inch size, we would not expect the >

torquing process itself to result in values significantly above 1/2 yield stress. Accordingly, while relaxation may, in rare instances, be a factor in larger pipes, generally i

torque will not be so high that significant relaxation will occur.

k 7 The vslue of 1/2 of yield stress is not a precise figure and should be interpreted to denote a level of stress in the  ;

general neighborhood of 1/2 of the yield stress, at which '

level of stress material stress relaxation stops.

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Hence, the effective torque for all pipe sizes will be the lesser of the value corresponding to a U-bolt stress of half-yield or the value achieved by a man with a torque wrench or impact wrench. Applicants believe that 240 ft-lbs. is a reasonable upper bound for the torque achievable by a man with a torque or impact wrench. On occasion, I higher torques may be measured because painting introduces a shear resistance to torquing the U-bolt nut. However, these higher torques would not correspond to the actual preload in l the U-bolt, which was achieved by torquing with the clean f (not painted) threads.

Q. Would you describe the tests that have been conducted and t summarize the results? l L

-A. A total of seven separate tests have been performed. The description of each test, including purpose, test configuration, instrumentation and results is provided in Attachment 1. Four specimens have been tested. They are:

a 4 inch Schedule 160 pipe with a 1/2 inch U-bolt; a 10 inch Schedule 40 stainless steel pipe with a 3/4 inch U-bolt; a [

10 inch Schedule 80 carbon steel pipe also with a 3/4 inch [

t U-bolt; and a 32 inch pipe (same size and schedule as the  ;

I main steam line) with a 2 3/4 inch U-bolt. We conclude from [

these tests that the U-bolt cross piece essentially can b

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[ perform effectively as a clamp provided that sufficient I l

l preload is established in the U-bolt. A brief summary of  ;

I' the' test results is set forth below: ,

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! (1) Torque versus Preload Test The objectives of this test were twofold. The first l objective.was to establish the relationship between torque applied to the U-bolt nuts and the resulting tension in the U-bolt as a function of pipe size. This information is needed to fully establish whether the particular connection behaves like the traditional bolted connection in which the .

l tension in the bolt is given by the following linear relationship:

l t = KTD where K is a constant (assumed to be 0.2 by Mr. Doyle, see

• CASE Findings at IV-14), t is the applied torque, D is the

! bolt diameter, and T is the tension in the bolt. The second ;

I objective of the preload test was to determine the strain in .

l the pipe as a function of preload. This latter information l is used to verify the adequacy and accuracy of the finite element analysis models.

7 The results of the torque versus preload test indicate  !

, that a roughly linear relationship exists between the torque t-imparted to the U-bolt nut and the tension developed in the +

I U-bolt. While the test does confirm the approximate  ;

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L linearity of the ~ formula t = KTD, it indicates that the values of K vary between 0.22 and 0.35 for the type of ,

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threads present in the U-bolts. Thus, for this type of ,

l l bolted arrangement, more torque is generally needed to develop the same tension in the bolt than would be estimated ,

i . t by using the conventional formula t = .2TD to which Mr. [

Doyle refers. It is noted that there is also a slight i

variation from linearity (which appears to increase with U- i t

bolt size) within the range of specified torque values for  !

! i each U-bolt size. This variation is not significant for the 1 I  !

range of torque values which exist in the field.

In addition, the test reflects that maximum pipe strains (and stresses) caused by preload are generally found '

r in the circumferential direction, are compressive in nature, and occur generally right below the cross piece. In the 10-inch schedule 40 stainless steel pipe, strains similar in magnitude are also seen near the U-bolt contact area, but in this instance they are longitudinally oriented. The magnitude of the stresses varies from pipe size to pipe size and, of course, varies with preload. The relationship between maximum stresses and preload is nearly linear for i

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' r all pipes. Refer to Figures 5 through 8 of Attachment 1 for the maximum pipe stresses caused by preload alone.

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(2) Friction Test The objective of this test was to determine the force on the U-bolt needed to cause slippage between the U-bolt / cross piece assembly and the pipe. Since the direction

- of slippage (which is of primary interest for stability) is in the plane of the U-bolt, i.e., rotation around the pipe, the slip force is determined by applying a tangential load, i.e., in the plane of the U-bolt, to the cross piece, shown

.in Photograph 5 of Attachment 1.

The-friction test produced two results. The first result is the force required to cause slippage between the U-bolt support assembly and the pipe in the plane of the U-bolt (i.e., the force that produces rotation about pipe axis). This force is that required to overcome the friction developed between the pipe and the cross piece at the line of contact between the two, plus the friction developed between the pipe and the U-bolt. The latter friction force develops asymmetrically around the U-bolt pipe contact area.

The lateral force applied to the cross piece produces a moment which causes unequal tension in the two U-bolt legs with correspondingly unequal normal forces and friction developed in their respective contact areas with the pipe, i

From this test, the lateral forces as a function of preload which overcome asymmetric friction were determined.

This lateral force required for slippage is indicative of whether the U-bolt / cross piece will rotate (slip) under the application of seismic forces. The seismic force applied through the strut can vary in inclination about a 5 0 cone from the strut axis. This 5 cone is allowed by CPSES's maximum permissible offset, which includes effects due to installation tolerances, thermal and seismic motion and pipe rotation. Therefore, there can be a lateral force (tending to rotate the U-bolt / cross piece around the pipe), axial force (tending to move the assembly along the axis of the pipe) or combined axial and lateral force acting on the assembly which equals the sway strut earthquake force times the sine of the inclination angle. The maximum value of this lateral or axial force is 8.7 percent of the sway strut force. (See Figure 1 for descriptive explanation of the forces mentioned above.)

If the lateral force required for slippage exceeds the maximum value of the lateral component of the sway strut force, then the U-bolt assembly cannot slip and rotate about the pipe. The U-bolt assembly would then be laterally stable, i.e., capable of transmitting and receiving the

l applied loads. This stability issue is the one upon which the Board focused its attention, i.e., rotation of the assembly around the pipe.

Although the assembly would not slip and rotate about the pipe under these circumstances, a slight roll with no slippage will occur to balance the moment created by application of the lateral component of the seismic force to l t

the cross piece. This roll in no way impairs the stability of the assembly. It is the roll which creates the asymmetry .

in the tension of the U-bolt legs and the consequent frictional force asymmetry, and it corresponds to the t

4 physical behavior of the assembly which wants to align with the line of application of the force. .

t In addition to the rotation about the pipe, insufficient axial friction force between the U-bolt assembly and the pipe would permit motion of the U-bolt assembly along the pipe axis. As will become apparent in later sections of this Affidavit, this axial motion, when occurring without rotational motion, can only result in an axial movement of the assenbly which reduces the strut inclination angle. This motion would continue until che

- inclination angle reduces to a value at which the axial component of the seismic force equals the axial, frictional ,

resistance force. This conclusion, which is borne out by i

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t the vibration and seismic test results, provides further assurance that axial motion alone does not cause concern {

with stability. '

The second result of'the friction test is the determination of the coefficient of friction which exists for the typical U-bolt / cross piece pipe assemblies. The I coefficient of friction is determined to vary between 0.12 and 0.225 for stainless steel pipes and 0.19 to 0.52 for carbon steel pipes. The variation in the friction coefficient could not be correlated to surface conditions.  !

With the exception of the 10-inch stainless steel specimen,  !

no polishing of the contact surface between the pipe and the f

i assembly was observed after all the testing. This is  !

i strongly indicative of little if any surface yielding at the contact points. Some surface polishing was observed on the i 10-inch stainless steel pipe used in the tests. In fact, an increase in the coefficient of friction for this specimen was observed when the friction test was repeated utilizing a f new U-bolt placed on an untested surface of the pipe.

(3) Load Distribution / Strain Measurement Test The objective of this test was to determine the stiffness of the assembly. The test was performed to P

provide information to refute Mr. Doyle's allegations that in addition to the preload, one should add thermal expansion i

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and mechanical loads directly to the support / pipe connection. Applicants do not disagree that thermal expansion loads are additive to the preload, but ran this test to demonstrate that the U-bolt / cross piece pipe connection behaves as a typical preloaded joint, whereby the total mechanical external loads are not directly additive to preload.

This test was conducted on the 10 inch stainless steel pipe /U-bolt assembly only. The choice of the test specimen was predicated on the fact that this specimen results in the  :

most flexible connection of tnose tested, and thus, is the least likely to behave as a preloaded joint.  ;

The test results indicate that for pretorquing levels [

ranging from 33 f t-lbs. to 100 ft-lbs. (these levels encompass most of the torques that are present in the plant

-on 10-inch piping), the behavior of the pipe assembly closely resembles that of an ideal preloaded joint. For the 1atter, no increase in compressive or tensile loads would be experienced until the applied load exceeds the preload.

Figures 17 through 20 of Attachment 1 indicate a relatively shallow slope of the U-bolt load / applied load relation (for the ideal preloaded joint the slope would be zero). The slope of the line is approximately the same for the joint

acting in tension as in compression. This indicates that regardless of the direction of externally applied mechanical l load, this load is not directly additive to the preload.

For instance, from Figure 17 of Attachment 1, an externally applied load of 7200 lbs. on a U-bolt assembly initially preloaded to 9600 lbs. causes only a 3600 lb.

increase in load, i.e., U-bolt final load is 13,200 lbs. and not 16,800 lbs. as one would obtain by directly adding the external load to the preload.

Thus, although the three mechanisms referred to by CASE on page IV-8 of their Findings are considered, namely preload, thermal induced loads, and mechanical loads, only two are directly additive (preload and thermal). The percentage of mechanical load addition depends on the pelative stiffness of the U-bolt / cross piece and pipe.

Measurement of the strains (stresses) registered by strain gauges placed in the cross piece during this test confirms the validit'y of the strains measured in the U-bolts. Finite element analyses of the cross piece (with the bracket welded on the two sides parallel to the long side of the plate) confirm that the plate behaves essentially as a cantilever. Correlation of the stresses that would result from application of the load applied by the U-bolt (U-bolt tension) with those measured via the strain gauges is good.

i l

For instance, the test stress derived from the cross piece  ;

strain gauge readings corresponding to a U-bolt tension of I 6,200 lbs. (per leg) is 17.1 ksi. Attachment 2 provides the I model and results of a finite element analysis of the cross -

piece. The model and analysis employed are linear. This model was executed for an arbitrary load cf 20 lbs. and hence the results are for a load equal to 20 lbs. in each leg of the U-bolt. Stresses resulting from higher (or lower) loads can be computed by linear ratioing. Thus, the stress computed from the finite element analyses, when a 6200 lb. load is applied, is 16.4 ksi.8 (4) Thermal Cycling / Thermal Gradient Test The objectives of this test were twofold. One i objective was to determine the additional load on the l

support and pipe (and resulting stresses) caused by '

! differential thermal expansion of the pipe with respect to l

the U-bolt. (The results of testing for this objective are set forth in Attachments 2 and 3 and were used in the finite

, element analyses.) The second objective was to assess the l ,

l relaxation of the U-bolt preload caused by long-term 8 This finite element analysis was performed prior to a final precise measurement of the cross piece dimensions. As a result, some. dimensions are not exactly equal to the dimension of the cross piece utilized in Attachment 3. The small differences in dimension do not affect the conclusion of this study.

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l temperature cycling to determine whether material relaxation effects would reduce the preload to the extent that slipping of the U-bolt / cross piece can occur. ,

i In this test, the specimens were cycled between room temperature and the maximum operating temperature. The  ;

- maximum operating temperatures are 560 F, 250 F, 250 0F, and 560 F for the 4 inch, two 10 inch, and 32 inch pipe .

1 specimens, respectively. }

Results of this test indicate that the temperature i i

distribution in the U-bolt is not uniform, regardless of ,

j. whether the assembly is insulated or not. Temperature  ;

distributions achieved in.the U-bolts at steady state are reported in Attachment 3. This Attachment reports on the results of the finite element analyses conducted in parallel I

with the testing program. The temperature distributions in i l

the U-bolts are used as input to the finite element -

analyses, r

Figures 21 through 24 of Attachment 1 show the loss of  ;

preload resulting in the U-bolts when the specimen "

assemblies are cycled between room and maximum temperatures.

f The thermal cycling was terminated after 10 cycles, as it became evident that the U-bolt had relaxed to a final stress state after a few cycles. The results of the thermal t cycling tests indicate that the 4-inch specimen had relaxed l

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from the maximum preload value (corresponding to a torque l

1evel of 60 f t-lbs. , which is the maximum. measured in the field) to a value which is approxLmately 64 percent of that value.

For the 10-inch schedule 40 stainless steel specimen, the relaxation is small, if any, as shown in Figure 22 of Attachment 1. Regarding the behavior of the 10-inch schedule 40 specimen, further explanation is necessary. As the specimen was heated during cycle one, the preload l immediately dropped to about 4000 lbs. from the initial value of 4500 lbs. This behavior is contrary to what was expected, (i.e., additional thermal load of approximately

+

700 lbs. per leg should have increased the preload from 4500 to 5200 lbs. per leg). This behavior is indicative of an i initial fit-up which caused the initial preload to revert to the lower preload corresponding to a better fit as soon as the heatup started. Problems with fit-up are also evident by the sudden relaxation of preload in one of the legs during the cycling. Thus, essentially no thermal cycling relaxation was experienced by the 10-inch schedule 40 specimen. This is not unexpected since the maximum temperature of the specimen is 250 F.

~- ,

n .

For the 32-inch specimeni the relaxation cannot be determined from the data (see Figures 23 and 24 of Attachment 1).

The 4-inch specimen is subjected to far more severe thermal ~ cycling than the 10-inch specimen. This specimen is cycled from 107 F to 560 F, with an initial maximum preload which induces stresses in the U-bolt body which are approximately 89 percent of the yield strength.9 After the final thernal cycle at ambient temperature, the stresses in

.the U-bolt are: reduced to about 54 percent of yield. This

< , is also not unexpectbd.; Materials with the characteristics

"= of A-36 exhibit relaxat' ion characteristics at low

- temperatures ~(low temperatures being defined as below 600 F) l~ which cause the stressas in the material to drop to about b 1/2 of. yield.10 Also, these materials exhibit the characteristic that most di the strain relaxation occurs very soon after the high stress 'is created. For instance, the C, Mn , Si steels (of which IU-bolts are made) tested for s

~

9 This value is based 'on 36 ksi minimum yield of SA-36 steel.

' , m 10 If the material initial stress state is above a nominal 1/2 yield stress, the material would relax even without thermal cycling, but probably at,a' slower rate. If the material initial stress' state is below the nom'inal 1/2 yield stress, thekmal expansion stresses might place it in a state of

~

stresm'above 1/2 yield, from which the material would relax.

4 m

'E

relaxation at initial strains corresponding to yield strains, relaxed to almost half that strain in the first hour and remained at -that level after 100 hours0.00116 days <br />0.0278 hours <br />1.653439e-4 weeks <br />3.805e-5 months <br />.ll This means that the maximum relaxation of each specimen, or conversely, the minimum preload that would remain in field applications, can be predicted with reasonable assurance. At the temperature and stress level existing in the 10-inch schedule 40 stainless steel specimens (250 F and 16 ksi, respectively), no relaxation was expected and none was noted (see remarks on fit-up.)

This is also true for the 32-inch pipe for which the stress data are not conclusive, but certainly indicative of low stresses. Further information regarding the relaxation characteristics of the specimens is provided by the creep and vibration tests.

(5) Creep Test The objective of this test was to determine whether long-term temperature exposure could result in material relaxation so that preload would be decreased or lost.

The creep test was conducted on all three specimens following completion of the thermal cycling test. The specimens were maintained at their peak temperatures for 1 " Compilation of Stress-Relaxation Data for Engineering Alloys," ASTM Data Service Publication DS-60.

over 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />. The results show that after the initial relaxation achieved during the thermal cycling test, no further relaxation occurred. This indicates that at these temperatures creep is not a concern.

(6) Accelerated Vibration Test The objectihe of thi's test was to determine whether normal vibration levels in the plant could cause material relaxation, and consequently, loss of preload. In order to simulate'40 years of accumulative effects of piping vibration, this test was'run as an accelerated vibration t test utilizing = vibratory forces : rying in frequency from 5 to 200 Hz at an amplitude equal to the maximum expected OBE force for the pipe ten'ted (4000 lbs.' as well as at lower forces (1000 and 1500 lbs.). The time duration of this test combined with the amplitude of the vibratory (sinusoidal)

' force: resulted in lui overall energy' input to the test

. specimen. exceeding.by orders of magnitude the energy that

' would be induced by an earthquake l2 (both operating basis and' design basis earthquake). An initial 4000 lb. test was run -in excess of 30 seconds, two t at- were run with a 1000 12 - Applicants recognize that the es. orgy 20 parted to the piping /U-bolt assembly during the 30 second duration of a design basis earthquake isslarger than the energy imparted during any 30 seconds of the test. Applicants therefore also performed a seismic test of the 10" Sch. 40S pipe which utilized a force equivalent to that of the SSE. This test is described later in the Af fidavit.

4 D

. - - - _ - - - -_-_-__-O

Ib. force for 2.5 minutes and a 1500 lb. test was run for 270 minutes as described in more detail later. This test simulates conditions far more severe than expected in the plant for normal vibration levels.

The major results of this test are noted below.

l Insufficient assembly preload will permit the assembly to rotate about the pipe and also to walk axially along the pipe axis in either direction with respect to the location of the strut. This can occur because the low preload permits the asymmetric tension in the U-bolt legs to relax sufficiently upon application of a sufficiently large compressive external load within the 5 inclination that the frictional force between the U-bolts and the pipe, and/or friction between the cross piece and the pipe, is overcome by the lateral or axial component of the externally applied force. This permits slippage of the U-bolt and/or cross piece along an erratic helical path (if the applied external load is inclined both laterally and axially). The axial i.

l motion away from the strut, i.e., the motion that would lead to an increase.in the angle of inclination, can only occur i' if the assembly is permitted to rotate. With rotation, the lateral-component of the force is increased as the size of i

the angle increases, and the axial component at the new s

location of the strut can also be higher, causing the cross piece to move away from the application of the compressive force.

With rotation inhibited by sufficient preload, any axial motion will be in a direction toward the applied load and will occur during the pull portion of the cycle at the cross piece contact point. The direction of travel is preordained since slip occurs axially only during the pull portion of the cycle. (During the push portion, sufficient friction is created at this point to resist slip by the component of the sway strut force increasing the cross piece / pipe contact load.) This behavior is verified by the accelerated vibration and the seismic tests.

In preparation for the full accelerated vibration test, a constant amplitude sinusoidal force of 4000 lbs. was applied to the pipe for a period exceeding 30 seconds with the frequency of excitation being varied from 5 to 10 Hertz.

(These frequencies are at or near the peak in the response spectrum of CPSES.) This pretest estimates the seismic excitation imposed by an OBE, since the maximum force from an OBE for a 10-inch pipe would be 4000 lbs. The pretest was interrupted to avoid damage to the hydraulic actuator in the test machine. At 20 ft-lbs. preload, the assembly was seen to rotate about and walk along the pipe. Although the

t l

r assembly was still capable of transmitting the same load, l i

approximately 4000 lbs., (thus acting as a stable support r i

for 30 seconds or more) Applicants consider this to be _ j

)

i insufficient preload. This was confirmed by a second two and one half minute pretest conducted at a sinusoidal force t t

i of 1000 .lbs. having a frequency sweep from 5 to 200 Hertz at 2 octaves 13 per minute. At the 20 ft-lb. preload, although the assembly was acting as a stable support, it still rotated and walked, although in a slower manner. The  ;

pretest was then rerun with the preload torque increased to  ;

35 ft-lbs. The input sinusoidal force had an amplitude equal to 1000 lbs. with the same frequency sweep. No motion r

of the assembly was observed during the 2 1/2 minute pretest. l To run the' official acceleration vibration test, the- l preload torque was increased to 50 ft-lbs. The official ,

' accelerated vibration test was run at 1500 lbs. (swept from 5 to 200 Hz and back at 2 octaves per minute) for 270 minutes. The assembly was observed to move initially  ;

axially (in the direction that reduced the strut angle),

13 An octave is'a doubling of frequency. Thus, from an initial  !

frequency of 5 Hz, the frequency of the excitation will be i 10 Hz in 30 seconds, 20 Hz in one minute, 80 Hz in 2 minutes and 160 Hz in 2.5 minutes (200 Hz takes slightly more than

'2.5 minutes).

t i,

r ase = w r ow-p , ywv.--- ---

=-,-v,. -,- m

. . then to stay in place for the duration (except for a sudden, but inconsequential cocking in the latter part of the test).

No rotation was obser7ed.

At the end of the test, the applied sinusoidal force was increased to 4000 lbs. With this preload, the assembly was also vibrated for a period of 30 seconds with the frequency being swept from 5 Hz to 200 Hz and back to 5 Hz at a rate of 20 octaves per minute. This rapid sweep was done to avoid damage to the hydraulic actuator. This last accelerated test conservatively hounds the seismic excitation imposed by an OBE, since not only is the force (4000 lbs.) equal to that which results from an OBE for 10-inch pipes, but r.he frequency range of the OBE is narrower than that swept in the test. Moreover, for the OBE, the forces at the high frequencies are much lower than the 4000 lbs. which was applied for all frequencies. No rotation of the assembly was noted during these 30 seconds, but there was the same axial motion of about 1/16 inch toward the strut that occurred during the 270 minute test. This motion reduced the angle of inclination.

The observed behavior confirmed that the assembly is stable at that preload torque, and also confirmed the theorized behavior of the clamping action.

At the end of the of ficial accelerated vibration test, the relaxation of the preload was also measured. The initial preload stress (based on U-bolt body area initial strain = 311 micro inch / inch) was equal to 9020 psi.

Significant relaxation was not expected since the initial stress was below half of the yield stress. After the initial repositioning of the assembly, which reduced the preload, no further decrease in preload was observed, indicating that the vibration per se had no effect on relaxation.

(7) Seismic Test The objective of this test (an auxiliary test to the accelerated vibration test, noted above) was to test the effect on the assembly of the peak SSE force, 7000 lbs.

Although the overall energy inputted in the tested system in the official accelerated vibration test (discussed above) is clearly much greater than that inputted by the design bases earthquake (SSE), Applicants wanted to test the effect of peak force on the specimen assembly and also the effect of inputting an energy rate comparable to or in excess of that expected from the SSE. There fore, a 39 second test with a sinusoidal force applied at a frequency of 9 Hz (roughly corresponding to the peak of the CPSES floor response spectrum) was also run to simulate the maximum response to

32 -

the SSE. Although the sinusoidal force was to have a magnitude of 7000 lbs., in reality the average force magnitude reached a maximum of 9500 lbs. initially, decreased linearly to 8600 lbs. after 21 seconds and remained at 8600 lbs. throughout the remainder of the test.

'(See Figure 1 of Addendum 1 to Attachment 1.) Preload was applied via a 50 ft-lb. torque.

The test was to have been repeated at the resonant frequency of the test specimen (established to be at about 75 Hz by sweeping the frequency range with a lower amplitude force).14 However, at this higher frequency, the hydraulic actuator was incapable of transmitting the required 7000 lbs. of force (due to play in the strut connections

~

exceeding the displacement output of the actuator at that frequency). The maximum force output of 75 Hz was approximately 1300 lbs. The SSE energy at this frequency is neglibible. While the 1300 lb. (75 Hz) test was aborted after 4 seconds, it provided useful additional qualitative information that the U-bolt assembly behaves stably. Prior to running the nominal 7000 lb. test, three trial tests were run to adjust test equipment and instrumentation. Although 14 The combination of the test run at a frequency corresponding to the peak of the floor response spectrum and of the test run at the resonant frequency of the test specimen would encompass the worst situation that can be encountered in the field.

\

m

• . none of the three trial tests can be considered official, we note that the performance observed was in all cases consistently the same as that in the official test. Thus, what was observed in the official test was systematic behavior of the assembly. In the official test, the assembly did not rotate, but moved axially toward the strut (confirming the theoretically predicted behavior) a distance of about 1/2 of an inch. After that, it remained in place.

The same motion had been observed for the three trial tests.

Although the 75 Hz test was also unofficial, it was observed to result in no-rotational motion and barely perceptible axial motion (toward the strut).

At the end of the nominal 7000 lb., 9 Hz test (actual force in excess of 8600 lbs.) the preload value in each leg of the U-bolt was measured to determine whether strain relaxation had taken place. Strain relaxation was expected since the applied force coupled with the initial preload stresses the U-bolt to above 1/2 of yield stress. In this instance, with an initial preload of 4484 lbs. in both legs of the U-bolt.and an applied external force of 8600 lbs.,

the peak tension in the U-bolt would be approximately 6600 lbs. (see Figure 17 of Attachment 1), resulting in a stress at the threaded area which is about 10 percent over 1/2 of l.

yield.15 Since prior to stabilization of the applied force, a higher amplitude force was seen, more relaxation could be expected and was in fact seen during the initial stabilization period. The remaining preload, measured after completion of the test indicates relaxation of approximately 12-13 percent to a stress level below 1/2 of yield stress.

Q. What do you conclude from the results of the tests regarding CASE's allegation concerning instability?

A. We conclude that the U-bolt / cross piece assembly can perform effectively as a clamp provided that sufficient preload is established in the U-bolt. (It should be noted that a clamp '

also requires preloading.) We further conclude that even if the preload level was insufficient, but still present in some amount, the U-bolt support would vibrate, but still be capable of supporting the necessary loads, thus behaving

" stably."

To provide further assurance that the preload on all affected cinched down U-bolts is adequate, Applicants will conduct a 100 percent inspection of the torque of all such U-bolts (380). At the time of the inspection, to remove questions regarding stability, Applicants will assure that 15 Using a yield stress equal to 36 ksi and recognizing once again that the 1/2 yield stress level at which relaxation ceases is not a prccise figure, but a more imprecise range of stresses near the 1/2 yield stress at which relaxation would stop.

l l

I

l such U-bolts are torqued to levels at which the assemblies will be stable in the absolute truest sense, i.e., no rotation, and axial movement, if any, is toward the strut.

(To check the torque of a U-bolt requires essentially no -

more effort than torquing the U-bolt to prescribed values.)

The levels to which U-bolts will be torqued are set forth in Table P of this Af fidavit.

The results of the tests conducted for vibration and for seismic response confirm the stability of the assembly when preloaded to these values. These values are set forth later in this Affidavit.

Q. What are your conclusions from the results of the tests regarding the stresses in the U-bolt ar.d piping?

A. The only conclusions that can be derived from the test program regarding the piping stresses are limited to data obtained during the preload test. The test results indicate no unacceptable stresses in the pipes for the preload conditions. Measurement of the strains on the pipe from which stresses can be obtained, is limited to locations near but not precisely at the location where peak stresses are expected to occur, i.e., under the cross piece and U-bolt at the points of contact.

_ - ___- - - - U

Because of this limitation, Applicants' plan to determine such stresses relies on measurement during tests of strains (stresses) at selected locations along the pipe axis and around the circunference near the cross piece and the U-bolt, to serve as reference data points for the finite element models. The stress distribution in the pip'e resulting from the application of preload, thermal expansion, internal pressure, and externally applied loads, is developed using the finite element models which will be described later. Detailed discussion of the stresses in the pipe is thus deferred to later sections of this Affidavit, which sunmarize the results of the finite element analyses.

i Several points need to be made regarding the potential high stresses in the U-bolt. First, the preload applied in the test is larger than that expected in the field.16 Hence, actual field stresses will be lower. Second, if initially the bolt is preloaded to stresses exceeding approximately one-half the yield stress, the material will i

relax, and the preload will consequently drop until a stress 16 The preload applied in the test is larger than that expected presently in the field. However, for large pipes such as

, the 32 inch tested specimen, Applicants plan to increase the l preload over that which is now present, to ensure adequate l margin for stability. For the larger pipes and large l diameter U-bolts, however, the U-bolt stresses are low. For instance, stresses in the 2 3/4 inch U-bolt of the 32 inch specimen prelo,aded with 250 and 500 ft-lbs. are l approximately 1000 and 2000 psi, respectively.

l

state of about 1/2 of yield is achieved. The deformation required for this relaxation is negligible 17 and sufficient preload still exists to assure the stability of the support.

The minimum preload necessary to assure stability is below the preload level which would exist when the stress _s about 1/2 yield. For the 4-inch specimen, for instance, the preload torque that would correspond to stresses equal to 1/2 yield is approximately 30 ft-lbs., and the minimum torque conservatively estimated to be for stability is about 25 ft-lbs. (finite element analyses would predict about 10 ft-lbs.), (see later discussion of results of finite element '

analyses). Third, normal plant vibration will not significantly affect the preload once it has relaxed to the

" final" condition. Here, final denotes the condition for which the stress in the U-bolt is about 50 percent of the yield stress and no further relaxation takes place. Fourth, seismic loads would not affect the " final condition" preload (as is evident from the accelerated vibration and seismic tests)18 Fifth, tests conducted by Applicants have demonstrated that there is adequate margin between yield and 17 Deformation is negligible since total strain is not a'ltered.

Stress relaxation occurs because part of the elastic strain is converted to inelastic strain.

18 Had the seismic test been run at 7000 lbs. (maximan seismic load for 10 inch line) little if any relaxation would have occurred.

J

. e f

failure of the U-bolts.19 There is thus no concern with failure for stresses that might be initially above the yield stress.

The stresses.in the U-bolts were measured during all of [

the tests. With the specimens torqued at the highest  !

preloads (60 ft-lbs., 100 ft-lbs., and 240 ft-lbs. for the  :

4", 10" and 32" pipe, respectively), the maximum stresses in !

the U-bolts occurred during the thermal cycling test.

First, we discuss the results for the 4-inch pipe. During I torquing, the stress in the U-bolt for the 4-inch pipe reached 35,360 psi (near yield) in the U-bolt shank.

(Within the first hour or so, due to relaxation, these i stresses would have decreased to about 1/2 yield.) During torquing, the threaded area of the bolt exceeded yield slightly; yielding of the threaded area was noted in the [

test. (However, due to the self-limiting nature of the load, this would have no adverse impact on the material.) i After initial material relaxation (if any, since the i

initial state of stress depends on actual preload and if  !

preload is icy there will be no relaxation) application of a mechanical load in the U-bolt tensile direction such as i

19 See Affidavit of R.C. Iotti and J. C. Finneran, Jr., l' attached to Applicants' Motion for Summary Disposition of CASE's Allegations Regarding One-Way U-bolts Acting as Two- i Way Restraints ( May 23, 1984).  !

f i

L

1 i .

l I

might occur during an earthquake would add some load.

Because the joint behaves as a preloaded joint, and because ,

the pipe is very stiff (4" Schedule 160), most of the load goes into the pipe and little into the U-bolt. (The finite {

element analysis conducted for a compressive load confirms this joint behavior). The increment in load experienced by the U-bolt would cause some further relaxation which would reduce the stress to 1/2 of yield stress again.

Secondly, we discuss the 10-inch pipes. For the 10" Schedule 40 stainless steel pipe, U-bolt strr.ases measured in the shank for preload conditions corresponding to the maximum torque (100 ft-lbs) were 12,838 psi. No appreciable I increase in load resulted from thermal cycling. (Actually, there was a decrease indicating release of some form of I mechanical binding). Externally applied loads, such as seismic loads, would increase the stress when directed so i that the U-bolt tension is increased. For the maximum i

postulated external load (7000 lbs.), the stress would l

1 increase by about thirty percent, placing it nesr 1/2 yield. I The 10" Schedule 80 carbon steel pipe was not tested during thermal cycling since the behavior of the corresponding stainless steel specimens is more severe. U-bolt stresses in the 10" Schedule 80 pipe test were measured )

9

- 40 _

to be 17,164 psi in the shank (22,650 psi in the threaded area), when the preload corresponds to a torque of 100 ft-lbs.

Finally, for the tests of 32" pipe, the preload value (240 ft-lbs. torque) placed negligible stresses in the 2-3/4 inch U-bolt.

Data on stresses in the cross piece are available only for the 10-inch and 32-inch tests. The geometric configuration of the 4-inch pipe prevented placement of strain gauges in the cross piece locations where their readings could be correlated to the readings of strain gauges placed on the U-bolt.

For the 10- and 32-inch cross piece specimens, no stresses exceeding 23,000 psi were observed. The stress of 23,000 psi occurred in the 10-inch stainless steel specimen cross piece when the maximum mechanical load was added (as a pull) to the maximum preload. (It should be noted that since / joint acts as a non-ideal preloaded joint, e

the exterhal mechanical load is not fully additive to the preload.) These stresses are less than the allowable, 0.75 times the yield stress.

In sunmary, we conclude that application of the maximum torques to the U-bolt pipe assemblies can potentially result in high but acceptable local pipe stresses and can further

- 41 _

result in high stresses in the U-bolts. Test results were not intended to provide, by themselves, sufficient information to assess the significance of the pipe stresses.

We defer discussion of that aspect to later sections of this Affidavit. We do not believe that the stresses produced in the U-bolts present a concern. To begin with, high stresses occur only if large preload values are applied (i.e., near the maximum used in the test) to small diameter U-bolts.

Large preload values are generally not present in the plant supports, nor are they needed to assure stability of the supports under seismic excitation (see note 15 in reference to large pipes and large diameter U-bolts). In those instances where high preload torques may be initially present, the characteristic relaxation behavior of the material employed ( A-36) will reduce the preload value, and hence, the stresses in the U-bolt, to acceptable levels.

Moreover, tests conducted for Applicants have demonstrated that there is adequate margin between yield and failure of the U-bolts. For instance, these tests showed that for the 1/2-inch U-bolt employed for the 4-inch specimen, the margin is about 2 to 1. Therefore, there is no concern with possible failure of the U-bolt even if it were to be initially pretorqued to the largest values noted in the field.

sim imi ieim . . . . . . . .. _ _ m . ...i

With regard to the cross piece, no unacceptable stresses have been noted. However, there was no test of the 4-inch pipe corresponding to the conditions in which the U-bolt experiences the largest load. Finite element analysis of this configuration indicates that bending stresses in the cross piece'are determined by stresses in the U-bolt.

Initial high preload values on the U-bolt could produce stresses in the cross piece that exceed allowable values on an elastically calculated basis. However, since U-bolt stresses ultimately revert to one half of yield, the cross piece would not have stresses above allowables at that point.

Q. Please restate why Applicants performed finite element analyses in addition to the tests.

A. The reasons for the development of the models and the execution of the finite element analyses have already been stated, but can be briefly summarized agains (1) without a theoretical model, explanation of results from tests may be impossible, (2) a theoretical model is needed to extend the information provided by test into the location of the tested specimens where the test provides no information, and (3) a theoretical model, verified by test comparison, must be used for any predictions / conclusions that one may have to make on configurations which are not tested.

Q. Please describe the finite element analyses used in your evaluation.

A. Each assembly tested was modeled utilizing MSC NASTRAN .

Version 63. This computer code was chosen because it is universally recognized and accepted by industry as having the capability of providing analytical solutions that accurately characterize the local stress, gap, friction effects, and plastic material behavior (if any) that are important for assessing the pipe and U-bolt assembly stress, and the support stability. A detailed description of the finite element models developed for each test configuration and of the modeling technique used is provided in Attachment 3.

Q. Why is not all the information essential to respond to the Board's concerns available from the tests?

A. Some of the information can be derived directly from the test program. For instance, the Board's concern regarding forces and stresses that are induced in the U-bolt itself can be obtained directly from the test results. The forces and stresses in the U-bolt which result from torquing, torquing plus thermal expansion, and oxternally applied loads are direct outputs of the preload, thermal cycling, and loading portions of the test program.

_ 44 -

On the other hand, to unequivocally answer the Board's concern with pipe stresses and how they are influenced by cinching the U-bolt and related stresses, a mix of information derived from test and analyses is required.

Where the peak stresses occur, they cannot be measured.

Also, it is not feasible to completely cover the pipe specimen with strain gauges (inside and out) to obtain the stress distribution in the pipe. In this instance, the finite element model is used to extend and complete the test data. Finally, there are concerns which can only be answered by test. Examples of these concerns are the relaxation characteristics of the assembly under long term vibration, thermal cycling, and preload. The thermal cycling, creep and accelerated vibration tests have provided answers to these concerns. No analytical tool could have done it.

Finally, test results are far more persuasive than any theoretical model with regard to demonstrating behavior. If the test shows that the assembly does not move, then it is impossible to argue that it did. This is different than having a static, theoretical model predict that it should not move. Endless arguments would ensue over the correctness of the prediction.

To answer the Board's concern on stability, we have conducted a test on assembly stability and 1. ave also correlated the prediction of the static model (finite element) to the dynamic results so that the Board would have confidence that the stable or unstable behavior of the assembly can be correctly assessed.

O. Can you describe the objectives and results obtained from the finite element analysis program?

A. The finite element analysis program was performed (1) to determine if the pipe would slip, thereby creating an unstable support condition when the hanger support was subjected to the preload, thermal, pressure and mechanical loads that could be expected in the Comanche Peak hanger assemblies; and (2) to calculate pipe and pipe support stresses that could be expected to be experienced at Comanche Peak and assess their significance. In order to perform the above evaluations, it was necessary to evaluate the U-bolt support-piping assembly, using finite element analysis, for the four loading. conditions noted below:

(1) Preload; (2) Preload + Thermalt (3) Preload + Thermal + Pressuret (4) Preload + Thermal + Pressure + Strut Applied Load (Push).

i

- .46 - '

f The thermal loading condition was based on temperatures associated with the normal operating condition. The pressure londs were based on normal operating pressure. The applied loads evaluated (representing the external mechanical loads) were generally higher than the loads that the support strut would be expected to carry. The strut -

loads noted below in Table A were dependent on the line size i considered.  !

TABLE A ,

Mechanical Loads on Sway Strut I l

4" Sch. 160 2,000 lbs.  ;

10" Sch. 40S 10,000 lbs. t 10" Sch. 80 10,000 lbs. l 32" Sch. Main Steam 100,000 lbs. l

. t This load was applied at the maximum permissible offset of  !

0 5 from the U-bolt axis parallel to the U-bolt legs. The )

I preload values, based on torque values appropriate for each  !

t pipe siza, are as given in Table B below. ,

I i

e l

l i

f

-_--,r-r-,-- -- , ..m, ..--n._n,-__,.m,,--n,--,,..nc. _ _ , - , , _ . , - , _ . . _n-,4+,_, -,-----.,,--nr,a.--,,,e -- , , . _ - . - -

_ 47 _

TABLE B Maximum and Minimum Analysis Torque Values 4" Sch. 160 Max. value 60 ft-lbs.

Min. value 9 ft-lbs.*

10" Sch. 40S Max. value 100 ft-lbs.

Min. value 46 ft-lbs.*

10" Sch. 80 Max. value 100 ft-lbs.

Min. value 11 ft-lbs.*

32" Main Steam Max. value 240 ft-lbs.**

• These minimum values are different from the minimum values employed for the test. They are values derived from iterative finite element analyses per-formed to help define lower bound values required for stability. The finite element analysis predicts that the assemblies would behave stably a: these and even lower values. (No optimization was done.)

• • No minimum value analysis was conducted for the 32 inch model.

The results of the analyses are given in Attachment 3. The maximum stresses in the U-bolt legs for the four cases considered were determined to be the following for the different torque values.

1 I

L ..

[

t TABLE C  :

Maximum Stresses in U-Bolts 4" Sch. 160 60.5* kai 10" Sch. 40S 22 ksi l 10" Sch. 80 27.5 ksi  :

32" Main Steam 7.5 kai i

• Calculated on an elastic basis. Actual i stresses will be lower and includes the I preload, thermal expansion restraint, external mechanical load and radial pres-sure expansion. Stresses are calculated in the threaded area. 1 These stress values are based on the U-bolt threaded area and compare generally very favorably with test results.

Deviations are exp'ainable l and are due to differences ,

between real configurations and idealized representations, e.g., fit-up, relaxation, and out of roundness. The loading case resulting in the highest U-bolt leg stress was Preload

{

+ Thermal + Pressure. When the push load was applied, the lag forces were unloaded. -

The stresses in the pipe for the maximum preloads f l

evaluated for each pipe size were obtained from the finite element model. The maximum preload torque values were used since this loading level produces the highest stresses in f

i

[

I

,- . -. - - , , _ ,- , , , , , . - - , , . ,- _ , , - . . , _ ,,-.,w. ,._----...-.-.-,.,..n.-,., , . . - - , - - - . _

_ 49 _

the pipe. Tabulated below in Table D are the maximum pipe stress intensities for the four load cases evaluated without the effect of mechanical piping stresses.20 TABLE D Maximum Calculated Stress Intensities

• 4" Sdi. 10" Sdt. 10' Sdi. 32" Main Steam Iomi Case 160 (ksi) 4 W (ksi) 80 (ksi) 336 (ksi)

Prdoad (P) 26.1 48.5 29.9 4.9 P + 'Dmmal (T) 39.3 60.6 34.9 21.6 P + T + Pressure (Pr) 42.9 58.6 32.9 34.4 P + T + Pr + Push 44.8 73.4 44.4 47.2

• Excludes external mechanica'l load.

It is evident from these results of the elastic finite element analysis that elastically calculated stresses can exceed yield. These stresses occur locally at the point of contact between the cross piece with the pipe, if the U-bolt is torqued beyond a certain value, which is generally in excess of that which can be expected in the field and is always above the value which is required for stability.

To investigate the effect of plastic deformation resulting from excessive U-bolt torquing, the elastic finite element model was modified to include the nonlinear material 20 The reader can determine how the mechanical piping stresses (i.e., those due to all effects other than the clamping action) affect the values of the stress intensities given in this table, from the table given on p. 59 of Attachment 3.

From this Attachment total maximum stress intensities are 64.14, 74.21, 54.69 and 47.17 kai for the 4", 10" Sch 40S, 10" Sch 80 and 32" pipes, respectively.

behavior. The finite element model chosen for this analysis i

was that for the 10" Schedule 40S pipe. An elastic / plastic analysis on this pipe /U-bolt assembly would therefore be the most indicative of the effects that local yielding would have on the conclusions reached regarding assembly behavior and distribution of stresses. The results of the '

elastic / plastic analysis show for an approximate 100 ft-lb.

preload, small plastic deformations of only those finite elements adjacent to the line of contact between the pipe and cross piece. The distribution of stresses for the elastic / plastic analysis is the same as for the elastic analysis and is discussed in Attachment 3. This result confirms what has been observed in the test, namely that no visible yieldings of the pipe take place. (Therefore yielding, if any, is inconsequential.) (Refer to pp.17-18 of the Affidavit and the observation that little if any surface yielding (polishing) was noted after all tests were completed.)

Q. Does the ASME Code provide direct guidance regarding i acceptance criteria for local stresses induced by external attachments such as these U-bolt clamp assemblies?

A. We are not aware of any direct quantitative guidance provided by the ASME Code regarding this issue. However, the ASME Code provides qualitative guidance in regard to

t 51 -  !

i consideration of the effects of local external attachments (Sections NB-3645, NC-3645, ND-3 645, and NF-3121). CYGNA, in its oral testimony presented in the May hearings corroborates our position. This qualitative guidance given in the Code leaves to the designer the responsibility for quantifying the acceptance criteria.

Q. In view of this lack of direct guidance, what did you use as acceptance criteria?

A. Applicants do not believe that it is proper on their part to establish firm acceptance criteria for industry regarding this situation. This does not mean that we are unmindful of t i

-our responsibilities to satisfy the intent of the ASME Code as expressed in NB-3111, NB-3624.1, NB-3645 and the '

corresponding articles in section NC and ND. We have in fact formulated acceptance criteria for CPSES to meet the intent of the ASME Code. The bottom line, however, is that i

the testing program and finite element analyses have  !

t demonstrated that cinching of U-bolts as done at CPSES and generally by the industry produce no adverse effects on  :

piping and supports for the range of pretorque values which are either representative of the worst conditions encountered at the plant or required to ensure stable behavior of the U-bolt assembly. [

-.p- , ,mm aw,- --n_,,,,,a ee,r-.m -,---,-.,.-.v ---- ,-e--

-m - - . . . , - , -. - - , , - ,-n-,,---e w ,.,re--,m, _w.. ,

6 In what follows, therefore, we will establish the logic of our interpretation of the intent of the ASME Code as expressed in the acceptance criteria which we have adopted,

.lar showing hos the conclusions reached using these accepta,nce criteria match those derived from the test and finite element analysis.

The following pages provide information from the results of the finite element analyses to enable interested parties to follow the logic of the conclusion reac'ed h by Applicants. To understand how the values given"in several tables are derived, it is necessary for the reader to refer beck to Attachment 3.

Q. What acceptance criteria have you adopted to assess the results of the finito element analyses?

A. To answer this question we need to define the type of stresses which have been measured and/or calculated by finite element analyses and compare them to our interpretation of the ASME Code.

Of the possible stress-type classifications, the only possible controversial one is that of the stresses caused by preload. Clearly preload is u one-time applied and maintained mechanical load. Tests as well as finite element analyses have shown that preload stress has some of the j

e . P characteristics of a secondary stress.21 Finite element analyses and tests indicated that very small deformation at the contact area between the pipe and the cross piece result in relief of the local stress without affecting the stresses in the adjoining areas. Therefore, the stress 22 it is clearly local. Finite element analyses show that the locally high stress reduces to the general level of stress within one meridional (along the circumference) element on either side of the contact area.

The finite element analyses further indicate that the stress is composed of both a membrane and a bending component with the latter being predominant. The membrane portion is a secondary mean stress which the ASME Code defines as the local primary membrane stress, and which the Code includes in equation (9),Section III. The bending portion is a secondary bending stress which the ASME Code does not single out, but includes in the general category of the secondary stresses of equation (12)Section III.

21 Secondary stress is defined as a normal stress or a shear stress developed by the constraint of adjacent material or by self-constraint of the structure. The above characteristic of a secondary stress is that it is self-limiting.

22 By stress here, we mean the stress caused by the preload proper and also stresses which result from the presence of the-preloaded U-bolt assembly such as radial thermal and pressure expansion constraint stresses and stresses induced locally by external mechanical loads at the line of contact.

- .

r . ..;.:.

, . *, :. ;q :_ .: s , ;. ; ;.y j .; .y y ,,

._,_..: _ , , ; n . ;

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. ,c, ..

... _ a c . . n ,. r , ._ .; . .

- a. z.

,. . . , y s .. s One way to look at the intent of the Code is to classify the entire local stress (membrane plus bending) as the primary local membrane stress to be inputted in a modified equation (9) with allowable stress intensity limits increased to 3S

m. to account for the presence of the local bending stress (which would have to be limited to 3S, by itself or in combint: tion with the primary membrane stresses). One can infer from the Code that for application of a one-time load, a doubling of the allowable limits of equation (6)~ is permissible without inclusion of secondary bending stresses.23 Since we are including secondary bending stresses in equation (9) (the secondary bending stress in. ny far the major component of the stresses caused by preload while the primary local membrane is very small),

and do not further raise'the allowable limit over that which ,

can be inferred from the Code for a one-time load application, we are copoervative with respect to'what the Code intent would allow.

An alternative approach which we believe to be not as conservative, is to place'the secondary bending stress into equation (12), and to evaluate the remaining, small local j 23 For example, equation 10a of the Code in Class 2 Rules for single, nonrepeated anchor movement allows a stress equal to 3S , where S is the material stress allowable for cold coEditions.

.jgl%:: % p[ Q yi!f f NpQ;y;__'?JY'kg,2 ;l 3 ?lO':l.) l ..'.':? a,. ~' Y

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• membrane stress with the regulation equation (9) (allowable limit = 1.5S,). Applicants have chosen to examine both alternatives in their evaluation of the pipe stresses.

Evaluation Using the First Alternative ( See p. 6 6 for second alternative.) Equations (9) and (12) are employed here, since the maximum piping stress due to the applied loadings is considered to be composed of primary and secondary stress. In the evaluation of the pipe, 'it is necessary to separate the primary and secondary stresses. The classification of the stresses is given below: o Pressure membrane stress - primary stress o Preload - local primary membrane stress (includes secondary bending) o Thermal and pressure pipe growth restriction - secondary or peak stress o Pipe hanger load (mechanical loads) - primary stress o Pipe hanger thermal load - secondary stress The rules of Subsection NB may be used to qualify a Class 1, 2, 3 or NNS (non-nuclear safety) piping component if the designer is willing to comply with all aspects of ASME Code Section NB. Subsections NC, ND and the ANSI B31.1 piping coda evaluate primary and secondary stresses only by limiting the principal stresses caused by pressure and moment loading in the. pipe. By evaluating stress

intensities (as opposed to principal stresses), detailed localized stresses, and the effects of cyclic loadings (fatigue), one is clearly enveloping the intent of sections NC and ND of the ASME Code and ANSI B31.1 by an NB evaluation. Based on the above, the evaluation of acceptability of pipe stresses induced by the U-bolt pipe support is generally addressed using Section NB3600 of the code. NB-3600 (which governs the design and qualification of piping systems) gives little specific guidance to the method that should be used to evaluate stresses similar to those caused by the U-bolt pipe support. However, using equation (9) and equation (12) of NB-3600, an assessment of these stresses can be made since they provide a means of evaluating the stresses caused by the loading applicable to piping in the vicinity of the U-bolt. A total stress limit equal to 3S (where S,is an allowable stress intensity for the material defined by the Code) is used for each equation. The loading considered for each of the code equations used are given-below:

1. Equation (9) of the_ Code with a 3S_ limit.

This equation must consider (a) primary membrane pressure stress 24 (b) piping moments at the hanger location due to pipe deadweight and seismic load, (c) maximum stress due to preload of the U-bolt, and (d) stresses due to the applied hanger load. The applicable stress limit is 3S, since preload of the U-bolt is a one-time applied load (e.g., similar to building settlement) and includes local, through-wall bending (secondary ben. ding).

2. Equation (12) of the code with a 3S_ limit.

This equation must consider (a) thermal stresses due to the restriction of pipe radial thermal growth at the U-bolt, (b) stresses due to the restriction of pipe radial pressure growth at the U-bolt,25 and (c) piping moments lat the hanger location due to longitudinal thermal expansion of the pipe. Tabulated below in Table E are ele 3S, stress limits for the pipes in question. 24 For convenience, these pressure stresses resulting from restraint of radial growth due to preosure expansion have also been included in this equation, although such stresses would be secondary. They have also been included in equation (12) so that these stresses have been accounted for twice. 25 See previous note.

       *

• TABLE E 3S at Normal Operating Pipe Size Temperature 4" Sch. 160 50.52 ksi 10" Sch. 40S 60 ksi 10" Sch. 80 60 kai 32" Main Steam 58.26 ksi A conservative estimate is obtained for the mechnical stresses in the pipe resulting from sources other than the U-bolt assembly itself. This is necessary since all local sources of load must be considered when comparing the pipe stress state at the U-bolt location to the above stress limit. Because of the multitude of supports involved in this evaluation, individual piping moments at the support locations, calculated from detailed stress analyses, have not been used. Rather, they have been considered in a general manner. The piping stress due to other mechanical loads has been conservatively estimated assuming that the stress was at its maximum allowable limit (1.5S ). This code allowable stress level (1.5S,) is used for the following three reasons:

1. SSE induced stresses need not be included in the secondary or peak stress evaluation.

2. SSE piping seismic levels are in the same order of magnitude as those associated with the OBE (operating basis earthquake) stresses.

3. In general, the stresses in the pipe will not be at the code allowable.

A

 %a

l - I The stress calculated using equation (9) that results from the maximum pipe moment has been determined by subtracting the pressure stress from the allowable stress (1.5S,) and multiplying by the ratio of stress index associated with high stress points (e.g., elbows). The maximum equation (9) moment stress is given in Table F below for the different pipe sizes. TABLE F Maximum Equation (9) Moment Stress M.LOWAS.E EQ (9) PRESSURE EQ (9) ICMENT STRESS EQ (9) TOT,9. STESS PIPE STESS 1.5S STRESS AT PIPE HANGER AT PIPE HANGER 4" Sch.160 25.25 ks! 4.8 ks! 12.146 ks! 16.95 ks! 10" Sch. 40S 30 ks! 4.4 ks! 6.048 ksi 10.45 ks! 10" Sch. 80 30 ks! 2.6 ks! 9.23 ks! 11.83 ks! 32" MS 29 13 ks! 7.1 ks! 6.7 ks! 13.80 ks! 9 The equation (12) maximum mechanical piping stress is determined in a similar manner; allowable stress (3S ,) times the. maximum ratio of stress indices values. The maximum equation 12 moment stress values at the pipe hanger locations are noted below. TABLE G Maximum Equation (12) Moment Stress EQUATION 12 PIPING PIPE SIZE ALLOWABLE STRESS MOMENT STRESS 4" Sch. 160 50.52 ksi 22.49 ksi 10" Sch. 40S 60 ksi 10.63 ksi 10" Sch. 80 60 ksi 15.15 ksi 32" MS 58.26 ksi 13.34 ksi l l I

l' The total stress intensity for each of the piping si2.es evaluated is given in Table H. TABLE H PRELOADED APPLIED TOTAL STRESS TORQUE STRUT LOAD INTENSITY 4" Sch. 160 60 ft-lbs. 2,000 lbs. 64.14 ksi 10" Sch. 40S 100 ft-lbs. 10,000 lbs. 74.21 ksi 10" Sch. 80 100 ft-lbs. 10,000 lbs. 54.69 kai 32" MS 240 ft-lbs. 100,000 lbs. 47.17 ksi Splitting the above total stress intensity into primary (equation 9) and secondary (equation 12) stress intensities results in the following (Table I). TABLE I Maximum Primary and Secondary Stress Intensities EQ. 9 Eq. 12 (PRIMARf STESS EQ. 9 (SECOIDARf STESS EQ.12 INTEN St TY) M.LOWAE.E INTEN Si TY) M.LOWAB.E (ksil (ksi) (ks!) (kst) 4" Sch. 160 31.60 50.52 32.54 50.52 10" Sch. 40S 60.61 60 13.6 60 10" Sch. 80 38.15 60 16.54 60 32" MS 30.57 58.26 16.6 58.26 As can be seen from a comparison of the above maximum stress intensities to the equation (9) and equation (12) allowable stresses, the 10" Sch. 8 0, 32" MS, and 4" pipes meet the stress criteria established herein. The 10" Sch. 40S pipe essentially meets all of the allowable limits. ig i ,, ,. . . .

O. In the preceeding discussion you have used " general" piping moments at the support location. How do you know that they are representative of the actual piping moments? Moreover, in the presentation of the mechanical stresses reported in Attachment 3, it is stated that these general piping moment stressec are " realistic" values. Please compar'e the Code primary piping moment stresses (NB, NC and ND Equation 9) and secondary thermal piping moment stresses (NB Equation 12 and NC and ND Equation 10) to actual, randomly selected, computer piping analysis stresses. A. In Table 3, summaries of stresses associated with straight runs of pipe for 4, 10 and 32 inch pipe are given. The piping lines and locations were randomly selected by Gibbs & Hill. The maximum stresses from this summary calculated using Equation 9 stresses for each of the pipe sizes given are compared to the Attachment 3 mechanical pipe stresses. This comparison is given in Table J. e i.. ~

TABLE J Comparison Between Equation (9) Stresses Computed by Attachment 3 Method and Stresses Calculated by Gibbs & Hill  : Attachment 3 Stresses l Mechanical Primary Piping Primary Pipe Pressure Stress Moment Plus Pipe Size Moment Stress (ksi) (ksi) Pressure (ksi) 4" Sch. 160 12.146 4.8 16.95 ' 10" Sch. 40 6.05 4.4 10.45 10" Sch. 80 9.23 2.6 11.83 3 32" MS 6.7 7.1 13.8 Gibbs & Hill Attachment - Stresses Randomly Selected Maximum Primary Piping Moment-Plus Pipe Size Pressure (ksi) 4" 7.37 10" 7.063 32" . _ . 10.6 As seen from comparing the primary piping moment stresses presented in Attachment 3 with the randomly l selected maximum primary pipe stresses presented in Table 3, the Attachment 3 stresses are all higher. The maximum (NC & ND) pipe stresses, calculated using equation 10, compiled at random straight run piping analysis locations (see Table 3) ' are compared to the Attachment 3 Code (NB) Equation 12 stresses in Table K below. I a } a I

o . TABLE K Comparison Between Attachment 3 Code (NB) Equation 12 Stresses-and Gibbs & Hill Calculated Stresses Attachment 3 Stresses Equation 12 Stress Pipe Size (ksi) 4" Sch. 160 22.49 10" Sch. 40 10.63 10" Sch. 80 15.15 32" Sch. MS 13.34 Gibbs & Hill Table 3 Code (NC & ND) Equation 10 Stresses Equation 10 Stress Pipe Size (ksi) 4" Sch'. 160 3.6 10" Sch. 40 1.5 32" Sch. MS 5.0 From the above tables, it can be seen that the Attachment 3 values calculated using equation 12 are higher than the randomly selected equation 10 stresses compiled by Gibbs & Hill. Therefore, the primary and secondary stresses used in Attachment 3 (Code Equations 9 ani 12) are conservative when compared to Gibbs & Hill's randomly selected cases. Q. In Attachment 3, " realistic" mechanical pipe stresses were '

                      . determined based on ASME Class 1 rules; how will these stresses vary considering ASME Class 2 and 3 rules?

l

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A. -Using Code Equations 9 and 12 requires that the mechanical  ! stresses in the pipe resulting from sources other than the i U-bolt assembly be included. Because of the number of  ; hangers involved, this information was not supplied to Westinghouse for evaluation purposes. - In order to conservatively estimate the magnitude of these mechanical stresses, a generic procedure was developed and is discussed in Attachment 3. In the development of these mechanical stresses, it is assumed that the pipe is stressed to its maximum allowable value at a high stress intensification point (e.g., elbow). It is assumed that the piping moment which results in this maximum allowable stress also occurs at the U-bolt hanger locations. To obtain stresses due to piping moments, Class 1 rules are used in Attachment 3. If Class 2/3 rules are used to determine the piping moment stresses, the mechanical pipe moment stresses would be different since the stress inches (stress intensification factors for Class 2 and 3) are different. Given below in Tables L and M is a summary of the mechanical pipe stresses based on ASME Class 2/3 (NC and ND of Section III) rules compared with those used to assess the pipe stresses in Attachment 3 based on Class 1 rules (NB of Section III).

65 - TABLE L Deadweight and Seismic Piping Moment Stresses Class 1 Class 2/3 Pipe Size Material Rules (ksi) Rules (ksi) 4" Sch.160 Stainless 12.146 14,3 10" Sch. 40 Stainless 6.05 8.6 10" Sch. 80 Carbon 9.23 11.2 32" MS Carbon 6.7 9.1936 As seen from the above comparison, the difference in primary pipe moment stress using Class 1 and Class 2/3 rules is at most only 2.5 ksi, which is less than five percent of the 3S, allowable limit used for piping stress assessment. TABLE M Thermal Piping Moment Stresses Mechanical Pipe Moment Stresses Class 1 Rules Class 2/3 Rules Pipe Size Material (ksi) (ksi) 4" Sch. 160 Stainless 22.49 26.5 10" Sch. 40 Stainless 10.63 10.7 10" Sch. 80 Carbon 15.15 12.29 32" MS Carbon 13.34 13.02 Only the 4" Schedule 160 pipe has a calculated stress that is higher (18%) following Class 2/3 rules than for Class I ru les. The Equation 12 piping moment stresses for the 10" Schedule 40, 10" Schedule 80 and 32" MS are lower or comparable to the Equation 10 stresses developed using Class 2/3 rules.

B

   -    .                                                         ~

In conclusion, the mechanical pipe moment stresses used in Attachment 3 are conservative or yield only slightly lower (<8% based on 3S,) stresses than those based on Class 2/3 rules. The mechanical pipe moment stresses given in Attachment 3 are adequate estimates for the mechanical' pipe stresses to be used in the assessment of the local pipe atress state in the vicinity of the U-bolt piping support assembly. Q. You have stated that Applicants adopted two alternate approaches to evaluate the acceptability of pipe stresses. In the first alternate approach, described previously, the total preload stresses were considered as a primary stress in NB code Equation 9 with a 3S, limit. What is the effect on the piping stress evaluation if a 1.5S, limit is used for Code Equation 9 and only the primary membrana portion of the U-bolt preload, push and pressure stress is considered? In other words what conclusions would you reach if you follow the second alternate approach? Here we present the result of Applicants' evaluation using-the second alternative acceptance criteria (see p. 54 for first alternative).

                                          -9 Evaluation Using Second Alternative A. To determine the primary membrane portion of the U-bolt preload, push and pressure stress, the stress state at the inside and outside of the pipe element surface is averaged.

To these average stresses, the mechanical piping stresses were added. The resulting total stress state is calculated using Code Equation 9 and compared to the allowable stress (1.5S,). The results are given in Table N below. TABLE N Eq. 9 Stress 1.5S, Allowable Pipe Size (ksi) (ksi) - 4" Sch. 160 20.99 25.26 10" Sch. 40 18.44 30.00 10" Sch. 80 15.33 30.00 32" MS 11.50 29.13 As seen from the above stress allowable tabulation, Code Equation 9 defined stresses are less than the 1.5S , stress allowable. To complete the evaluation of the piping stress, Code Equations 10 and 12 (if Equation 10 is exceeded) must be evaluated using an allowable of 3S,. Code Equation 10 calculates the stress intensity which occurs in mechanical or thermal loadings which take place as the system goes from one load set, such as pressure, temperature, moment and force loading to any other load set which follows it in time. Loads which are noneyclic in nature need not be considered in Code Equation 10. The

n \^ - maximum Equation 10 primary plus secondary stress intensity range for the four pipes evaluated is tabulated below in Table O. TABLE O Eq. 10 Stress Eq. 10 Allowable j Pipe Size (ksi) (35S_) - & 4" Sch. 160 50.8 50.57 '- 10" Sch. 40 44.33 60.00 10" Sch. 80 39.83 60.00 32" MS 43.58 58.26 As seen from the above comparison, only stresses for the 4" Schedule 160 pipe slightly exceed (0.5%) the 3S limit. Two items must be noted concerning the conservEtism of the stress state calculated using Equation 10 for the 4" Schedule 160. They are:

1. The dead weight stress of the pipe (1 to 1.5 ksi) is included. This should be removed from Equation 10.

2. Conservative values for the mechanical piping stresses (primary 12,146 ksi and secondary 22.49 ksi) have been used. The conservatism was previously demonstrated by comparison with piping moment stresses calculated and randomly selected by Gibbs & Hill.

If the conservatism reflected by these two items were removed, the Equation 10 stress for the 4" Schedule 160 pipe would be below tie 3S, limit. Since the rquation 10 stress allowable (3S,) is not exceeded for the four pipe sizes, it is not necessary to consider stresses calculated using NB Code Equation 12. In conclusion, as demonstrated from the above, the piping stresses in the vicinity of the U-bolt piping assembly will be acceptable.

Q. Please discuss the effect of U-bolt cinching on the fatigue usage factor associated with the pipe.

                                                                                                          ^

A. To ascertain further the possibility of failure under cyclic stresses caused by the clamping action, a. fatigue assessment was performed to determine the possibility of initiating a crack in the pressure boundary at the vicinity of the U-bolt. This assessment included a calculation for Class 1 piping and a calculation for Class 2 and 3 piping. The fatigue calculation for Class 1 piping utilized the U-bolt stresses from the 4" Schedule 160 case. The calculation for Class 2 and 3 piping utilized the worst case peak stresses from the four pipe /U-bolt assemblies evaluated. For the Class 1 piping, locations having the relative highest usage factors, excluding U-bolt considerations, were identified so that the maximum overall usage factor including the effect of the U-bolt could be determined. These locations were the 10" accumulator injection line and the 3" normal charging line. These two Class 1 piping locations represent the highest straight run fatigue usage factor (due to extremely severe thermal transient loadings) of all the Class 1 auxiliary piping. The fatigue calculation performed included all design transients for these sections and added the maximum clamp induced stresses as represented by the 4" Schedule 160 case. I

s For the charging and accumulator lines the final usage factors considering U-bolt local stresses were .04 a nd .06, re'spectively. Based on an ASME Code allowable usage factor of 1.0 or 0.1 for pipe break requirements, the usage factor requirements are met. For the Class 2 and 3 fatigue assessment, a simplified worst case fatigue calculation was performed. In this calculation, a conservative peak stress (Sp) was determined using the maximum total stress intensity without preload. The effect of preload is not included since it is noneyclic and not required to be considered per the code in a fatigue evaluation. The 4" Schedule 160 pipe has the largest peak stress: Sp = 50.57 ksi Based on this stress, the shakedown criteria (allowable

     = 3S,= 50.52 ksi) is met recognizing that S p               also includes deadweight stress which is noneyclic.                Therefore, the simplified elastic / plastic damage factor equals one (Ke =

1.0). The alternating stress (Salt) is: S = P = 25.3 ksi 2 The allowable number of cycles associated wit.h S Alt is 22,000 (Code Figure I-9); Ultimate Tensile Strength = UTS < 80 ksi).

V - s' To determine the number of loading cycles, the worst of the following two cases were considered. For Case 1, 200 cycles were considered based on 200 cycles of heatup and cooldown, or for Case 2, 200 cycles were considered, based on 10 earthquakes, 20 significant cycles per earthquake. The incremental usage factor is: Ui = 22 000 = .009 The fatigue assessment just described was based on both a representative Class 1 method and a Class 2 and 3 method. It provides a measure of the fatigue damage that would be expected at the U-bolt location. This assessment indicates that the integrity of the pressure boundary, based on fatigu.e considerations, would not be significantly affected by the localized U-bolt effects. This assessment is over and above the ASME Code requirement for Class 2 and 3 piping components. The Code does not require any specific fatigue usage factor calculation for Class 2 and 3 piping other than controlling stress levels by the factor f. Q. How would you interpret the results from the tests and finite element analyses regarding the stability of the U-

              . Bolt assembly?

h- i. iiii..i--

9 - A. Stability of the U-bolt assembly is an issue that was addressed using the " push" applied load. For a " pull" load, the U-boat assembly, if it slipped, would be self-aligning and therefore not a stability concern. Stability for'a " push" force requires that sufficient friction exists between the cross piece and pipe, and the U-bolt and pipe to " balance" an applied load which is not directed through the pipe centerline. Attachment 3 addresses the method by which the U-bolt assembly balances the " push" load. The results given in the report show that the U-bolt assembly must be evaluated as two coupled assemblies (U-bolt and cross piece) rather than one assembly which behaves rigidly. This is because slippage can occur at the pipe and cross piece interface and not at the interface of the U-bolt and pipe. The results from the finite element analyses (using the minimum torque values) indicate that the U-bolt / pipe assembly will be stable. The minimum preload value used in the analyses is not the absolute since a determination of the absolute minimum value is an iterative process. No attempt was made to further define the absolute minimum torque value which could assure stability since the reported values are already generally representative of the lower bound value found in the field (with the exception of

e - 73 - the value of the 3 2-inch pipe) . The minimum preload value that would insure stability can also be lower than the values given by finite element analysis since the pipe can roll on the cross piece until sufficient normal forces are developed between the U-bolt and pipe, along with friction forces, to balance the cross piece moment. This physical mechanism in the cross piece and U-bolt interaction, not represented in the finite element analyses, would be self stabilizing. If the U-bolt assembly was pushed in the circumferential, direction, the tension would increase in one U-bolt leg and decrease in , the other one resisting the applied load. As long as the load is not applied to the cross piece at an angle which exceeds the inverse tangent of the friction coefficient, the cross piece /U-bolt assembly will not slide. As stated previously this mechanism is not incorporated in the finite element model and therefore the minimum torque values

   ,   determined by finite element analysis are larger than those                               ,

that would be needed in reality. From the finite element analyses performed and reported in Attachment 3 and observation of the tests, torque values to which cinched down U-bolts will be torqued during the previously mentioned inspection were established e - - - , . , - . , -+r.--- . , - - - - .

for various sizes (Table P). (For other sizes, torque values in line with those set forth in Table P will be established.) TABLE P Torque Values for Inspection Pipe Size Torque Value 4" Sch. 160 25 ft-lbs.* 10" Sch. 40S 50 ft-lbs. 10" Sch. 80 50 ft-lbs. 32" Main Steam 250 ft-lbs.

• 25 ft-lbs. is a compromise between 9 ft-lbs. at which the finite element analysis predicts that the 4" assembly would be stable and tie 35 ft-lbs. at which the test indicated percent contouring of the U-bolt l to the pipe was achieved. Torquing of the U-bolts in l the field indicates that for this size lines and U-l bolt diameters, perfect contouring of the U-bolt to the pipe occurs at 20-25 ft-lbs.

The U-bolt piping support assembly can be stable for smaller torque values. A good practice for U-bolt installation is to insure that there is a good fit ( " Snug" tight) between the U-bolt and pipe at torque values given above. Q. Since the torque value for inspection (Table P) torque value for the 3 2-inch pipe is larger than the value at which the analyses for the piping stresses have been conducted, would there be an adverse effect on those stresses from this increase? l i l

i A. No. The piping stresses will remain within the established allowable since the piping stresses due to preload are small t in the 32" main steam pipe (4.9 kai). Therefore, the total primary stress would be about 36 ksi, which is still below the established allowable. Q. What overall conclusion would you draw from the testing and analyses program? i A. As mentioned previously, the Applicant committed to an analytical and testing program that would address the , Board's following concerns: [ 1. U-bolt cinching may not be an appropriate means of eliminating potential local instability without introducing adverse effects in the piping and the U-j bolt itself.

2. There may be adverse long-term effects from U-bolts' due to heat-up and cooldown and related friction on the pipe.

l

3. Stresses on the pipes caused by thermal expansion in local areas around cinched U-bolts may not be J

acceptable. These concerns were addressed within this Affidavit, and it was shown that cinching of the U-bolts is a viable method of eliminating potential local instability of the U-bolt strut supports. There are no adverse effects in the 4 piping or U-bolts themselves from the cinching process. This statement is based on the fact that the U-bolt is generally cinched to a preload value which is less than the ] W

value used for test and analysis. Therefore, stresses i existing in the field will be less than those obtained by

                                                                                                                                                                                                       ,\

the finite element analyses, which are shown to be acceptable, and also will be less than those obtained during  ; testing, both of which were based on the highest value of  ; pretorque.  ! , With regard to stability, there is some concern that if the pretorque level is insufficient, i.e., the U-bolt is not  ; snug tight, that the assembly might work loose and walk. i

                                                                                                                                                                                                       ?

This doesn't mean that the support is unstable in the sense of not being able to carry load. (The test program showed, ) + for example, that when the 10-inch Schedule 40S assembly was torqued to 20 ft-lbs., the U-bolt / cross " walked" both axially and radially,"but was still transmitting and carrying load.) However, as previosuly noted, Applicants i will inspect every U-bolt on a single strut or snubber (a , total of 380) and assure that each U-bolt is torqued to the  ; torque values set forth in Table P. ' , The specific points that were made pertaining to the f t Board's concerns demonstrating that cinching of the U-bolts were the following:  !

1. Issue - U-bolt cinching will eliminate potential suppart instability.

I

                                                                                                                                                                                                     ~
                                                                                                                                                                                                       )

k [

r. i Response - It was demonstrated by both testing and finite element analyses that when the U-bolt is cinched to a minimum preload value, the support will be stable.

2. Issue - Cinching of the U-bolt will cause adverse effects in the piping and U-bolt itself.

Response - Stress results obtained from the finite element analysis of the U-bolt support piping assembly associated with the anticipated support and piping loads as well as recommended preload values are within acceptable limits. The support as well as pipe will not experience any gross

       ' distortion or loss of function.

3. Issue - There may be adverse long term effects from U-bolts due to heat up and cooldown.

Response - Both test programs and analyses have demonstrated that this is not a valid concern. 9

78 - f f

                                                                                          ' Y]

Robert C. Iotti i Sworn to before me this day of June, 1984. ,

                    ,, o s i ,a_              u         e4 Notary Public My Commusion Expires May,31,1987 wf         A-dohn C. Finneran, Jr.        f Sworn to before me this                        day of June, 1984 Opf,e ,'s ut * [ ] /
           /

A s /?*t k Notary Public

            . My Commmion Exgircs Msg 31.198Z-

c. -

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r TABLE 2 Observation of Torques in U-Bolts Pipe Size No. of U-bolts Torque Sampled Range (ft-lbs.) 3/4" 4 4-8.5 1" 2 5-5.5 2" 14 4.5-20 3" 15 9-30 4" 31 7.5-60 6" 16 10-75 8" 18 8-70 10" 14 10-83 12" 8 t4 26-87 14" 4 s 19.5-95  ; 16" 13 %25-126 1 18" 3 35-117 I 20" 2 55 108 4 l 24" 9 30-05 26" 1 22 30" 4 55-112.5 40" 2 61-82.5

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\ h1E3- fr.re.undSrae<rw-aw-wenew;;r Gibbs 8 Hill. inc. Job No. 2.3 2 3 Client n a ce./ cns unn } ~ Subject S 7* * ' ' * " " ' WG " &"- d ito A * ? '" c " D "' u \ ~ Calculation Number - SheetNo. ~-

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o- . REPORT NO. EQ&T-EQT-860 REVISION O COMANCHE PEAE STEAM ELECTRIC STATION 0-BCLT SUPPORT / PIPE TEST REPORT WESTIN0HOUSC ELECTRIC CORPORATION P. 0'. B01 355 PITTSBURGH, P4 15230 kM[eafM f/s7/9+ PREPARED BY: MARE EAMENIC

                                                     .)   .              W"Y REVIEWED BT.                    J.M. 3           R r

APPROVED BY: ly. L

                                                         . bLE!3pp
                                                                   )

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i . . Table of Contents h tien , Descrintion 23gg I Introduction and Scope 3 II- Objective 3 III Test Items 3 IV Sumasry - 4 Y Test Description and Results 6 Y.A Torque Versus Preload Test 6 V.B Friction Test 24 V.C Load Distribution / Strain Measurement Test 41 V.D , Thermal Cycling / Thermal Gradient Test 50 Y.E Creep Test , 64 V.T Normal Vibration Simulatio'n Test 72 Appendix I Engineering Calculations 77 Appendix II U-bolt Identification 82 Addendum.1 Conianche Peak Steam Electric Station 85 U-bolt Support / Pipe Test Report e 2

f o . I. INTRODDcTION AND SCOPE Tests were carried out by Westinghouse under contract with Texas Utilities Generating Company (TUGCO) to investigate friction and stresses in piping /U-bolt support conrigurations in use at the Comanche Peak Steam Electric Station. More specifically, in accordance with TUGCO's request, tests were performed,to investigate the following:

a. The range of friction coefficients developed between the pipe and the U-bolt support while the U-bolt is preloaded to specified levels.

b. The effects of hanger size and pipe satorial on the friction coefficient for a constant U-bolt preload,

                                      ~

c. The stresses and strains generated in the U-bolt.

d. The stresses and strains generated in the pipe.

e. The ext $nt of long ters effects such as normal piping vibration and cyclic (hermal expansion on U-bolt preload.

To address the above itess, Westinghouse developed a test pregram which consisted of six separate tests as shown below:

1. Torque versus preload test

2. Friction test 3 Load distribution / strain sensurement test

4. Thermal cycling / thermal gradient test

5. Creep test

6. Normal vibration simulation test The tests were performed under test conditions defined by TUGC0 at the facilities of Westinghouse Advanced Energy Systems Division (MAESD) in Large, Pennsylvania during the time frame of March 1984 through May 1984 in accordance with nrocedures developed by Westinghouse and approved by TUGCO. These tests are described and the results are presented in Section Y.

II. ORJECTrrE The test program had two basic objectives: (1) Confirm whether a nonslip condition exists between the U-bolt support and the piping and (2) Provide test data to facilitate the correlation of test results with analytical models of the U-bolt support / piping. III. TEST ITEMS . Test items were provided by Texas Utilities Generating Company and were delivered to the Westinghouse Advanced Energy Systems Division Engineering Laboratories located in Large, Pennsylvania. The following test specimens were provided by TUGC0 for use in the test program: 3

l I o o Number of Pipe Size Material Lenath U-Bolt Sunnarts 32= Carbon Steel 10' (10'-7=)e 1 108 Stainless Steel 10' (9'-9 3/4")e 2 10" Carbon Steel 10' (9'-8 1/4")e 2 48 Stainless Steel 10' (9'-9 1/2")e 2 U-bolts were identified by a unique number for their use on a specific pipe size prior to the initiation of the test program. (See Appendix II) Insulation, struts, pins and clavises were also provided by TUGC0 so that test configurations could simulate as closely as possible the actual field

     ~~~

inats11ation configurations. IV.

SUMMARY

Six separate tests were performed under this program as identified below: Torque versus preload test: Preloads produced in the U-bolts were determined for the four pipe test specimens as a function of U-bolt torque. Pipe stresses resulting from U-bolt proloads were also

                                                                          ' determined. Results from this testing are presented in Section Y. A.,

pages 6 throut.h 23 Friction test: The forces required to cause the U-bolt supports to slip on the pipes were determined for the four pipe test specimens at various increments of U-bolt torque. The coefficients of friction between the U-bolt supports and pipesLwere also determined over a defined range of U-bolt torques. Results from this testing are presented in Section Y.B., pages 24 through 40. Load distribution / strain measurement test: U-bolt loads and crosspiece stresses were determined for given U-bolt torques and support loads on the 10" stainless steel pipe test specimen. Results from this testing are presented in Section Y.C., pages 41 through 49 Thermal cycling / thermal gradient test: The effects of thermal cycling at elevated temperature on U-bolt preload were investigated on the 48 stainless, 10' stainless and 32a carbon steel test specimens. Results from this testing are presented in Section Y.D., pages 50 through 63. Creep test: The effects of high temperature as a function of time on the U-bolt preload were evaluated on the 48 stainless, 10" stainless and 32' carbon steel test specimens. Results from this testing are presented in Section V.E., pages 64 through 71. Normal vibration simulation test: The effects of piping vinration on U-bolt preload were evaluated on the 10" stainless steel pipe test specimen. Results from this testing are presented in Section V.F., pages 72 through 76.

• Actual test specimen lengths are shown in parentheses.

4

O O The friotion test was performed to confirm whether a nonslip oonditiod exists between the U-bolt supports and the piping. The other five , tests provided results used as a basis for development of analytical models of the U-bolt supports and piping. 9 e 1 5

m a l I V. TEST DESCRIPTION AND RPRULTS V.A. Toroue Versus Preload Test 1.0 Purname

                          ~ The purpose of this test was:                 (1) measure the preload produced in the U-bolt below the U-bolt support crosspiece as a known amount of U-bolt nut torque was applied, and (2) measure the stresses in the pipe caused by the preload in' the U-bolt.                This testing was performed on four test specimens (pipes) using three designs of U-bolt type supports as specified and provided by TUG CO.

Pine Test Some4mann 32 inch carbon steel pipe 10 inch stainless steel pipe 10 inch carbon steel pipe 4 inch stainless steel pipe U-bolt Desi an a 32 inch U-bolt type pipe support to inch U-bolt type pipe support' 4 inch U-bolt type pipe support 2.0 Test confienration (see Photogr'aphs 1 through 4) The four' pipe test specimens described in Section 1.0 had a mounting plate velded to each end. These plates were then fastened to two floor mounted brackets using high strength fasteners. The carbon steel test specimens had a Steel Structures Painting Council No. 2 finish applied on a four foot section in the center of the pipe. Strain gauges were installed on the U-bolt, pipe and support crosspiece. (Note: No strain gauges were installed en the crosspiece of the 4" specimen.) The U-bolt suppsrt was then installed in the center of the test pipe with the nuts hand tightened.

                                        \

• This design of U-bolt support was used on both the 10 inch carbon and 10 inch stainless steel pipes.

6

D 6 30 Instrumentation Instrumentation used for these tests is identified below.

1. Strain gauge signal conditioning and readout devices

2. Seventeen strain gauges The strain gauges were installed on each test specimen prior to testing (see Sketches 1 and 2). The above instrumentation was calibrated prior to use.

4.0 Loading Devices Loading devices used for these tests are identified below.

1. Torque wrench

                .The torque wrench was calibrated prior to use.

5.0 Land in nli cation The load induced into the test specimen was due to the U-bolt nut torque. No other loads were applied. Tests were performed at laboratory ambient envircumental conditions. 6.0 Test Descrintion After a specimen had been installed and the instrumentation calibrated, testing cassenced. The test steps were the same for the four test' specimens. The testing variable was the U-bolt nut torque range. The torque ranges for the U-bolts were defined by TUGC0 as: 4 inch test specimen: 5 to 60 foot pounds in increments of 5 foot pounds. 10 inch test specimens: 10 to 100 foot pounds in increments of 10 foot pounds. 32 inch test specimen: 20 to 240 foot pounds in increments of 20 foot pounds. With the test specimen U-bolt nuts hand tight, the strain gauges were zerood. The U-bolt nuts were then torqued to the lowest value in the ranges defined above. Readings were recorded from each strain gauge. When strain gauge data had been recorded, the U-bolt nuts were alternately (torque Leg 1 and then Leg 2) . torqued to the next increment. Again the strains were recorded. This process was repeated for each test s;.ecimen up to and including the maximum torque as defined above. 7

i 70 Data Reduction l l The test output consisted of U-bolt and piping strains. Using  ! the strain data, U-bolt proloads were determined. The analytical method for determining U-bolt preload is contained in Appendix  ! I. Longitudinal and circumferential pipe stresses were also determined using the elastic stress-strain relationships t contained in Appendix I and the strain readings obtained from the l

         -        pipe strain gauges,                                                j 8.0 Data Descrintion/ Presentation                                      l Test output was reduced and plotted so results could be presented graphically. U-bolt preload versus bolt torque was plotted for     ,

each of the four specimens in Figures 1 through 4. From the , piping strain measurements, the maximum longitudinal or  ; circumferential stresses in the pipes were determined. These values are plotted versus the U-bolt preload or bolt torque for each of the specimens in Figures 5 through 8.  ; 9.0 Results Results show that U-bolt preload increases proportionally to bolt torque. Proloads associated with the maximum applied bolt [ torques are: l o 4 inch stainless steel specimen: 5918 pounds at a bolt  ! torque of 60 foot pounds. "See Figure 1. A conforming fit between the U-bolt support and the pipe was h achieved at a U-bolt torque of 35 foot pounds, conforming i fit is defined as the absence of visual gaps between the j U-bolt support and the pipe. e f o 10 inch carbon steel specimen: 8057 pounds at a bolt  : torque of 100 foot pounds. See Figure 2.  ! j A conforming fit between the U-bolt support and the pipe was achieved at a U-bolt torque of 70 foot pounds. j o 10 inch stainless steel specimen: 5444 pounds at a bolt  !

       .               torque of 100 foot pounds. See Figure 3                       l A conforming fit between the U-bolt support and the pipe was achieved at a U-bolt torque of 50 foot pounds, o    32 inch carbon steel specimen:     4651 pounds at a bolt      :

torque of 2'40 foot pounds. See Figure 4. ' A conforming fit between the U-bolt support and the pipe was not achieved during tae Torque versus Pr_eload Test even at the bolt torque of 240 foot pounds. A conforming fit [ 8

l B 6 between the U-bolt support and the pipe was achieved during the Friction Test at a bolt torque of 1170 foot pounds. Preload variation between the two legs of each U-bolt ranged from 165 for the 4 inch stainless steel specimen to 265 for the 32 inch carbon steel specimen. The pipe stresses associated with the maximum applied bolt torques were: . o 4 inch stainless steel specimen: The nazimum stress was circumferential with a value of -7217 psi measured at strain gauge S8. See Figure 5 o 10 inch carbon steel specimen: The maximum stress was circumferential with a value of -8063 psi measured at strain gauge 39 See Figure 6. o 10 inch stainless steel specimen: The maximum stress was circumferential with a value of -19.386 pai measured at strain gauge S12. See Figure 7. o 32 inch carbon steel specimen: The maximum stress was circumferential with a value of -1562 psi measured at strain gauge 39 See Figure 8.

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l I I l i Y.B. Friction Test i ! i 1.0 The purpose of this test was to measure the coefficient of i friction between the U-bolt type support and pipe while the torque in the U-bolt was varied from a minimum to a maximum i

                      . level. The maximum U-bolt torque level was defined as the one that could develop a U-bolt thread stress of ninety percent of                                                                     i yields during the test. This limiting criterion was established                                                                  l to preclude yielding of.the U-bolts. This permitted use of the                                                                   .

same U-bolts in subas+2ent -tests. The data obtained from this , l l test was also used to define the relationship between the force f l that causes the U-bolt supppor.t to slip on the pipe and U-bolt  ! l torque. Tests were performed on four test specimens (pipes) } using three designs of U-bolt type supports as specified and , provided by TUGCO. - Pine Test Soecimens i 32 inch carbon steel pipe , 10 inch stainless steel pipe 10 inch carbon steel pipe  ; 4 inch stainless steel pipe l U Bol t Danteen , I 32 inch U-bolt type pipe support l 10 inch U-bolt type pipe support l l 4 inch U-bolt type pipe support  ; Note: Two (2) U-bolt type supports were tested on both 10 inch l pipes and the 4 inch pipe. The second support was used to  ; evaluate polishing phencuena. Only one 32 inch specimen was i available for testing. The effects of polishing on the results [ obtained for this specimen are discussed in Section 9 0 below. [ f l t t efield was defined as minimum yield for A-36 1.e. 36,000 pai. l I 24 j

                  - _                 , ,. _ _ _ . - , , _ . .    ,    _m.,

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• a  !

i i i i i 2.0 Tant conr<,uration (See Photographs 5 through 7). , The four pipe test specimens described in Section 1.0 had a l mounting plate welded to each end. These plates were then l fastened to two floor mounted brackets using high strength j fasteners. The carbon steel test specimens had a Steel  ! i l Structures Painting Council No. 2 finish applied on a four foot  ! section in the center of the pipe. Strain gauges were installed j on the U-bolt and crosapiece. The U-bolt was then installed in the center of the test pipe with the crosapiece perpendicular to the floor, the U-bolt nuts were initially hand tightened. A i loading apparatus consisting of a hydraulic cylinder and a load cell was then installed. 30 in=trunentation  ! Instrumentation used for these tests is identified below. j

1. Strain gauge signal conditioning and readout devices  ;

2. Load cells  ;

3 Magnetic tape recorder and brush recorders l

4. Eight strain gauges  ;

' Strain gauges were . installed on the U-bolt support (see Sketch 3). The above instrumentation was calibrated prior to use. i l 4.0 fandine nevices , Loading devices used for these tests are identified below. \ l

                                                                                                               -                               l

1. Torque wrench

2. Hydraulic cylinder l i

l The torque wrench was calibrated prior to use and the hydraulic ' cylinder load was controlled by the above referenced load cells. l 5.0 t n=d An ni t enti on } The test specimens were subjected to two types of loading during " this test. Torque loads were applied in increments to the U-bolt - nuts and the hydraulic cylinder / load cell was used to apply a load at the crosspiece clevis hole centerline tangentially to the  ! pipe and parallel to the crosapiece. Tests were performed at laboratory ambient environmental conditions, j 6.0 Tant neuerietion j After a specimen had been installed and the instrumentation j calibrated, testing commenced. The test steps,were the same for

• the four test specimens. The testing variables were the U-bolt l nut torque increments and the load applied at the clevis to cause I

25 l

 -       o                                                                                                       t i

I l alippage between the U-bolt support and pipe. The torque  ! increments were defined by 1UGC0 as:  ! 4 inch test specimen: 5 foot pound increments. , 10 inch test specimens: 10 foot pound increments. j 32 inch test specimen: 20 foot pound increments. With the test specimen U-bolt nuts hand tight, the strain gauges  ! were zerced. The hydraulic cylinder / load cell was connected and I positioned so that it was not loading the U-bolt in either the upward or downuard vertical directions. The load cell reading l vas then zerood. . The U-bolt nuts were then to'rqued to the first incremental value i and the U-bolt strain recorded. With the recording equipment activated, the hydraulic cylinder load was increased until , slippage occurred. Slippage was defined as a negative change in l slope of the load time history. i The U-bolt strain in the leg opposite the loading was checked against the following maximum levels: 4 inch test specimen: maximum strain = 839 u inches per inch. i 10 inch test specimens: maximum strain = 877' u inches per inch. . l 32 inch test specimen: maxistia strain = 964 9 inches per inch. I ! t ! The above strains correspond to U-bolt thread stresses equal to j 905 of the acterial yield. l t If the above maximum strains were not reached, the U-bolt nuts l l were loosensd and the pipe support roturned to its original ' i . position. The U-bolt nuts were then alternately retorqued to the l next incremental torque value and the load applied until slippage [ occured. This sequence was repeated until the maximum strain levels were reached. , once the maximum U-bolt strain levels ware reached the U-bolt was  : loosened and returned to its original position. The U-bolt nuts  : ! were then retorqued to the same value and thi friction test was repeated two more times. This procedure was performed to establish data repeatability. , After ocepletics these tests, the U-bolt support was removed and a second U-bolt support was installed on the pipe several inches , away from where the original had been positioned. The friction f testing was repeated with the U-bolt preloaded to a strain value , equal to one-half of the value at which the repeatable tests were  ! performed. This was done to evaluate polishing phencuena.  ; Polishing is defined e+ tha smoothing of the U-bolt l support / piping contact surfaces due to wearing action.  !

                                                               \

26 , r

l - i

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l t i E 9 i 7.0 Data Reduction The test output included U-bolt strains and maximum slip friction  ! forces. This information allows for a comparison of the maximum l alip friction force versus bolt torque and also determination of l the coefficient of friction between the pipe and the U-bolt  ;

          - support. The analytical method for determining the coefficient               ;

of friction between a U-bolt support and pipe is presented in l Appendix I. , , 8.0 Data Descrietion/ Presentation [ Test output was reduced and plotted so results could be _ presented graphically. Maximum slip friction force, that is the force applied to the support at which friction was overcome, was plotted versus bolt torque for each of the four test specimens in Figures 9 through 12. coefficients of friction were determined and plotted versus bolt torque for each of the four specimens in Figures 13 - 16. 9.0 nasuits Results determined for each test specimen are as fol16va: o 4 inch stainless steel - Coefficients of friction ranged  ! from .12 to .193 Slip friction forces ranged from 270 , pounds at 15 foot pounds of.U-bolt torque to 1050 pounds at , 35 foot pounds of U-bolt torque. See Figures 9 and 13 Polishing was evaluated by performing a final friction test ' on a virgin pipe section. Results indicate that this specimen did not experience polishing while subjected to the specified test conditions. o 10 inch carbon steel - Coefficients of friction ranged from 1,

                  .196 to .279 The trend indicates that the coefficient of              I friction increases along with bolt torque, and consequently the normal forces between the pipe and support. Slip                  [

- friction forces ranged from 320 pounds at 10 foot pounds of , U-bolt torque to 4725 pounds at 160 foot pounds of U-bolt- . torque. See Figures 10 and 14. Polishing was; evaluated by t performing a final friction test on a virgin pipe section. Results indicate that this specimen did not experience  ! polishing while subjected to the specified test conditions. j o 10 inch stainless steel - Coefficients of friction ranged from .160 to .225. There is a slight trend for the  ; coefficient of friction to decrease as the bolt torque ' increases. This is attributed to the polishing phencoena occurring between the pipe and support due to the repeated  ! test steps. Polishing was also verified by performing a , final friction test on a virgin pipe section. This test i yielded a higher coefficient of friction, .231, than did the 1 6

                         .                                                              I 27                                            ,

r 1 l l polished section at comparable strain levels. Slip friction t forces ranged from 330 pounds at 10 foot pounds of U-bolt  ; torque to 3500 pounds at 130 foot pounds of U-bolt torque. See Figures 11 and 15.

             ~

o 32 inch carbon steel - coefficients of friction ranged from  !

                  .220 to .516. Slip friction forces ranged from 3200 pounds at 200 foot pounds of U-bolt torque to 11,000 pounds at 1170                                  l foot pounds of U-bolt torque. See Figures 12 and 16. Data obtained at bolt torques of less than 200 foot pounds was                                      ;

not included in the results. At these low torque levels the i dead weight of the support caused normal forces between the i U-Dolt support and piping that were far more significant  ! than the normal forces caused by the bolt torques, is a  !

           ,     result, unrealistic coefficients of friction in excess of 1.0 were determined. From the results of the 10 inch carbon                                    ,

steel specisen where friction coefficients ranged from .196 l to .'279, it was verified that coefficients of friction exceeding 1.0 were unrealistic, l L Because of the load capability limitations of the torque ' wrench, the marimum strain of 964 inches per inch was  ! never achieved. The trend was for coefficients of friction to decrease with  ! increasing bolt torque. Even though a final friction test  ; to evaluate polishing could not be performed on this ' specimen, the decrease in:. coefficients of friction was characteristic of the polishing phenomena. i l l I i f f r I 28 i t

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                                           ,                Y.C. Less bistribution/stenin Mammurement Test                                                                                                                      !
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                                                                              .1.0                 2Er2E.ak                                       s
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j This test _was dcaigned to determine the load in the U-bolts and. t ' 4 1 l y ,, ; . crosspiece stresses for given U-bolt nut torque values and j n-  ;, t

                                                                                                 ' support loads. With the U-bolt nuts torqued to three different                                                               !

values and'the U-bolt support subjected to fcur different values L of support lods ac each value of torque, strains in the [ crosspiece and U-bolt were measured.' This test was perfermed on  !

                                                                                                ;tne 10 inca stainless steel specimen as specified and provided                                                                 ;
                                                                                                 'by TUGCO.                                             ' -

s - i

                                        ~

0 i.0 Test confinuvation (Me Photograph 8)

                                                                                                                                     -                                                                                           i,
                       )!                                                                          The test specissa had nounting plates welded to each end                                                             These     ,
        '~

y 4 ,' plates were then ftstorad to two floor sour.ted brackets using j l . high stre gth fasteners. -Strain gauges rere installee on the i

       ^                     '

c.U-bclt and support croc:spiece. The U-boltfsuppart was then  ;

     '                     . 'V '

f.nwelled in the 'centay ;ct the test pipe vitt the crosspiece ; I

                                                                                                ~ ts.allel to the flocr. ~ Ttie U-bolt nuts tore hand tightened. A                                                              !

s i 1Jadtr.g device consisting ,of-a hydraulic! cylinder and load cell, , ehich providtd loading tu the vertical upward and downward f

- ), . , directions, Nas iratalled.

tq ,, s c i l_ , lP* ll'~ f 3.0' 1pO*:InsanLuinn j {'~' I Instrumentation used for this testing is identified below.

                                                                                                   %,                                                    [,                                                                     ,

1. i Strain gauge elsnal canditioning and readout devices-  !

! , j. , 's , , T. Losd cella , - .- Q'f p/

                                                                                                -3.          ;Cighteen atrain gauges                                                  -
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                                                                                                 .The; atrain grarge.i were instialled on the test specimen prior to                                                            l testing (see Sketuh'4). The above instrumentation was calibrated i                                                                                                                                                                                                               l
                'I                        ];

prior to use.

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                                                                                                                                                                                          ,                                   l,
                                                                                                                          ,                  y 'i                                                                                 7 iN               r                                                                          f LoadirJt devices used fo: this teac are identified below.^
                                                                                                                                                                          ^

l

  .                                                                                 .'                 -                                                                                                                         l Torque wrench
                                                           , ,                                    ,1 .
                                 '- $1 ,

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2. Rydreille cylinder -

i

                                                                                                ' The' torque wrecch was. ca153;ra",ed prior to use and the hydraulic e                                cilinder load ves ccatrolled by the above referenced load cella.

i 5.0* Land Anni

• cation I f " r. ,

The tes't specimen was subjected to two types of loading during  ; e ,. 7'" , 2- this test (' Torque J oods were applied to the U-bolt nuts and the 6r - '5bydraulic'cy'.inder/ load cell was used 'to apply a vertical upward t(> _.-

                                                   "                                        . 4                                                                                                      e
            '               .                                                           '+                        ,                             .

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                                                                                                                                                                             .                                                  y

and downward load at the U-bolt support clevis. Testing was performed at laboratory ambient environmental conditions. 6.0 Test Descrietion After the specimen had been installed and the instrumentation

            , calibrated, testing commenced. The loading apparatus was used to induce the following loads on the test specimen as specified by TUGCO:

7208 lbs. 5406 lbs. 3604 lbs. 1802 lbs. These loads were applied bot'h vertically upward and downward at each U-bolt torque value. The U-bolt was torqued to the following values as specified by TUGCO: 100 foot pounds 67 foot pounds 33 foot pounds These torque values were applied in increments of 10 foot pounds prior to each load test. The U-bolt strain gauge readings and crosapiece strain gauge readings were recorded at each of the three torque levels without the. support loads applied and then recorded again with the support Eloads applied. 70 Data Reduction The test output inclutsd U-bolt and crosspiece strains. This information was used to determine the load in the U-bolts and crosspiece stresses for each U-bolt nut torque and support load. The analytical methods for determining U-bolt load and crosspiece stresses are contained in Appendix I. 8.0 Data Descrintion/ Presentation Test output was reduced and plotted so results could be presented graphically. U-bolt load was plotted versus support load for bolt torques cf 33, 67 and 100 foot pounds. Results due to the support load applied vertically upward are shown in Figure 17

Figure 18 contains results for the support load applied i vertically downward. Maximum crosspiece stresses were plotted versus support load for bolt torques of 33, 67 and 100 foot pounds. These results are included in Figures 19 and 20 for the support loads applied vertically upward and downward,

respectively.

42

                    -                                                       . .. . _ .- . . . . 1

w e . 9.0 Results , With the vertical load upward the U-bolt load increased as the support load increased, thus adding to the U-bolt preload. Conversely, the U-bolt preload decreased as the support load was

              - applied in the vertically downward direction. The maxicum U-bolt load occured with a bolt torque of 100 foot pounds and an upward . support load of 7208 pounds. The resultant U-bolt load was 8327 pounds. See Figure 17   This was an increase of 1858 pounds over the unloaded condition at 100 foot pounds of torque.

The decrease in U-bolt load due to a 7208 vertical downward l support load with the bolts torqued at 100 foot pounds was 2306 pounds. The resultant U-bolt load was 4163 pounds versus the

6469 pounds for the unloaded condition. See Figure 18.

The maximum crosapiece stress occurred with a bolt torque of 100 l foot pounds and an upward support load of 7208 pounds. The resultant crosspiece stress was 22,4 87 psi. See Figure 19 This was an increase of 5,384 pai over the unloaded condition at 100 foot pounds of torque. 4 I l 1 e l l 43

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l 5 e 51-S18 0.25" S6  ! 1 1 LOAD DISTRIBUTION / STRAIN MEASUREMENT TEST  ! i STRAIN GAUGE LOCATIONS i

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10 INCH STAINLESS STEEL 9.0 - 100 FOOT POUNDS 8.0 . . 7.0 - 8 O 5 6.0 -- 67 FOOT POUNCS a

n. 5.0 - -

STRAIN GAUGES S3, S4

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3.0 - - 33 FOOT POUNOS 2.0 -- 1.0 -- NOTE: RESULTS PRESENTED ARE FOR A SINGLE U-BOLT LEG ONLY. i j  :  :  ;  ;  ; . - 0 1.0 2.0 3.0 4.0 5.0 6.0 7.0 8.0 ! VERTICAL LOAD UPWARD - P00NDS/1000 1 i l ! FIGURE 17

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  ;                                       O 1.0 2.0       3.0        4.0       5.0       6.0 7.0   8.0 1

VERTICAL LOAD UPWARD - POUN05/1000 FIGURE 19

       . . ~ . , . . ~ . . . .

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( llf Il , l l l 11 ll1

l i I l i e i Y.D Thermal Cvel i na /The rmal Gradient Test l 1.0 Purpana The parpose of this test was to investigate the effect of thermal i cycling at elevated temperatures on U-bolt preload. This test , was performed on three test specimens (pipes) using three designs . of U-bolt type supports as specified and provided by TUGCO. l Pine Test Snacimens f 32 inch narbon steel pipe, insulated  : 10 inch stainless steel pipe, uninsulated . 4 inch stainless steel pipe, insulated  ! U bol t batt enn 32 inch U-bolt type pipe support  ! 10 inch U-bolt type pipe support , 4 inch U-bolt type pipe support

        .                                                                                                        i 2.0 Tant co nfi ruvation (See Photographs 9 through.11).                                          ,

The three pipe test pecimens described in Section 1.0 had a  ! mounting plate welded to each end. These plates were then [ fastened to two floor mounted brackets using high strength fasteners. The carbon steel ta'st specimen had a Steel Structures [ i Painting Council No. 2 finish applied on a four foot section in  !

i. the center of the pipe. Strain gauges were' installed on the U-bolt. Thermocouples were installed on the U-bolt, crosspiece l and pipe. The U-bolts were then installed with the nuts hand  ;

tight. The 4 inch and 32 inch test specimens were insulated. Strip heaters were installed inside the total length cf the pipe  ; to provide the internal heat scarce. I 30 Instrumentation j Instrumentation used for these tests is identified below. j

 \                  1. Strain gauge signal conditioning and readout devices                                   "

2. Theraccouple signal conditioning devices 3 Five strain gauges l

4. Eleven thermocouples i 5 Temperature data recorder See Sketches 5 and 6 for locations of strain gauges and thermocouples. The above instrumentation was calibrated prior to us..  !

4.0 Loading Devinam [ I Loading devices used for these tests are identified below. i 50 .

                                                                                      ,,..v.. ..-,--m .,,-.wy- y

8. 6 F

i i t i

1. Torque wrench

2. Strip heaters The torque wrench was calibrated prior to use.

50' r.nea teeliention The test specimens were subjected to two types of loading during [ this test. Torque loads were applied to the U-bolt nuts and the i strip heaters were used to provide the thermal cyclic loads. 6.0 Test Descrietion After each specimen had been installed and the' instrumentation calibrated, testing commenced. The testing method was the same j for the three specimens. The only variables were the U-bolt nut  ; torques and maximum test temperatures. These values were specified by MGCO. t t Nut Tercues; l l 4 inch test specimen: 60 foot pounds; in increments of 5 foot pounds. i 10 inch test specimen: 100 foot pounds; in increments of 10 foot pounds. l 32 inch test specimen: 600 foot pounds; in increments of 20 foot pounds. Mer1=nm Test Temneratures: 4 inch test specimen: 10 inch test specimen: 560250 {F

• 32 inch test specimen: 560F ,

With the test specimen instL11ed, the U-bolts were tightened to the above torque valtss. With the test specimen at ambient , tamperature, the recording equipment was activated. The strip heaters were then used to bring the test specimen to the test temperature. After maintaining the peak temperature for one ' hour, the strip heaters were deactivated to allow the test specimen to cool to the laboratory ambient temperature. This thermal cycling was repeated a total of ten times for each specimen as specified by 20C0. 70 Data Reduction

                                                                                                    ~
        .       The test data recorded included U-bolt strains and temperatures of the pipe, U-bolt and crosspiece. Using the atrain data, U-bolt preloads were determined. The analytical method for r

51

I  ! l determining U-bolt preload is contained in Appendix I. This 5 information was used to evaluate the effect of thermal cycling on the U-bolt preload. - l 8.0 Data Descrintion/ Presentation i Test output was reduced and plotted so results could be presented graphically. U-bolt preloads were plotted versus thermal cycle , for each of the specimens while at maximum pipe temperature, i This information is presented in Figures 21 through 2'4 . "A 90 Results i , o 4 inch stainless steel The U-bolt proloads at ambient temperature (105 F) with a torque of 60 foot pounds prior to cycle 1 were: l Leg 1 (Gauges 32. 35) = 5903 pounds.  : Leg 2 (Gauges S3, S10) = 6292 pounds. The U-bolt preloads at ambient temperature (107 5 F) with a  ! torque of 60 foot pounds prior to cycle 10 were:  ; 1 Leg 1 (Gauges 52, SS) = 3271 pounds. . Leg 2 (Gauges S3, 310) = 4529 pounds.  ! The U-bolt proloads on the 4 inch gtainless steel specimen, while at a pipe temperature of 560 ? and a torque value of  ; 60 foot pounds, dropped during the first few cycles but l stabilized at approximately 7000 pounds for Leg 1 and 6400 l pounds for Leg 2. See Figure 21. , Since strain gauges are sensitive to temperature, direct strain readings may include temperature induced or l

                                    " apparent" strains. The portion of the atrain readings that      ;

are " apparent" were detersined by calibrating the gauges j l over a wide temperature range. " Apparent" strains were determined for the strain gauges used on the 4 inch , stainless steel specimen by calibration testing. The

                                    " apparent" strains were combined with the direct strain readings to assure that accurate strains were determined           ,

over the entire temperature range. The 4 inch stainless steel specimen reggired 2 hours to  ! reach the nazimum pipe temperature (560 F) during each ,

                               -    Cycle.

o 10 inch stainless steel  !

The U-bolt preloads at ambient temperature (78.2 F) with a  ; torgt.e of 100 foot pounds prior to cycle 1 were* l 52

Leg 1 (Gauges S4, 311) = 4515 pounds. Leg 2 (Gauges S8, S1) = 4503 pounds. The U-bolt preloads at ambient temperature (76*F) with a torque of 100 foot pounds prior to cycle 10 were:

                                                                              ~

Leg 1 (Gauges S4, S11) = 3209 pounds. Leg 2 (Gauges 38, S1) = 3516 pounds. The U-bolt preloads on the 10 inch stainlesg steel 0 specimen, while at a pipe temperature of 250 F and a torque value of 100 foot pounds, dropped slightly 'during the first few cycles but stabilized at approximately 3500 pounds and 3900 pounds for Legs 1 and 2, respectively. See Figure 22. Apparent strains were determined for the strain gauges used on the 10 inch stainless steel specimen by calibration tasting. The apparent strains were combined with the direct strain readings to assure that accurate strains were determined over the entire temperature range. The 10 inch stainless steel specimen reguired 20 minutes to rea:h the maximum pipe temperature (250 F) during each cycle. o 32 inch carbon steel The U-belt preloads at ambie'nt temperature (75.7 F) with a torque of 600 foot pounds prior to cycle 1 were: o Leg 1 (Gauges 13,18) = 15,158 pounds. Leg 2 (Gauges 15, 17) = 15,675 pounds. Because of the large mass of the 32a pipe test specimen there was a temperature lag between the pipe and the U-bolt support during cooldown between cycles. As a result, the U-bolt support never reached ambient temperature coincident with the pipe. Therefore, while the pipe had reached ambient temperature between cycles as specified by the test procedure, U-bolt proloads at ambient temperature prior to cycle to could not be determined. Apparent strain data was not directly available for the gauges used on the 32 inch specimen. Since the same type of gauges were used on the 10 inch specimen, a statistical evaluation was performed and a confidence interval was determined for apparent strain at a 95 percent confidence t level. This confidence interval was applied to the apparent strain curves obtained for the 10 inch specimen so that apparent strain ranges were established for use with the 32

                         , inch specimen.

l 53

i

  .,   t f

i Apparent strain is a function of gauge type and material to  ; which the gauge is attached. Both the 10 and 32 inch U-bolts are of the same material. As noted previously, the i same type of gauges were used on both. U-bolt size has no j

                            -     effect on apparent strain.                                     ;

The U-bolt preload on Leg 1 (Gauges 13,18) of the 32 ineg l carbon steel specimen, while at a pipe temperature of 560 F  ! and 600 foot pounds of torque, went from 30,144 pounds i during cycle 1 to 29,799 pounds during cycle 10. The , difference between the highest and lowest recorded values i over the course of 10 cycles was 2756 pounds. See Figure 23 , The U-bolt preload on Leg 2 (Gauges 15,17) of the 32 ineg  ! carbon steel specimen, while at a pipe temperature of 560 F i and 600 foot pounds of torque, went from 25.579 pounds during cycle 1 to 24,373 pounds during cycle 10. The  : difference between the highest ana lowest recorded values , over the course of 10 cycles was 3359 pounds. See Figure

24. .

The 32 inch carbon steel specimen requ15'd " /2 hours to , reach the maximum pipe temperature (560 F) during each cycle. . I e T

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                                                                                                                                    ..                                                51-518 0.40" FOR 4" AND (j
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t J 10". SPECIMENS m/ 0.80" FOR 32" m3 / ~ % SPECIMEN armre nar THERMAL CYCLING / THERMAL GRADIENT TEST AND CREEP TEST STRAIN GAUGE LOCATIONS SKETCH 5 58 _ _ - _ _ _ _ _ _ _ _ _ - _ _ _ _ _ - _ _ _ _ _ _ _ _ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ ~ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - _ _ _ _ _ _

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THERMAL CYCLING /fHERMAL GRADIENT TEST AND CREEP TEST THERMOCOUPLE LOCATIONS e SKETCH 6 59

7-I ^I $ 8.0 PIPE TEMP = 560*F , U-BOLT TEMP = 450*F* - i l LEG 1 GAUGES S2, SS

                                       ~

7.0 8 S s E N g x o LEG 2 GAUGES S10, S3 6.0 E 8 $ E ,

• Representative temperature. Actually, temperature varies at different loca-3 g tions in the U-bolt. Refer to Attachment S 3, page 16, for actual temperature 5*0 -

distinction. a 4.0  :  : : . .  : .  : . I 2 3 4 5 6 7 8 9 10 CYCLES FIGURE 21

i i ! 10 INCH STAINLESS STEEL

                                                                                                                                  ~

PIPE TEMP

• 250*F i U-BOLT TEMP = 150*F*

1 l

'                                       5.0 "

LEG 2 ~ i GAUGES 58, Si ! o 8 4.0 . - t . E z , a -

    -                              O

' " LEG 1 3.0 e GAUGES S4, Sll i E l m 3 w i E 2.0 . -

• Representative temperature. Temperature j g varies with location in U-bolt.

m ' l $ l 1.0 . l ,

i

:  :  :  :  :  ? O

' 8 9 10 0 'l 2 3 4 5 6 7 1 1 j CYCLES FIGURE 22 J l

 ~

40 32 INCH CARBON STEEL PIPE TEW = 560*F U-BOLT TEW = 310*F

• MAXI M

o. 30 ,.

8 R , E -

-   g                                                                                .

2 . E MINI M g E b ,

  . 2 S

• Representative temperature.

10 LEG 1 - Temperature varies at different GAUGES 13, 18 locations in the U-bolt. Refer to Attachment 3, page 16, for

                      ,                                                     actual temperature distribution.

1 2 3 4 5 6 7 8 9 10 CYCLES FIGURE 23

e 9 40 - - 32 INCH CARBON STEEL PIPE TEMP = 560*F U-BOLT TEMP = 310*F * ,

                             'o      30 -

8 MAXIMilt t - i 8 . E

2 -

e a 20 - i S .I O E "s [ . g NINIMUM - 1 A i 10 . - LEG 2 -

• Representative temperature.

Temperature varies at difTerent GAUGES 15' 17 locations in' the U-bolt, Refer to Attachment 3, page 16 for j actual temperature distribution. ll

                                                   ;    ;         :                :           :     :           :  . l t                                           1      2    3         4                5           6     7           8  9 10 l

l CYCLES i l FIGURE 24 1 -

l-j' . . Y.g. Cream Test 1.0 Bir.ggat The purpose of this test was to evaluate the effect of high temperature on the U-bolt preload as a function of time. This test was performed on three test specimens (pipes) using three designs of U-bolt type supports as specified and provided by TUGCO: Pine Test Soecimens 32 inch carbon steel pipe, insulated 10 inch stainless steel pipe, uninsulated 4 inch stainless steel pipe, insulated U bol t Dan 4 rne 32 inch U-bolt type pipe support 10 inch U-bolt type pipe support 4 inch U-bolt type pipe support Tant confirur ation (See Photographs 9 through 11) 2.0 The Creep test was performed immediately after the Thermal Cycling test with the same t'est configurations as that used during the Thermal , Cycling test. 3 0 Instrumentation Instrumentation required for this test is identified below.

1. Strain gauge signal conditioning and resdout devices

2. Thermocouple signal conditioning devices 3 Five strain gauges

4. Eleven thermocouples

5. Temperature data recorder See Sketches 5 and 6 for strain gauge and thermocouple locations. The above instrumentation was calibrated prior to use.

4.0 Leading Devines Loading devices used for these tests are identified below.

1. Strip heaters 5.0 Load Annlication The test specimens were subjected to two t pos of loading during this test. Torque loads applied to the U-bolt nuts prior to the 64

l l Thermal Cycling / Thermal Gradient Test were maintained. Strip heaters were used to provide the thermal load. 6.0 Tant namerietion , _ After each specimen had been installed and the instrumentation  ; calibrated, testing commenced. The testing method was the same for all three specimens. The U-bolt nut torques were maintained frog the Thermal Cycling test. The maximum test tgaperature was 560 F.for both the 4 and 32 inch specimens and 250 F for the 10 ., inch specimen.  ; i With the test specimen at the maximum test temperature, U-bolt strains were recorded at periodic time intervals over a 24 hour period. Af ter 24 hours, the strip beaters were deactivated and the test was concluded. 70 Data naduetton  ; I The test data recorded included U-bolt strains and temperatures  ; of the pipe, U-bolt and crosapiece. Using the strain data, U-bolt proloads were determined. The analytical method for ' determining U-bolt preload is presented in Appendix I. This

      . information was used to evaluate the effect of elevated                          .

temperature on the U-bolt preload as a function of time. 8.0 Data Damerietion/ Presentation Test outptt was reduced and plotted so results could be described graphically. U-bolt proloads were plotted versus elapsed time at , maximum temperature for each of the test specimens. This > information is presented in Figures 25 through 28. 4 90 Ranulta The variation in preload over the 24 hour time period at maximum temperature was minimal for all three test specimens. c 4 inch stainless steel - The maximum preload variation war 85 pounds, or 1.25 of the tuitial value of 6728 pounds at maximum pipe temperature. See Figure 25. ,i The U-bolt proloads determined at ambient temperature (77 F) and 60 foot pounds of torque prior to creep test host up were: 1 Leg 1 (Gauges S2, SS) = 4871 pounds. , Leg 2 (Gauges S10, $3) = 4410 pounds. j The Ugbolt proloads determined at ambient temperature (91.4 F) and 60 foot pounds of torque af ter creep test  ; completion and cooldown were: r 65 l

s

  . e                                                                                                      :

f f [ Leg 1 (Gauges S2, SS) = 4854 pounds. Leg 2 (Gauges $10, S3) = 4398 pounds. , i The 4 inch stainless steel specimen exhibited no loss of ambient condition U-bolt preload as a result of the creep

                    ~                                                                                       '
'                                  test.

t o 10 inch stainless steel - The maximum preload variation was 103 pounds, or 2 75 of the initial value-of 3862 pounds at maximum pipe temperature. See Figure 26'.- , The Ugb olt proloads determined at ambient.-temperature [ (75.8 F) and 100 foot pounds of torque prior to creep test heat up were: Leg 1 (Gauges 38, 31) = 3625 pounds.  : Leg 2 (Gauges 54, 311) = 3587 pounds.  ! I The U-bolt proloads determined at ambient temperature , (66.9'F) and 100 foot pounds of torque after creep test l completion and cooldown were: r Leg 1 (Gauges S8, 31) = 3593 pounds.  ! Les 2 (Gauges S4, 311) e 5316 pounds. [ The 10 inch stainless steel. specimen exhibited a small loss  ; I l of ambient condition U-oolt' preload as a result of the creep i j test.  ! As described in Section 9 0 of the Thermal cycling test  ; description, apparent strains were determined for the gauges used ' on the 4 and 10 inch stainless steel specimens by calibration ' testing. The apparent strains were combined with the direct strain readings to assure that accurate strains were determined at the marinum temperatures, o 32 inch carbon steel - The maximum preload variance was 517 , pounds or 3 8% of the initial value of 13,436 pounds at } anximum pipe temperature. This is based on the lower limit i of the 95 percent confidence level for Leg 1. See Figure  ! t 27 bolt proloads determined at ambient temperature t The (86 7 ,F) and 600 foot pounds of torque after creep test U l 1 completion and cooldown were: Leg 1 (13,18) = 8526 pounds. l

                                        . Leg 2 (15, 17 ) = 9904 pounds.                                   [

l r l 66

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                                                                      ~

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s- . t ',s ._ Apparent, strain ' data was, not directly avf.ilable for the b%' , / \, asuJes md , an the 32' isch, specimen. Since the same type of

         \                                                                               -    4                . gauges wore used on the 10 inch specimen, a statistical evaDtation was performed and a confidence interval was
                                     ,                                                     'U,                determined for apEtrent strtin at a 95 percent confidence N,,                                                                          \' level. his cocridence interval was applied to the apparent S strain cudos obtained for the 10 inch specimen so that

( D' s appsrent strain.rtngen vera'astr.blished for use with the 32 ' ~ '

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                                                                                                                                                       '                                                           *         ' ' '..               I PIPE TE M = 250*F                                                                                                     'O'~~                                    "'

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The purpose of this test was to evaluate the effects of piping  !

                 - vibration on the U-bolt preload. The U-bolt was preloaded to the                                     l value of torque that provided a conforming fit between the U-bolt                                    ;

and pipe. This test was performed on the 10 inch stainless steel , pipe specimen as specified and provided by TUGCO. l l 2.0 Test Confinuration (See Photograph 12) [ The pipe test specimen described in Section 1.0 had a sounting i plate welded to each end. These plates were then fastened to two 3 floor mounted brackets using high strength fasteners. Strain gauges were installed on the U-bolt. The U-bolt support was then i instdled in the center of the test pipe with the crosapiece

• perpendicular to the floor. The U-bolt nuts were hand tightened. An electrodynamic shaker was fixtured so that a ,

forcing function could be induced to the U-bolt support clevis ' via a strut. The strut was arranged such that the pipe test specimen experienced triazial loading. , 3.0 Tnatru=entation 1 ) Instrumentation required for this testing is identified below. 2 l

1. Strain gauge signal conditioning and readout devices

2. Five strain gauges The strain gauges were installed on each test specimen prior to ,

testing. See Sketch 7 for location of strain gauges. The above instrumentation was calibrated prior to use. l 4.0 Loading Devices Loading devices used for this test are identified below. i

1. Torque wrench .

2. Electrodynamic shaker l The torque wrench was calibrated prior to use.

5.0 t aad Annliention The loads experienced by tha test specimen were due to U-bolt nut torque and the electrodynamic shaker. Testing was performed at laboratory ambient environmental conditions. 6.0 Test Descrintion L L 5 72

                                                                                             ~

. e- [ After the specimen had been~ installed and the insteunentation f onlibrated, testing commenced. With the U-bolt located at the  ! senter cf th) pipe with the crosspiece perpendicular to the

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floor, the nuts were torqued to 50 foot pounds. With the test  ! specimen in the unloaded condition, resi41ng equiposat was l activated. The electrodynasio shaker.was then activated to i induos a sinusoidal load of 1500 pounds on the specimen as l

             .specified ty 1U000.        The electrodynamic shaker was calibrated to            '

assure tbo load magnitude was 1500 pounds. The frequency varied  ! from 5 to 200 to 5 herts at a rate of 2 octaves per.ainuts. The test duration was 270 minutes. This is equivalent to 90. minutes , of vibretion in each of the three ares. .I888 382-1980 was used I as a basis for establiabing test pecanetern. IEE8 382-1980 is an  ! industry accepted standard for simulating piping vibration in a  ! plant. ' 70 nata neduction [ The test data recorded included U-bolt steains. Using the strain f data, 0-bolt proloads were determined. The analytical method for i determining U-bolt proload is presented in Appendix I. This information was used to evaluate the.effect of vibration en the '

            . D-bolt preload,                                                                   i l

8.0 .Duta DameriDtion/Prasantation l

             . Test output was reduced and plotted so reaults could be presented graphically. U-bolt preload was plotted in Figure 29 as a                         :

function of test' time et vibration. , 90 namul t_a l The initial preloed resulting from the U-bolt nut torque.cf 50

          . foot pounds was 4285 pounds in both Les .1 (Gaugas 31, S2) and Les 2 (Gauges 33, S4). During the first few seconde of vibration,                      ,

the U-bolt support moved in the axial pipa. direction aliantly  ! (41/14 inch) but repositioned itself to a location where it i remained during the rest of the test. This repositioning r resulted in a doorease of U-bolt preload of approximately 600 pounds. l The. maximum. variation in preload, af ter repositioning occurred, , was 1083 pounds in Leg 1.and 621 pounds in Leg 2. The obanges in 7 proloads beoed on data recorded 5 25.ainutes into the . test and  : 270 minutes into the test were 660. pounds in Leg 1 and 564 pounds  ; in Leg 2. See Figure 29. l l l F p 73 I

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i r s APPENoix 3

ENGINEERING CALCULATIONS l

e 0 l I 77

• 5

, o 1 l Following are the engineering calculations used to reduce test data - and graphically present results. A. U-bolt preload (or load) Preload (or load) = microstrain x Ex bolt shank area l Where: E = Young's modulus = 29 x 106 lb/in2 and: bolt shank areas are 4 inch specimen = .196 in 22 l 1r. inch specimen = .442 in,  ; 32 inch specimen = 5.94 in" Strains were recorded as test output. Stains recorded were in units of 10-6 in/in. Stains of 10-6 are defined as microstrain. . B. Pipe and crosspiece stressas , Stresses were obtained from the elastic stress-strain relationship: i e = cE' Where: c = microstrain - E = Young's modulus . 2 .

                  = 28.3 x 10 lb/in for stanless steel
                                                 ~

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                  = 27.9 x 10 lb/in for carbon steel l

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I. C. Friction coefficient The coefficient of friction between the U-bolt support assembly  : , and the pipe is determined as follows-r f t 6 I

                                                                                               "                             X f                           v, l

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J 1 e S 8 k 79

l I Summing moments about the pipe centerline:  ; PL = 1 (V1 + V2) h=V1+V2 Suming forces in the Y direction: P = V2 -V1  ; By combining the last two expressions: L  :' P (1 + I) = V2 Upon determination of V1 and V2, .the average coefficient of  ; friction between the U-bolt support and pipe is expressed as: , u=f(h+h) Where: N = T1 + T2 P = Slip force u = Coefficient of friction

                                                                                                                                +

D. 95% confidence level determination for apparent strains of 32"

                                                                      ~

specimen S a l Max, Min = Y +f t Y n Where: Y = average = b Xi/n ' i =1 " - 1/2 S = standard deviation =. b (Xi - Y)2 /(n-1) l

                                                                   .i = 1                       ,

r n = number of samples = 4  ; t{=3.182for95%confidencelevel 4

                                                 \

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E. Friction test strain criteria Maximum strains measured on the bolt shank were determined such that resulting stresses in the threads are 90% of material yield.  ;

                 - For A36 material, Sy = 36,000 psi
                                                 .95y = 32,400 psi First determine strains that would cause shank stresses of 32,400 psi.

o 32,400

                   *

• E " 27.9 x 106 -

c = 1161 u inches / inch  ; i Where: c = microstrain E = Y6ung's modulus . Since there is a difference in area between bolt shank and thread, i the microstrain that causes a stress of 32,400 psi can be determined u:;ing a ratio of the shank and thread' areas. Shank areas T'hread areas  ! 2 2  ! 1/2 inch: .196 in 1/2 inch: .1419ig 3/4 inch: .442 in 2 3/4 inch: .334 in 2  ! 2 3/4 inch: r 94 in ' 2 3/4 inch: 4.93 in

    ~

The shank microstrain that' would cause a thread stress of 32,400 psi is determined by: l Threaa area  : microstrain = 1161 x Shs,k area l l ,

                                                                                                                              \

The required microstrains are: . 1/2 inch: 839 y inches / inch 3/4 inch: 877 y inches / inch 2 3/4 inch: 964 u inches / inch 81

t 8 e O O APPENDIX II U-BOLT IDENTIFICATION e w 82

i e Following is a summary of which U-bolts were installed on the test specimens for each specific test. Torque versus Preload 32 inch carbon steel pipe: 32 inch U-bolt - since only one 32 inch U-bolt was supplied. itewas used on all testing conducted on-the 32" carbon steel specimen. 10 inch stainless steel pipe: 10 inch U-bolt #1. 10 inch carbon steel pipe: 10 inch' U-bolt #1. 4 inch stainless steel pipe: 4 inch U-bolt #1. Friction 32 inch carbon steel pipe: 32 inch U-bolt. 10 inch stainless steel pipe: 10 inch U-bolt #2 - Polishing test used 10 inch U-bolt #1. 10 inch carbon steel pipe: 10 inch U-bolt #4 - Polishing test used 10 inch U-bolt #3. 4 inch stainless steel pipe: 4 inch U-bolt #2 - Polishing test used 4 inch U-bolt #1. Load Distribution / Strain Measurement 10 inch stainless steel pipe: 10 inch U-bolt #2. 83

     .-           e Thermal Cycling / Thermal Gradient 32 inch carbon steel pipe:       32 inch U-bolt.

10 inch stainless steel pipe: 10 inch' U-bolt #4. 4 inch s'tainless steel pipe: 4 inch U-bolt #1. Creep 32 inch carbon steel pipe: 32 inch U-bolt. 10 inch stainless steel pipe: - 10 inch U-bolt #4. 4 inch stainless steel pipe: 4 inch .U-bolt #1. Vibration 10 inch stainless steel pipe: 10 inch U-bolt #3. The U-bolts were identified by a unioue number for their use on a specific pipe size prior to.the initiation of the test program. O 84

a ADDENDUM 1 TO

                                                                           ~

REPORT NO. EQ&T-EQT-860 REVISION 0 t CCHANCHE PEAK STEAM ELECTRIC STATION U-BOLT SUPPORT / PIPE TEST REPORT WESTINGHOUSE ELECTRIC CORPORATION P.O. BOX 355 PITTSBURGH, PA 15230 7 % & 4 c./v,4 PREPARED BY: MARK KAMENIC 6-$-8[ REVIEWED BI: J.M. SNIDER F APPROVED BY: L.I. WAL R 85

d Seia=4e Loadine Simulation Test 1.0 Purpose The purpose of this test was to evaluate the effects of seismically induced piping vibration on U-bolt preload. This test was performed on the 10 inch stainless steel pipe specimen as specified and provided by TUGCO. _ 2.0 Test contiauration (See Photograph 1) The pipe test specimen described in Section 1.0 had a mounting plate welded to each end. These plates were then fastened to two floor mounted brackets using high strength fasteners. Strain gauges were installed on the U-bolt. The U-bolt support was then installed in the center of the test pipe with the crosspiece perpendicular to the floor., The U-bolt nuts were hand tightened. A hydraulic shaker was fixtured so that a forcing function could be induced into the U-bolt support clevis by a strut. The strut was arranged such that the pipe test specimen experienced triaxial loading. The strut was instrumented with strain gauges to form a load cell. 30 Tn=tm =antation Instrumentation required for this testing is identified below.

1. Strain gauge signal conditioning and readout devices

2. - Magnetic tape recorder 3 Six straic gauges.

The strain gauges were installed on t_he test specimen prior to testing. See Sketch 1 for location of strain gauges. The above instrumentation was calibrated prior to use. 4.0 Leadine Devices Loading devices used for this test are identified below.

1. Torque wrench

2. Hydraulic shaker The torque wrench was calibrated prior to use.

5.0 Lead Aneliention The test specimen experienced U-bolt preload and seismically induced vibration resulting fron the U-bolt nut torque and hydraulic shaker, respectively. Testing was performed at laboratory ambient environmental conditions. 86 6..-. .

J 6.0 Test Descrintion After de specimen had been installed and the instrumentation calibrated, testing commenced. With the U-bolt located at the center of the pipe with the crosspiece perpendicular to the floor, the nuts were torqued to 50 foot pounds. This value of torque provided a conforming fit between the U-bolt *and the pipe. Seismic loading simulations were then performed. The first simulation was performed using a sinusoidal input at 9 hertz. The input force as a function of test time is shown in Figure 1. The second simulation was performed using a maximum sinusoidal input of 1300 pounds at the test specimen natural frequency. Using the-strain gauges installed on the strut, the hydraulic shaker was calibrated to assure the proper load magnitudes were achieved. - The first seismic loading simulation was performed at 9 hertz with full amplitude input lasting 39 seconds. A magnetic recorder was used to record the U-bolt strains as well as the input load from the strut load cell gauges. The second seismic loading simulation was conducted at the natural frequency of the pipe test specimen in the fixture. The natural frequency was determined by exciting the pipe with a 1000 pound sinusoidal input and sweeping upward from 5 hertz until the natural frequency was found. An accelerometer mounted on the U-bolt support crosspiece was used to identify the natural frequency. Once the natural frequency had been determined, an input level of 7000 pounds was to be introduced at the system natural frequency. However, at the system natural , frequency the hydraulic actuator was not capable of transmitting the 7000 pound force due to movement in the strut connections (resulting from design and manufacturing tolerances) which exc:sded the displacement output of the actuator. As a result, the marimum input level that was achievable from the actuator was used for

 '                this test. This level,1300 pounds, was applied at the natural frequency. The minimum duration of this test was defined as 300 cycles divided by the natural frequency in cycles per second. A magnetic tape recorder was used to record the U-bolt strains as well as the input load from the strut load cell.

70 Data Reduction U-bolt strains and load input were recorded. Using the strain data, U-bolt proloads were determined. The analytical method for determining U-bolt preload is presented in Appendix I of EQ&T-BQT-860, Revision 0. This information was used to evaluate the effect of seismic level piping vibration on the U-bolt preload. 87

  .___                                                                   o            __

8.0 Dain Descriotion/ Presentation Test output was reduced and plotted so results could be represented graphically. The input force was plotted in Figure 1 as a function of test time for the 9 hertz seismic loading simulation test. This figure shows the input forces in the push (U-bolt compressive) and pull (U-bolt tensile) directions in addition to the average force magnituda. U-bolt preload is shown in Figure 2 as a function of test time for the 9 hertz seismic loading simulation test. 9.0 Results The initial preload resulting from the U-bolt nut torque.of 50 foot pounds prior to initiation of any vibratory input was 4484 l pounds in both Leg 1 (Gauges S1, S2).and Leg 2 (Gauges 33, S4). , Four seconds elapsed from the time vibration was initiated until  : the full amplitude input was reached. .During that time the U-bolt l preloads decreased to 4375 pounds and' 4426 pounds -in Legs 1 and 2, i respectively. These values are indicated at the zero timo point f on the abscissa of Figure 2 and represent the U-bolt preload at

   '          the time the vibratory input reached full amplitude.

After completion of the seismic loading simulation test (with no external cyclic load applied), strains remaining in the U-bolt [ corresponded to a preload of 3914 pounds in each of the two legs. . Thus'a relaxation of 570 pounds from the original 4484 pound j preload occurred in both U-bolt legs. The U-bolt support was  : observed to move approximately one-half inen in the axial (toward the strut) direction. No rotational motion was perceptible. l Prior to running this test, three trial tests were run to adjust  ! teet equipment and instrumentation to obtain the desired input.  ; Although none of the three trial tests can be considered official,  ;

             ,their observed behavior was the same as that of the official test.                                                                     j The natural frequency of the test specimen as mounted in the                                                                           {

fixture was found to be 75 hertz. As stated in Section 6.0, a l etismic test was to be performed at the test specimen natural j i frequency using an input load magnitude of 7000 pounds. However, at this higher frequency, the hydraulic actuator was not capable of transmitting the required 7000 pounds of force because the l movement in the strut connections resulting frca design and l manufacturing tolerances exceeded the displacement output of the actuator at 75 hertz. . ii The maximum actuator force output at 75 hertz was 1300 pounds.  ! The duration of the 1300 pound, 75 hertz test was in excess of  ; four seconds. During the 75 hertz test, the U-bolt support was i observed to exhibit barely perceptible axial action (toward the l strut) and no rotational motion. i Because the desired 7000 pound load was not achievable during this i test, the test was considered unofficial. For this reason no j other test data is presented. 88 j

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oo 12.0.

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;                                                                                                                                                                                               WAS REACHED. TOTAL DURATION OF THE TEST AFTER i                                                                                                 10.0-                                                                                          REACHING FULL AMPLITUDE INPUT WAS 39 SECONDS.

12.0- . l < FIGURE 1

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I i ATTACHMENT 3 t COMANCHE PEAK STEAM ELECTRIC STATION U-BOLT FINITE ELEMENT ANALYSIS WESTINGH0CCE' ELECTRIC CORPORATION P. O. BOX 355 , PITTSBURGH, PA 15230 l l June 12. 1984 I i i r l \ l AUTHORS l l M. J. FISHE1 W. S. LAPAY i I l l l -_ , , . , , . . . - . , - - . - . . - - , - . . - - - . . - - , . - - - . -

l

l l

r 1 n . , l C(MANCHE PEAK U-BOLT ANALYSIS REPORT

TABLE OF CONT 2NTS l

l l l I. Introduction ( II. Summary and Conclusions  ! l l

• l III. Input te Analysis l

     .                                                                                                                     l IV.         Analytical Model Description and Analysis Technique Y.          Correlation of Analysis Results to Test Results VI.         Behavior of Pipe /U-Bolt Assembly for " Slip" t

l VII. Acceptance Criteria for Pipe Stress VIII. Stress Summary Tables APPENDII I - U-bolt Hanger Drawings , l , ! APPENDII II - Nastran Element Description l l l i i I ! i (

                     -,. _     . _ .         - , _ _  . - - .   -,-.-------,--,_,m . _ . - , ---. . . _ . . . _ _. - - - .

      's    .

COMANCHE PEAK U-BOLT ANALYSIS REPORT I. INTRODUCTION , During the Atomic Safety and Licensing Board hearing for Comanche Peak Unit No. 1, concerns were introduced regarding the pipe support assembly shown below. , SRW

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! Specifically, the concerns are that sufficient testing and analysis has not been conducted to ~ ensure that:

1. Adequate frictional forces exist at the pipa/ pipe support interface to balance a moment created when the U-bolt legs are not parallel to the strut so that the >

U-bolt strut assembly support is stable.

2. Pipe and pipe support stresses are within acceptable limits.

Texas Utilities Generating Company (TUGCO) requested that Westinghouse perform a finite element analysis and testing to l investigate these concerns. - l

o .  ! < , The tests performed are discussed and results presented in Westinghouse Report Number EQ&T-EOT-860, Revision O, " Comanche Peak Steam Electric Station U-bolt Support / Pipe Test Report." The finite element analyses are discussed herein. i The intent of the analysis program was (1) to determine if

       ' the pipe would slip, thereby creating an unstable support condition when the hanger support was subjected to the preload, therma l, pressure and mechanical loads that would be e xpected in the Comanche Peak hanger assemblies and assess its significance, and (2) to calculate pipe and pipe support stresses that could be l

l e xpected to be ezperienced by the Comanche Peak U-bolt support as semblies . To obtain the analytical information necessary to reach the conclucions related to the scope of the analysis pro gram, finite element models of the U-bolt, cross piece and strut attachment, and pipe were de'veloped. These models are discussed in Section IV. They were verified by comparison to analytical data as well as test data. This is discussed in t Sections IV and V. The results from the analysis program related l to slip are discussed in Section VI and the pipe stress conditions in Section VII. i i i l

 -   . .   . . . .        . . . .              . . . _ . ~.

f l II.

SUMMARY

AND CONCLUSIONS The finite element analysis performed on the U-bolt pipe support assembly was performed to evaluate: (1) the stability of the U-bolt support system with respect to sliding, and (2) to evaluate the stress intensities in the piping caused by the cinching of the U-bolt. In order to perform the above e va luations, finite element analysis models were developed to

                   ~

r .nalyze the U-bolt piping assembly for four loading conditions:

1. Preload

2. Preload and thermal

3. Preload and thermal and pressure

4. Preload and thermal and pressure and strut applied load (Push)

From these loading conditions, it is possible to determine the forces in the legs of the U-bolt, the stresses in the pipe, and the minimum torque required to maintain stability against l l sliding when the U-bolt assembly is subjected to all of the i loadings with the push force having a 5 permissible offset from ! r ! the vertical a xis. Four U-bolt support piping assemblies were studied consistent with the configurations tested. They are (1) 4" Sch 160, (2) 10" Sch 40s, (3) 10" Sch 80, and (4) 32" Main Steam 1 piping systems. The finite element models developed were shown i to provide results that are in agreement with the test results. The analytical results showed similar magnitudes and stress trends in the vicinity of the location of strain gauges in the

     . t                                     4     4 test.      There were differences between the tests and analytical results; however, this is due to the difference in the fitup between the pipe and U-bolt in the test and in the dinite element model.      The difference in fitup is attributed to the pipe out-of-roundness and tolerance in the U-bolt bend.                      The finite element analysis model was not intended to model details such as
       ' gap nonlinearities but rather to evaluate as-designed conditions.

The U-bolt leg force summaries are given in Tables II-1, through II-4, As seen from these tables, as one would intuitively expect, 1) the U-bolt preload torque is indicative of the U-bolt leg tension, 2) the thermal expansion of the pipe increases the U-bolt leg tension, 3) the pipe expansion due to internal pressure increases U-bolt leg tension, and 4) the applied hanger " push" load decreases the U-bolt leg tension. Recommended torque values are'given below based on consideration of the maximum and minimum preload values evaluated in the finite element analysis. The stress in the U-bolt, based on the threaded area, will be under yield at these preload . values. Recommended torque value 4" Sch 160 25tt-lb 10" Sch 40S 50ft-lb 10" Sch 80 50ft-lb 32" Ms 250ft-lb

   .              i s             .

Further, it is necessary to torque to these values to assure proper U-bolt fitup and therefore stability of the U-bolt assembly. Hanger stability at these preload torques was demonstrated by the analyses, reported herein. As a guide to assuring that the U-bolts will function properly, they should be installed with torque values equal to the recommended values so that a " snug" uniform fit is maintained between the pipe and U-bolt. a The upper bound torque value is in the ultimate, (1) the maximum torque to which the piping can still be shown within acceptable stress limits, and (2) a value tnat will not yield the crosspiece, U-bolt, or pipe locally to such an e xtent that preload is lost and the hanger loses stability. This upper torque range value was not determined herein. The stress state in the pipe. with ma ximum preloads, therma l, I pressure and push, as defined in Tables II-1, to II-4, was found to be acceptable. It is concluded that the U-bolts can be cinched to prevent slidin g, thereby assuring the transfer of loads (force, moments) to the pipe and maintaining the stability of the U-bolt support system. The pipe will not have stresses that exceed acceptable limits if preload torquing iz restricted to the recommended va lues.  ! i I I f L e e , (

                          ,'                          U-BOLT LEG FORCE 

SUMMARY

YA Bt.c 2C -1 4" SCH 160 Pash 5* t U-Bal+ Lea . ELEMENT ~ y [ U-Bol& p EL M T Le3 f - Tension for Tension for Loading Condition Element 81 Element 100 Preload (60 ft-lbs.) 5.41 kips 5.41 kips Preload + Thermal 8.31 kips 8.31 kips Preloac+ Thermal + Pressure 8.58 kips 8.58 kips Preload + Thermal Pressure + Push at 5* 8.21 kips 8.33 kips (2,000 f) Minimum Preload (9 ft-lb.) .80 kips .80 kips Minimum Preload+ Thermal + Pressure + Push at 5* 2.80 kips 2.89 kips

l a t U-BOLT LEG FORCE

SUMMARY

TA S tX IC -2 10" SCH 405 lush

                                                                           ~

u-Bol+Q t U-Bol+ Lq ELEMENT ELEMENT 1 7 42 Tension for Tension for Loading Condition Element 1 Element 42 Preload (100 ft-lbs.) 5.62 kips 5.64 kips Preload + Thermal 6.91 kips 6.93 kips Preload+ Thermal + Pressure 7.22 kips 7.23 kips Preload + Thermal , Pressurc + Push at 5* 4.43 kips 4.78 kips (10,000 #) Minimum Preload (46 f t-lb.) 2.56 kips 2.57 kips Minimum Preload+Ther. mal + Pressure + Push at 5' 1.44 kips 1.78 kips ip i i i . . _ . . . .

i

• U-BOLT LEG FORCE

SUMMARY

7 A 6 t.tr .2I. - 3 10" SCH 80

                                                                                           ,Pask 5

U-BaltLt3 0-BoMLE) ELEMENT ELEMENT

                                                                                    '~

861 - 880 1 - Tension for Tension for Loading Condition Element 861 Element 880 Preload.(100ft-lb.) 7.51 kips 7.54 kips Preload + Thermal 8.91 kips 8.91 kips Preload+ Thermal + Pressure 9.19 kips 9.19 kips Preload + Thermal Pressure + Push at 5* 7.68 kips 8.11 kips (10,000 f) Minimum Preload (11 ft-lb.) .83 kips .83 kips Minimum Preload+ Thermal + Pressure + Push at 5* .09 kips .57 kips

     .                                                                                               ~

,' ,' U-BOLT LEG FORCE

SUMMARY

TAnter IL-4 32" Main Steam Rsh s' U-BoM Leg ELEMENT

                                                                                  '            u-8.I+ Zrg ELEMENT 133                                                                             134 e

Tension for Tension for Loading Condition Element 133 Element 134 Preload (240 f t-lb.)* 6.04 kips 6.05 kips Preload + Thermal 27.44 kips 27.45 kips Preload+ Thermal + Pressure 37.13 kips 37.15 kips Preload + Thermal + 12.16 kips 12.11 kips Pressure + Push at 0* Preload + Thermal + Pressure + Push at 5' 8.13 kips 16.23 kips (100,000 #)

• Because U-bolt leg force and pipe stresses are small for preload only, no attempt was made through analysis to exactly duplicate the 240f t-lb test preload. The analysis results are closer to a 380ft-lb preload. See stress results, test gage K, p. 39. However, the conclusion reached for the 240f t-lb (as discussed later) remains valid since the U-bolt leg forces for the various load' combinations would only be about 2.2 kips lower at the 240ft-lb preload than the 380ft-lb preload. Thus, there would still be sufficient margin to maintain preload in each U-bolt leg under any load combination.

III. INPUT TO ANALYSIS 5 Westinghouse was requested by Texas Utilities to perform a finite element analysis of each of the four hanger assemblies shipped to the WAESD Engineering Laboratory (located ati the Westinghouse Advance d Energy Systems Division, Large, Pennsylvania) for testing. 'As input for the development and verification of the finite 'elemant models Westin@ouse used the following information: -

1. Ranger Drawings MS-1-001-005-572R, Rev. 4 (Appendix I)

RH-1-024-007-822R, Rev. 3 (Appendix I) RC-1-018-016-C81R, Rev. 7 (Appendix I) These drawings were supplied by Texas Utilities. Dimensions from these drawings were used in developing the models. They were' verified and supplemented as necessary by measurements of the principal dimensions on the test specimens used in the testing program. This was done so that the model used in the analysis corresponded to the actual support assembly used in the test. No

                                                                              ~

verificatien was performed for secondary effects on dimensions (out of roundness, bowing, etc. ) since the finite element model represents the ideal situation. See Figure III-1 through III-4 for dimensions used in the analyses.

2. Additional information about the piping is tabulated below.

Pipe size, schedule, material, normal operating temperature and pressure were providad'by TUGCO. Pipe Sise Schedule Material Teep.'F Fressure g 4' 160 A312 Grade 304 559 2485 10' 405 A312 Grade 304 210 600 10' 80 SA 106 Grade B 210 600 32' (0.D. ) T = 1.45' 3A 155-EcF70 55T 1285

WINCH PIPE /U-BOLT ASSEMBLY h

                                                     ~

5 N[,%$$ - N\ 33I 4 .$' & m D FIGURE III-l

                              /OINCH STNNLESS PIPE /U-BOLT ASSE MBLY c        de Nf%ds!      .       ~

I.2

                                               ~

l bN & * & c FIGURE III-2

                 /OINCH CARBON PIPE AJ-BOLT ASSE MBLY
                     &~,

( -

                                       */

s 3

j. >-

se t/

                                            -. .n ~

9 FIGURE III-3

j i f E .' ~14- t L t 32" P AIN STEAM -l PIPE /U-BOLT ASSEMBLY j

                                                                                                                      ~

i 4 l l l C .. s44p. {b}hS l

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4  ! v\ s Og$ g R %s N V'

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, l i FIGURE III-4 f

E .' Pipe support materials were obtained from the pipe support drawings in item 1 above.

3. Information from the preload test, load distribution test and thermal cycling test given in the " Comanche Peak Steam Electric Station U-bolt Support / Pipe Test Report," Report Number EQ&T-EQT-860, Revision 0, were compared to the analytical results t to verify the analytical models.  ;

4. Steady state pipe support temperature distributions were obtained from tests conducted at WAESD engineering laboratory. _

The thermocouple temperature distributions for the 4" and 32" pipe supports are listed below. The thermocouple temperature distributions were used in the analysis to obtain pipe and pipe support stresses resulting from the restriction. of the ! thermal expansion of the pipe by'the U-bolt assembly. l

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og ,, g fe ... J _............ 1 . . . ( 4, / l (1 ,/ s - 54tLT

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                                                      ~

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                                   )                                            )

T N /. 8 BEETMB YHK EITRT!W MW

                           , . - . _ - - - - .     --        -      - - . , - . - - - - , - . .        - . - - , , - - - - . , - ~ , , , . . . - , . . . . . - ,               - - , - - - . . _ . - , - - - - . - - - , , . , - . . - . . . . - - . . . - , - - - - - - - - - - , - - - ,

     . - _ . .              =
                                               ,          The above diagrams are for the 32" pipe /U-bolt assembly thermocouple locations; the 4" pipe /U-bcit assembly has thermocouples at the equivalent 32" assembly locations.                                 l THEIM0 COUPLE LOCATION     4" PIPE SUPPORT                 32" PIPE SUPPORT        !

THERMOCOUPLE READING THERMOCOUPLE READING I l 0 563*F 563 F 1 304D 149"F 2 326*F 150*F I 3 342 F 166"F  ; 4 320*F 150 F  : 5 329 F 174*F 3 6 442 F 350*F 7 332 F 185 F - ( 8 497 F 440*F  ! 9 '343 F '171 F f 10 367"F 248*F Thermal distributions for the 10" pipe supports were conservatively assumed to be at ambient (70 F) since the maximum line temperature is 210 F and the line is uninsulated.

5. TUGCo advised that the maximum permissible misalignment f l

between the strut and U-bolt legs to be used in the analyses was 5 degrees, see Figure III-5. t 6., The strut loads used in the finite element analysis are as l followss , [ 4" 2000 lb. l

        .       10" Sch. 40      10,000 lb.                                                       I 10" Sch. 80       10,000 lb.                                                       j 1                                                                                                  \

_ _ _ _ _ _ - _ _ . _ _ _ _ . -_ J

o . l 32" M.S. - 100,000 lb. These values were determined to be reasonable maximum loads in comparison to loads compiled by ITT Grinnell, and agreed to be Texas Utilities. .- 7. U-bolt gap condition was defined by Texas Utilities, see Figure III-5. STRUT

                                                                                                                               ..            j A                        A      .R. ATE Y 9tn ygpg          t- >SDCT               :
                                                                                                                  -

• 6 -j*g ap;  ;

Gap Varies Between 8"$ar f

1/16" and 0" due to " . "

i dffference in pipe  ! and U-bolt radus Flocatt s 3r .5 r A summary of the pipe /U-bolt finite element analysis loadings is  : given below: [ Load CASE 4" sch 160 10" sch 40s 10" sch 80 32" M.S.  ; Preload T'srque 60 Ft.-lb. 100 ft.-lb. 100 ft.lb. 240 ft.-lb. Thermal Plpe Temp. 559'F 210*F 210*F 557'F  ! Hanger Temp. 304-497'T 70*F 70*F 149-440*F  ! Pressure 2485 psi 600 psi 600 psi 1285 psi l Strut Load 2000 lb. 10,000 lb. 10,000 lb. 10,000 lb. f

8 8 o . t t IV. ANALYTICAL MODEL DESCRIPTION AND ANALYSIS TECHNIQUE i P Analysis Model Description l t TUGCO requested that Westinghouse perform finite element i analyses for four U-bolt assemblies. Each assembly was modeled i on MSC NASTRAN, Version 63, a widely used and verified finite  ; element code. NASTRAN was chosen since it is a computer code that is accepted and recognized by industry as having the i theoretical capability to provide analytical solutions that j reflect the local stress, gap, friction effects, and plastic  !, material behavior that are important in assessing pipe and U-bolt support assembly stress and support stability issues. j The fc11owing NASTRAN elements were chosen for the U-bolt analysis model based on their compatsbility with the desired non-linear solution and solution' accuracy, see Figures IV-1 through IV-4 and Appendix II for more detail.  ; ITEM NASTCAN I ELEMENT ELEMENT SPACING Pipe QUAD 4_ Circumferential1y one QUAD 4 element was l used to represent each 10 of the pipe l

                                                                                                                                                           ~

wall, i.e., 36 elements were used for the pipe circumference. Longitudinally, f near the area of interest, i . e . , th e U- [ l bolt, the element aspect ratio was ul ( l and increased with the distance from the U-bolt. k i

o . This Figure is applicable to 4",10" SCH 405, and 10" SCH 80 pipe support finite element'models.

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   .         .                                                                                                                                                   t
 .         ,                                                                                                                                                          t This Fiure is applicable to 4" 10" SCH 405, and 10" SCH 80 pipe support finite element models.                                                 '

l i l t 4 -0UAD4 l I

                                                            's-meas                                                                                                  [

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                                                                       +-GAP                                                              .

g ,I <---B EAM . n t QUAo4 i . .

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                                                                                                                                  .                                   l GAP                                                                                                      .

l r l t f l SECTION,A-A , t FIGUREE-2, [; i E t l l 1 i 5 I P e

                                                                                                          - - - - - .     --,--m----,---n              - - - - , , -
                                              ,,,,c-    ,-,..,-,.---.,---,,-..,.-y.m-
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b*$ IMAGE EVALUATION

                                                          //// /     /IS%

d kf77 i f"f/ TEST TARGET (MT-3) /4 y,,, + /Q,,,, <[*[///$

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                                  -              l! l.8 11 -

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l n , , J f P This Figure is applicable to the 32" pipe support model. l l 6 L

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                                                                                                                                                                  . FIGUREE-[                                                                                                                                                            .
 ---,----r--.,-.                 ,,r-    ..                            -._..,,__#                 ,,,r.,,                              m,,,.~_ . , _ , , . , , _ , , _ , . , , . _ , _ , .                                                                        _ _ _ _ , _ _ _ _ ___ _ _ _ _ _ _ _

g , 22-  ! l l This Figure is applicable to the 32" pipe support model.  ! 1  ! F I  ! l <- QUAD 4 l 9 K  ! y=m t&4, F ( QUAD 4 j i i . .

                                                                                                                                                                                                                    ,                               j 6w                                                                                        ;

gp % j 1 jp QUAD 4 T , BEAM i  !

• l t

BEARING PLATE QUA04 l P

                                                                                                                                                                                                                /                                    '

s.a -. / .. + L' -

                                                                                                                                                                                                  /                                                  !

i i  ! i i l' e i l i SECTION B-B . - e

                              ,                                                                                            nGURE E-9                                                                                                                !

i I f 6 h a l l I i wwemw- wm. + , ---~,w. w- a,-- --w-~,,,-m--,-,-,-n-,,- - s--ew,m - ww_ pg w w y-vg e s, na m,w w g-mmmm--m,e-q.m m-- =-e---wr----- w* w - - 'w'vww w

  . ..: . . .- :. = .                                    .-    . - - - . - . -:-..:-                _ ~.

l i 4 U-bolt Beam The curved portion of the U-bolt has one i beam element "for each 10 arc. f Gap /  ! Friction Gap Gap elements were used in the 4", 10"  ! I stainless and 10" carbon steel pipe /U- I 1 bolt models at two locations (1) between  ! the pipe and cross piece and (2) between  ! l the pipe and U-bolt. t i Three rows of gap elements were used between the pipe and cross piece. The  ; inner row represents the point contact  ! I between the pipe and the cross piece } (gaps closed) . in the unloaded condition. I The tso outer rows of gap elements [ connect the pipe nodes 10 from the top l i of the pipe to the c oss piece and are gapped for the distance between the pipe [ outside diameter and cross piece at this i location. { i i Gap elements connecting the pipe and U-bolt were spaced at 10 intervals . In f the unloaded condition, the bottom gap j l

                  -,-.,---,.------,,---_----,--,y-,                                     r-. - ---    ww.w--..sg.,y    .y   --4   .---r.--w   e-    ,g.-cmgy.--

(at the apex of U-bolt) is closed and the gaps vary from 0" to 1/16" at the pipe center line elevation. Gap

• elements were used in the 32" model at three locations (1) between the pipe and cross piece, (2) between the pipe and bearing plate, and (3) between the bearing plate and pipe.

Gaps between the pipe and cross piece in

     .                                 the 32" model are the same as for the 4" and 10" models.

Gaps between the pipe and bearing plate w in the 32" model are spaced and gapped as in the 4" and 10" models with one exception. Instead of one gap element every 10 , there is a row of gap elements the width of the bearing plate. Gaps in the main steam model were represented with beam elements. I

Gaps between the U-bolt and the bearing plate in the 32" model are spaced at 10 intervals, and all gaps are closed in the unloaded condition. A friction coefficient (3p=.16) is specified for all gaps on all models. This is of course only active if the gap is closed or closes. Cross Piece QUAD 4 The element spacing used is defined by the spacing of the gap and rear bracket elements. The analytical model discussed above is used for elastic non-linear gap and friction effect' solutions. In order to assess the effect of plastic material behavior on preload and pipe stresses, non-linear material behavior was included in the 10" Sch 40 pipe model, as discussed further in Section VII. The remainder of the 10" Sch 40 pipe /U-bolt model for the plastic analysis was unchanged. Adequacy of Modeling Technique Three methods were used to demonstrate that the models are (a) accurate for the idealized situation, and (b) representative of the real condition existing in the field. 7 tem (a) was addressed by: l j

1. Comparison of element modeling size to recommended l modeling limits.

l

                                                                                          -i'G t

l 2. Comparison of analytical results obtained from the l model to results calculated by other methods. Item (b) was addressed by:

3. Comparison of analytical results to rest results. .

Items 1 and 2 are discussed below, Item 3 is discussed in i l ! the ne xt section.  ! . 1. Element Size , of primary importance to the solution accuracy of this model is the element selection and element circumferential spacing which represents the pipe wall. The following information j reproduced from the MSC NASTRAN Application Manual, Volume II, I

              ,       confirms that the selection of QUAD 4 elements a't 10 degree                                                                             :

circumferential intervals to represent the pipe wall is l l t l appropriate.  ; 1.4 ' " j g Q tana) l l 1.s - - l

                                                                                                     ",,.. p gMMP e.
                                                                                                             -                                                i

f. .

l ERROR . .

                                                                   .                          No MESH. SIEE                                                                            ;

I"I * ' (Saal

                                                                                                -.~ ===> TH roRETIc.R L                                       '

5.n SOLUTION l A 5 x 5 mesh of QUAD 4 plate elements, one element every 16 degrees yields a result within approximately one percent i i of the theoretical result. l

t i i Although the QUAD 4 is not a shell element in the strict {

                                                                              )

sense, it should be noted that excellent results can be obtained when the angle subtended by individual elements is J limited to some 10 degrees. ' Additional examples and user information on the QUAD 4 element are presented 'in the MSC/NASTRAN Application Manual.  ! Some of the important features of the QUAD 4 are indicated I below.  ! QUAD 4 may be used as a membrane element, bending i element, or as a combined membrane-bending i e lement. , The accuracy of the QUAD 4 element is effectively independent of aspect ratio. . L From the above, it can be stated that the element size used in the finite element model will yield results that are in close agreement with the theoretical solution. l 2. Comparison of Analytical Results l l The U-bolt bending moment calculated with the finite element i analysis model was compared to hand calculations. Assuming the pipe is infinitely rigid, as the U-bolt is preloaded the gap between the U-bolt and the pipe will close with no pipe

     , distortion.      The theoretical moment resalting from the change in U-bolt c'urvature would bei M = EI (1/R - 1/R )

E = 27.9 x 10 6 p,y I= 4 7T D /64 (D is the U-bolt diameter)

i i

                                                                                                                                                                                                                         }

l l l R and R are the U-bolt radii of curvature before and after preload The U-bolt moments calculated with the finite element model and the theoretical U-bolt moments calculated by the above equation are compared below. i

                                                                                                                                                                                                                         }

r i PIPE ANALYSIS DIAMETER R I 2 l THECRETICAL SIZE MGIENT 'U-BOLT INCRES INCHE3 MGtENT IN-IIPS INCHES IN-KIPS 4' .818 .5 2.5 2.5625 .835 1c'-403 .738 .75 5 75 5.8125 , .810 l 108-80 .847 75 5.75 5.8125 .810 328 4t3 13.890 2.75 18.375 18.4375 14.449 l [ / . As seen from the above table, the moments calculated with the i finite e'lement analysis model compare favorably to those  ; calculated with the theoretical U-bolt bending moment equation. t I I Analysis Technique L The Westinghouse analysis of the pipe /U-bolt structure [ t included four load cases which represent the loading conditions for the pipe /U-bolt assembly. A description of each load case l and the analysis simulation technique used to evaluate the effect , i is given below. i t I i

l t

                                                                                                             ,.                  L l

t LOAD CASE SIMULATION TECHNIQUE NASTRAN i LOADING l j Preload/" Cinching" - Torquing of the U-bolt nuts TEMPRB f due to torque on shortens the U-bolt legs l i U-bolt nuts (straight portion). Preload ( was therefore simulated by j z imposing a negative temperature [ r I on the U-bolt legs which  ! I results in shortening of the  ; U-bolt legs.  ; l - Radial expansion To simulate the effect of pipe PLOAD 2 l of the pipe internal pressure, internal from internal pressure was defined on the  ! pressure inside surface of each pipe l element.  ! l l I I Thermal expansion To simulate the thermal load, TEMPRB i of the pipe / pipe' the pipe wall was specified to

• TEMPP1

               .                                                                                                                 I support                 be at the normal operating I

temperature and, the pipe ,

support was given the temperature distribution determined from the test thermocouple readings. Applied strut Engineering and construction FORCE load

• tolerances are such that the ma ximum strut udsalignment -

which results in pipe torsion is 5 de gree s, see Figure IV-1. Maximum strut forces were applied as shown to determine - pipe stresses and checked for slippa ge.

   *All four finite element models had 10 ft. of pipe length (corros-ponding to test speciment) with pinned end conditions. Therefore the applied strut load, at pipe midspan, would create a moment equal to PL/4. An equal and opposite moment was- applied to the ends of the beam so that at the pipe midspan, the moment would be zero. This was done so that the results from the finite element analysis would reflect only the local effect of the U-bolt. The magnitude of piping moment stresses is determined in section VII of this report.

s s . V. COMPARISON OF ANALYTICAL RESULTS TO TEST RESULTS In this section, the analytical results are compared to the test results obtained from the tests performed at the WAESD Engineering Laboratory. Three test conditions are compared. They are preload, thermal, and load distribution. The other tests performed (torque versus preload, frict' ion test and creep) were not used for comparison since they were performed to provide data on physical properties and behavior.

1. Preload Test The pipe, U-bolt and cross piece were instrumented with strain gauges at the locations shown ,in Figures V-1 and V-2.

The U-bolt nuts were torqued to the maximum value specified j by Temas Utilities and strain readings were taken and converted to stress. ' Analysis The U-bolt legs were " shrunk" using the previously described technique until the analytical U-bolt leg force corresponded to the preload test leg force. A stress comparison was then performed between test and analytical i stresses. Table 1 provides the strain gauge versus analytical element correlation. Tables 2, 3, 4, and 5 correlate I stresses from the test and analysis for the 4", 10" f stainless,10" carbon, and 32" pipe /U-bolt assemblies, respectively . L

 . . _ _ _ .   ._       - ___  _ a                     ,
                                                                                   ~ _ _                _                                              _ ,

r ' l ,' P 1 E . t D d  !

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• F .

l FIGURE V-1 l

                                                                                                                                                          'a PIPE INSTRUMENTATION                                                                            l l

l l

                  *A. Circumferential mounting on pipe as close as possible to U-bolt contact with pipe.                                                    ..

I i

                  *B. Circumferential mounting on pipe midway between cross piece                                                                       r and U-bolt contact points.
                  *C. Circumferential mounting on pipe as close as possible to point of cross piece contact with pipe.                                                                                            l
                  *D. Longitudinal mounting on pipe as close as possible to cross                                                                       I piece contact with pipe.                                                                                                           i
                  *E. Longitudinal mounting on pipe two inches from D.                                                                                   .

[

                  *F. Longitudinally mounting on U-bolt shank opposite cross piece.

l G. and *H. Longitudinal and circumferential strain measurements 45 degrees below horizontal center line on pipe ao close as , possible to the.U-bolt contact area. t T

i I . l l . l 33-l I. and *J. Longitudinal and circumferential strain measurements i l on pipe opposite cross piece as close as possible to the U-1 i bolt contact area. K. Longitudinal strain measurement on U-bolt shank. Two gauges were placed on each U-bolt shank 180 degrees apart. The strain readings were averaged to remove the effects of

                     ~

bending. L. Longitudinal strain measurement on cross piece as close as possible to strut rear bracket. l l

• Actual strain gauge locations are detailed on Figure V-2.

1 w MM e 4 4 s 5

 ,      .a                e- ,,<-.n   -
                                                 .sa.  ~.-         ,n m a - -         -s.        - - - - - - -   --a      a-    --.s-  w._ ..             -

i TESTING STRA/N GAGE Lo cA T/oNS

                                         \ Cross Pirci                    /                                                    Y I'5

• D E

                                                               ,           8                                                   E- l%~44- 2' -b
                                                                        .s                                                   -

l 7 a 0*fgod 16 } s e,1o* \ \ U! G H \ egg y , I g*1 a +- v _ _ _ _ _ _l > . 1 ItT - I- y

                                        \ enoss prect                         /                                              Y-24 y           g c                          r g                                              N- 2. "*4-2'
                     ,                                      e8                  A to                                                       ;,.                    .
                                                                                           /

___ __ ____s_ O 9o H i, __) __ __ G_l (; Vi1 a +-h: ( I ~7 s cross virce

                                               ^
                                                                              /                                             Y 3"*     D       E i
                                           /                                  g                '
                                                                                                                             &-3fl4- 2'
                                                                                                                                                     \

e*& e '5"

                     "                                           - -- n h                    c
                                                                                               ---                     --       - _ -. _ a
                                                                                                                                                   /
                                                                                           }

e too e es*} H __y Gl j _ 2_~ 1 G +-f (> . i T T FIGURE V-2 V

o . e . TABLE 1  ; I Analysis Elements Corresponding to Strain Gauge - l Lccations on Previous Page i i

        ,                          4" Schedule 160                10" Schedule 405               10" Schedule 80 32" Main Steam                         l Test Gage              Model Element                 Model Element                Model Element                         Model Element   l c_-.                                                                                                                                                     >

A 152/151 - 155 78 215  ; 8 154 152 82 219 C 157/156 149 85 223 D 734 723 590 295 E 806 759 626 295 i F Not used G 145 125 73 210 i H 145 125 73 210 f I 141 129 69 206 f J 141 129 69 206 K 81 3 861 133 l L Not used/4" test ~ 1132 7 60 NOTE: Two elements are given where the strain gage location falls between f elements.  ; h i O a l l

                                                               - . .. .                         ..  . - - - - -       .-.    . _. :.                . - -        ... :==.                     :.

l TABLE 2 - 4 Inch Pipe /U-Bolt Assembly l l comparison of Test Results to Analysis Results for Max Preload l Max Preload = 60 ft-lb. f Test Test Strain Test Stress lb./in2 Analysis I 2 Gage c(uin/in) a = cE Stress (ib./in ) ' A +158 4,424 10,014/9,026 8 + 86 2,408 3,413  ! l C -255 -7,140 -8,009/-16,265  : D -107 -2,995 -3,878 E + 53 1,484 1,020 F Not used on 4" et'st i G -15 - 420 -10,410/-1066 H -184 -5,152 -14,519/-5100 I - 95 -2,660 -9,827/-3815 ' l J -209 -5,852 -12,151/-7769 { K +951 26,628 27,544  ; L Not used on 4" test > i l E = 28 x 10 6 lb./in2 Note: Two element stresses are given where the strain gauge falls between elements. , 9 I T l l

                 . . _ . . _ ~ - _ . - - . - - - - _ . - - _ _ _ _ - - . . . . . - , ,                                            - . _ . _ _ - _ _                  _ _ _ -  - _ _ _ - . _ -

  - . . .                    . . . .    .   -    . . .     . . . - - . = - . . - - . .    . - .                      .   . = -     - - . . =- - = . - .
            ,                     ,'                                                                                                                      i i

i  ! i TABLE 3 i t 10 Inch Schedule 40S Pipe /U-Bolt Assembly f comparison of Test Results to Analysis Results for Max Preload " i 1 Max Preload = 100 ft-lb.

i Test Test Strain Test Stress Ib./in2 Analysis I 2

Gage (uin/in) cr = eE Stress (ib./in )  ;

                                                                                                                                                        )

A +425 11,900 13,568 f

                                                     +486                         13,608                           11,536 B

( C -561 -15,708 -20,040 - l< D -270 -7,560 -20,561 l l

;                                  E                   -10                          -280 '                       -8,764/-6,135                           l F                                 Not used on 10" test                                                               I G                 -275                         -7,700                          -10,212                               I H                 -685                       -19,180                           -13,684                                I I                   +30                            840                          -6,489                               !

J +70 1,960 - 8,816  ! K +458 12,838 12,726  : L +579 16,212 16,702 I i 6 l E = 28 x 10 lb./in 2 [ l 1 - l i I 4

          - -- .-._ - _. _ _ - ..                                      .- - - -                 _ . - .. - ._           - - _ -. - - _- _ _ - - ._ I

TABLE 4 10 Inch Schedule 80 Pipe /U-Bolt Assembly comparison of Test Results to Analysis Results for Max Preload Max Preload = 100 ft-lb. Test Test Strain Test Stress 1b./in2 Analysis Gage (vin /in) a-E Stress (ib./in2) A +213 5,964 8,910 8 +200 5,600 6,261 C -289 -8,092 -14,170 0 -24 -672 -8,291 E 20 560 . -2,094 F Not used on 10" test G -76 2,128 -6,665 H -92 -2,576 -10,875 l I -47 -1,316 ,

                                                                    -5,755          -   -

l J -44 -1,232 -6,888 K +613 17,164 17,008 L +754 +21,112 23,678 6 2 E = 28 x 10 1b./in l l

                                                                                ~

9 L

e ' TA8LE 5 32 Inch Pipe /U-Bolt Assembly Comparison of Test Results to Analysis Results for Max Preload Max Preload - 240 ft-lb. Test Test Strain Test Stress Ib./in2 Analysis Gage s(uin/in) cr sE Stress (ib./in2)

A 0 0 1,113 l 8 +15 +420 789 C -56 ,

                                                   -1,568             -4.859 0               -11                       -308               -4999       .

, E 4 +112 '493* F Not used on 32" test G 0 0 -1,170 H 0 0 -1.446 I -3 -84 -1,075 J 0 0 -1.208 X +23 +644 +1.018 L +25 +700 +1.242 6 E = 28 x 10 lb./in 2

        *0ne Analysis element encompasses both these strain souge locations.

4 L

DISCUSSION Given below is a comparison of the test and analytical results for each pipe size evaluated. 4" Pipe Sch 160 Good agreement 1/ exists between the measured and calculated stresses at B and C. Stresses measured at test points A, G. H, I and J do not compare as well with those calculated with the finite element analysis model. This is due to the difference between the pipe and U-bolt boundary conditions (U-bolt fit-up). In the test, gaps existed between the U-bolt and pipe which did not completely close due to out-of-roundness of the pipe and tolerances in the U-bolt band. This resulted in point contacts between the pipe and U-bolt which cannot be simulated exactly in the finite element analysis. The' stress in the U-bolt (point K) is comparable in both the test and analysis. Good agreement between stresses exists at points D and E. In general, there are differences between the test and finite element preload analysis, but they are within the accuracy of (1) the location of the strain gauges, (2) the finite element location with respect to the strain guages and, (3) the simulated fit-up between the U-4 bolt and pipe in the analysis. 1/ Agre'ement between tests and analysis will be considered good, if the finite element analysis stress within (+) 10 degrees of the test strain gauge in the circumferential. direction matches the test stress. Longitudinal stress agreement will be considered good if the finite element stress matches the test stress within one longitudinal analysis element of the test strain gauge location.

                                   = s .-               -
                                                                                     -                   - =-               -   - -     -

r i  !

                    ,                                                                                                                j 1

i l

 ,                  10" Pipes Sch 40 and 90 t

Good agreement exists between measured and calculated stress } 4 at locations A, B and C. There are differences near the contact h point of the cross piece and pipe, but this is due to the large i increase in stress over a short distance (small angle). P: eload .- b test stresses for points G, H, I and J do not compare well to . calculated stresses at the same point. This is again attributed l to the difference in U-bolt fit-up between the test and the finite element analysis model. Analytical results at points D  ! l and E do not agree with the test results. The differences are l attributable to the following:  ;

First, the same cross piece was used on both the 10" Sch 40

                                                                  .-                                                                       3 and the 10" Sch 80 preload test and it was observed to be                                                              ;
;                   " bowed".       The edges of the crass' piece, in the direction that bears on the pipe, are approximately 1/64" higher than the cross piece centerline.                 This is probably due to welding of the rear                                          l bracket.        The deflection of the pipe in the analysis at the                                                      ;

l contact point of the cross piece is approximately 1/32" for the . 10" Sch 40 pipe. Therefore, the offact of " bowing" would be to l l concentrate the contact force at the center of the plate which  ; l would lower the stresses at points D and E making the test and I I ! analysis comparison more favorable. [ The,second observation wi'.ch would also result in forces 2 being concentrated under the center of the plate, thereby  : l f reducing the calculated stresses at points D and E,is the fit-up l I o f the U-bolt. If the unloaded gap condition in Section III is l L

 .           -                   ,      . .-                                 .               c-_       . . . . _    -.     -

r

                                      -                                                                                                                                                                       i

, i

                *                   .e                                                                        .

I L assumed between the pipe sidewalls and the U-bolt, contact forces  ; I concentrate at the edges of the cross piece plate. If the U-bolt is assumed to conform to the pipe in the unloaded condition, the F bearing force under the cross piece will be concentrated at the I l center of the cross piece plate, thereby reducing the calculated i stress at points.D and E. This observation is supported by the [ l analysis performed for the boundary conditions described. The l

                                                                                                                                                                                                              )

l fit-up of the U-bolt and pipe was observed to more closely ( I approximate conformance to the pipe in the unloaded condition j

than the 1/16" gaps at the pipe sidewalls input in the analysis. (

r 32" Pipe  : The stresses measured in the test and calculated for the 32" f pipe, cross piece, and U-bolt are comparable. They are very low,  ! l almost immeasurable. This is due to the fact'that the U-bolt [ I r preload torque was very small (240 fb=1b). It sho.1d be noted however that the radius of curvature of the main steam line U- , bolt was too small, as shipped to the WAESD Engineering Laboratory, to fit on the pipe. The U-bolt legs were therefore j i

                                   " spread" at the Laboratory so that fit-up to the pipe could be                                                                                                            l accomplished.                                                 This spreading resulted in pipe /U-bolt fit-up not                                                           l consistent with the U-bolt fit-up as modeled.                                                                                             Specifically, (1)                l i

point contact existed in the unloaded position between the pipe sidewalls and U-bolt, and (2) the apex of the U-bolt which should ( have been the only contact point in the untorqued condition had a gap which did not close even at the maximum preload torque value. i r

l l , 6

2. THERMAL ,

, Test t Thermocouples were placed on the pipe /U-bolt assembly at the  : locations defined previously in Section III, and high temperature . i strain gauges were placed on the U-bolt legs. All four U-bolt - supports were assembled and insulated in accordance with Texas Strip he'aters were then sealed in.the Utilities' instructions. l pipe and heated until the pipe skin reached the specified normal  ; operating temperature. The skin temperature was held constant i 4 until the pipe hanger reached equilibrium. Thermocouple readings ; were recorded at equilibrium and U-bolt leg strain readings were recorded and converted to force. Analysis The pipe elements were given temperatures equivalent to the normal operating tamperature defined by Texas Utilities and used in the test. The pipe hanger thermal distribution was input from i thermocouple data recorded during the thermal test. The forces ! l in the U-bolt legs were calculated and are ccmpared to test , i results in the following table. THERMAL THERMAL  ! PIPE SIZE TEST FORCE ANALYSIS FORCE l 4" 1.7 kips 2.9 kips j 10" stainless .5 kips 1.'3 kips f r 10" carbon Not tested 1.4 kips 32" Main Steam 10 kips 21 kips I - -

                            . .  ..    .   . . - - - -      .. - - - - - .                   .~
        . e
     *

• Comparison of 32" Pipe Thermal Test and Analysis As stated previously, the actual test fit-up between the pipe and U-bolt for the 32" pipe pipe /U-bolt assembly was totally different than that for the 32" finite element analysis.

Therefo're, dif ferences in results would be expected. Comparison of 4" Pipe Thermal Test and Analysis The measured forces were 70 percent smaller than those calculated with the finite element analysis model. During the preload test, high stresses and some yielding in the U-bolt and hanger cross piece were noted for the 60 ft-lb preload torque. This observation is substantiated by the finite element analysis that was, performed. See Stress Tables in Section VIII. In addition to increasing the hanger loads and stresses, the elevated temperature would decrease the yield point of the hanger material. The expected result would be that more yielding would occur and the increase in U-bolt leg load predicted by a linear finite element analysis would not be duplicated in the test. This conclusion is supported by the figure below developed from the test results. ,

                                                                  , ,, ,,,33 , ggt Pipt Tt*

• 940*F g.acLT 7tw e 490*F
• Las ,

                                                                                       '"EI 33' 38 0 Representative tempera,ture.           7,,

Refer to page 16 for actual g tcmperature distribution in _ th:2 L-bolt. , Las t

                                                                                        '"E3 58' 33
                                     *     ..o                                                    '

I i I S.O i , . . . . , , , ,, CTCLt1

l . The sharp decrease in the U-bolt preload during the first few thermal cycles indicates the U-bolt assembly did not behave elastica 11y during the thermal testr i.e., yielding occurred. Comparison of 10" Stainless Steel Pipe Thermal Test and Analysis Although the elastic finite element model predicts piping stresses above the pipe material yield point test results shown below indicate linear material behavior. 10 INDI sTAINL!ss sittL

                                        ,      PIPt Tt e e ite*f U 4GLT few = 190*F

• 6 .0" l

Lf4 2

                                "                                     sausts se, si

'

• Representative temperature. 4.0 ._

                                        \        '

3.0 < , ggg y

                        ,                                             causts $4, sti
                            ..o                              .

to

3 4 s a 7 e s cuts Because U-bolt preload as a function of thermal cycles is essentially constant for both U-bolt legs, the difference between measured and any calculated thermal loads is attributed to hanger

            " binding". Both U-bolt legs were torqued at ambient temperature to approximately 4,500 pounds. At temperature (pipe = 250 F) the U-bolt leg forces decreased to 4,000 pounds but no further decrease was noted for the second and third thermal cycles.                    Leg i decreased approximately another 600 pounds between thermal cycle 3 and 4. These step decreases in U-bolt load indicate some

form of mechanical binding existed in the hanger assembly following the 100 ft-lb preload torque which was relieved during thermal cycling. The results between test and analysis are therefore not comparable because of this phenomenon.

3. LOAD DISTRIBffTIONS

Test In this test, the strut was given an applied force (" push"),

i see figure III-5. This test was performed on the 10" stainless pipe only. The test consisted of torquing the U-bolt nuts to 100 3 ft-lb and then pushing on the rear bracket with a force of 7,000 lb. The test resulted in a 2,000 lb. reduction in U-bolt preload. Analysis In the analysis corresponding to this test, a preload of 100 ft-lb and a 10,000 lb. push force was applied to the rear bracket. 1 The analysis resulted in a 2,700 lb'. reduction in the U-bolt i preload value. This is equivalent to a reduction in U-bolt preload of 1890 lb. for a 7000 lb. push force. As can be seen from the above comparison of test and analysis i , l the finite element model closely predicts the test behavior. g 1

                                                             -47 l

VI. BEHAVIOR OF PIPE /U-BOLT ASSEMBLY FOR " SLIP" This section describes the characteristics of the U-bolt assembly whJn subjected to a push load which produces a moment about the pipe centerline.

                                                              -    F(Applad rew)          -
                                                                -       suvT
                                                  $                  "I purr I                                     hamn y                  *-nos 1

1 The stiffness equivalent diagram below can be used to show that the normal force (Fg ) for an externally applied load (F) will be distributed between F p (pipe contact force) and FD (U-bolt leg force) in a manner that depends on the relative magnitudes of the f hanger stiffness (!gUBE' "" ** "***

• U-BOLT

( PIPE

• 7 l

F

                                       &                                                        Fu .                                   1
                  .                               war erressee earnesee<

i fl

                             <-> cra                       ss,rm pi Af
                                                                                    )w,,                g gra t                 *
                                                         ~ _ ,          .-                                                             l storness                             k. % s'!
                                          -        l 1             Equivam/7"       S, Kvar
                      \ ,\ N                      bhl                                                                                  \

 ,.    .                                  .48-  .

If the pipe is very stiff compared to the hanger, the external force component (F y ) will increase the pipe contact force (F ) by essentially F y with no decrease in the U-bolt preload value (F,). If the pipe is not very stif f compared to the hanger stiffness, the external force component will be split between an increase in F,and a decrease in the U-bolt preload F, which dependson the relative stiffnesses of the hanger and pipe. The change in F g and F, due to the applied external force component (Fy ) is tabulated below for the four pipe /U-bolt models evaluated. A minus (-) sign indicates a decrease in load; a plus (+) sign indicates an increase in load. . Change Resulting Fross FM PIPE SIZE Fy MEYSIS F g AN EYSIS F, 4" 2,000 lb. - 134 lb. + 1,630 lb. 10" stainless 10,000 lb. - 2,415 lb. + 4,770 lb. 10" carbon 10,000 lb. - 1,295 lb. + 7,410 lb. 32" Main Steam 100,000 lb. -24,400 lb. +50,800 lb. The distribution between F g and F ,of the applied external force component determines to a large extent the ability of the hanger to resist. This is explained below. For the sinrple example below, it can be shown that the block will not slide until the force (F) is applied at an angle (4) which exceeds tan,p"I . where/4.is the coefficient of friction, 12,3,4,,Fg is equal to jaAFg at the initiation of' slip.

                                  ,        k    I em Ale h
                         ' /    /   /  /   /4k/    / /    /  /

 -. u.-   ---               - . . . . . . . .                      -                                                        -           . . . . - - -        . - -

l

        =

l l I i, The cross piece for the pipe support assembly in question also j obeys the same principle. l b I

                                                                                                  ~

sinvT l

• t ,

I A A mir I _W4 Mm I

                                                                                                                              '19                                                         i w                                                                           w FaL From the free body diagram of the cross piece shown above, it can be seen that the cross piece will not " slide" until the applied strutforcecomponent(F)                                                                                                                  Using We d n hum 3 exceeds /F.                                                  p                                               ;

coefficient of friction determined by the testing program ( p = i

               .16) and Fp

• from the analysis, the following information is j i

given and conclusions reached regarding " slippage" of the j hangers. , l PIPE SIZE Fg F* gFp " SLIP" OCCURS l p 4" 174 lb 15,280 lb 2,444 lb No  ; 10" Stainless 871 lb 19,210 lb 3,073 lb No i

                                              ~

871 lb 25,790 lb 4,126 lb No e 10" Carbon  :

                                   ~                         ~

32" Carbon 8710 lb 105,000 lb 16,800 lb . No x s r

                       .,                                                                                                                                                               t Fp   is the calculated load due to preload, thermal pipe expansion pressure pipe expansion and push.                                                                                                                         ;

       ,.                                                                                                      l L

As can be seen from the table above, Fg is in all cases is much smaller than and therefore slippage of the cross piece will jAAF p not occur. Although all four hanger cross pieces can be seen not to " slip", the external force component (Fg) creates a moment about point A which must be balanced by a difference in U-bolt leg tensions (FBI "" B2 r - The difference in U-bolt leg tensions is balanced by two methods, the friction between the U-bolt and the pipe. The second and a mechanism of resistance that would be in effect if the friction capacity between U-bo,lt and pipe is exceeded. The mechanisms of resistance would be as diagramed below. l l .I Y // TE I [K A It has previously been demonstrated that the cross pieces will not " slide" on the pipe. However, the cross piece will " roll" on the pipe surface until a difference in leg tensions exists

 . _ _ . _ _ _ _ _                                       ..                     . .___ m              m._       -

s j i sufficient to balance the moment created about point A by force i F. In other words, it can be shown that as the cross piece , g rolls , point A moves to A', point Y moves to Y', and point Z moves to Z', and that the distance from Y' to Z' (around the pipe) is greater than the distance _Y to Z. The increase in l length creates additional tension in the U-bolt. The cross piece. will continue to " roll" until (friction forces, uFn) sufficient normal forces between the U-bolt and pipe are created to balance ' i the cross piece moment created by F g. fs r I F t l . l i l i l l i ! \ l l i ll t l e

                   - . - - -   ,,,--._,_----,,__---y              __-.--_.-_,w,   ,     , - .   , _ .     .c _,

_e__--. -

x.. - . . . . . - . f

                           ,a                         l                                                                                                                                  i i

t VII. ACCEPTANCE CRITERIA l f l The rules of Subsection NB given in the 1979 ASME Boiler and I Pressure Vessel Code, Saction III, (referred to as code hereaf ter) l may be used to' qualify a Class 1, 2, 3 or non-nuclear safety piping components if the designer is willing to comply with all requirements of Subsection NB. Subsections NC, ND and the ANSI B31.1 piping code provide criteria for evaluating primary and j secondary stresses only by limiting the pressure plus principal l stresses caused by moment loading in the pipe. By evaluating stress f intensities (as opposed to principal stresses) , detailed localized j r stresses, and the effects of cyclic loadings (f atigue) , one is j clearly enveloping the intent of NC, ND and ANSI by an NB evaluation. Based on the above, the discussion for acceptability of pipe stresses induced by the U-bolt pipe. support will generally be i addressed herein using NB-3600 of the code. NB-3600, which governs the design and qualification of piping systems gives little specific  ! guidance to the method that should be used to evaluate stresses similar to those caused by the U-bolt pipe support. However,

                                                                                                                                                                                              )

equations 9 and 12 of NB-3600 may be used to assess the piping stresses as discussed herein. This will provide a means for evaluating piping stresses in the vicinity of the U-bolt caused by l the loading assoc*2d with the U-bolt.

                                                          'In applying code equations 9 and 12 to assess the significance                                                                      j of the piping stresses it is necessary to classify the stresses that                                                                     .

exist in the pipe due to the different loading conditions associated with the U-bolt. These piping stresses are classified following , l' ASME nomenclature.

                                                   ,a   l Using code equations 9 and 12 requires that the mechanical stresses'in the pipe resulting from sources other than the U-bolt assembly be included.       Because of the number of hangers involved, this information was not supplied to Westinghouse for evaluation purposes.

In order to conservatively estimate the magnitude of these mechanical stresses, a gener-ic procedure was developed and is discussed herein. In the development of these mechanical stresses, it is assumed that the pipe is stressed to its maximum allowable  ; value at a high stress intensification point (e.g., elbow). It is assumed that the piping moment which results in this maximum allowable stress also occurs at the U-bolt hanger locations. The maximum piping stress is assumed to be at the stress limit of 1.5 S,. 2/ This code allowable stress level is used instead of the

 -     higher Level D limit (3 S,) for the'following three reasons.

1. SSE induced stresses need not be included in the secondary or peak stress evaluatidn.

2. SSE piping seismic levels are in the same order of magnitude as those associated with the OBE (operating basis earthquake) stresses.

3. In general, the stresses in the piping will not be at the code allowable.

2/ Sm is the allowable stress intensity defined by the Code. l

 --    . . u- -                 :                     _
                                                                    ;__ ; _                      _;    _      ___ , 3-Therefore, based on the above, using a 1.5 S , limit will                                                  l produce mechanical piping stresses that are realistic.                                                             j In the assessment of the piping stresses, the allowable stress                                             i used for code Equations 9 and 12 is 3                     S,.              It is realistic to use                 j this allowable for evaluation of primary stresses in the vicinity of                                              !

the U-bolt since the preload stresses in the pipe are predicted to j be very high by elastic analysis due to local stress  ; concentrations. In reality, these stresses will not result since small insignificant plastic deformations will occur in the region of  ! I the high stress concentration resulting in stress redistribution in I a small local area of the pipe. Further, a fatigue evaluation was l performed in which this high stress state is considered. l l In the sections that follow, the piping mechanical stresses are determined and the piping stress state in the region of the U-bolt i support assessed.  ; 1 ! Code Classification of Stresses l l  ? Pressure Membrane Stress - Primary Stress j t Preload - Local Primary Membrane Stress j Thermal Pipe Growth Restriction - Secondary or Peak Stress Pressure Pipe Growth Restriction - Secondary or Peak Stress . Pipe Hanger Load - Primary Stress I Applicable Code Equation Stresses l From the above stress definitions and code requirements, the l appl,1 cable code equations for these stresses and limits can be i l selected.

i c . , i l i t

1. Equation (9) of the code.  ;

This equation must consider (a) primary membrane pressure stress, i (b) piping moments at the hanger location due to pipe deadweight and Seismic, (c) stress due to preload of the U-bolt and (d) stresses , due to the deadweight and seismic portion of the applied hanger load. t

2. Equation (12) of the-code.

i This equation must consider (a) thermal stresses due to pipe radial  ; thermal growth restriction, (b) pressure growth restriction at the [ U-bolt, and (c) piping moments at the hanger location due to pipe  ! I longitudinal thermal expansion.  ! Development of Mechanical Pipe Stresses  !

                                                                                                                                  ?

Tabulated below are the 3 S, stress limits for the pipes in . question.  ; PIPE SIZE 3 S, AT NORMAL OPERATING TEMPERATURE [, 4" Sch 160 STAINLESS 50.52 KSI 10" Sch 40 STAINLESS 60 KSI f 10" Sch 80 CARBON 60 KSI 32" MS CARBON 58.26 KSI e I All sources of load must be considered when evaluating the pipe l t stress state at the U-bolt location. Because of the number of l t hangers involved, detailed information was not supplied to j Westinghouse for the piping moments at the hanger locations. However, these moments are considered in a general manner herein.  ! i Gene,ric piping stress intensification values are developed in the

             )                                                                                                                    I l           l              sections that follow.

l  ; I f

 .---r--   -
               ,e-,-.      .,v----rttv-------------em--r-v               -v--'--

l

1. Piping Moment Stresses for Equation 9  :

Piping stresses caused by sources other than the U-bolt can be l approximated assuming that (1) the pipe meets the service level f i limit of 1.5 S, (pressure + moment) , (2) pipe supports are not  ! located at high stress index locations and (3) the high stress index j locations generally control the piping system stress qualification l (elbows , tees , branches, etc.) . The magnitude of moment stress can . be determined at the hanger locations assuming (1) the maximum - system stress is at a long radius elbow (high stress I intensification), (2) the long radius elbow is stressed to the code i allowable (1.5 S,) and (3) the piping moment at the pipe hanger is l the same as the piping moment at the long radius elbow. The result  ! of this evaluation is presented below. I l l PIPE SIZE ALLOWABLE STRESS LONGITUDINAL EQ. 9 PIPING i FOR SERVICE PRESSURE MOMENT STRESS [ ) AT PIPE HANGER * [ LEVELS A AND B STRESS l l 4" Sch 160 25.26 ksi 4.8 ksi 12.146 ksi i 10" Sch 40 30 ksi 4.4 ksi 6.0477 ksi ( 10" Sch 80 30 ksi 2.6 ksi 9.23 ksi ' 32" Main Stm. 29.13 ksi 7.1 ksi 6.7 ksi l

• Piping Moment Stress = (Allowable Stress - Pressure Stress) l l x Ratio of Stress Indices i

2. Piping Moment Stresses for Equation 12 ,

l

                                                                                                                                                                                \

The qualification of a piping system for thermal growth moment i stresses is known to be controlled by the elbows. This is due to the high stress index at elbows and the fact that the piping thermal l expansion moments are generally highest at elbows. Assuming again that (1) the maximum system stress is at a long radius elbow (high  ! L i I

            -,        ...~,-                        - ,   . . -   . . - - - - . . . - , - -             -- .- -_.-.                      . . - - . - - . - - - .            ,          -- -

     . a :-.      z..                                                                     ..             -                         .     : = . =.: . 2 - -        ._.:.-.=.=               ..,

l

              .                  .                                                                                                                                                              t l

stress intensification), (2) the long radius elbow is stressed to the code allowable (3 S,) and (3) the moment at the pipe hanger is , the same as the long radius elbow, the magnitude of the piping ' moment stress can be determined at the hanger locations. PIPE SIZE ALLOWABLE EQ. 12 PIPING MOMENT  ; STRESS STRESS AT PIPE HANGER

• l 4" Sch 160

                                                                                     '~

50.52 ksi 22.49 ksi l 10" Sch 40 60 kai 10.63 ksi  !- 10" Sch 80 60 ksi 15.15 ksi [ 32" Main S tm. 58.26 ksi 13.34 ksi  : L

• Piping Moment Stress = Allowable Stress x ratio of stress indices. [

EVALUATION OF PIPING STRESSES i From the finite element analysis performed for the four different U-bolt assemblies, the highest stressed piping element  ! (see Stress Tables In Section VIII)' correspond to the same relative location, directly under the cross piece. At this location, the circumferential and longitudinal stresses correspond to the major and minor principal stresses (i.e., no shear i stress). The longitudinal, circumferential, and the major and minor principal stresses at the inside and outside pipe surface at this {

                                                                                                                                                                                               \

location are tabulated below for the maximum load case (Preload + [ Thermal + Pressure + Push) . i 4 I k

i Principal Stress (ksi) (ksi) (ksi) (ksi) , Long. Circ Mai Minor r 4" Sch 160 (inside) 10.49 44.79 44.78 10.50 l (outside) -26.65 -34.07 -26.63 -34.08 10" Sch 40 (inside) 10.77 72.71 72.71 10.77 ' (outside) -48.02 -73.46 -48.02 -73.46 l 10" Sch 80 (inside) 10.24 43.15 43.15 10.23 l (outside) -30.22 -44.38 -30.22 -44.38 l 32" M.S. (inside) 19.58 47.17 47.22 19.52  ! (outside)

                                                                                                                                 -31.01      -34.10       -30.89     -34.22 (negative is compressive stress)                                                       l It is evident from the values given above that the icngitudinal and circumferential stress are similar to the principal stresses.

P i Adding the longitudinal pressure effects results in the following: Long. Circ. [ ksi ksi  : 4" Sch 160 (inside) lDFT79 44.79 (outside) -21.85 -34.07  ! 10" Sch 40 (inside) 15.17 72.71 l (outside) -43.62 -73.46 l 10" Sch 80 (inside) 12.84 43.15 (outside) . -27.62 -44.38 t l 3 2" M . S . (inside) 26.68 47.17  : (outside) -23.91 -34.10 l The relationship between the longitudinal and circumferential  ! stresses and the stress intensity is illustrated in Figure VII-2. It can be seen from Figure VII-2 that the primary and secondary j piping moment stresses, which only affect the longitudinal stress, will not increase the stress' intensity provided: f I

1) The longitudinal stress does not exceed the circumferential j stress.

2) The longitudinal stress does not decrease to the extent it t

l changes sign, in which case the stress intensity would be the absolute sum of the circumferential and longitudinal i stresses.  ! l .

                                                                                                                                                                                         ?

I

l l The mechanical piping stresses developed earlier for the primary (Equation 9) and secondary (Equation 12) effects must be included with the longitudinal stresses. The total stress intensity for each [ of the piping sizes evaluated are: . TOTAL STRESS INTENSITY 4" Sch 160 64.14 ksi , 10" Sch 40 74.21 ksi l 10" Sch 80 54.69 ksi I I 32" M.S. 47.17 ksi i Splitting the above total stress intensity into primary (equation 9) and secondary (equation 12) stresses results in the following: t t [ Eq 9 Eq 9 Eq 12 Eq 12 I Allowable Allowable l

                            ~

ksi - ksi' ksi ksi 4" Sch 160 31.60 50.52 32.54 50.52 t 10" Sch 40 60.61 60 13.6- 60 t 10" Sch 80 38.15 60 ~16.54 60 3 2" M.S. 30.57 58.26 16.6 58.26 As can be seen from a comparison of the above maximum stress  ! I intensities to the equation 9 and equation 12 allowable stresses, l the 10" Sch 80, 32" MS, and 4" pipes meet the stress criteria established herein. The 10" Sch 40 pipe essentially meets all of l I the allowable limits. ' Further, the following elastic / plastic analysis and test results i are used to demonstrate the acceptability of this pipe /U-bolt l assembly for existing stress states. I

 - --.c,     ,      . - ~ _   _ . _ . . . . , . -_ _ _ _ _ _ _ . , . _ _ , _ _ _ , . . , . . . _ _ _ _ . , _ _ _ ,        _ _ _ _ _ _ . _ , . , . _ _ _ _ _ _ . . _ _ . , _ _ _ _ , , .

Because the acceptance stress criteria used (3 S ,) exceeds the material yield strength, the magnitude of yielding must be addressed, and the effect on U-bolt preload determined. Two methods were used to investigate the yield effect on preload. First, the nonlinear material properties (stress-strain curve) were input into the finite element analysis, and second, a test was performed at the WAESD Engineering Laboratory. The elastic / plastic finite element analysis was performed on the 10" Sch 40 pipe /U-bolt assembly for the preload load case (100 f t-lb) . The 10" Sch 405 pipe was selected for this analysis because the stresses in this assembly were the highest, and therefore would yield the highest strains. The stress-strain curve input is should in Figure VII-1. The results from the analysis showed that yielding occurred only locally around the contact point of the pipe and cross piece. Further, the strains were ve'ry small and do not significantly affect the linear stress distribution within the pipe. This is seen below in the stress comparison table. TEST GAUGE ELASTIC ANALYSIS PLASTIC ANALYSIS

• PSI PSI 13568 15358 A

11536 14068 B

                                                   -20040            -17567 C

i D- -20561 -19674 E -8764/-6135 -9232/-6536 F G -10212 -10667

• See footnote next page.

l

 ,        l                                                                                                                                                         !

i B -13684 -14152 l I -6489 -7265 g J -8816 -9677 ) K 12726 14335 i L 16702 18640 l

• l U-Bolt 5.62 kips leg tension 6.33* kips leg tension By comparison of the above elastic and elastic / plastic analysis f

results, it can be seen that for the preload case, the difference in j deformation between the elastic analysis, that shows stresses equal } to twie yield locally under the cross piece, and the j elastic / plastic analysis, that shows yielding and redistribution of the stress under the cross piece, is negligible, f This analytical result is substantiated by test results. The f test was conducted by preloading the hanger to 100 ft-lb, and then l pushing with a 7000 lb force. The loss of preload as a result of j j the 7000 lb applied push load was an accurate indication of the ) Plastic deformation encountered. The U-bolt leg forces before the e load was applied were 5393 lb and 6469 lb, when the load was removed the leg loads were 5060 lb and 6168 lb, respectively. The 300 lb l loss in leg force is an indication of very small plastic deformation. It is possible, under certain conditions, for progressive radial > distortions of the pipe wall to occur at pipe clamps. This would [ occur only if two conditions existed. , ( l. Maximum pipe stress intensity exceeded 3 S,and

• Slightly higher (13%) U-bolt leg tension applied than in the elastic l

analysis. No attempt made thru iterations to make them identical I i

since they are close in magnitude.

! l l l

   .      l

2. The clamp is periodically retorqued.

This is not a problem for the typical U-bolt type restraint because U-bolts are not periodically checked for preload and, therefore, not retorqued. A fatigue assessment was performed to determine the possibility  ; r of initiating a crack in the pressure boundary at the vicinity of the U-bolt. This assessment included a calculation'for Class 1 piping , and a calculation for Class 2 and 3 piping.~ The calculation for  ; Class 1 piping utilized the U-bolt stresses from the 4" SCH 160 case. I The calculation for Class 2 and 3 piping utilized the worst case stresses from the other three cases. For the Class 1 piping, two locations were chosen for the evaluation. These locations were the 10" accumulator injection line I and 3" normal charging line, assuming the U-bolt was located some-where between the first check valve and the reactor coolant loop nozzle. These two locations represent the most difficult section to qualify for fatigue, (due to extremely severe thermal transient loadings) of all the Class 1 auxiliary piping. In this evaluation, t several load cases were modified to conservatively include the clamp l induced stresses. l This evaluation included all design transients as well as the l i clamp induced stresses. For the charaing nnd accumulater line the usage factors were .04 and .06 respectively. Based on an ASME code allowable of 1.0 or .1 for pipe break requirements, the usage factor requirements are met. i i l

l For the Class 2 and 3 assessment, a simplified fatigue calculation  : i was made using the worst case data. Based on maximum stress,  ! i excluding preload (Which is not a cyclic event), the 32" main steam line was selected. Based on a simplified calculation, the incremental usage factor was less than .01. This assessment was based on a l representative Class 1 method. It provides a measure of the fatigue

damage that would be expected at the U-bolt location. This assessment i

indicates that the integrity of the pressure boundary, based on fatigue consideration,would not be significantly affected by the - localized U-bolt effects. From the above discussion, it can be concluded that the pipe f f stresses induced in the pipe due to cinching of the U-bolts will not exceed acceptable limits. Further, the loss of preload due to any yielding in the pipe material will be small as demonstrated by the test and plastic analysis performed. [ l . b ! L ( , l  ! L r e t l l

uI

                                       ~

O _ Y f o o y, y r s )d a (" 0 3 S S ev I 1 l M R n o

                                                                                 /

3 - T C g S g a i o n r 3

                      ,                                            s

{ S 6 f i

                   )

g , s n , 7 s i s' e r O ^)s'( )t.s

                                 "')r                                      S+

d ep*(%,p1 $ $W K - ( s o

       /r 6 A             b ' [7f'a 7
                                  =
                                                   /               5 bN GC                         l'          1

5. *,

i". - a 5 4 i.sJ2 2 2 l'. 2 i 7 4 5 L1 h s

                             ~

4 4 5 ob&$D i

l l I i l l SHEAR STRESS i MINOR PRINCIPAL MAJOR PRINCIPAL i STRESS EOUALS THE STRESS EQUALS THE LONGITUDINAL STRESS CIRCUMFERENTIAL  : STRESS

                                                ~

NORMAL STRESS t t STRESS INTENSITY: [ M0HR'S CIRCLE r s l FIGURE VII-2 l MINOR PRINCIPAL STRESS EQUALS THE LONGITUDINAL STR:

            .                                                                                                                                                                        I r

r

    -- -- -   , . _ , , , .     -,,.------n,      _--,---,--,,,,-.--,--n-                 -------,-,,,...,,,c- ,-- ,,_       - - . , _ , , _ , . ,

t

   .                      I I

VIII. STRESS

SUMMARY

TABLES l This section presents the stress results from the four load 4 [ cases defined in Section IV. j Terms used are defined below. - MAX Longitudinal Strass: The longitudinal direction of the pipe is parallel to the pipe centerline; for the cross i piece it is parallel to the long axis of the cross piece, and for the U-bolt is. parallel to the U-bolt centerline. The maximum longitudinal stress (at the inside or outside t surface) at the centerpoint of the analysis element is l tabulated. , MAX Circumferential Stress: The circumferential stress is { in a direction parallel to the pipe circumference; for the cross peice it is parallel to the shortest axis of  ! E the cross piece. The maximum circumferential stress I (at the inside or outside surface) is tabulated at the centerpoint of the analysis element. 1 MAX Stress Intensity: This is the maximum absolute difference between the major prinicpal stress or minor principal'  ; i stress and zero. The maximum surface stress (inside or l outside) at element center point is tabulated. Note, r l that because max surface stresses were tabulated for the 1 longitudinal and circumferential stresses the stress j intensity cannot in all cases be derived from the longi-

                                               .tudinal and circumferential stresses tabulated.

f l l i 1 i r

                                                                                                                                                   ?

i

    ,'            ;                                                                                                                                                                           l i
  -                                                                          4" Sch 160                                                                                                            !

STRESS

SUMMARY

TABLE ELEMENT DIAGRAM f l l l

                                       \        N \b N(   b\                                 NN Cross Piece Max Stress @ Element 1 i

619 663 699 73 6 77/ Go7 \ Top l N \ b 621 658 6'19' 73 0 76 6 801 ' i 418 454 490 72f 74 2. '??8 i l ( l t l l l I / d 6l3 66 685 22/ 73 7 793 )1ss- l 3  : I 409 fo%5 68I 7/7 753 784 1 Bottom i A Li-Bolt Leg Element is 81 - l NOTE: Not to scale, for relative location of stress summary elements only. f T SIDE VIEW OF PIPE  : i 1

j  ; 4" Sch 160 l

      ,                                                                 Preload                                                                     }

l Max Max Max l Analysis Longitudinal Stress l Element Stress lb./in2 StressCircum.2 lb./in Intensity lb./in2

                                                                                                                                                    ?

627 -13480 -26091 26091 j 663 -13646 - 24736 24736 j i 699 -10669 -20639 20644  ! l 735 -3879 a -13042 13047  ! 771 441 7770 7787 l 807 1021 4198 4201 622 2921 4 363 8396 l ! 658 2423 -7392 7674 694 1728 -5973 6573 j 730 852 -4006 5049 i 766 406 -1602 , 3411 f 802 -343 216 1486 618 8610 -14544 14552 l 654 7112 -12789 12834  ! 690 5405 -10828 10875 726 3645 -8923 - 8946 762 1461 4 510 6511 { 798 -985 3941 3947 [ 613 -10410 -14519 14535 [ 649 -1066 -5100 5833 l 685 -1967 -1824 4640 l 721 1533 1071 3886 l 757 741 730 2629 f 793 229 232 1422  ! 609 -9827 -12151 12153  ! l i 645- -3815 -7769 7776 i 681 -2165 -5590 5889 l 717 -1746 5756 6169 753 -1189 5476 5484 789 965 3810 3815 [ 1 (Cross Pc.) -34187 3838 34187 [ 81(U-Bolt) 27544

j.

                                                              ,.     ;                                   4" Sch 160 Preload + Thennal l

Max Max Max l Analysis Longitudinal Stress i Element Stress Ib./in2 StressCircum.2 lb./in Intensity lb./in2  ! 627 -19970 -39305 39305 i

            '663                -20250                   -37204            37204 699              -15779                   -30939            30947          :

j 735 -5510 -19338 19348 ~ 771 918 11174 1,1199 i 807 1578 5915 5919 622 4930 -13667 13706 f 658 4018 -12023 12373 694 2790 -9661 10426 l 730 -1325 -6454 7680 766 -638 ,

                                                          -2568             4934          ;

802 -590 339 2077 618 12011 -22022 22050 l 654 9965 -18753 18888 [ 690 7686 -15433 15564 726 5256 -12563 12626 , 762 2137 -9105 9108 798 -1369 -5507 5518 613' -14775 -20100 20179 l 649 -1141 -8065 8414 [ 685 -2999 -

                                                          -2946             6596          l 721                2351                     1840             5598 757                1198                      1236            3742 793                  400                       377            2020         !

609 -13728 -17332 17334 ( 645- ,

                                 -4498                   -10389            10396          !

681 -2393 -7170 7600

                           ~

717 -2207 7300 8142 753 -1770 7346 7356  ! 789 1336 5289 5297 [ l (Cross Pc.) -50841 5714 -50841 81 (U-8olt) 42320

 .,7.__
       *     *                                                   -70_                                                        '

4" Sch 160  ; Preload + Themal + Pressure l Max Max Max Analysis Longitudinal Circum. Stress Element Stress 1b./in 2 Stress 1b./in 2 Intensity 1b./in2 f l

                                                                                                                            \

l 627 , -20675 42905 42905 l 663 -21146 41286 41286 i 699 -16266 36879 36879' 735 -5526 28795 28795 . 771 968 20854 20854 l 807 1616 15399 15399 622 5075 15120 15120 658 4160 14510' 14510 694 2911 13637 13637 730 1396 12413 12413 i 766 -652 10792 10792 l 802 -

                                             -599                    9600                   10199                            [

618 12433 23645 23645  ! 654 10345 22256 22256 l 690 8005 20834 20834 i 726 5482 19210 - 19210 762 2235 16628 16628  ; 798 -1416 13580 13955  ;, 613 -15730 14803 15730 [ 649 -1130 11666 12796 ) 685 -3161 11296 14457 l 721 2455 11120 13273 y l 757 1243 10523 10523  ; 793 405 9642 9642 l 609 -14403 9228 14403 , I 645 4613 11944 11944 r l 681 -2379 15176 15176 , 717 -2214 16829 16829 753 -1809 16887 16887 [ 789 1382 14747 14747 j i 1 (Cross Pc.) 52255 5607 52255 81 (U-8olt) 43685 --- ---

               ~              _         . . . . _ . . . .        _ _ . .     - ---.                 . . - . -

l 4" Sch 160 Preload + Thermal + Pressure + Push Max Max Max i

      ~

Analysis Longitudinal Circum. Stress  ; Element Stress 1b./in2 Stress 1b./in2 Intensity 1b. fin 2 j i 627 , -26648 44786 44786 j 663 -27411 43125 43125 l 699 -22498 38753 38753 l 735 -10773 29820. 29820  ; 771 -2929 21973 24902 .  : 807 -3567 15274 18523 622 -6045 15419 15419 l 658 -5625 14790 14790 f 694 -5045 13895 14801 730 -4430 12596 15024 766 -3912 10904 14816 { 802 -4290 9603 13893  ! l "618 11641 24341 24341 j 6 54 9525 22929 22929 l 690 7132 21459 21459 [ 726 4531 19767 19767 5 78 2 1 62 13 613 15003 15043 15043 649 2745 11903 11903 685 5540 11471 11471  : 721 5055 11252 11252 i 757 3783 10551 10551 793 2874 9705 9705 609 12220 10173 12220 l' 645 . 3159 12824 12824 681 1879 15808 15808  ; 717 2913 17629 17629 753 3552 17004 17004  ! 789 5787 155f8 15598 1 (Cross Pc.) 52824 55S3 52824 81 (U-Bol t) . 41797 --- ---

l i

     ..   <           .                                                                                           l 4" Sch 160 Partial Preload + Thermal + Pressure + Push                                           .

i Max Max Max l Analysis Longitudinal Circum. Stress l Element Stress 1b./in2 ' Stress 1b./in2 Intensity lb./in2  ; 627 , -15591 24998 24998 663 -15566 23857 23857 f 699 -11032 21198 21198 I

                    -                                                                                            i' 735               -5380                   16877                       21435 i

771 -3314 14662 17631 l 807 -4393 11483 15876 l 622 -4202 11749 13791 i l 658 -4039 11447 13812  ! 694 -3779 11031 13819 l l l 730 -3541 10475 13854 1 766 -3856 9842 13698  ; 802 -3894 9354 13248 l 618 -3032 15649 15649  ! 654 -2782 15066 15066 f 690 -2528 14443 14443 I 726 -2236 13722 . 13722 .f 762 -1856 12574 12574 i i 798 -1570 11211 12609  ! 613 6003 10560 10560 , 649 2608 9671 9671 [ 685 3480 9876 9876 [ 721 3338 9922 9922 f 757 2912 9697 9697 793 2589 9449 9449 609 8078 - 12990 12990 f 645 , 3711 12786 12786 { 681 3350 13068 13068 i 717 3677 13487 13487  ! 753 3173 12528 12528 l 789 5172 12277 12277 { 1 (Cross Pc.) -20235 2702 20235 i 81 (U-8ol t) 14275 --- --- j

I

                                                                                                                            .-             ;                                                                                                                                                         l 10" Sch 40                                                        .

I STRESS

SUMMARY

TABLE ELEMENT DIAGRAM k t l t I

                                                                                                                                                                            -   i NN                            N N                                                   \'>-N                                            Cross Piece Max Stress 9 Element 107 Nb \

652

                                                          \

688 72N 76 O l Top i j

                                                                                                                                                               )                l 1

656 l 692. 72 8 74N < s= b i  : I  ! l l -

                                                                                                                                  ,                                           . l l                                                                                                                                                                 !
                                                                 -                                                                                              l               \

(c i

                ,                460                          6%                           732.                  76 8                                    s          90*

l 5 i j t i  ; i I / J f29 665 70/ 73 7 1 ' 5' l 9

                                                                                                                                                              \

I l

                                                                                                                                                              ]                 l 633                           66'l                         705                 7W                                           ?     Bottom      [

A  ! U-Bolt Leg Elenent is 3 .

                                                                                                                                                                                )

t ETE: Not to scale, for relative location of stress sumary elenents only. SIDE VIEW OF PIPE h i

10" Sch 40 Preload Max Max Max Analysis Longitudinal Stress  ; Element Stress Ib./in2 StressCircum.2 Ib./in Intensity Ib./in2 652 -25252 -48528 48529 688 -26833 -46576 46577 724 -20567 -37190 37300 760 -8764 25535 25913 656 8779 -14247 14280 ' 692 7692 -13375 13642 728 5930 -11909 12529 f 764 3754 -9904 10916 660 10263 -14157 14194  ! 696 9222 -12958 20700 l 732 7985 -11702 18916 , 768 6907 -10990 11310 l 629 -10193 -13672 13752 i 665 -1779 7056 7214  ! 701 -1434 4953 6393 I l 737 -1334 4999 . 5034 633 -6487 -8826 8826 669 -3209 -4414 4414 705 -2709 -3093 5445 ' 741 -3300 3955 5230 107(CrossPc.) -16933 160 - 17093 3 (U-Bolt) 12726 l l I i l , I

10" Sch 40 Preload + Thermal Max Max Max Analysis Longitudinal Stress Element Stress Ib./in2 StressCircum.2 Ib./in Intensity Ib./in2 652 -33218 -60636 60637 688 -34015 -57528 57595 724 -22554 -44359 44613 i 763 -9875 30795 31222 656 11344 . -18210 18250 i 692 9818 -17050 17374 728 7410 -15115 15862 764 4553 -12502 13707 660 11423 -17027 17390 696 10612 -15001 15740 732 9352 -13250 14109 768 8086 -12454 13086 629 -12267 -16841 16856 665 -2850 78036 8130 701 -2248 5208 7457 737 -1543 5738 5788 633 -10070 -13948 13948 669 -4804 -5884 5884 . 705 -4095 -3267 6911 741 -3434 4517 6291 107 (Cross Pc.) -20153 191 20344 i 3(U-Bolt) 15645 - i 4 e f l- [

   ~

e' 10" Sch 40 Preload + Thermal + Pressure Max Max Max Analysis Longitudinal Circum. Stress Element Stress Ib./in2 Stress Ib./in2 Intensity Ib./in2 652 -34425 58585 58589 688 -35642 56732 56911 724 -23452 49436 50182 760 -10214 -21518 40903 656 11943 24013 24037 692 10327 23023 23193 728 7781 21414 21692 764 4766 19256 19508 660 11402 21929 21930 696 10887 21546 21571 732 9680 21195 21222 768 8371 20680 20683 629 -12759 -8981 12770 665 -3117 l'624 0 13747 701 -2467 13643 16112 737 -1582 14316 14403 633 -10557 -6608 10588 669 -5255 7577 12833 705 -4377 11465 15844 741 -3501 13253 15094 107(CrossPc.) -20897 190 21087 3 (U-Bolt) 16346 e

                                                                                    --           .                                               . . = . . .
                ... .. ,. y -                               .-         -
                                                                                                                           --- c .. .
                  ,n                                                       -                                    -

y I k

        .             3[              '

10" Sch 40

                                  +

Preload + Thermal + Pressure + Push

                                               ,                 3                                                              -
                   \

( - l I

                    .s                                         Max                                 Max                                       Max Analysis !N~                     T Longitudinal                                                                                     Stress Element                               Stress lb./in2                                       Circum.2,.

Stress-lb./in , Intensity lb./in2 - 652 -48024 -73462 ,- 73462  ; 688 - '

                                                        -50094                                   70533                                    70737 l.

724 -35933 61070 N 62038 760 . s1 s -16968  ;: 48720 53286 i

                                                                         ,, ~                                                              .                               ;

656 s - 12386 28408 28468  ; g ' 692 10499 27111

                                                                                                                      ^

27466

                                    ~

728 7431 24987 25578 764 -5410 22154 22755  : 660 ;13686' 23408 23414 ) 696 12995 23154 23158

                                          '                                                                                                                                l 732              l                          ' ' 12033                                   22897 ,                            v 22913 768                                               11017             .                   22635                                    22702 629                                                8949
                                          .                                                      11275                                    11425                            !

665 1701 12442 12454  ; 701 2029 14188 - 14189 737 - 3438 14422 14424 l 633 7057 10646 10649 669 1062 12144 12144 705 ~ 638 s 14577 14578 74 1 2150 '15729 15703 ( 107 (Cross Pc.) -14242 449 , 14694 i 3 (U-Bolt) ,t 10026 l t

                                                                                                                                        %                                  l s     s                                  !

N s f s

• h i

[ I _ _ . - __:.__-~.____._.,.___.____.___________. . _ _ _ . _ _ _ . _ _ _ _ _ _ . _ _ _ . _ _

 .'                 .                     10" Sch 40                             ,
     ,                  Partial Preload + Thermal + Pressure + Push Max                     Max              Max Analysis           Longitudinal                                 Stress Element           Stress Ib./in2        StressCircum.2 Ib./in    Intensity Ib./in2 652              -34782                   52415             52418 688              -35733                   50846             50977 724              -25655                   44150             44756 760              -12735                  35559              40018 656                 7365                  21878             21920
                                                   ~

692 6115 20977 21215 728 -4931 19501 19904 764 -4330 17547 17975 660 8923 17762 17775 696 8671 17783 17806 732 8305 17828 17875 768 7845 17873 17965 l 629 6046 12770 12770 665 3090 12121 12125 701 3125 12165 12171 73 7 3486 11936 11942 633 4171 12251 12251 669 2376 12874 12874 705 2528 13776 13776 741 2952 14278 14278 . 107 (Cross Pc.) -5953 340 6300 3(U-Bolt) 3269 l l l l t l l i l

10" Sch 80 STRESS

SUMMARY

TABLE ELEMENT DIAGRAM I i

                                         \            (          Cross Piece Max Stress 9 Element 7         f bh        W82.

h 5/8 N364 390 fo2(o \ Top .\ i L ) D z/78 5/4 550 586 421 ' 45- l r J I  ;

     /                                                                                                   -

l, (> . l - c  :

         ,        47Y       5/O               5%        682.          4/8                      j     so-S                                                                                         t            i i

, l  : l $ /  ! i 969 -5a5 SYI 577 4G 1135' l 4 3  ! l

                                                                                                )            i V65        Sol              537        6 73         (,09                     ) Bottom      i A                                                                                              '

U-solt Leg Element is 861  ! i NOTE: Not to scale, for relative location of stress sumary elements only. l I SIDE VIEW OF PIPE

                                                                ~

10" Sch 80 Preload  ; Max Max Max  : Analysis Longitudinal Stress Element Stress lb./in2 StressCircum.2 lb./in Intensity Ib./in2 l t 482 -16270 -29008 29880  : 518 -17270 -29510 29520 f 554 -16130 -27490 27530  ! 590 -8291 -18660 18710 . 626 -2094 11030 11100  ! 478 4976 -9296 9315 ' 514 4594 -4848 9025 + 550 4012 -8165 8567 l, 586 2632 -4534 7310  : 622 -800 -3751 5302 t 474- 7611 -11110 11120  ; 510 6403 -10030 10130 l 546 5751 -9041' 9242 582 4793 -7896 8127 618 3257 -6633 6757 f 469 -8840 -10900 - 11000 505 -2808 -6819 6900 l 541 -1091 -4377 4488 [ l 577 -923 -2650 2834 l ! 613 50 2249 2454  ! t < l 465 -4324 -7122 7173 , 501 -2143 -4749 4751 l 537 -1912 -3272 3274 573 -1481 -2634 2639  ! 1 609 -2194 3283 3487 I l 7.(CrossPc.) '

                                        -4108                     -14310                                    -18490 861 (U-Bolt)              17008 l                                                                                                                             l l

l . i I

10" Sch 80 I

          '                     \

Preload + Thermal Max Max Max i Analysis Longitudinal Stress  ! Element Stress Ib./in2 StressCircum.2 Ib./in Intensity lb./in2

                                                                                                                                                           )

482 -19410 -34890 34890 518 -19330 -33550 33560 554 -17850 -30970 31000 [ 590 -9655 -21460 21530 626 - -2258 12650 12730 [ 478 6040 -11390 11410 l 514 5493 -10750 10950 j 550 4720 -9848 10290  ! 586 3020 -7808 8657 622 -964 -4412 6069 f 474 , 7230 -11570 11600 i l 510 6933 -10730 10900 I 546 6555 -9801 10080 , 582 5545 -4658 8960 f 618 3766 -7398 7548 l 469 -10100 -12380 12390 l 505 -2839 -7737 7780 i 541 -1522 -4775 4863 I 577 -1255 -2791 2979 f 613 9 2446 2683 I 465 -8426 -10240 10240 I 501 -2424 -6388 6388 537 -2636 -4127 4128 l 573 -2034 -3087 4039 [ 609 -2392 3809 4117 [ 7,(CrrassPc.) -4847 -15850 20730  ! 861 (t)-Bolt) 20175 1 I I

 ~ _ _ _ _ - - . _ _ _ _ _ _ - . _ . , _ _ . .                  . . . _ . _ _ _ _            __      _ . ____.. _ _ _ _ _                     __ _ _ _ _ J

                                                 ,.    ,                                 10" Sch 80 Preload + Thermal + Pressure Max                    Max              Max Analysis          Longitudinal               Circum.          Stress Element          Stress Ib./in2       Stress 1b./in2     Intensity lb./in2 482               -19940                  32960            .32960 518               -19960                  32210            32230 554               -18460                  30720            30830 590                -9937                  25260            25500 626                -2309                   18120            18180 478                  6249                 12990            13000 514                  5684                  12640            12680 550                  4884                  12140           12310 586                  3123                  10900            11000
,      622                  -995                   8604             8652 474                  7329                  13620           13620 510                  7084                 13410            13410 546                  6726                 13240             13240 582                 5698                  12820            12820 618                 3871                   11880            11880 469               -10390                  -7606             10400 505                -2885                    5875             5876 541                -1604                    6746             6758        -

577 -1319 7611 9009 613 8 7615 7700 465 -8832 -5672 8832 501 -2471 3636 af373 537 -2778 5070 5071 573 -2147 7125 7126 609 -2432 9001 9001 7*(CrossPc.) -5006 -16200 21250 861(U-Bolt) 20803

10 Sch 80 Preload + Thermal + Pressure + Push Max Max Max Analysis Longitudinal Stress Element Stress lb./in2. StressCircum.2 lb./in Intensity Ib./in2 I 482 -30220 -44380 44380 518 -31010 -43430 43430 f 554 -29230 41070 41310  ; 590 . 16560 32650 32930  ! 626 -5343 22890 25022  ! 478 -5915 15580 I 15600 514 -5665 15100 15210 550 -5284 14440 14640 586 -4488 12770 13030 622 -3248 9659 11430 474 8729 16040 16060 510 8344 15840 15860 f 546 7844 15650 15670 582 __6670 15190 15240 618 4559 14120 14240 i 469 8234 7742 8249 [ 505 2592 7536 7542 l 541 767 8035 9401 l 577 1299 8435 8447 613 1464 8000 8050 465 5638 6329 6329 f 501 1708 6783 6790 l 537 464 7794 7805 r 573 1062 9521 9527 609 1833 10600 10710 7.(CrossPc.) 4475 -14420 18955 , 861(U-Bol't) 17388 I i I

__ _ _ _ _ _ _ _ .m__. ._m... . . _ _ .

              .                   ,                                            10 Sch 80                                                                    ,

Partial Preload + Thermal + Pressure + Pusn > Max Max Max Analysis Longitudinal Circum. Stress  : Element Stress Ib./in2 Stress Ib./in2 Intensity lb./in2 , L 482 -14810 21520 21530 518 -14960 21620 21630 554 . -14370 20950 20980 l 590 -9178 17110 17140 626 -3606 12800 15012 l 478 -3821 9266 -9285 i 514 -3689 9056 9111 550 -3482 8781 9082  ! 586 -3105 8071 9269  ; 622 -2455 6777 8995 474 2901 9358 9382 510 2821 9354 9385 546 2701 9322 9378 1 582 2379 9241 9324 618 1684 8978 9096 469 2114 5751 5751 505 1717 5889 5893 l 541 1505 5954 5957 l 577 1479 5967 5970  ; 613 1390 5810 . 5811 465 3122 - 7825 7827 501 2576 7970 7975 537 2328 8071 8080 ( 573 2440 8410 8425 609 2056 8076 8223 l 7-(CrossPc.) ,

                                         -518                                             -2546                                   3139 861(U-Bolt)                  1294                                                                                                              ,

l

t 32" MAIN STEAM ST.1ESS'

SUMMARY

TABLE ELEMENT DIAGRAM l

                                                                                                                           <                        Cross Piece Max Stress @ Element 60 s                            588            (ol9                                                            foCS                  69/            \          Top
                                                                                                                                                                                   \

579 65 45/ 687 ' 45* i

                              !                            5 75            4//                                                                  49'7              4 83              l 90*

570 406 4% 4 78 h 13

                           )   '
                                                                                                                                                                                     }i l

566 602. 638 I 6 7 +< ' Botta a ,

                                                     \

U-Bolt Leg Element is 133

                            ,                            NOTE: Not to scale, for relative location of stress suninary elements onh SIDE VIEW OF PIPE

{

                                                                                                                                                                             .                   l

32" Main Stream

           'i                                 Preload Max                    Max                  Max Analysis          Longitudinal                                 Stress Element          Stress Ib./in2     StressCircum.2 Ib./in       Intensity Ib./in2 e

583 -2966 -4859 4859 619 -3412 -4918 4921 655 -2568 -4116 41'40 691 -1092 -2717 2733

                                                    -1147 579                668                                        1148 615                633                -1113                   1126 651                548                -1028                   1066 687         -

390 -863 947 575 1001 -1154 '1154 611 989 -1155 1156 647 ~957 -1160 , 1161 683 891 -1166 1168 570 -1170 -1445 1455 606 -650 -1134 _1148 642 -209 ' #758 761 ] 678 -120 -419 . 697 566 -1074 -1208 1209 602 -702 -1009 1011 638 -492 -827 830 674 -432 -691 713 60(CrossPc.) 1242 1255 133(U-Bolt) 1017 s 1 4 e h e

       -    . = . = . = -      ..    .
    .'                                           32" Main Steam Preload + Thennal Max                    Max                    Max Analysis           Longitudinal                                        Stress Element         _

Stress Ib./in2 StressCircum.2 1b./in Intensity lb./in2 583 -12786 -21306 21305 . 6i9 -14810 -21570 21582 655 -10966 -17917 18030 691 -4249 -11542 11618 579 3664 -5991 5997 615 3487 -5833 5880 651 3050 -5439 5575 687 2227 -4660 4957 575 3469 -5418 5425 611 3420 -5235 5291 647 3295 ,

                                                           -4865                    4998 683                   3030                   -4348                    4545 570                  -4082                   -4942                    4945 606                  -2508                   -4030                   .4042 642                   -996                   -2841                    2870 678                    -592                   -1733                    1792 566                   -3609                   -4397                    4397 602                   -2117                   -3518                    3519 638                    -867                   -2459                    2459 674                   -1150                   -1583                    1710 60 (Cross Pc.)       3679                   6092                     6178 133 (U-Bolt)           4619 e

          .                                   32" Main Steam Preload + Thermal + Pressure max                                                           MAX                                 v1X Analysis            Longitudinal                                              Circum.                          States Element           Stress ib./in2                Stress 1b./in2                                             Intensity lb./in2 P

583 -17149 33973' 33960 619 -19883 34387' 34411 655 -14676 32016 32296 , 691 -5588 27716 27887  ! 579 5078 19409 19410 615 4830 19290 19299 I 651 4218 18981 19003 687 3066 18337 18371 575 3934 18018 18019 ( 611 3971 17973 17973 647 4030 17993 17993  ! 683 3901 18028' 18028 , 570 -5105 13159 13159 606 -3058 13166r 13166 642 -1101 13599 13684 l 678 -851 14439 15290 5 66 -4972 12830 12830 602 -2959 12834 12834  ! 638 1187 13300 14160 674 -1526 14264 15790 l 60 (Cross Pc.) 3679 7979 8076 133(U-Bolt) 6251 -- , l l l

              .                                                                                                                                                                l

                                                                                        .'              '.                                                           32" Main Steam Preload + Thermal + Pressure + Push Max                 Max             Max Analysis                                         Longitudinal                           Stress Element                                       Stress Ib./in2     StressCircum.2 Ib./in    Intensity Ib./in2 583                                            -26304              46413            46413 619                                            -31915             47170             47223 655                                            -22391             43296             43916 691                                             -7224             36269,            36633 579                                              6752             22707             22713 615                                              6383             22517             22559 651                                              5477             22021             22129 687                                              3788             20994             21176 575                                              3255             20452             20453 611                                              3218             20454             20456 647                                               3116             20473             20478 683                                              2892             20530             20544 570                                              2291             15172             15172 606                                              1594             15184,            15186 642                                               954             15344             15346 678                                                655             15604.            15606 566                                               3497             16896             16896 602                                              2861             16907             16907 638                                               2341             17091             17091 674                                               2182             17471             17471         .

60(CrossPc.) 8924 3865 8926 133 (U-Bolt) 2048 I

APPENDIX I e e

1'CO FOR SGT AND 12 > - O.'y nm, m.....___ L,es_ s,_ BLUELINE roxt4Ruf JJgg [6cs.uo**A", O*SENTE'[b * - g L g latS a o,E.AnoNS SETSaeic sway stairr AssEsuit.Y costsisTIsso Ors cese:

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