ML19305B776
| ML19305B776 | |
| Person / Time | |
|---|---|
| Site: | Trojan File:Portland General Electric icon.png |
| Issue date: | 03/15/1980 |
| From: | Broehl D PORTLAND GENERAL ELECTRIC CO. |
| To: | Schwencer A Office of Nuclear Reactor Regulation |
| References | |
| TAC-12369, NUDOCS 8003200315 | |
| Download: ML19305B776 (22) | |
Text
W-March 15, 1980
=m eon!CCd GONTEj EOChiC CCn':Eny Trojan Nuclear Plant
.h s
Docket 50-344 License NPF-1
' n :. ?w r - +. m~
1 Director of Nuclear Reactor Regulation ATTN:
Mr. A. Schwencer, Chief Operating Reactors Branch #1 Division of Operating Reactors U.
S. Nuclear Regulatory Commission Washington, D. C.
20555
Dear Sir:
In accordance with our letter of February 29, 1980, our com-ments on the " Report on Design Criteria for Masonry Walls in the Trojan Power Plant" by Dr. James Colville, consultant to your Staff, are provided herewith.
This material was not provided last week as predicted in that letter in order to allow the opportunity to <:onsider additional information supplied by Dr. Colville in a meeting between PGE, Bechtel and your Staff on March 8, 1980.
Dr. Colville's recommendation of a 12 psi allowable for shear bond strength of the mortared collar joints relies in part on a report
- published by the Alberta Masonry Institute.
Licen-see's consultant, Professor Boris Bresler, has reviewed that report and provided Licensee with his comments.
Professor Breslers comments are provided as Attachment 1 to this letter. is a review of the basis for Bechtel's determina-tion that 18 psi is a conservative value for the allowable collar joint shear stress which we believe continues to be an appropriate value.
As mentioned in our letter of February 29, 1980, we are under-taking a short-term program to quantify the margins in the in-situ shear capacity of the mortared collar joints in Trojan double _wythe masonry walls.
Description of the test program and results will be provided on completion of the tests.
In Section B and Appendix B of his report, Dr. Colville de-scribes his bounding analysis of the Auxiliary Building wall on column line 46.
Bechtel has reviewed his calculations and has noted that a number of the assumptions used by Dr.
Colville do not reflect the specific circumstances applicable to that wall. Bechtel has recalculated the collar joint shear
- Hatzinikolas, M.,
Longworth, J.
and Warwaruk, J.,
" Evaluation of Tensile Bond and Shear Bond of Masonry by Means of Centrifugal Force" Alberta Masonry Institute, Edmonton, Alberta.
A MCH3/
AL-6 S
///
8008200 3 ) {
1 m
_ :2 J
EC.#dEnd GOCC%l ECCUiC CCrrpeny Mr. A. Schwencer, Washington, D. C.
Page 2 of 2 stresses with more realistic assumptions than used by Dr.
Colville and has shown that these stresses in wall 46 would be less than even Dr. Colville's suggested allowable of 12 psi. provides Bechtel's comments on Dr. Colville's calculations and presents a more realistic evaluation of wall 46.
Based on this analysis, Licensee concludes that wall 46 has sufficient capacity to resist out of plane loads during an SSE event.
Sincerely, S
Donald J.
Broehl DJB/LWE/dd Attachments c:
Mr. R.
H.
Engelken, Director U.
S. Nuclear Regulatory Commission Region V Mr. Lynn Frank, Director State of Oregon Department of Energy
)
AL-6
A. SCHWENCER MARCH 15, 1980 ATTACHMENT 1 Review of
" EVALUATION OF TENSILE BCND AND SHEAR BOND OF T
MASONRY BY MEANS OF CENTRIFUGAL FORCE,"
1 S
s.
M. Hatzinikolas, J. Longworth, and J. Warwaruk, J
Alberta Masonry Institute, Undated (1978-1979) a n
n e
Y-t E
Prepared by B. Bresler i
S t
n e
This review is primarily focused on the relationship between I
shear bond and tension bond strength values reported in this publi-a n
d cation.
In preparing this review, additional information regarding A
s the test program and test results was obtained from Refs.1, 2, and s
0 9
3.
I f
A preltninary review of the publication revealed the following:
a e
8' (1)
The calculated ratio of mean value of shear bond to the 1
{
mean value of tensile bond was 0.63.
This value normally is expected to be greater than 1.0 and the reason for the low ratio was not immediately apparent.
(2)
The scatter of test results for shear bond strength (with type M mortar) appeared excessive.
The results ranged from 28.9 psi to 71.6 psi, with coefficient of variation of 27%.
(3)
Reported values of the size of specimen were not con-sistent with normal unit weights of masonry.
(4) Detailed data on test apparatus, test procedures, material control test values, etc. given in the publication of the Alberta Masonry Institute under review were insufficient to assess the validity of E
the test results and particularly the use of these 1
S s.
results in assessing shear bond strength values in J
a the masonry walls ~ of the Trojan Control Building.
n n
Further investigation led to the discovery of additional E
Publications dealing with.these tests (Ref.1 and 2) and to a 1
s brief discussion of the test program with one of the co-authors t
n e
(Ref. 3).
Further review of the additional information revealed r
a the following:
n d
(1)
The size of the center portion of the shear bond Aj specimen was 4x4x8 inches.
O t
(2) The radial distance from the center of the centrifuge I
a t
to the center of the specimen was 18 inches.
e 8'
(3)
Shear stresses reported in the publication were based
[
on sidple beam theory and were calculated assuming full gross section (4 x 4 in.) effective in resisting shear at time of failure.
(4) The mortar used at the interface had a mean compressive strength of 2540 psi with a coefficient of variation of 8.3%.
The thickness of the mortar layer was not reported.
j (5)
Failure was defined as initial slip (or deformation) on the shear plane. A control switch terminated the test when a predetermined (small but unspecified) slip had
+
taken place.
(6) Modes of failure and significance of variation in the
" weight of separated portion" reported in Table A4 of the publication could not be ascertained from the
\\y additional References.
Although Ref. 2 states that i[
specimens were supported to avoid bending at mortar 3
joints, it is not clear that with the test setup a
n n
used this was possible to achieve.
No reference e
Y.
is made to provisions for eliminating effects of E
I tangential components of end reactions, s
t H
Based on the information available to date, the following r
conclusiens were reached:
a n
d (1) Variability in the data is due to such factors as
'f large size of the specimen relative to the size of 5
a the centrifuge, variability in the end restraint i
a conditions for individual specimens, arbitrary e
s.
definition of " failure", and neglecting to account I
n for the influence of end restraints and failure t
modes on the calculated values of shear bond strength reported in this publication. Variations in the effects could substantially change the mean value of the actual shear bond strength.
(2) Because tension bond strength depends on adhesion at mortar-block interface, and because pure shear bond l
strength depends on both adhesion and friction, data on tension and shear failures in cementitious materials indicate that pure shear strength (in the absence of
1 secondary normal stresses) is equal to or greater than tension strength.
Preliminary calculations indicate that prior to failure, local tension stresses at the l
mortar-block interface were approximately equal in Y
magnitude to the shear stresses.
The low shear stresses I
s S.
reported in this publication are partly due to the 1
a presence of these local tension stresses, which are n
k characteristic of the type of test used here. Normally, Y.
E such high tension stresses would not be present in 1
S realistic walls where shear bond stress is generated n
e by out-of-plane loading.
I a
(3) Thus, the ratio of shear to tension bond of 0.6 reported n
d in this publication is characteristic of the test A
s procedure, of the definition of failure, and of the O
C approximate manner of calculating shear bond strength I
a t
of failure. However, in the light of information reviewed, e
this ratio is not applicable to masonry walls used in the I
E TrojaIn Control Building.
References 1.
M. Hatzinikolas, J. Longworth, and J. Warwaruk "The Use of Centrifugal Force for Determining the Tensile Bond Strength of Masonry", Proc.
of the 6th International Masonry Conference, London, England,1977.
2.
M..Hatzinikolas, J. longworth, and J. Warwaruk " Concrete Masonry Walls", Str. Eng. Report No. 70, Department of Civil Engineering, University of Alberta, Sept., 1978.
3.
J. Warwaruk, Private Communication, March,1980..
-l A.
SCHWENCER DOCUMENT.TTION SUBSTANTIATING 18 PSI MARCH 15, 1980 ALLOWABLE COLLAR SHEAR STRESS ATTACHMENT 2 Licensee's selection of an allowable collar joint shear stress of 18 psi for mortared double block walls is supported by test resul.ts which were described in Licensee's response dated De-cemcer 22, 1979 to NRC Staff Request No.
3.
The allowable collar joint shear stress recommended by Dr. Col-ville is in part dependent on his derivation of shear stress values from the results of tension stress tests.
In arriving at a relationship between shear stress and tension stress Dr. Colville relied on tests performed at the Alberta Masonry Institute to determine a ratio of shear stress to tensien stress of 0.6.
The appropriateness of using the reported values from the Alberta tests for this purpose is addressed in the report prepared for Licensee by its Consultant, Professor Boris Bresler.
The following discussion reviews the information which sup-ports Licensee's selection of an allowable collar joint shear stress, and also provides comparisons which support Licensee's view that the ratio of shear stress to tension stress is greater than 1.0.
1.
In the tests performed by Testing Engineers of Oakland, California, a pre-cast concrete cover was attached to a concrete slab using Portland cement mortar without steel ties.
Direct shear tests were performed on three valid samples which showed an ultimate shear strength ranging from 200 to 286 psi.
The tests also indicated that the ultimate shear strength is about 3 times the tensile strength of mortar.
DL-15/1
l Page 2 2.
Qualification tests were performed for Stucco Stone Products of Napa, California by Structural Tasting, Inc. of Santa Rosa, California to qualify the adhesion of artificial stone veneer to masonry backing.
The " stone" is made of lightweight concrete and is attached to the backing with type S mortar without mesh or metal ties.
The direct shear tests on the mortar joint showed an ultimate shear strength ranging from 56 to 519 psi.
The minimum shear strength to qualify under ICBO certification is 50 psi.
3.
Tests similar to item 2 above were performed for Eldorado Stone Corp. of Kirkland, Washington by the Seattle office of Pittsburgh Testing Laboratory.
This artificial stone is also attached to masonry backer without metal ties but using mortar similar to type M.
Ultimate shear strengths from these qualification tests range from 113 to 127 psi.
4.
A series of unreinforced brick-mortar couplet tests were run by Benjamin and Williams to determine shear strength under various conditions of tensile and compressive normal stresses, using three different qualities of mortar.
The most relevant are those in pure shear.
The average ultimate shear stress of 5 specimens for each of the 3 mortar qualities ranged from 109 to 148 psi.
Multiplying by 2/3 to convert from a parabolic to a rectangular stress distribution yields a conservative ultimate shear strength of 72 to 98 psi.
The average ultimate tensile strength DL-15/2
j l
Page 3 of S specimens for each of the 3 mortar qualities ranged from 30.7 to 38.8 psi.
The above indicates a shear to ultimate strength ratio of greater than 2.0.
Although these tests were not-performed on masonry units, Licensee considers them applicable because they do provide results of tests of mortar collar joints.
Therefore, we conclude from these examples that the minimum ultimate mortar shear strength in a collar joint between two wythes would be expected to be at least 50 psi (this is the code value for veneer units in the UBC) where mortar is present.
This provides a safety factor of about 3 over the allowable vertical shear stress acceptance criterion of 18 psi.
The nature.of global vertical shear stresses permits redis-tribution around areas which may have low strength due to poor workmanship without affecting overall structural per-formance.
Thus, the wide range of values above 50 psi in the test data provides additional conservatism.
The above tests also indicate that mortar shear strength exceed the tensile strength.
UL-15/3 l
l l
l
\\
COMMENTS ON APPENDIX B OF DR. COLVILLE'S REPORT OF 2/13/80 ON TROJAN MASONRY WALLS A. SCHWENCER MARCH 15, 1980 I.
Introduction ATTACHMENT 3 The following discussion comments on Dr. Colville's evaluation of the mortared double wythe wall on column line 46 of Trojan Auxiliary Building for out of plane inertia loads at elevation 61'-0".
II.
Description of Wall 46 Between el. 61'and el. 77', wall 46 consists of mortared double wythe blocks supported by the slabs at those eleva-tions and by steel columns on either side. (Figure 1 shows the size of these columns, the smallest of which is W14X142).
Each wythe of the wall is vertically reinforced by #6 rebar l
at 16".
(Dr. Colville's analysis assumed #6 at 24").
The blockcellsarefilledwithgroutwithfh=5000 psi.
Dr. Colville's evaluation does not specifically account for the presence of this cell fill.
III. Discussion As Dr. Colville's report shows, the inertia loads are sensitive to the natural frequency of the wall, which in turn is sensitive to the wall's boundary conditions and the rigidity of the wall.
For example, his calculations show that as the assump-tion about the boundary conditions and the degree of cracking in the wall are varied, the natural frequency changes from 49.9 cps, (beam action, both ends fixed, concrete uncracked, I, = 3815 in4=Ig) to 6.8 cps (beam action, both ends pinned, CZ-33/1
/
Page 2 4
concrete completely cracked, I, =347 in ).
For these values, 2
the calculated inertial loads varied from 79 lbs/ft (corres-ponding to 0.44g) to a maximum of 252 lbs./ft2 (corresponding to 1.4g), resulting in collar shear stresses of 4.46 psi and 16.1 psi, respectively.
Thus, it is apparent that unless all parameters are appropriately considered in the analysis, an unrealistic appraisal can result.
]
~Cn order to predict the shear stress levels in the collar joint, one must make reasonably accurate assumptions of:
(i) the boundary conditions of the wall and (ii) the moment of inertia of the concrete section Because the wall panel is supported at el 61' and el 77' by floor slabs, the top and bottom edges will have partial fixity, between a fully hinged and a fully fixed condicion.
The wall panel is supported elastica 11y on the two sides by the structural steel columns.
CZ-33/2
i Page 3
)
Bechtel considers the moment of inertia of the section to be somewhere between the uncracked condition with I
- I e
g 4
= 3815 in and the fully cracked condition with I
=I e
cr 4
= 490 in.
(This latter value reflects consideration of actual rebar distribution).
As discussed in Appendix A hereto, Bechtel considers the I value proposed in Supplement 1 to e
LER 79-15 (I, = (Ig+Icr]/2) appropriate for this evaluation.
Table 1 shows calculated wall frequencies for various assumed boundary conditions and moments of inertia.
Column 6 of Table 1 shows results considered to be the most appropriate simulation of boundary conditions.
Column 5 shows the re-sults for the panel if one neglects the support provided by the columns.
This was merely used as a base case and the results were modified to account for the stiffening effect of the edge steel columns.
Frequencies were determined'using the techniques in References (1) and (2).
From this evaluation, it can be noted that the calculated frequency change between the uncracked and fully cracked conditions is much less for a panel with steel column supported edges than for a panel with-out edge. supports.
The accelerations, inertia loads and collar joint shear stresses corresponding to the various frequencies calculated for the wall with edge supports (column 6) are provided in Table 2.
IV. Conclusions It is obvious that the influence of the edge supports is CZ-33/3
Page 4 significant in affecting the frequency of the wall panel.
This effect becc.es more pronounced as the concrete in the wall cracks.
Based on what Bechtel considers to be realistic assumptions of I
= 1/2(Ig+Icr), with conservative assumption e
of hinged top and bottom, we obtain a natural frequency of 17.6 cps.
The corresponding inertia loads are 108 lb/
ft2 (0.6g), resulting in a collar joint shear stress of 6.9 psi which is quite low and would not result in wythe separation.
Even if I, = 0.2 I is assumed as an extreme g
case, collar shear stress is still only 9.2 psi and would not result in any wythe separation.
Therefore, Bechtel does not agree with the assumptions made by Dr. Colville in his evaluation, and thus his conclusion that this double wythe wall is incapable of resisting out of plane inertia loads due to a 0.25g SSE.
As shown above, when what Bechtel considers to be realistic boundary condi-tions and the correct cracked section moment of inertia are considered, the wall is adequate to resist the SSE loads.
Similar conclusions are drawn for walls at el.
77'-0".
V.
References (1)
Timoshenko and Krieger, " Theory of Plates and Shells",
McGraw-Hill Book Co.
(2)
W.
H. White, " Free Vibrations and Buckling of Edge Stif fened_ Square _ Plates", _ Doctor.al._ thesis submitted to University of Colorado (1970).
CZ-33/4
Page 5 TABLE-1 NATURAL FREQUENCY OF WALL 46 l
Boundary l l
l l
l 1
l l
Conditioni 1
l 2
l 3
l 4
l 5
l 6
I I
I I
I I
I i
l I
I I
I I
I 88 I
88 I
su
,s
,r I
I I
I I
I i $ pg I
I I,
I I
I i
I I
I y
y I
I I
l Partiall I
ss l
ss l
1 1
I Fixity I I
I l
l l
l l
l l
l l
l l
l l
l l
l 1
I I
i 19.2 l
42.9 1
31.3 l
29.9 l
19.5 l
21.0 I
g I
l-1 I
I l
l I
I I
I I
I I
I I
i 1/2(Ig+Icr)l 14.4 1
32.2 1
23.5 1
22.5 l
14.6 l
17.6 I
I I
I I
I I
I I
I I
I i
I I
i i
i l
i I
I I
I I
I 0.2r I
8.6 l
19.2 1
14.0 l
13.4 l
8.8 l
12.5 I
g I
I I
I I
I I
I I
I I
I
,1 1
I I
I I
I I
I I
I I
l I
6.9 l
15.4 l
11.2 l
10.7 l
7.02 l 10.3 I
cr I
I I
I I
I I
I I
I I
I I
I I
I ss = simply supported CZ-3S/S
i
\\
.l Page 6 TABLE -2 COLLAR JOINT SHEAR STRESS IN WALL 46 AT EL 61 FOR BOUNDARY CONDITION 6 I
I I
I I
I I
I I,
I FREQUENCY l ACCELERATION I INERTIA LOAD l VO I
V I
I I
(CPS) l (g) l (lbs/ft.)
l I I
Dd t
l l
l l
l l
l l
1 I
I I
I I
I I
I I
I I
I I
I I
I I
l 21.0 1
0.45 I
81 1 5.2 1
4.5 I
g I
I I
I I
I I
I I
I I
I I
I l
l l
I I
I I
Il_(I +Ier)I 17.6 1
0.6 l
108 l 6.9 I
6.0 l
g
-12 I
I I
I l
l l
l l
l l
l l
l l
l l
l l
l 1 0.2I i
12.5 1
0.8 l
144 1 9.2 I
8.1 1
g I
I I
I I
I i
I I
I I
I I
I I
I I
I i
l l
l I
l 10.3 l
1.4 l
252 116.2 1
14.2 I
er I(0.128Ig)l l
l l
l l
I I
I I
I l
CZ-33/6
N x
Sk
=
2[
8 w
s:
W/Mx/4f 4
Ss H
W/4 s 2/9 wt4 a ei9 p
s-wt4n at9 wiqa287 g __
aux. 8LO4 WALL 4G, pl AU @ EL U OL
P/Guee 1
A.
Schwencer APPENDIX A March 15, 1980 DETERMINATION OF Appendix A.
MOMENT OF INERTIA As stated in Supplement 1 to LER 79-15, the moment of inertia, I,.
used to calculate the frequency of walls is taken as 1/2(Ig+Icr)
I,
=
where moment of inertia of gross section I
=
g I
=
moment of inertia of cracked section cr This approach is also used in reference 1.
Further justification for the above stated approach is based on the ACI 318-77 method (Chapter 9) for deflection calculations.
The double wythe walls in the Complex consist of blocks where the core is filled with grout with a compressive strength of 5000 psi.
Since the grout fill in the blocks is significantly stronger than the blocks, only the grout cell fill will be considered in this evaluation.
The cross sectional area of the core is one half of the total block area.
The uncracked moment of inertia of the cell fill is:
3 I
12.8 5
2 2 + 5 x 12.8 4 x2 1
4
= 144.7 in /in 12 16 g
(
EA-2 I
Page 2 A.
Schwencer March 15, 1980 Appendix A.
Then the section modulus ist 3
S = 144.7/6.5 = 22.3 in /in taking f
= 7.5 (reference: ACI 318-77, EQ. 9-8) r The cracking moment M is:
er M
= 7.5.[5000Tx 22.3/1000 11.8 ftk/ft
=
er The maximum moment, M, at the stage the deflection is computed, a
is taken as the sum of the cracked thermal moment, the moment due to interstory displacement and the moment due to the dynamic response.
The thermal moment acting on the section based on a 50*F gradient, 3
using an estimated relaxed section modulus of 15.38 in, is calculated below: (The modulus of elasticity is determined based on section 8.5 of the ACI 318-77 code):
50/2 x 6.5/8 x 6.0 x 10-6 x 4.3 x 10 x 15.38/1000 = 8.1 ftk/ft 6
The moment due to interstory displacement at the top of the wall caused by a shear stress of 4 psi, as obtained from the dynamic analysis, is:
4 x 12 x 16 x 8/1000 = 6.1 ftk/ft However, when considering the maximum combined moments resulting from inertia and interstory displacement, only 2.8 ftk/ft is contributing to the maximum moment.
EA-2 l
A.
Schwencer March 15, 1980 Page 3 l
Appendix A.
i Using an initial assumption of a moment of inertia of 85 4
6 2
in /in and a modulus of elasticity of 4.3x10 lbs/in, the frequency of the wall is 14.5 cps based on one way, pinned-pinned end conditions.
The corresponding moment due to the seismic inertia load on the wall is:
2 M=.152 x 16 /8 = 4.9 ftk/ft.
Therefore, the total moment is:
M
= 8.1 + 2.8 + 4.9 = 15.8 ftk/ft.
a According to ACI Std. 318-77, (9-7),
[M
[M 3
3 er er
,I
+
1-I
=
I M
M cr
\\.* /
g
(* ) _
e (11.8/15.8)3 144.7 + [1-(11.8/15.8)3] 41.4 = 85 in /in 4
I
=
e Since the resulting I, is the same as that initially assumed, the dynamic response was appropriately calculated.
As the calculation of the effective moment of inertia according to the ACI 318-77 reflects the nature of the deformation of a generally cracked reinforced concrete member, there is no reason to assume that in case of repeated loading the above result would need any correction.
EA-2 l
A.
Schwencer March 15, 1980 Page 4 Appendix A.
When the natural frequency is calculated using 4
I, = (Ig+Ier)/2 = 185 in /in as shown in Supplement 1 of LER 79-15 and a modulus of elasticity for the block 6
2 equal to 1.5 x 10 lbs/in we obtain a frequency of 12.7 cps for pinned-pinned end conditions.
By comparing the two values of the natural frequencies, it can be seen that the present more detailed analysis provides results very close to the previously established response.
Also, as a lower natural frequency results in higher response, the previous value was conservative.
Therefore, it is our conclusion that the use of I
= 1/2 (Ig+Icr) is appropriate and justified.
n
Reference:
(1)
ASCE - Manuals of Engineering Practice - No. 42 " Design of Structures to Resist Nuclear Weapons Effects" prepared by the Committee on Structural Dynamics of the Engineering Mechanics Division, 1964 Edition.
EA-2 3
~
UNITED STATES OF AMERICA NUCLEAR REGULATORY COMMISSION BEFORE THE ATOMIC SAFETY AND LICENSING BOARD In the matter of
)
)
PORTLAND GENERAL ELECTRIC COMPANY,
)
Docket No. 50-344 et al.
)
(Control Building Proceeding)
)
(Trojan Nuclear Plant)
)
CERTIFICATE OF SERVICE I hereby certify that on March 15, 1980, Licensee's letter to the Director of Nuclear Reactor Regulation dated March 15, 1980 with comments on a report by Dr. James Colville, consultant to the NRC Staff has been served upon the persons listed below by depositing copies thereof in the United States mail with proper postage affixed for first class mail.
Marshall E. Miller, Esq., Chairman Joseph R. Gray, Esq.
Atomic Safety and Licensing Board Counsel for NRC Staff U.S. Nuclear Regulatory Commission U.S. Nuclear Regulatory Washington, D.C.
20555 Commission Washington, D.C.
20555 Dr. Kenneth A.
McCollum, Dean Division of Engineering, Maurice Axelrad, Esq.
Architecture and Technology Lowenstein, Newman, Reis, Oklahoma State University Axelrad & Toll Stillwater, Oklahoma 74074 1025 Connecticut Ave.,
N.W.
Suite 1214 Dr. Hugh C.
Paxton Washington, D.C.
20036 1229 - 41st Street Los Alamos, New Mexico 97544 Frank W. Ostrander, Jr.,
Esq.
i Assistant Attorney General Atomic Safety and Licensing Board State of Oregon Panel Department of Justice U.S. Nuclear Regulatory Commission 500 Pacific Building Washington, D.C.
20555 520 S. W. Yamhill Portland, Oregon 97204 Atomic Safety and Licensing Appeal Panel William Kinsey, Esq.
1 U.S. Nuclear Regulatory Commission
,Bonneville Power Admin.
Washington, D.C.
20555 P.O.
Box 3621 Portland, Oregon 97208 Docketing and Service Section (3)
I Office of the Secretary U.S. Nuclear Regulatory Commission Washington, D.C.
20555 l
CE-13
CERTIFICATE OF SERVICE Ms. Nina Bell Mr. Eugene Rosolie 728 S.E.
26th Avenue Coalition for Safe Power Portland, Oregon 97214 215 S.E.
9th Avenue Portland, Oregon 97214 Mr. John A.
Kullberg Route 1, Box 2500 Columbia County Courthouse Sauvie Island, Oregon 97231 Law Library Circuit Court Room Mr. David B.
McCoy St. Helens, Oregon 97051 348 Hussey Lane Grants Pass, Oregon 97526 Dr. Harold I. Laursen 1520 N.W.
13th Ms. C. Gail Parson Corvallis, Oregon 97330 P.O. Box 2992 Kodiak, Alaska 99615 A
^
Ronald W. p nson Assistant General Counsel Portland General Electric Company Dated:
March 15, 1980 i
I CE-13