ML20077C392
| ML20077C392 | |
| Person / Time | |
|---|---|
| Site: | Dresden  |
| Issue date: | 06/18/1983 |
| From: | Azzazy S, Mcfarland D, Olson D SARGENT & LUNDY, INC. |
| To: | |
| Shared Package | |
| ML17194B657 | List: |
| References | |
| EMD-043846, EMD-43846, NUDOCS 8307260028 | |
| Download: ML20077C392 (39) | |
Text
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SARGENT & LUNDY EN GIN EE R S CHICAGO O
1 DRESDEN - 2 MAIN STEAM MONITORING PROCEDURE TEST-ANALYSIS CORRELATION f
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CECO., 5486-23 Report No.: EMD-043846 June 18, 1983, Rev. 00 Page 1 6f 39 O
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Issue Summary Report No.: EMD-0 4 3 84 6, Rev. 00 File No.: EMD-04 384 6 Rev.
Date Signatures 00 6/18/83 Prepared:
$ C-Date: [-Jo -f 3 S.
E. Azzazy 7/f'[l Date:[!20
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D. M. McFarland
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ate: / M/3 w-r D. E. Olson Reviewed: b
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6 /h /93 M. H. Bhatti 6
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Date:
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!$ Mahendranathan*
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Approved:
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[hT[Kitz#
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- Reviewed only Section 3.1 portions concerned with finite element analysis i
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Table of Contents i
i Page 1.0 Introduction 4
2.0 Piping Thermal Expansion Response 5
5 2.1 Field-Measured Responses 9
2.2 Design-Predicted Responses 2.3 Field-Measurements vs. Design-Predictions 13 16 3.0 SRV Discharge Loads I
3.1 Measured Loadings 17 l
I 3.2 Test-Analysis Correlation.......................
21 I
23 4.0 Conclusions..........................................
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26 5.0 References...........................................
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!1,0 ~ Introduction'
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The purpose of the Dresden-2 main steam monitoring'provadure
-SP 83-4-52 (Reference 1) was to. determine the cause of.five main steam line snubber failures.
The procedure specified
-visual and instrumented monitoring of the inside-containment main steam system.
Thermal expansion and dynamic ' transient events were monitored.
A summary of the acquired data-is
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provided in Reference 2.
Analysis of the data revealed no measured response sufficient to cause snubber: failure.
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- The availability of thermal expansion and SRV discharge dynamic r
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transient data prompted interest in two additional: issues.
The e
first of these concerned proving the measured. piping responses to be'within Code'allowables; the second involved ~ demonstrating the correlation between the measured responses and the analytically-i The analyses of Reference 4 and 8 conservatively predicted values.
i demonstrated that'the measured system responses are within Code allowables for both thermal expansion and dynamic transient events.
The purpose of this document is to demonstrate the. correlation
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l between the field-measured system responses and the analytically-
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predicted responses.
The correlation is illustrated for both jl piping thermal expansion and SRV discharge dynamic transient events.
Section.2.0 and 3.0 of this docune'ht demonstrate the test-analysis
-correlations for. thermal expansion and SRV dynamic transient, F~g 4
respectively.
Section 4.0 provides a summary of tte methods and conclusions of.this report.
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^O tv 2.0 piping Thermal Expansion Response There is good agreement between the analytically-predicted and the field-measured main steam piping -thermal expansion responses.
This agreement is demonstrated in the three following subsections.
~A side-by-side comparison of field vs. design displacements is shown in Table 3.
In subsection 2.1, the field data acquisiti5n nothods of Reference 1
are. reviewed, and that subset of data most accurately and completely describing the piping thermal expansion response is determined.
An estimate of data acquisition error is also developed.
pd Subsection 2.2 addresses the piping thermal' expansion analytical design calculations.
Conservativ'e'and simplifying assumptions af fecting design calcuidtiion accuracy are evaluated, and quantitEtive estimates are
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derived.
Finally, in 2.3, the design-predicted and field-measured piping displacements are compared; their differences are analyzed in light of the design'aEsumptions and errer ;
estimates addressed in the previous sections.
2.1 Field-Measured Responses The main steam piping displacements were measured by a.valks down visual inspection of support settings, and by hardwired instruments mounted on the piping and supports.
These two.
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data' acquisition methods are described below.
It is concluded that the walkdown visual inspection method provides the most reliably accurate thermal displacement data.
2.1.1 Walkdown Data Description.
- The walkdown visual inspection method is a simple, accurate way to measure piping thermal expansion response at.many points throughout a piping system.
The data is obtained by recording the cold and the hot support positiori settings of a number of supports on a piping system (Figure 1 illustrates typical load / position scales on the main steam supports).
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The piping thermal response to the cold-to-hot transition is determined by taking the difference of the recorded cold and hot support position settings.
During the Dresden walkdown, the cold walkdown was done with the piping at 105'F; the hot walkdown was done when the piping was at 450*F.
The difference between the cold and hot settings is equal to the thermal expansion i
displacement of the piping at the location and in the r
direction of the inspected support.
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Accuracy. - The accuracy of the visual inspection method is estimated by. analyzing the support position scales of the Dresden main steam lines.- The two types
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CARCENT O LUNDY ENOINEERO Drcedsn-2 cmcco EMD-043846 Ceco, 5486-23 June 18, 1983, Rev. 00 Page 7 Of 39 l'
>,v are shown in Figure 1; the first of these is the PSA-10 snubber position scale which has gradations every half inch; the second is a typical spring hanger scale which has gradations every quarter inch.
The snubber scales were read to quarter-inch accuracies during the walkdown; the hanger scales were read to quarter-gradation accuracies, corresponding to sixteenth-inch accuracies.
Since each scale had to be read twice to derive the resultant thermal displacement at the support, the reading accuracies are multiplied by two to derive the displacement accuracies.
There-r~T
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fore, the snubber displacement accuracy is one-half inch; the spring hanger accuracy is one-eighth inch.
Summary.
- The visual inspection method provides the thermal expansion response of the main steam headers at each support location.
For the Dresden main steam headers, there are seven to nine supports for each of the four lines.
The data accuracy of this method is approximately inch for snubbers, 1/8 inch for hangers.
2.1.2 LVDT Data Description. - Linear variable differential transformers (LVDT) were placed at various points on the main steam esb
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system to allow-remote monitoring of piping thermal expansion displacements.
The LVDT in its two typical mounting configurations is shown in Figure 2.
Eleven LVDT's were used:
five of 'he eleven were mounted along the five previously-failed snubbers.
At these l
five locations there is a direct comparison between J
4 RLVDT and walkdown thermal expansion data.
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Conservatism. - Analysis of the LVDT data showed it to be consistently and significantly conservative, in terms of recording larger magnitude displacements, relative to the walkdown data.
Table 1 shows the
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comparison of walkdown and LVDT results for the five measured snubber displacements.
The reason for this conservatism is currently being investigated.
Summary. - The LVDT hardwired instrumentation recorded conservative results.
The results were used in Reference 4 calculations which conservatively demon-strated that the piping thermal expansion responses were within Code allowables.
However, for the purposes of accurate system. response description, the LVDT results are considered to be inferior to the walkdown data.
2.1.3 Summary The piping thermal expansion responses of the main steam headers were monitored by hardwired instruments
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and by the walkdown and visual inspection of piping support settings.
The hardwired LVDT displacement data-is conservative relative to the walkdown data.
It was used to conservatively prove Code allowable behavior of the piping system, but it does not provide the most reliably accurate description of the piping thermal expansion response.
The reasons for this consistently conservative deviation from the walkdown results is currently being investigated.
The hard -
wired data has the further disadvantage of having only eleven monitored displacements rather than the more
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_than thirty recorded walkdown displacements.
The walkdown visual inspection data is therefore used for the purposes of this report to represent the piping thermal expansion response.
The accuracy of the data is estimated to be inch for snubbers and 1/8 inch for hanger support data.
The agreement between the walkdown and the analytical predictions that is addressed in Section 2.3 serves to further validate the use of the walkdown data.
2.2 Design Predicted Responses Reference 3 contains the design calculations for the inside-containment main steam system piping.
The calculations O.
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incorporate simplifying, conservative assumptions to avoid unnecessary. analysis complexity and cost.
Two important examples'offthese assumptions are analyzed to estimate' their
.effect on thermal. expansion-design accuracies.
The first
. assumption concerns the use of conservatively high RPV nozzle
. thermal expansion movements.
The second assumption concerns support variability.
The effects of these assumptions on design calculation accuracy are addressed in the following subsections.
2.2.1 RPV Nozzle Movements The thermal' expansion movements of the RPV at the main
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' steam line nozzles largelyfdetermines the behavior of the piping system.
For Dresden-2, the RPV nozzle move-ments are very.close to.those calculated
- using an equation supplied by General Electric.
GE states that their equation..
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(Reference 7) overpredicts nozzle movements.
Using overpredicted vertical nozzle movements causes the design' calculation to predict displacements too large in the positive vertical direction.
The effect is.particularly-pronounced on the piping closest to the RPV nozzle connection.
The.overpredicted vertical
- stretching'~of'the system causes a corresponding design under-prediction of horizontal displacements.
This " stretching" also results in a conservative design overprediction of piping stresses.
'~* See. Reference 5-
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GI An analysis using RPV movements chosen to simulate walk-down spring hanger displacements is in Reference 5.
The results of'the analysis are contrasted with design: move-ments in Table 2; the~walkdown data for the system is also included.
2.2.2 Support Variability The force exerted on piping by a spring hanger is a function of the hanger spring constant, its current displacement from equilibrium, and internal friction.
A parameter called variability is used to quantify the effects of these hanger qualities.
It is defined as the ratio of the required applied force (to displace the hanger spring a standard dictance) to the designed weight load.
An ideal, constant-applied force hanger has zero variability; a rigid restraint has a near infinite variability.
Design calculations conservatively assume that spring hangers have zero variability; i.e.,
that they are ideal constant-force hangers.
This idealization is based on the assumption that the effect of variability is negligible if it does not exceed ten percent; maintaining less than 10% variability is thus a major consideration in the selection of spring hangers, n
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By. neglecting the variability effect the thermal analysis is conservative.
This is because hanger variability tends to oppose movement; hence, the simplifying design assumption of zero variability causes the design to overpredict thermal expansion displacements, and therefore to overpredict forces, moments and stresses.
To assess the variability effect on the main steam piping system, a sensitivity analysis (Reference 5) was performed on main steam line C.
The effects of
()
10% variability were simulated by applying forces at the spring hanger locations.
The study demonstrated that " actual" piping thermal movements are less than design (zero variability) values.
Simulation of 10%
variability changed the predicted displacements by as much as 3/16".
2.2.3 Summary The thermal expansion design analyses incorporate two conservative assumptions that adversely affect the displacement prediction accuracy.
The effect of these two assumptions will be used to explain the difference between field-measured and design-predicted piping
/'N displacements in Section 2.3.
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The design RPV nozzle movements are conservatively large, resulting in design displacement predictions too.large in theup' direction.
A further result of this overpredicted vertical " stretching" of the piping is a corresponding design underprediction of horizontal displacements.
These effects are demonstrated in Table 2.
The design prediction error is particularly evident close to the RPV nozzle.
The assumption of zero variability spring hangers ignores the hangers' general resistance to displace-ment.
Thus, design calculations tend to overpredict O
displacements.
This effect is most pronounced for the piping furthest from the RPV enforced nozzle displacement.
2.3 Field-Measurements.vs. Design-Predictions A side-by-side comparison of field-measured and design-predicted therma] displacements is provided in Table 3.
The field measured displacements are derived from the Reference 1 walkdown visual inspection data.
The walkdown cold-to-hot temperature differential was 345*F.
The data accuracies are estimated at i 1/2 inch for the snubber displacements, and i 1/8 inch for the hanger displacements.
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The design-predicted displacements are derived from the Reference 3 piping stress analyses.
The displacements in the analyses correspond to a 480*F temperature dif"ferential; they were linearly scaled down to the walkdown 345'F differ-ential before being entered into the table.
Most of the main steam header support displacements are included in the comparison table.
The displacements are listed in groups of four, with each group indentified by its common support location on the four header lines.
The first group corresponds to the spring hangers closest to the RPV nozzle connections; the last group corresponds to the spring hangers closest to the containment penetrations.
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Thus, the first listed displacement groups are the most affected by the RPV noz-le movements; the final displacement groups are more affected by support variability.
The RPV riser spring hanger field data, shown in Part 1 of Table 3, consistantly underpredicts design displacements by approximately one-quarter inch (in the vertical direction).
Note that the field data variance within each of the two displacement groups is less than the estimated data accuracy of the spring hangers.
Thus the " test-analysis" correlation for Part 1 of the table is shown to be strong.
The Part 2 horizontal snubber displacement comparisons show that the design-predicted displacements are consistantly
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less than the field-measured displacements.
This design-underprediction of horizontal displacements is another
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Note that the difference between field data and design pre-dicted snubber displacements is in every case less than the estimated snubber data accuracy of one-half inch.
The Part 3 containment riser spring hangers demonstrate the effects of spring hanger variability.
The measured movements are consitantly less in magnitude than their corresponding design predictions.
The effects of RPV nozzle movement overprediction are not evident for these displacements, just as the effects of spring hanger variability are not apparent in the RPV riser displacements.
Thus it has been demonstrated that there is a strong correl-ation between the design-predicted and field-measured displace-ments when the design-conservative assumptions are considered.
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i 3.0 SRV Discharge Loads Loads in the Main Steam Safety Relief Valve (SRV) discharge piping are experienced immediately after the opening of the SRV valve.
These loads result from the rapid pressure transient in the piping.
The loads act primarily on the SRV
' discharge piping.
Secondary effects are also experienced by the main steam header piping.
Loads on the main steam header piping mainly result from the deflection of the dis-charge piping during the valve opening pressure transient.
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The deflection of the discharge-piping will cause a reaction m
force in the header piping.
These secondary forces'are typically small and do not affect the design basis of the header piping.- The Dresden-2 design basis analyses also predicted small loads in the MS header piping.
The instrumentation located on the header piping for the 1
main steam monitoring procedure was intended to detect and quantify loading of the magnitude that would be capable of damaging the installed snubbers.
A monitoring program cap-able of verifying the accuracy of the design basis SRV dis-charge load analysis would req'uire additional transducers with different calibration ranges than those used, different s
' types of transducers would be required, and different loca-tions would be monitored.
However, the information obtained
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3.0 SRV Discharge Loads (Cont'd) during the monitoring program can be used to make an approx-imate assessment of the secondary loading effects on the MS header piping.
The responses measured during the monitoring program confirm that the header piping does not experience significant loadings as a result of SRV discharge.
3.1 Measured SRV Discharge Loadings Strain gauges were located on the cylinder end plugs of seven snubbers located on the main steam header piping.
The gauge output was correlated to snubber tension and compression loadings by means of static calibrations that were performed for each of the snubbers.
These instrumented snubbers allowed both tension and compression loadings, resulting from movement of the main steam header piping, to be monitored.
However, for purposes of correlating the I
test data and analysis results, only the tension loadings are used.
The tension loads are considered i
to be most representative of the actual system response.
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3.1 Measured SRV Discharge Loadings (Cont'd)
The compressive loads, as determined from the cali-bration curves, are considered to be overpredictions of the actual snubber force.
(Note that for confirm-ing that the piping remained within Code limits, the larger of the tension and compression loads were used.)
The detailed logic for considering the tension loads to be most representative of the actual system response is included in a separate report, Reference 6.
.The major points made in Reference 6 are: the stress dis ribution in the cylinder end plug makesit a less reliable force transducer for compres-()
sion than for tension; the analytically predicted relationship between piping tension and compression loads disagrees with the relationship determined using the calibration curves; and the measured piping dis-placements indicate that the calibration curves over-predict the actual loading.
These points are summarized below.
Cylinder End Plug Stress Distribution - A finite ele-ment model was made of the cylinder end plug to deter-mine the stress distribution in the vicinity of the strain gauges under tension and compression loadings.
Figures 3,
4 and 5 illustrate the geometry of the gS cylinder end plug and its corresponding finite ele-
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ment model. Figure 6 illustrates the stress intensity
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Cylinder End Plug Stress Distribution (Cont'd) distribution in the end plug under tension. loading.
Figure 7 illustrates the stress intensity distribution in the end plug under compression loading.
Under tension loadings, the stress distribution in the vicinity of the strain gauges is uniform with a relatively constant stress magnitude.
However, under
' compression loading the stress distribution has a high stress gradient.
As discussed in Reference 6, the stress distributions under tension and compression loadings are responsible
()
for the variance experienced in the tension and'com-pression calibration curves of the various snubbers.
The non-uniform rapidly changing stress distribution
- under a compression load suggests that this type of instrumented arrangement would not be very reliable as
-a compression force transducer.
The uniform stress distribution under a tension load indicates that the arrangement would be a reliable tension force trans-ducer.
Analytical Predictions - The analytically predicted SRV discharge loadings indicate that the tension and compression loadings should be of. equal magnitudes.
Using the calibration curves, the larger of the es b-measured compressive loadings are indicated'as being over twice the magnitude of the corresponding tension
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EMD-43846 Ceco, 5486-23 SARGENT O LUNDY June 18, l!
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7T loadings.
Therefore, the analytical predictions also
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suggest that.the compression calibration curves are not reliable.
Measured Piping Displacement - The measured piping dis-placements are the most convincing evidence that the compression calibration curves overpredict the-actual-compressive loadings.
The measured piping displace-ments are of equal magnitudes
- in the positive and negative directions.
Since the snubber loadings result directly from the pipe motion, the relationship between the positive and negative displacements must be reflected in the snubber loadings.
In other words, for the com-pressive loadings to be larger than the tension loadings the negative piping displacements would have to be larger than the positive displacements.. Since the dis-placements were equal in both directions, the compres-sion loadings as indicated by the calibration curves are considered to be unreliable.
In summary, tension loads in the cylinder end plug re-sult in a uniform stress pattern which will result in a i.
. reliable calibration curve.
The stress pattern for the compression. load, in.the vicinity of the strain gauge, is not uniform and the resulting calibration curves are not considered to be reliable.
Therefore, for the pur-
. poses of correlating test and analysis results, only
()
-the tension calibration curves are used.
The conservatism of LVDT's should not affect the ratio of their measured plus and minus transient displacement magnitudes.
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3.2 SRV Discharge Test - Analysis Correlation The SRV discharge design' loads are compared to the measured loads in Table 4.
The design loads.at 940 psig are less than the design basis loads.
The design loads at 940 psig. represent values that were reduced from design basis loads to facilitate direct compari-son with the measured load.
The design loads-typically bound or approximate the measured tension loads.
The design load is slightly exceeded at two snubber locations, corresponding to snubber number 50 and 52.
The maximum exceedance is 1100 lbs.
Note that snubber number 50 and 51 are
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paired at the same location on the piping, as are snubber numbers 53 and 52.
At snubber pair locations a resultant load is divided amoung the two snubbers.
The resultant design load at the location of snubbers.
50 and 51 is larger than the resultant measured load.
The resultant-design load at the location of snubbers 53 and 52 is exceeded by approximately ECO lbs.
This is not considered to be significant.
The design basis-analysis of the main steam header t
piping predicted that piping predicted that SRV opening would result in relatively small loads on the main steam header piping.
The results of the monitoring program O,_
indicate that small loads were in fact experienced during the SRV discharge.. Both the design and measured load
2 ARGENT C2 LUNDY Drocdsn-2 rNolNEEDS EMD-43846 c"'c^c Ceco,.5486-23 June 18, 1983, Rev. 00 Page 22 of 39 are small and they have no significant effect on the design basis of the header piping.
The snubbers were installed on the header piping to accommodate the postulated seismic loadings.
Both the design basis and measured SRV discharge loads are too small in magnitude to affect the design basis of the MS header piping.
In summary, the design basis analysis predicted that the SRV discharge loadings on the MS header piping would be small.
The measured tension loads resulting from SRV discharge were generally smaller than the predicted loads.
There was only one exceedance of
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predicted loads, and this exceedance was not signifi-cant.
The measured SRV discharge loadings confirmed the analytically predicted small loadings.
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4.0 Conclusion The purpose of this report is to demonstrate the correlation between the field measurements of SP 83-4-52 and the i
analytical predictions of Dresden~2 main steam system response to thermal expansion and SRV discharge dynamic l
transient events.
Thermal expansion. -Good agreement is demonstrated between the thermal expansion field data and the analysis predic-tions.
The walkdown visual inspection data of support I
displacement is shown to provide the most comprehensive and reliably accurate description of system thermal expansion response.
The design basis calculations are
()
shown to include two simplifying and conservative assumptions
(
which adversely affect displacement prediction accuracy.
The first conservatism is the use of GE's conservative reactor pressure vessel thermal expansion equation; GE states that the equation yields thermal expansion predic-tions in excess of expected values.
Thus, design analyses I
conservatively over-predict vertical movement of the piping f
near its RPV nozzle connection.
Likewise, this vertical
" stretching" of the piping causes the design calculations to underpredict horizontal displacement.
The second con-L servatism regards hanger variability; design thermal ex-pansion calculations disregard the effects of spring l
drs.
hangers.
This simplification is shown to result in design calculations that conservatively overpredict piping thermal expansion displacement.
I
CAR 2 ENTO LUNDY-E N GIN E E RO lDresden-2 EMD- 04 384 6 c
Ceco,.5486.-23 June 18,.1983, Rev. :00 Page 24 of 39 The comparison of field-measured'and design-predicted thermal expansion displacements is provided in Table 3.
The effects of the too-high RPV thermal movements and the support-variability on theLdesign predictions consistently explain ~
the variance between the two' sets of displacements.
In this manner, the strong' correlation existing between empirical and design displacements is established.
SRV-Discharge Hydraulic Transient -
A correlation between design-predicted main' steam header response and field measured-response is demonstrated.
The correlation centers upon the-forces experienced by seven instrumented snubbers restraining the main' steam header piping.
Design-predictions and field measurements agree that the forces. experienced by.
the= main steam header snubbers are small, their maximums not exceeding two thousand pounds.
The field-measured snubber forces are reviewed.
It-is deter-mined that, of the measured tension and compression loads, the latter are not reliably' accurate.
This is demonstrated in Reference 6.
A summary of the Reference'6 analyses is as follows:.the stress distribution in the' snubber cylinder end plug (Figure 3) upon which the force strain gauges are. mounted make it a less reliable force transducer for compression than
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for tension; secondly, the analytically predicted relationship
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Page 25 of 39 between piping tension and compression loads disagree with the relationship derived using the calibration curves; finally, the measured piping displacements indicate that the calibration curves ove predict the compression-side loadings.
It.is thus demonstrated that the tension loads provide the most accurate representation of piping response to the SRV The design basis analyses of the main discharge events.
steam header piping predicted that SRV discharge would result The field in relatively small loads on the header piping.
data demonstrates that small loads were experienced during The measured tension loads are compared the SRV discharge.
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to their equivalent design loads (940 psig) in Table 4.
The design loads typically bound or approximate the measured tension loads.
The maximum exceedance of design load is Both the design basis and measured SRV loads 1100 pounds.
l are too small to affect the NS header design basis, hence the exceedance is not considered to be significant.
In summary, the correlation between empirical data and design predictions has been demonstrated.
The agreement between them has been shown to be within reasonable limits.
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Page 26 of 39 5.0 References
~
1)
Dresden Special Procedure SP 83-4-52, "Dresden-2 Main Steam Piping System Monitoring - Phase I,"
Rev. O.
2)
"Dresden-2 Main Steam Monitoring Procedure:
Seven-Day Data Evaluation,"'S&L Report No. EMD-043449, Rev.00, 5/9/83.
3)
Sargent & Lundy Piping Stress Analysis for Dresden-2 Inside-Containment Main Steam Lines A, B, C and D; S&L Accession No.'s EMD-040825, EMD-041326, EMD-041341, EMD-041333.
t
.V 4)
" Thermal Analysis Assessment with LVDT's Recorded Movements for Subsystems MS-C and MS-D," S&L Calculation EMD-043388, Rev. 00, 5/9/83.
5)
" Thermal Sensitivity Analyses to Simulate In-Plant Test Data-for Subsystem MS-C," S&L Calculation EMD-043692, Rev. 00, 5/19/83.
f l
6)
"Dresden-2 Main. Steam Monitoring Test: Snubber Calibration I
and Test' Analysis Evaluation," S&L Calculation EMD-044006, Rev. 00, 6/16/83.
- 7) 1GE RPV Expansion Formulae, Spec. 22A3828, Rev. 01, MPL No. A42-3670.
8)
S&L Calc. EMD-043399, 5/9/83, Rev. 00.
- =.
CARGENT O LUNDY EN NEERG Dresden-2 EMD-043846
.o 4
Ceco, 5486-23 June 18, 1983, Rev. 00 Page 27 of 39
,lO TABLE 1 COMPARISON OF LVDT AND WALKDOWN THERMAL EXPANSION DISPLACEMENTS Snubber Walkdown*
LVDT*
Ratio e
- 44 1.00 1.33 1.33
- 46 1.50 2.11 1.41 i
- 50 1.50 2.14 1.43 1
- 51 1.75 3.02 1.73
- 53 1.00 1.64 1.64 average 1.50 1
Displacements taken at approximately 450'F, from l-baseline temperature of approximately 105'F.
1 i
d O-i
CAROENTQ LUNDY l
ENGINEERO Dresden C"'C^*
EMD-043846 CECO,.5486-23 June 18, 1983, Rev. 00 Page 28 of 39
/
TABLE 2 EFFECTS OF RPV VERTICAL NOZZLE MOVEMENTS ON MS LINE C THERMAL DISPLACEMENTS Support Thermal Displacements Number Design Modified Measured Comments B2-3001 0.88 0.56 0.56 Y-direction, close to RPV nozzle B2-3002 0.36 0.00 0.00 B2-3004 0.24 0.16 0.20 Y-direction, far from RPV nozzle B2-3005
-0.19
-0.29 0.06
- 50 1.18 1.48 1.50
- 51 1.63 1.94 1.75 Horizontal
'/~T
- 44 0.32 0.44 1.00 (J
. Notes:
All displacements given in inches; All analyses correspond roughly to 345'F temperature differential; Modified analysis used RPV movements about
" less than design; Measured data from visual walkdown of Line C supports.
Drc= don-2 SARGENT Q LUNDY EMD-043846 CECO, 5486 EN GIN EERs June 18, 1983, Rev. 00
- "'C^
Page 29 of 39 r"g D
TABLE 3_
FIELD-MEASURED VS. DESIGN-PREDICTED THERMAL EXPANSION DISPLACEMENTS
- PART 1:
RPV RISER SPRING HANGERS Support type:
vertical spring hanger Location:
on RPV riser, about 20 feet below RPV nozzle MS Design Field Probable Reason Line Prediction **
Measured Difference For Difference A
0.88 0.63 0.25 Design RPV B
0.88 0.56 0.32 movements are conservatively C
0.88 0.56 0.32 high, by
(])
approximately D
0.86 0.56 0.32 one-quarter inch average 0.88 0.58 0.30 Support type:
vertical spring hanger Location:
near bottom of RPV riser MS Design Field Probable Reason Line Prediction Measured Difference For Difference l
A 0.37 0.19 0.18
[
B 0.41 0.20 0.21
- as above -
C 0.36 0.00 0.36 D
0.33 0.13 0.20 average 0.37 0.13 0.24 All displacements are stated in inches.
Design predictions are scaled to the cold (105 F)/ hot (450'F)
{w/i support field walkdown temperature differential of 345*F.
w-
1
.,Dracdsn-2
' "$E LUNDY EMD-043846 g g g
June 18, 1983, Rev. 00-i gCEco-- 5486-23 cmcaso
. Page 30 of 39
- O:
TABLE 3
PART:2:
HORIZONTAL SNUBBERS Support type:. horizontal snubbers Location:
near bottom of.RPV riser
~
l MS Snubber Design Field Probable Reason Line Location Prediction Measured
- Difference For Difference' i
'A
- 48 0.74 0.75
-0.01 Conservatively high design A-
- 47 0.87 1.25
-0.38 RPV movements predict-too-
.B-
- 45 0.49 0.50
-0.01 small horizontal movement;'in
- 46 1.19' l.50
-0.31 effect, stretching B
system in the C
- 50' l.18 1.50
-0.32 vertical direction i
C
- 51-1.63 1.75
-0.12 i,
D-
- 53 0.78 1.0
-0.22 D
- 52 1.00 1.0
-0.00 i
Support type:
horizontal snubber Location:
near.. top of riser nearest cont. wall
'MS
- Snubber Design Field Probable Reason L
' Line Location Prediction Measured Difference For Difference-as above -
A-
- 42 0.82 1.25
-0.43 B
- 41-0.31 0.25 0.06 C
- 44 0.32 1.0
-0.68-
- D-
- 43 0.83 1.0
-0.17
.'v**
=p'e
+ var +>-rw
- -*r
--s-er
- -ensev
+e,=--
-e r r---
s e e --
e-+e-.,o.-=vvae
.wa-
,.w--w--,w
-we o
e.-we-r-r--**-
e-=t9.+-=+=+drt wt***-sr----
---P" F-
'+F *'t
Dresdsn-2 CARCENTO LUNDY EMD-043846 CECO, 5486-23 EN GlN EEns June 18, 1983, Rev. 00 Page 31 of 39 0
TABLE 3
PART 3:-
CONTAINMENT RISER SPRING HANGERS Support type:
vertical spring hanger Locations near top of containment riser MS Probably Reason Line Prediction Measured Difference For Difference A
0.14 0.10 0.04 Differences are well within B
0.23 0.25
-0.02 accuracy estimates C
0.24 0.20 0.04
(
D 0.08 0.00 0.08
~
(2)
Support type:
vertical spring hanger Location:
on horizontal pipe near bottom of cont, riser MS Design-Field Probable Reason l
Line Prediction Measured Difference For Difference 1
A
-0.26 0.19
-0.45 These hangers are the most remote from B
-0.18
-0.13
-0.05 RPV nozzle movement effects; these hanger C
-0.19 0.06
-0.25 movements are therefore the most affected by D
-0.27 0.00
-0.27 hanger variability --
see text.
O
Dresden-2 EMD-043846-
-CECO, 5486-23 June 18, 1983, Rev. 00 Page 32 of 39 TM 4
SRV DISCHARGE IDAD COMPARISON
- Sni4+*r MS:
Design Measured Ntaber Header Ioads -(lbs) at Tension Ioads (lbs)
Remarks 940 psig
- at 940 psig 45 B
708 80
)
Electromatic 4
Valve B
'46 B
736 400 Actuation
'45 B
1382 80 Electromatic Valve E-46 B
1529 1500 Actuation
.j 50 C
1361 1700 b -
51 C
1453 550 44 C-1491 100 1
D 1283 550 52 D
900 2000-l Design loads at 940 psig represent values that were reduced from design basis l
loads to facilitate direct cm parison with the measured loads.
l L
1 t ~
.0L L
Dresden-2 SARGENT & LUNDY EMD-043864 Ceco, 5486-23 ENGIN ERS June 18, 1983, Rev. 00 Page 33 of 39 Figure 1.
Support Position' Scales Position indication 000000 D
0000 i
ow to m
00000E D
000
-o
+-
i 1/2 inch gradations PSA SNUBBER O
f M
i position d
indicator f
1/4 inch gradations 75' I
h/2
.1
,rl 1- /6 N
t'
-~.
'. j.
(C g @-d-a-
l 3
O O
TYPICAL SPRING HANGER.
Dresden-2 CARGENTO LUNDY EMD-043864
- ~
Ceco, 5486-23 ENG QERS June 18, 1983, Rev. 00 Page 34 of 39 O
Figure 2.
LVDT Mounting Configurations Brass rod LVDT
& LVDT core transformer
^
body d
J l
Steel pad and bracket t
/
/
Tape Structural support Main Steam header Y
PIPE-MOUNTED LVDT l:
i LVDT Brass rod transformer
& LVDT core
\\
body s
p rt clamp
[l LVDT support clamp O
lJ l
i 1
l Telescoping
-- Snubber support ylinder housing ' cylinder O
SNUBBER-MOUNTED LVDT l
l l
e Dresden-2 SARGENT & LUNDY EMD-043864 CECO, 5486-23 LNGINEERs June 18, 1983, Rev. 00
}
.Page 35 of 39 C"'C^*
L e n el cap Fa c.e -
6 c a. r-i n j a 35 e m bly
$Yi~
+..e J. v,J ~y s\\
<,....u u o
e n cl C o P
- t. g4lJ.I f b*d G
,,g=
l a
x
\\
Efi
.8a \\s ljjd
\\\\
\\
D N
U Strmin l
Rosette n
. p/"~
l f
- f. t. a S'
- l
/. c ; 5 "
- l
?igure 3.
Snubber End Cap Assembly O
l l
Dresden-2 CARGENT Q LUNDY EMD-_043864 Ceco, 5486-23 E N GIN E Eno June 18, 1983, Rev. 00 Page 36 of 39 C"'C^*
(O Figure 4.
Three-D View of End Cap Assembly O
b I
1 O
Dresden-2 SARGENT & LUNDY EMD-043864 CECO, 5486-23 ENLiNE June 18, 1983, Rev. 00 RS Page 37 of 39 3
O Figure 5.
Finite Element Model of Snubber End Cap w
0 a
=
I eg a
a I
i E
e w
k 3
I i
3 g
r g
O
~
- r e
9_ e
~
@n @
5 e
4 1
't.V ~
k
=
2 g
E s
1 Os a
l 1
d 3
I h.,
.i
=
3 3
e e
I 4.
O 3 3 1
e 7
't
-g i
\\
e n
Dresd:n-2 CARGENTQ LUNDY EMD-043864 CECO, 5486-23 EN NEERS June 18, 1983, Rev. 00 Page 38 of 39 4
O Figure 6 Snubber End Plug Stress Intensity Distribution:
--- TENSION LOADING ---
Stress levels:
i A = 2000 F = 12000 B = 4000 G = 14000 C = 6000 H = 16000 D = 8000 I = 18000 E = 10000 J = 20000 K = 22000 O
R
~
v n
l.
l I
I i
a dAA i
O
-,w-
- -e-~,w+-
e ow
---~---w
Dresdan-2 EMD-043864 CARGENT & LUNDY Ceco, 5486-23 June 18, 1983, Rev. 00 ENo:Norna Page 39 of 39 cwicaoo 4
iV Figure 7 Snubber End Plug Stress Intensity Distribution
--- COMPRESSION LOADING ---
Stress Levels:
A = 1000 H = 8000 B = 2000 I = 9000 C = 3000 J = 10000 D = 4000 K = 11000 E = 5000 L = 12000 F = 6000 M = 13000 G = 7000 N = 14000 0
h_ = _
r
)
)
- se_,
O
-