Text

FENOC-Davis-Besse Nuclear Power Station, Unit 1 Docket No. 50-346, License No. NPF-3 Submittal of Contractor Root Cause Assessment Report-Section 7
	ML12138A089
Person / Time
Site:	Davis Besse
Issue date:	05/14/2012
From:	; FirstEnergy Nuclear Operating Co
To:	; NRC/RGN-III
References
	L-12-196
	Download: ML12138A089 (145)
	v • d • e

Exhibit 65 Toledo resides at approximately 42 degrees northern latitude. Since the solar declination angle as well as insolation undergoes the greatest variations in magnitude between the solstices and equinoxes, these would represent the most extreme changes in solar heating that could be expected during the course of a single year. Cons thermal anal es were Aside from variations in the solar flux, the length of day also varies through the course of the year, being longest at the summer solstice and least at the winter solstice. Solar heating would only be present when the sun is visible in the sky. This wi" occur when the zenith angle is 90 degrees or less. Sunrise and sunset wi" occur when the zenith angle is precisely 90 degrees.

Using a latitude of 42 degrees along with the appropriate solar declination angle, the corresponding hour angle that marked sunrise and sunset was calculated. In this way the number of hours when daylight was present was determined. This enabled precisely locating not only the solar position relative to the zenith but also relative to east and west. Using this methodology, it was determined that at the Summer Solstice, the longest period of daylight was 15 hours1.736111e-4 days 0.00417 hours 2.480159e-5 weeks 5.7075e-6 months whereas at Winter Solstice there was only 9 hours1.041667e-4 days 0.0025 hours 1.488095e-5 weeks 3.4245e-6 months of daylight. At both Equinoxes, there was precisely 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months of daylight. These daylight hours were checked online for accuracy at the US Naval Oceanography Portal (http://www.usno.navy.mil/USNO/astronomical applications/data-services/rs-one-year-us). Here the number of daylight hours as well as the time of sunrise and sunset can be determined for any location worldwide and for any month of any year.

The solar heating was treated as a directional flux of constant intensity during the hours of daylight. The direction cosines relative to east-west, north-south, and the zenith were derived as a function of the time of day. These definitions were used by NASTRAN to determine the direction of the solar flux as well as those portions of the outer concrete shield building that would be lit as we" as those portions that would be shaded. in addition to the shading attributed to suns position in the sky, additional shading provided by any adj acent structures had to be taken into account. This was especially true of the auxiliary building that resides adjacent to the eastern half of the concrete shield building. Con se quently, a Boolean solid representation of the auxiliary building was derived that approximated its outer shape. This was coarsely meshed and included in the model solely for the purposes of including any shading effects. Temperatures were not calculated or recovered for this structure. Rather it was assumed to have the same temperature as the surrounding ambient environment. in Figure 3, Page 5 Page 5 of 23

Exhibit 65 the 9 :00am temperatures contours that were calculated for the summer solstice are plotted along with an outline of the auxiliary building. The shadow cast on the concrete shield building is clearly visible by the lower temperatures on the southeast and northeast walls due to the lack of solar heating during the early morning hours.

Palra,., 201 0 4'BII (MD Enabl0d) 29-J~n-12 07 2 1.13 FMgg 58. RECOR D._NOWlliO DAY 3. l-5 Tlme = 0' 507. Temperalures ,. (NON L4 YER ED) 1 15+00' 1.10<002 1 0 H ),1(

105-00' 1 2+00::

9 90+00 1 g 8*001 892+00 1 866 00 1 8 40*00 1 defalolLFringe .

Max 1.23+002@No 21071 1 Min 8 .40+001 @Nd 1 FI GU RE 3. SU M MER SOLSTICE 9:00AM TE M PERATURES SHOWI NG TH E AUXI LLIARY BUI LDING AN D TH E SHADI NG OF THE CONCRETE SHIELD BUILDING In addition to solar heating, another source of radiant heating is attributable to the inner steel containment that houses the reactor. This steel containment operates at an elevated temperature that is well above the surrounding ambient air as well as the annular air.

Unfortunately, no direct temperature measurements were available to determine the spatial distribution, if any, in the temperature on the exterior of the steel containment. Consequently, a uniform temperature of 120°F was used . This was the same operating temperature that was cited in the Davis -Besse thermal analysis (CALC 27 014 Shield Bldg Temp Analysis page 4, dated 12/20/2000). It was further assumed that this temperature remained invariant in response to any seasonal changes in the external ambient conditions. For the purpo ses of incorporating the Page 6 Page 6 of 2 3

Exhibit 65 radiant heat transfer to the interior face of the concrete shield building, the steel containm ent was treated as a simple constant 120°F ambient source with a view factor of 1.0 since the entire steel containment has an unobstructed view of the interior of the concrete.

For all of th e thermal conditions examined, radiation with the external ambient was included. This was especially important at night and in winter when the cooler night sky and surroundings would result in an appreciable heat loss even in the absence of any convective losses due to wind . For all radiant heat transfer effects, the typical published and CR3 measured properties for concrete were assumed. Namely, an emissivity of .93 and an absorptivity of .6 were used.

2.2 AM BIENT TEM PERATURE The Davis-Besse facility resides at 42 degrees north latitude and as a result experiences dramatic seasonal variations in the ambient temperatures and conditions. The average and record monthly temperatures for Toledo are shown in Tables 1 and 2. These temperatures were obtained from The Weather Channel and are based upon NOAA historical records that extend back to 1870.

Month Jan I Feb I Mar *I Apr May iJun I

Jul Aug Sep Oct Nov Dec High (OF) I 33 1 38 I 48 I 61 I 74 I 83 87 84 77 6451 38 DC ' 3 1 9 16 23 1 1 28 30 29 25 18 10 3 Low rF) 22 . 24 33 42 53 63 68 ' 66 59 48 38 27

-5 ' -4 5 12 17 20 19 15 9 3 -3 TA BLE 1. MONTHLY AV ERAGE TEM PE RATU RES AT TOLEDO, OHIO Month Jan Feb Mar Apr May JunJul Aug Sep Oct Nov Dec Record high (OF) 68 . 71 82 90 99 104 106 103 101 91 80 70

C 18 ' 22 28 32 37 40 41 39 38 33 27 21 Record low (OF) -16 -5 4 11 28 39 47 36 36 24 6 -13

C 21 12 -2 I 3 8 2 2 -4 25 TABLE 2. MONTHLY RECORD DAYTIME TE MPERATURES AT TOLEDO, O HIO The monthly average high and low temperatures were used to define the excursion in daily average temperatures for the winter and summer solstices and the vernal and autumn Page 7 Page 7 of 23

Exhibit 65 equinoxes. The same daily variation in average temperature was also used for either record high or low conditions. An examination of the record temperatures reveals that the highest temperatures always occur in the summer and fall and the coole st in the winter and spring. To ascertain the effect of thermal stress over the maximum range of annual temperatures, thermal transient analyses were performed using the average and record high temperatures for the summer solstice and autumn equinox. Conversely, the average and record low temperature s were used for the winter sol stice and vernal equinox .

To simulate the thermal lag that typically occurs in the ambient temperature relative to the solar transit, it was assumed that for the first 90 minutes after sunrise, there was no appreciable change in the ambient temperature. The sunrise temperatul"e was set to the minimum nocturnal temperature. After the first 90 minutes, the temperature was gradually increased in a sinusoidal fashion to its daily maximum at 1:30 pm in the afternoon . The temperature was then allowed to gradually loose 90 % of the daily temperature rise at sunset.

After sunset, the temperature exponentially decayed to its nocturnal minimum at 11:30pm. The temperature then remained invariant till 90 minutes after sunrise on the following morning.

The average and record temperature profiles that were used are shown in Figures 4 through 7.

LEGEND

- - SF"R'N GAVER.o\G PR i'lLE

- - SPRING RECORD LOW PROfiLE 500 40.0 iL 300 0

w Q.

w 20.0 tr

l

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tr 10 U W

Q ill f

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-/ 0 0

- ~o.o

o. /6 0 to, TI ~iE FRO SUNRI<>E (hOU S)

FIGURE 4. VER NAL EQUINOX AVERAGE AND RECORD LO W HOURLY TEMPERATU RES Page 8 Page 8 of 23

Exhibit 65 105 95 0 u:

UJ 8 B50 tlJ a:

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'"w f) 750 ;--------+,~-----I -~~~-+--~---~------~----

~

65.0

.. I 550 4.00 BOD ; 20

6.0 ~OO TIM [ FR OM SUNRISE (IiOURS )

FIG URE 5. SUMMER SOLSTICE AVERAGE AND RECORD HIGH HOU RLY TEMPERATURES LEGEND

- - AUTlJ I...IN AVEFV\ GE PROFltE

- - AU"Tl.IMH RECORD IlI GI~ PROPLE

1:..

- - -1*

-,--,------t----i---r-r--t---.----,-,--. *-It-.-r-~i--r

'i 0 8 00 12.u 16 J ?O0 11 E FROM 5U IJR eE (HOURS) fiGURE 6. AUTU!'V1I1'J EQUINOX AVERAGE AND RECORD HIGH HOURLY TEM PERATUR ES Page 9 Page 9 of 23

Exhibit 65 LEGEND

- - l'>tNTE1', ,'\VEPAOE P!'<OFILE

- - IMNr!:F.I'£(;ORD LOW P"IOFILE

100

. ~

.~, " , : . -I JU0 G: 200 ..

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J .... . ... . .... ..

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- - i ** *******

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-20 0 .... *... . ",

-300 1 o.

I 400 8.00 120 16 0 20.0 TI I-IE FROf,1SUNRISE (H OURS)

FIGURE 7. WINTER SOLSTICE AVERAGE AND RECORD HIGH HOURLY TEM PERATU RES These transient temperature profiles defined the ambient temperatures that were used to compute the heat transfer to the exterior of the concrete shield building due to either radiation, free convection or forced convection. The base of the concrete shield building was assumed to be grounded to a large thermal mass whose temperature remained constant and fixed at the daily minimum. This behavior is consistent with temperatures at deep subterranean levels which typically are unaffected by short term hourly changes in air temperature or solar heating.

2. 3 CO NVECTION The principal source of heat loss from the concrete shield building is due to wind. Winds inthe Toledo, Ohio and surrounding area are predominantly southwesterly as shown in Table 3, especially in the winter and spring. During the summer and fall the incidence of winds from the west and northwest increases, but they are far less intense. The average winds are quite light, with winds speeds of less than 10 mph. In comparison the record wind speeds are considerably higher, often in excess of 60 mph, and generally are from the southwest. This can be seen in Figure 8 where the record high wind speeds for the period of 1998 through 2011 complied by Page 10 Page 1 0 of 2 3

Exhibit 65 NOAA are shown plotted as well as in Table 3 where the corresponding record and average wind speeds as well as wind directions are listed for the same time period.

FIGURE 8. TOLEDO AVERAGE ANNUAL RECORD HIGH WIND SPEED (1998 TO 2011) 201 ' t~ .. 64 S 52 ;- ~!) I, ~ ,! t ._f .,

2" 10 Ss. S.. -!: .! ~ It ~ r 11"':'-":}o"iI,lt

co s .. S ~i) U l o I,I~ .~ II * ** ,,~ t ,"

\:- 5 .~ S CS J :C i c.,,. .~. ic:t.~l-:vU:

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.. "UI s.~ t*j , 10 t 1,... t

.:7,..t! 9 :5~ oM 1ft"'" -I t . tl..:~'Wl": I

~7 1S :; ' ';0' to}, ." :1" .f

" ;' ,l,: 3 ' :J . It. . ..... 1.'-. *1 '

TABLE 3. TOLEDO A 1\1 NUAL WIND DATA FOR 1998 TO 2011 Page 11 Page 1 1 of 23

Exhibit 65 The primary intent of this thermal analysis effort was to determine the maximum range in temperatures that occur during the course of the year as well as during a given season.

Consequently, the decision was made to use wind speeds that were more frequently encountered but not necessarily equal to the record wind speed. Consequently, the following wind speeds and directions were chosen at each of the solstices and equinoxes (based upon historical wind data available from http://weathersource.com) .

Summer Solstice No wind with average and record high temperatures 34 mph Northwest wind with average and record high temperatures Winter Solstice No wind with average and record low temperatures 76 mph Southwest wind with average and record low temperatures 105 mph Southwest wind with record low temperatures (1978 Blizzard)

Autumn Equinox No wind with average and record high temperatures 34 mph Northwest wind with average and record high temperatures Vernal Equinox No wind with average and record high temperatures 34 mph Southwest wind with record low temperatures The choice of a northwest wind during the summer and autumn would minimize the convective heat loss on the southwest wall which is always the warmest. Conversely, a southwest wind would maximize the heat loss on the southwest wall . This would maximize the effect that wind has on the temperatures in the southwest wall during the course of a single year.

To maximize the effect of radiant heat loss at night as well as to better simulate the effect of large bodies of water on local wind patterns, it wa s assumed that winds diminished by 80 % after midnight. Typically onshore winds prevail during th e day and gradually diminish after sunset. This is primarily due to the high specific heat of water which causes any large body of Page 12 Page 1 2 of 23

Exhibit 65 water to heat up more slowly than the su rrounding coastal areas during the day and cool off more slowly at night. Consequently air over th e coastal land will hea t up faster and rise, creating a daytime onshore convection pattern. Conversely, after sunset the gradient in air temperature over water and land will weaken and the winds will abate.

One event that was of particular interest was the 1978 blizzard which struck the Toledo area on January 25th through Janua ry 27th 1978. This blizzard is listed as one of the worst on record to hit the Great Lakes area . According to NOAA records, wind speeds in the Toledo area attained 105 mph and were from the southwest. The 850mb temperature contours for this event are shown in Figure 9 for midnight January 27, 1978. This plot reveals a temperature of approximately -20 of in th e Toledo area . This matches up very favorably with the winter solstice temperature profile shown in Figure 9 which had a -24 of nighttime low with a -14°F daytime high .. This is approximately equal to the -16 of daytime record low shown in Figure 9 for January. Consequently, this temperature profile was used to simulate this event. It must be emphasized that this is not the surface temperature but rather the temperature associated with elevation where the pressure is 850mb, or approximately 4500 feet. This was done since the primary mechanism for heat loss at night in the absence of any strong wind will be radiation to th e night sky and not the surface air which is transparent to radiation. To ensure an accurate calculation of the radiant heat loss, it is imperative that a value for the temperature of the night sky be used which is more indicative of a higher elevation.

Page 13 Page 13 of 23

Exhibit 65 T~iPprs, HGTprs at OOZ F ri 27 jan 1978 J:~.;:;- i r:.

, Iff' i

" 1 ' 1 '/ *

-: ~":::,-'

.... :'2~[, .

\

I ' i

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- 30 20 10 -5 o 5 10 15 20 25 FIGURE 9. 850mb TEMPERATURES FOR JANUARY 27, 1978 (OF)

Another event of significance that was examined was the January 1977 blizzard which struck the Toledo area on January 28th and 29th. The minimum daily and noontime temperatures for January 29th are shown in Figure s 10 and 11. The peak wind gu st and direction recorded for thi s blizzard is shown in Figure 12. An examination of these figure s reveals that the minimum temperature for this event wa s: approximately -8°F and the daytime high temperature was between 10 and 20 OF. This was acc?mpanied by peak wind gusts of 73.5 knots, or about 84mph. For the purpo ses of modeling this event, a -8 OF minimum temperature w as assumed which occurred after midnight. A max daytime temperature of 20 OF wa s used which wa s once again as sumed to occur at 1:30 pm in the afternoon. The ambient temperaLure was gradually decreased to it s minimum at abou t 2:00 am and remained constant until sunrise when it gradually increased to its pea k daytime valu e. Since t he ma x sustained wind speed will be less than the peak gust speed, a slightly lower 76 mph southwe st wind was as sumed. The daily temperature profile that was used to model t hi s blizzard is shown in Figure 13 . In compari son to the winter sol stice record low profile shown in Figure 7, it can be seen that thi s blizzard had con Siderably warm er temperatures.

Page 14 Page 14 o f 23

Exhibit 65 FIGURE 10. JANUARY 28,1977 BLIZZARD MINIMUM DAILY TEMPE RATURES (OF )

n~psig995. PRMSLmsl at 122 Fri 28jon 19 77 20 - \Q 0 i., 20 30 ~o FIGURE 11. JANUARY 28, 1977 BLIZZARD NOON TEMPERATURES (OF)

Page 15 Page 15 of 23

Exhibit 65 FIG URE 12. JANUARY 28,1977 BLIZZARD PEAK WINO GUST DIRECTI ON AND SPEED LEGEND

- - WINTER 1977 EUZ z/*RD PROFILE 210 HO LL

(!)

IJ.I Q 7 GO I.U 0::

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-7 ~O

-14.0 TIM E FR o tA SUNR SE( HO RS, FIGURE 13. JANUARY 1977 BLIZZA RD DAILY TEMPE RATURE PROFILE Page 16 Page 16 of 23

Exhibit 65 To determine the surface convection loss due to wind, a detailed FLUENT based computational fluid dynamic (CFO) model was generated. This model incorporated not only the concrete shield bUilding but also the auxiliary building as well. The auxiliary building was needed since its close proximity would strongly affect the flow field . Individual analyses were run for each of the wind conditions that were being examined. FLUENT has the ability to export the surface mesh that was used for concrete shield building CFD analysis along with the corresponding pressures and convective heat transfer surface film coefficients in a standard NASTRAN bulk data format. This permitted importing the surface mesh with the pressures and heat transfer coefficients directly into the PATRAN database containing the thermal model. The pressures and convective heat transfer coefficients could then be mapped onto the exterior of the NASTRAN thermal model as well as the NASTRAN and ABAQUS structural models using mapping routines specifically designed to accommodate dissimilar meshes. An example of the convective heat transfer coefficient that was calculated for a 105 mph southwest wind is shown plotted on the concrete shield building in Figures 14 and 15. Here the convection coefficient on the windward southwest face and the leeward northeast face are plotted. Note the drop in the magnitude of the coefficient on the leeward side as well as the swirling pattern induced by flow separation. In addition for the flute located approximately 90 degrees from the flow direction, there is a sharp augmentation in the heat transfer coefficient. This marks the onset of flow separation. The extreme non-uniformity in the heat transfer coefficient clearly illustrates the high degree of inaccuracy that would have been introduced by assuming a simple uniform value obtained from an empirically derived relationship for a cylinder subjected to simple cross flow.

Page 17 Page 1 7 of 23

Exhibil 65 Patlan 201 0 64-811(MD Enab ed) 29-J "n- 2 15.40.1 5 3.12-001 2 9 -001 2 .7 -00 1 2 51-0 01 2 30-00 1 2 10-00 1 I 90-001 159 01 1<1 9-001 12lH)01 1 0800 1 83H)02 6 .78-<l02 4 .i 5-0o;>

27 H)02 6 SO- 03 2o FIGURE 14 . W IN DWARD SIDE FORCED CONVECTION CO EFFICIENT (BTU! HR-IN P) FOR A 105MPH SOUTHWEST WIN D Page 1£l Page 18 of 23

Exhibit 65 ralran 20 I 0 6 -811(I'>ID Enaoled) L9-Jar.- 2 155306

-138-001

'1 00-001 3 "2-<lOI 2 94-001 2 >7<)01 2 03 001 1 ~O-<JO I 1 5 1-<l 1 1 22-001 l' 38-00, S Fi l -CO~

3 6lj-0Q~

7 85-0C3 2

FIGU RE 15. LEEWARD SIDE FORCED CONVECTION CO EFFICI ENT (BTU! HR-IN °F) FOR A 105MPH SOUTHWEST WI ND In the absence of any wind, the free convection will occur along the exterior of the concrete shield building. Using empirical derived relationships for free convection from vertical walls and horizontal plates (see Heat Transfer by J.P. Holman, 1972 McGraw-Hili), the free convective coefficients were calculated. Initial estimates were made for the difference in ambient and concrete surface temperatures in deriving these coefficients . In general, the free 2

convection coefficients were on the order of .004 Btu/hr-in OF for the dome and .006 Btu/hr 2

in OF for the lateral sides. The small magnitude or these coefficients resulted in only a minor amount of heat transfer to and from the exterior of the concrete shield building. Rather radiation heat transfer effects dominated in conditions or little or no wind. Similarly, the air flow in the interior annular air passage between the outer concrete shield building and the inner steel containment was included, but had little if any effect on the interior temperatures of the concrete. Based upon a 55,000 cfm flow rate (Bechtel CALC 68-011r1, January 1968) and an 2

approximate 1900 ft cross-sectional annular area (based upon a 834 and 780 outer and inner radius) resulted in an average flow speed of only .48 ft/sec. Using the properties for 80 OF air, 2

this gave a convective heat transfer coefficient of only .00054 Btu/hr-in OF. Once again the Page 19 Page 19 o f 23

Exhibit 65 radiant heat transfer from the 120 of inner steel containment would dominate and dictate the temperature at the inner face of the concrete shield building.

2.4 THERMAL PROPERTIES Initially the Crystal River CR3 measured thermal properties were used. These properties are summarized in Table 4. These properties compared closely with the published generic properties for concrete. Based upon these thermal properties, transient thermal analyses were conducted for both the solstices and equinoxes using the average and record temperature profiles with and without wind . The time of day when temperatures peaked on the southwest wall were subsequently mapped onto both the NASTRAN and ABAQUS structural models as well as selected detailed ABAQUS featuring a highly refined mesh to improve stress recovery. These temperatures were then used to conduct thermo-structural analyses to determine the radial stress in the outer concrete shield building. Based upon these initial analyses, those ambient conditions were identified that produced the highest tensile radial stresses in the southwest wall of the concrete shield building.

PClge 20 Page 20 of 23

Exhibit 65 From "Davis-Besse" Drop Bo x "Materla'_Propertles_Comparislon xis" dated 10/31/20,. 1._1 _ _ _ _ _ _ _ _ _ _ _ _ _.....,

Davis-Besse Concrete Thermal Data per Or. )(1 Me chanical Properties Davis-Sessa Calc

( -CSS -099.10-OS4 nJ Te mp O!(,HU '~ O.2rtslty,w co nduct ivity. K 01 tustvity, a Specific HeJt, {~ l mi~ sivrty, ( Tt,~tm4J f): ~ruio ~

rCi Ikalm') ,W/ m*' K) 1m I~ec. x 10' KJ/kg**K It)l!f 'c 43.0 2**0 0.73 1.310 0,3597 0 .928 0.93 6.8<<.00 (f) (ftI/II'l ') (Btu/hr- ln.* f) (lnt/hr) B ruJr b-~F I" / Iflj-F (l0" ) IbsJml (xl 0')

109 .4 0.0867 0.091 4.797 0. 2 22 1i!J1 1.80 3.tiO 0.25 1].7';/l w* ~ I I J A.:."",~j !ono/M;)IUI r . ~':lr"'~""1bic:4 ;O Genf.'ric St eel - Rebar & liner Ecke rt & Drak e , "Heat & Mass Transfer" Mechanical Prope t'ties Oavl s-aQ,sse COllie

( -CS S-099 .l0-0S.:l ,0 L Te:rr.. !r Jtur~. I l)en\ it y. w J Cond ud ivHV. K I OllhaMty.a I s~ ~ l nc H~at, c. I

Emi ssIVI ty, r: h~ tM .&I- xoam.ian '- 'l'oun a.~" M odul:.l.. l I P(,tiislon s RtltJo....

,'r) I lIl/ll'l/-r(:d O~ Ibs/li')~ (d Q'l I I l a t lJ/h r-in "~': ) ~H u/lb_ O F I I I

{Ib/I!) I) Ilo1jh r) 1).:).0 1'-281 2 .61':: 1 .-<1',,) I e l5 . 07'~ 0.111) 0.25

11) As$u~ LCS I) I : iU Q81) t' _"J",",\~

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~ When mat£:t1a !.:! enslty IS In SI.e:6..0-Bdtlw.llID \~ ~u.nr f..." R.ad..;.IIII"I:S Mas> Unit; (~ = c** 336.4 Ini.io?(l t .1 ~)E- l l Dru*w rJl~ '~'

TABLE 4. INITIALLY A SSUMED DAVIS-BESSE THE RMAL PROPERTIES During the later phases of this project, the actual thermal properties for the Davis-Besse concrete were measured. These are shown in Table 5. Compared to the initial CR3 based properties, the new properties were con siderably different and atypical of concrete. The thermal conductivity and specific heat were considerably higher and exhibited an increased sensitivity to temperature. The higher conductivity would result in a greater depth of heat penetration characterized by less steep thermal gradients through the wall of the concrete shield building. The higher specific heat would result in a lower rate of temperature change.

This was reflected in subsequent thermal analyses conducted with these revised properties.

With the initial CR3 properties, repetitive 24 hourly daily temperatures could be achieved in about 2 days after beginning from a uniform soak temperature eq ual to the seasonal minimum .

In comparison, with the revised properties, it took nearly 3 days to achieve repetitive 24 hours2.777778e-4 days 0.00667 hours 3.968254e-5 weeks 9.132e-6 months temperatures.

Page 21 Page 21 of 23

Exhibit 65

\bterial Proporlies for Da\i,-B.ss. Onr.1I Global fiuito Elomonl ~lod.llfDII Id.,lizalions US8R r~ sr Va(tJes per (' -l'IIoj( memorondum I ! Jonuory l Ol l Themrot Properr;es Mechonical Propertic) Foiluli! lriuria

~1t 1l -~11 1 ~1l u~.t:

u*"""U "*11 ~HtN-U !l~U U-J'IboJJ fer.lp.ut"". Oon.,,*, ..,,' "" ' '""",,, ,." Diti" Mty, "" ~ pE"c lfl c H" .I. 0;,

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ThcrnulE..tarwon , r I Younr; "1 M l)du~a. E . POfUO"l ih IIO. ';'" Fll!l.tfSirolU,r, FraCllltewq:V. G, r FI I'b/" I (Blu/tu 'In-"f) l"ih'l B;u{lb-" "1 _fFl,lO I I M/ I'I'!x10 , {'biro'l (1:"~1

~ 50 MaGS O~ 7632 0..478 O~I 510 4.>1 C20 £00 os.;

100 Q.}( 3 '614 ;:: 421

~ 1')0 o.,as Slto OllS

1"" ....... '" Ib J CCIInmlI to ca

KJtw*c,l USH~l'CItt 1.-. 7.soo~"'~t latK'!'~pI!Icr. )j ltCXO'..*... jI<<'Drlt

... =r.,NUJMU ~lo.c!1 Generic Steel - Rebar & Inner Steel Containment Thermal p(ope(l;~ Mechanical Properoe)

I I Te-m petltl.lre' I Co Illn! (:M1.Y, ~ I DtrftollMry.CI J 'P<<rl'', ~ J "(:nIJ.J iy1f'lf, t J 'u;" etM .I .Uq)1I'WOn I VOUBC(MadUM ,l. I I ODJ.OJ1'J"I1Io, ~

I

' !eMily, Yo In/~rF ldO' 1 lbi/in 1<10~

I iFI 80 II>I<nl 0.2.82 I {Bfu/hr'!O-~1 2Ei70 I (in/hr) 86.074 I Blu/lb: f u 0..110 0.25

! l ) ~\J3 I 680 I 2900 I OJO T obi. 1. Connote & ~teel 4.~ ylhtn mlteriildel'lllty13 in s..fz..&=mC"""", r.. b!w.oo Mm Uniu '" =c., * "6.~ InI1t( 11\Gl!*1I B..IJ ",.-'8.'

Notes on ibld~lion WN on T~itnllhe fm31 Analysis for tflellf\il tranSient be.tlfinsf~, mode ls ridQtion berwtt n lh~ site/ ind tht conaele" uflud IS I simple ~diition 10 ilmbient temptflt"rt 01 120'"F. :t ovn Illumed th~ neeltemp YlU In'JilriilnL Trw " UhOl hctor W.iIi Y!t to 1 lind lfIl!!

i1 bsorptr,ity of ti1b concrete W.illlel 10 0.6. Thii ilvoids ftefiltl!! on IN> elflntnt by E:emint \jjfwflnor \\"h~d1 tpEEcll up Iflr trinlient Inilli'JiI. AI\O tbis ilvoidl h.il";inrto (ofuidrr the proptttit.s of the S1Cootl for rldiition.

'i.lt: Maleflll prOperly values fOl 5\1bwodels represent:"ivco{ speCific regions ""Y "Otclloc alized mea~"elllent>.

TABLE 5. FINAL MEASURED DAVIS-BESSE THERMAL PRO PERTIES Thermal analyses were repeated using the measured properties. A t emperature dependent conductivity and specific heat were substituted that matched the measured data.

For temperatures less than or equal to 50 OF , the lowest temperature where measurements were made, constant values were used equal to the 50 OF values. Similarly, the 250 OF values were used for all temperatures greater than or equal to,250 OF. Rather than rerun all of the original cases, only those previously identified six cases :were rerun with the new properties where tensile radial stresses were greatest. The winter so'istice with a 105mph wind and record low temperatures as well as with no wind and average temperatures were also rerun . These last two runs corresponded to the 1978 severe blizzard conditions and the warmest typical winter conditions where free ze/thaw could occur daily. An additional run was also made of the January 1977 blizzard, the second worst blizzard on record in the Toledo area. This blizzard was characterized by high winds but significantly warmer temperatures compared to 1978 event.

Initially, all winter runs produced compressive radial stresses when the expansion effects due to the freezing of entrapped water were ignored. The intent of the winter runs was to determine if the compressive radial stresses were eliminated with the new concrete properties .

Page 22 Page 22 of 23

Exhibit 65 To facilitate the determination of the time of day when radial stresses maximized, a highly detailed NASTRAN plane strain model was generated of the entire cross-section of the concrete shield building that corresponded to an elevation 8202 inches. This model featured a highly refined meshed that greatly improved the recovery of stresses. Initial thermo-structural analyses conducted with this model revealed the presence of local maxima in the radial stress in the vicinity of the shoulders on the southwest wall where cracking was observed.

Consequently this provided confidence that the plane strain model could be used to determine when radial stresses peaked.

Temperatures that were obtained from the 3D transient thermal analysis model were mapped onto the plane strain model at 1 hour1.157407e-5 days 2.777778e-4 hours 1.653439e-6 weeks 3.805e-7 months intervals that spanned an entire day beginning and ending at sunrise. Since this model used a plane strain formulation, it could be quickly run, typically less than a few minutes. The resultant time when the radial stresses peaked was determined and the corresponding temperatures were subsequently mapped onto the ABAQUS and NASTRAN l1li models to determine the radial stresses in the concrete shield building.

Subsequent thermo-structural analyses were performed using a temperature dependent CTE that attempted to simulate the expansion due to freezing of entrapped moisture in the concrete. For these analyses, a negative CTE was assigned at those temperatures where the freezing of water and expansion of ice would be the most pronounced.

Results performed with freezing effects indicated the presence of high radial stresses located in regions adjacent to the shoulders where cracking was observed. This condition occurred when the temperatures associated with both the 1977 and 1978 blizzard were used with the revised thermal properties. This insinuated that the observed cracking may have been a product of the low temperatures that accompanied these events coupled with high moisture content in the concrete.

To ascertain if existing cracks would have a propensity to propagate, the NASTRAN l1li model was modified to incorporate a 30 x 30 foot crack located just below the ring girder and that spanne? a flute located on the southwest wall where cracking was observed.

Additional thermal analyses were performed with this cracked model using those conditions that produced the highest and lowest temperatures:

The temperatures obtained from these analyses were mapped onto the cracked NASTRAN _ model in order to ascertain if the stresses induced by thermal effects at the crack tip were sufficient to induce further crack propagation.

-~~~ ~~- --~ -~---- ---- -~~-- - -~- ---~---~--

r d d Page 23 Page 23 of 23

Exhibit 66: Toledo 1978 Weather

Exhibit 66 1978 TOLEDO BLIZZARD WEATHER

SUMMARY

The blizzard of 1978 occurred on January 25-27, 1978 and has been classified as the storm of the century for the Great Lakes and Ohio Valley. This storm was characterized by severe blizzard conditions with 12 inches of snowfall reported at Toledo and 22 inches at Saginaw, Michigan. Constant wind speeds of 45 mph with gusts to 75 mph were recorded in Toledo.

Larger snow amounts and higher winds speeds in excess of 105 mph were recorded just west of Toledo.

This blizzard was characterized by a rapid drop in temperature. This can be seen in Figures 1 through 6 where the surface temperatures for this event are shown beginning at 12 noon on th th January 25 and ending at midnight January 27 .. An examination of these temperature maps th reveals that as late as 6:00pm on January 25 temperatures were at the seasonal daily average of approximately 30 to 40°F. However, by midnight the temperatures had fallen to _50 F in the Toledo area. Daily temperatures were below 20° F by noon the next day. Surface temperatures th continued to drop to near zero degrees by midnight on January 27 .

The corresponding 850mb high altitude temperature contours show an even more dramatic decline in temperature. These are shown in Figures 7 through 10. These temperature contour maps reveal that just prior to the onset ofthe blizzard, the 850mb temperatures were between o and -5 0 F. However, by noon on January 26 th , the 850mb temperatures were less than -10* F and approximately -15 to -20 0 F by midnight on January 2ih. The 850mb are of special significance, especially with regards to radiant heat transfer at night. In the absence of any appreciable cloud cover, any solar heating will be radiated to the distant night sky. The 850mb temperatures are more indicative ofthis temperature, not the surface air temperature.

Conversely, in the presence of heavy low lying cloud cover, heat will be radiated to the sky at a temperature that is more indicative of the surface temperature.

Near the winter solstice, the sun in the Toledo area is very low in the sky, more than 60 degrees from vertical at noon. In addition, during winter at the Toledo latitude, the solar flux is reduced more than fourfold from the summer solstice value. Coupled with this is amount of daylight, approximately nine hours of which the southwest face of the shield building receives less than six hours of direct sunlight. In the presence of a 105mph southwesterly wind as recorded during the 1978 blizzard, the majority of the solar heat would be convected away. The peak temperatures on the exterior ofthe southwest face of the shield building would increase

Exhibit 66 slightly during the day. Instead the exterior temperatures would be defined more by the temperature of the night sky since the primary mechanism for heat loss would be radiation.

For the purposes of assessing the worst case effects of thermal gradients in the absence of any freezing and expansion of entrained moisture in the concrete, an ambient temperature profile was selected that was more indicative of the 850mb temperatures recorded toward the peak of th th the blizzard late on January 26 and early January 27

For this reason an ambient temperature profile was chosen that had a -14 F daytime max and a _24 F nighttime minimum. This would 0 0 be the same as assuming a clear sky prevailed at night. This would necessarily make this analysis extremely conservative with regards to just thermal gradient effects.

Had the night sky been extremely cloudy, the ambient temperatures as well as the exterior concrete temperatures would have been significantly warmer and more representative ofthe surface air temperatures that were recorded for this event. Since radiation at night would still dominate, the peak temperatures in the exterior of the concrete would be approximately 15 to 0

20 F warmer. In the presence of entrained moisture and its expansion, selecting this higher temperature would make any thermal stress analysis more conservative.

Unfortunately, the amount of precise meteorological data available from NOAA, especially with regards to hourly cloud cover, is somewhat sparse for this blizzard. Consequently, the decision was made to choose those ambient conditions that would produce the worst case results when assessing thermal effects.

Exhibit 66 TMPsig995, PR~ASLrnsl at 12Z Wed 25jan 19 78

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r

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-20 -10 10 20 30 40 50 50 70 80 90 FIGU RE 1: JANUARY 25, 197812:00pm SURFACE AIR TEMPERATU RES

Exhibit 66 TMPsig995, PRMSLmsl at 18Z Wed 25jan19 78 j

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- 20 - 1U o 10 20 30 40 50 60 70 80 90 FIGURE 2 : J NUARY 2 5, 1978 6:00pm SU RFACE AIR TEMPERA URES

Exhibit 66 TMPprs, HGTprs at 002 Thu 26jon1978 1530

-30 -25 -20 -15 -10 -5 Co 5 10 15 20 25 FI GUR E 3: JANUARY 26, 197812:00am SURFACE AIR TEMPERATURES

Exhibit 66 TMPsig995, PRMSLmsl at 06Z Thu 26jan 1978

-20 -10 o 10 20 3D 40 50 60 70 90 FIG URE 4: JANUARY 26,1978 6:00am SURFACE AIR TEMPERATUR ES

Exhibit 66 TMPsig995, PRMSLmsl at 12Z Thu 26jan 1978

--.... -- 1():!-j 1011' * ~ , . . ..

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- 20 -1 0 a 10 20 30 40 50 50 70 80 90 FIGURE 5: JANUARY 26,1 978 12:00pm SURFACE AI R TEMPERATURES

Exhibit 66 TM Psig995, PRMSLmsl at OOZ Fri 27jan1978

. t1iJl) ."

. 93.1. ,".

1Q fi I -;-- -~-

10 20

- 20 -10 () 10 20 30 40 50 50 70 80 90 FIGURE 6: JANUARY 27,1978 12:00am SURFACE AIR TEM PERATURES

Exhibit 66 TMPprs, HGTprs at 122 Wed 25jan1 9 78

, 5,O O *-------~- -*-~

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_- 1 ~ U l------------'-,

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-30 - 25 -20 - 15 5 o 5 10 15 20 25 FI GUR E 7: JANUARY 25, 19"78 12:0 0pm 850mb AIR TEMPERATURES

Exhibit 66 TMPprs. HGTprs at GOZ Thu 26jon1978

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- 30 -25 -2D -15 -10 -5 I] 5 10 25 FIGU RE 8: JANUARY 26, 1978 12:00am 850mb AIR TEMPERATURES

Exhibit 66 TMPprs HGTprs at 12Z Thu 26jan1978 j

. 13:.!ll l2'~ 1 1 'ii":

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1530

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-300 -25 -20 10 -5 () 5 10 15 20 25 FI GURE 9: JANUARY 26, 1978 12:00pm 850mb AIR TE MP ERATURES

Exhibit 66 TMPprs, HGTprs at OOZ Fri 27jan 1978

- 30 -25 - 20 -15 - 10 -s o 5 10 lS 20 25 FIG URE 10: JANUARY 27, 197812:00am 850mb AIR TE M PERATU RES

Exhibit 67: CFD Analysis Summary

Exhibit 67 DAVIS-BESSE SHIELD BUILDING ANALYSIS The _ performed for this report includes:

Na surrounding buildings 34mph from the Northwest(summer) 34mph from the Southwest (winter)

72mph from the Southwest (winter)

Surrounding buildings

34mph from the Northwest (summer) 72mph from the Southwest (winter) 10Smph from the Southwest (winter)

Tornado Category F2 Traveled from the Northwest to Southeast Boundary Conditions for the problem consisted of:

Winter Ambient temperature of -13 OF.

Temperature of the shield building remained at a constant 7°F.

Summer Ambient temperature of 104°F.

Temperature of the shield building remained at a constant 130°F.

Results extracted from the CFD:

Pressure distributions on the surface.

Heat transfer coefficients.

Vorticity shedding calculated on the 72mph case.

Model Generation The CFD mesh consisted to create the air volume. The total size of the air volume was a 2,500 ft. diameter with a height of 670 ft. Using a large air volume eliminates any wall effects that might be caused from flow interactin~

~ 1). The element mesh size used for the model was _ _ _ _

_ . Using a . . mesh size allows the vorticity shedding to be captured more accurately.

Page 1 Page 1 of 21

Exhibit 67 FIGURE 1: AIR VOLUME FOR CFD MOD EL FIGURE 2: SHIELD BUILDI NG CFD SURFACE MESH Solution Method Page 2 Pa g e 2 of 21

Exhibit 67 The CFD program used for this analysis was FLUENT version 13, an industry standard and proven analytical code. Some documentation from th e website states "To address the stringent quality requirements of the nuclear industry, ANSYS h as well-documented devel opment processes and verified software rel eases. The compre hensive, best -in class software solutions comply with NQA-l standards for developing software for the nuclear industry, including ANSYS WORKBENCH , ANSYS FLUENT , ANSYS CFX a nd ANSYS MECHANICAL. The capability of FLUENT is us ed throughout the Aerospace and Defense, Nuclear, Automotive and Ma terials and Chemica l The code has been certified The energy and turbulence model was turned on for the solution. With the wind speed being b e low Mach 0.55 the incompressible ideal gas law was used. The shield building analysis without the buildings was done using a steady state solution. The reason for this method was there is not a lot offluid interaction between structures. The shield building analysis with the building was done using a transient analysis solution.

Analysis The first analysis performed was the 34mph case from the Northwest with summer conditions. Figure 3 below show the pressure loads on the windward side ofthe shield bUilding.

-1.33Hl2

-1 B6e-02

-2.40e-02

-2 .94e-02

-3.47e-02

-4.01 e-02

-5 .08e-02

-5 .62e-02

-6.15e-02

6.6ge-02

-7.23e*02

-7.76e-02

8.30e-02 FIGURE 3: 34MPH NW SURFACE PRESSURES (Psi)

Page 3 Page 3 of 2 1

Exhibit 67 1.60...e1 1Age+Ot 1.37e+01 126e+01 1.15e+01 1.03e-t01 9.16e+OO 8.02e+00 6.87e+OO 3.44e+00 2.2ge+OO 1 .15e+00 O.OOe+OO FIGURE 4: 34MPH SURFACE HEAT TRANSFER COEFFIC IENT (BtujHr-FtI\2-0F)

Heat Transfer Coefficient Analytical Comparison TOWER = 130°F (S4.4°C)

AIR TEMP = 1 04°F (40°C)

TEMP AVERAGE = 117°F (42.22°C) v = 0.1693 cml\2js k = 0.027 wjm*k Pr =0.71 U = 15.20 mjs (34mph)

D = 44.73m Re = U*D j v Re = 40,159,244 Nu = h*Djk Nu =0.3 + (0.62*ReI\0.5*PrI\0.33)j([1 +(0.4jPrYO.67] 1\ 0.25) *

[1 + (Rej282,OOOYO .625] 1\0 .8 Nu = 38,092 h = (38,092*0.027 w jm*k) j 44.73m h = 22.99 wj m l\ 2*k

0.1761 BTU j h*ftI\2* of h = 4.05 BTU j h*ftI\2* OF (This numb er compares to the front surfac e of the tower.

Region of compariso n is the light blue and cyan). This in d icates the CFD model has predicted th e correct surface heat transfer coefficients shown in Fi gure 4.

The cross section pict ure below (Figure 5) shows th e ve locity contours as the wind impacts the shield building. At slow wind speeds, th e flow m ainl y stays attached exce pt along the top front and aft edge. The flow tries to stay attached while passing by the Page 4 Page 4 of 21

Exhibit 67 tower, but flow separation happens at the bottom half due to a lower pressure region.

The top dome has a profound effect on the flow separation caused from the raised lip.

Another contributor of flow separation is the architectural flute located on the side of .

the building.

FIGURE 5: CROSS-SECTION OF 34MPH NW VELOCITY CO NTOURS (Ft/Se c)

The second analysis performed was the 34mph case from the Southwest winter conditions. Figure 6 below shows the pressure loads on the shield building.

2A21H12

327a.()2

-4 .12.'()2

-4.98.'()2

-5.83e'()2

-6 .68e'()2

-7 .54e'()2

-8.3ge'()2

-9 .25.-02

1 .01e'()1

-1 .10e01

-1.18e-01

-1 .27e-01

-1.35. 01 L

FIGURE 6: FRONT VIEW FIGURE 6: 34MPH SW SURFACE PRESSURES (Ps i)

Page 5 Page 5 of 21

Exhibit 67 1.51 e+OO L

FIGURE 7: 34MPH SURFACE HEAT TRANSFER COEFFICIENT (Btu/Hr-FtJ\2-0F)

Heat Transfer Coefficient Analytical Comparison TOWER = -13°F (-25 °C)

AIR TEMP = rF (-13.9 °C)

TEMP AVERAGE = -3 °F (-19.4°C) v = 0.1168 cm"2/s k = 0.02248 w/m*k Pr = 0.72 U = 15 .20 m/s (34mph) 0= 44.73m Re = U*D / v Re = 58,210,273 Nu =h*D/k Nu = 0.3 + (0 .62*Re " 0.5*Pr"0.33)/([1+(0.4/Pr),,0.67]"0.25) *

[1 +(Re/282,OOO),,0.625] "0.8 Nu = 55,111 h = (55,111

0.02248 w/m*k) /44.73m h = 27.7 w/m"2 *k

0.1761 BTU / h*ft"2

of

=

h 4.87 BTU / h*ft"2* of (This number compares to th e front surface of the tower.

Region of compari so n is th e light blue and cyan). This indicates the CFD model has predicted th e correct surface heat tran sfer coefficients shown in Figure 7 .

Th e cross se ction picture below (Figure 8) shows the velocity contours as th e wind Page 6 Page 6 of 2l

Exhibit 67 impacts the shield building. A cold dense air has a tendency to shed from structures more easily due to a higher Reynolds number. During winter conditions, the flow separates completely from the tower at 34mph. A result of this is sever vorticity shedding. An effect offlow separation at lower speeds will cause a cyclic pressure load on the containment tower. The top dome has increased the effect offlow separation.

Another contributor of flow separation is the architectural flutes located on the side of the building.

3.540'42 27~2

.115..02 U8..o3

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-1.18.-01

127,-01

-1.35.-01 1.1 FIGURE 8: CROSS-SECTION OF 34MPH SW PRESSURE CONTOURS (Psi)

The third analysis performed was the 72mph case from the Southwest with winter conditions. Figure 9 below show the pressure loads on the windward side of the shield building.

-1 .06e-02

-345e-02

-583e-02

822e-02

1 06e-01

1.30e-Ol

1.54e*Ol

1.78e*Ol

2.01e*Ol

2.25e-Ol L -2.49e-Ol

-2.73e-Ol

-2.97e-Ol

-3.21e-Ol

-3,45e-Ol

-3.6ge-01 Page 7 Page 7 of 21

Exhibit 67 FIGURE 9: 72MPH SW SURFACE PRESSU RES (Psi) 3.48~01 3 . 27~01 3 . 05~01 2 . 83~01 2.61e+01 Hx AREA OF 2.40e+ 01 COMPARISON L

FIGURE 10: 72 MPH SURFACE HEAT TRANSFER COEFFICIENT (Btu / Hr-FtJ\2 -0F)

Heat Transfer Coefficient Analytical Comparison TOWER = -13°F (-25°C)

AIR TEMP = 7°F (-13.9° C)

TEMP AVERAGE = -3 °F (-19.4°C) v = 0.1168 cm"2/s k = 0.02 2 48 w/m*k Pr = 0.72 U = 32. 63 m/s (72mph)

D = 44.73m Re = U*D I v Re = 124,9 60, 6 0 7 Nu = h*D/k Nu = 0.3 + (0 .62 *Re" 0.5 *Pr"0.33)/([1 +(OAIPr},,0.67J "0.25) *

[1 +(Re/ 282, 000)" 0. 62 5J " 0.8 Nu = 97,0 32 h =(97,032

0.02 248 w/m*k) 1 44.73m h =48.76 w/m"2 *k

0.1761 BTU I h*ft"2* OF h =8.587 BTU I h *ftJ\2* OF (This numbe r co mp ares to the front surface of th e tower.

Region of com p ariso n is th e light blu e and cyan) . This indi cates t he CF D model has Page 8 Page 8 of 21

Exhibit 67 A\

p redicted the correct surface heat transfer coefficients (Figure 10).

The pressure contours have not dramatically changed from the 34m ph, but the pressure load and suction has increased (Figure 11). Cold dense air has a tendency to shed from structures more easily due to a higher Reynolds number. During winter conditions, the flow separates completely from the tower at 72mph (Figure 12). An effect of the flow separation at higher speeds will cause more cyclic pressure loads on the containment tower. The top dome has increased the effect of flow separation. Another contributor of flow se paration is the architectural flutes located on the side of the building.

132.02

-1060-02

-3.45e-02

-583e-02

-a22.ol!

FIGURE 11: CROSS-SECTION OF 72MPH SW PRES SUR E CONTOURS (Psi)

Page 9 Page 9 of 2~

Exhibit 67 FIGURE 12:CROSS-SECTION OF 72MPH SW VELOC ITY CO NTOURS (Ftj Se c)

Creation of Combined Shield and Auxiliary Building CFD Model The CFO mesh for the combined model consisted of 3.26 million cells to create the air ht of 670 ft.

Using a small mesh siz e will allow vorticity shedding to be captured more acc urat ely.

FIGURE 13: AIR VOLUME FOR COMBINED SHIELD AND AUXILIARY BUILDING CFD MODEL Page 10 Page 1 0 of 21

Exhibit 67 FIGURE 14: CONCRETE SHIELD BUILDING WITH ADJACENT AUXI LIARY BUIL DING The fourth a nalysis consisted p erforming the same previou s an alysis with the au xiliary building present. This analys is w as run at 34mph from the Northwest with summ er conditions.

FIGURE 15: 34MPH NW SURFACE PRESSURES (Psi)

With th e a ddition of the surrounding buildings, the pressure has increas ed by O.027ps i Page 11 Page 11 of 21

Exhibit 67 from the previous analysis without the buildings (Figure 15). With the addition of the surrounding buildings, the heat transfer coefficients on the backside of the tower are not as uniform (Figure 16).

I lire-OJ 1 64e'-Ol 1:;Je-Ol 1370-+01 1 23e+lll 1100-+01 9 se.*O!l IIlle-+OO 684...130

, ,4Be- OO

~ 110-+00 274. +130 1 37e"' OO 000.*00 FIGURE 16: 34MPH SURFACE HEAT TRANSFER COEFFICIENT (BtujHr-Ft"2-0F)

Pressure contours have dramatically changed with the addition of surrounding buildings. There is a large low pressure region located above the buildings on the aft side of the containment tower (Figure 17). The velocity vectors are disrupted from the buildings causing the flow to separate at lower wind speeds (Figure 18).

FIGURE 17: CROSS-SECTION OF 34MPH NW PRESSURE CONTOURS (P si)

Page 12 Page 12 of 21

Exhibit 67

')"0'

r O' Large area of 11 ' disrupted flow.

FIGURE 18: CROSS-SECTION OF 34MPH NW VELOCITY CONTOURS (FtjSec)

The fifth analysis consisted of setting th e boundary conditions to 72mph from the Southwest with th e buildings at win te r conditions. With the addition of the surrounding Au xiliary building, the pressure has increased by 0.OS4psi (Figure 19).

With the addition of the Auxiliary building a nd the resultant increase d wind speed, the heat transfer coeffici ents are more disrupted (Figure 20).

.1 HIe-or

., 58e-OI

.19Se-Ol

.2380-01

2 ;eo-or

-3190-01

359.*0 r

399..01

439.*01

-4 7ge*01

5 190*01

-5590-0 I

-5 gge-0 1

-6 .*0e-0 I FIGURE 19: 72MPH SW SURFACE PRESSURE CON TO URS (Psi)

Page 13 Page 13 of 2l

Exhibit 67 I SoI** OI 141le'O!

1 2S~'Dt I I~.*DI 101 -..00 FIGURE 20: 72MPH SURFACE HEAT TRANSFER COEFFICIENT (B tu j Hr-FtJ\2-0F)

The stagnation pressure region has shifted up towards the top of the containment tower (Figure 21). This is a result of the surrounding Auxiliary building being partially in front of containment tower. The flow on the aft side of the tower is turbulent compared to the case with no Auxiliary building (Figure 22).

FIGURE 21: CROSS-SECTION OF 72MPH PRESSURE CONTOURS (Psi)

Page 14 Page 14 of 21

Exhibit 67 FJGU RE 22:CROSS-SECTION OF 72MPH SW VELOCITY CONTOURS (FtjSec)

The sixth analysis consisted of setting the boundary conditions to lOSmph from the Southwest with buildings at winter conditions. The pressure on the windward side of the containment tower has significantly increased (Figure 23). In addition, the heat transfer coefficient stagnation area for the lOSmph case has dramatically decreased on the windward side (Figure 24).

26<1e-U1

-3 53e 01

..j 43e-01

S33e-Ol

en...ol

711e*Ol

8 Ole-O1

8.91.*01

9 ale*Ol

I07e+00

I 16e+00

. I ]5e+00

1 34e +OO

- I 43e+OO FIGURE 23: 105MPH SW SURFACE PRESSURE CONTOURS (Psi)

Page 15 Page 1 5 of 21

Exhibit 67

__ '01 15,.,-01 135;: - 01 11&e - UI 9.£610 -00

~~ -00 5 79<- 00 JS6e

ijij I~..m GUo.*w FIGURE 24: l05MPH SURFACE HEAT TRANSFER COEFFICIENT (BtujHr-FtI\2°F)

The stagnation pressure region has shifted up towards the top of the containment tower (Figure 25). This is a result of the Auxiliary bUilding. being partially in front of containment tower. The flow on the aft side of the tower is unsteady and turbulent (Figure 26). The addition of the Auxiliary building has caused the flow to rise due to the pressure increase just before the shield building. This results in a higher pressure region at the midpoint causing a larger overturning moment. The two figures below show pressure and velocity contours at l05mph during winter conditions.

FIGURE 25: CROSS-SECTION OF lOSMPH SW PRESSURE CONTOURS (Psi)

Page 16 Page 16 of 21

Exh ibit 67 FIGURE 26: CROSS-SECTION OF 10SMPH SW VELOCITY CO NTOURS (FtjSec)

To show the flow disturban ce created by the containment towe r and Auxiliary building, two figures are shown below of the velocity streamlines (Figure 27). A large wake is created downstream and the flow does not reattach within the fluid boundary defined for the analysis model (Figure 28).

FIGURE 27: 10SMPH SWVELOCITY STREAMLINES (Ft j Se c)

Page 17 Page 17 of 21

Exhibit 67 FIGURE 28: TOP VIEW OF l05MP H SW VELOCITY STREAMLI NES (FtjSec)

The last analysis that was performed for the Davis-Besse shield building was of the 1998 tornado. This tornado was classified as F2 tornado with wind speeds between 113 and 157 mph. The base of the tornado was estimated at 100 yards wide and touched down just west of the Davis-Besse facility between 8:45 and 9:00pm on June 24, 1998. The tornado traveled just past the shield building in a southeasterly direction for a distance of3.5 miles.

The tornado was a much more complicated analysis. To prop erly model the tornado, a transient analysis was performed using the FLUENT large eddy solver. This allows the ture data at selected time st Ie stea state solution.

The tornado was created using a 300 foot wide circular face on the ground and top boundary surfaces of th e CFD model control volume. This allowed a ro tational bound condition to be used Page 18 Page 18 of 21

Exhibit 67 3 17e-+02 2.88e-+02 2.60e-+02 231e-tm 2.02e-+02 1.73e+02 1.44e+02 1.15e+02 8.65e+01 5.n e+01 2.8ge+01 6.93e*03 FIGURE 29: VELOCITY VECTORS SIMULATlNG A TORNADO (FtjSec)

Once the tornado was stable, the process began to move the domain with the tornado towards the buildings. The tornado characteristics were obtained from the NOAA website where the direction and forward velocity were listed , These boundary conditions were used to move the tornado pass the containment tower. To capture 9 seconds of data, the FLUENT CFD analysis ran for about 30 hours3.472222e-4 days 0.00833 hours 4.960317e-5 weeks 1.1415e-5 months . During this time, FLUENT was programmed to capture pressure contours every second so a MPEG video could be created at the end of the analysis. Also during the analysis, data was exported to capture the pressure loads as well as the heat transfer coefficients.

During the analysis, one could observe when the tornado got closer to the buildings as the vortex structure of the tornado started to break down, This results when the vortex flow at the base of the tornado is disrupted by the presence of the buildings. Since the shield and auxiliary building are comparable to th'e width of the base of the tornado, they constituted a major impediment to the flow of the tornado. Consequently, at its closest approach, which was assumed to be 100 yards, the pressures induced by the tornado were attenuated. If there is enough energy, the tornado will reform once any obstruction is gone. This behavior is evident below in the images of the pressure contours as the tornado passed by the shield and auxiliary buildings (Figure 30 and 31)

Page 19 Page 19 o f 21

Exhibit 67 11.36sec 12. 36sec 13.36sec FIGURE 30: SURFACE PRESSURE CONTOURS DURING THE PASSAGE OF THE 1998 F2 TORNADO 17.36sec FIGURE 31: SURFACE PRESSURE CONTOURS DURING THE PASSAGE OF THE 1998 F2 TORNADO Page 20 Page 20 of 21

Exhibit 67 Estimation of Surface Pressures and Convection Coefficients for Other Wind Speeds CFD analyses were performed only for 34, 72 and 105 mph speeds. However, thermal stress analyses were also performed for 36 and 76 mph south and north westerly winds as well. If there is no change in the wind direction, then the surface heat transfer coefficients derived from the CFD analyses can be scaled to estimate the corresponding values for slightly higher or lower velocities.

For a free standing cylinder subjected to a cross flow, the convection coefficient is proportional to Res where Re denotes the Reynolds number and is given by the expression:

Re = pv 0

~1 where:

p = density v= velocity

[l = viscosity Since the Reynolds number is proportional to the velocity, the surface convection coefficient will scale as the ratio of the velocities to .8 power. In this way, the convection surface coefficients at other velocities could be derived from the CFD analyses for winds traveling in the same directions analyzed.

Similarly, the surface pressure on shield building will vary as the square ofthe velocity.

Consequently, the surface pressure at slightly different wind speeds was derived by scaling the CFD derived pressures by the square of the velocity ratio. Once again, this methodology will only work if the direction of the wind remains unchanged and the change in wind speed is not appreciable.

Page 21 Page 21 of 21

Exhibit 68: Debonding of Rebar/Concrete Lab Testing

Exhibit 68 Pedormance Improvement Intemafional Providing a competitive advantage through research and applications To:

From Date: 02/27/2012

Subject:

Laminar Cracking of the Davis-Besse Shield Building - Concrete Sample Testing for Debonding Based on my observation and examination of concrete-core samples received from the Davis Besse Shield Building, my findings for Debonding are detailed in what follows.

Page 1 of 10

Exhibit 68 Concrete/Rebar Interaction Concrete/Rebar Interaction Laboratory testing and examination was conducted on the only two concrete/rebar interaction samples received from Davis-Besse, F2-790 .0-4.5 and S13-633 .08. The evidence indicates that there is potentially a lack of bonding between the concrete and rebar on both samples.

Testing and examination of the concrete/rebar-interaction samples shows the imprint from the rebar onto the concrete as well as iron oxide transfer.

Concrete/Rebar Interaction: Lack of Adhesion The following plant photo shows the imprint of rebar onto concrete . The photo documents the measured distance between the crests as 1.0 inches. In addition, it also points out possible rust transfer. To confirm that the imprint originated from concrete/rebar interaction , Figure A2 is a confirmatory photo of a 1.25 inch diameter rebar that shows the same pattern as the imprint. The figure also illustrates the 1.0 inch crest to crest measurement.

Page 2 Page 2 of 10

Exhibit 68 Figure A 1* Shield Building Concrete Sample Illustrating Rebar Imprint Figure A 2: 1. 25" Diameter Rebar with 1.0" Crest-to-Crest Measurement Page 3 Page 3 of 10

Exhibit 68 Figures A3 - A5, which show different views of Core F2-790.0-4.5, illustrate the lack of adhesion between the concrete and rebar, as rebar rib groove imprints are found on the sample .

Figure A 3: Rebar Rib Groove Imprints (Core F2-790.&-4.5)

Page 4 Page 4 of 1 0

Exhibi t 68 Figure A 4: Rebar Rib Groove Imprints (Core F2-790.0-4.5)

Figure A 5: Rebar Rib Groove Imprints (Core F2 -790.0-4.5)

Page 5 Page 5 of 10

Exhibit 68 Concrete/Rebar Interaction: Iron Oxide Transfer

. Figures A6 - A8 and Figures A9 - A 13, respectively show in sequential order

1. Possible iron oxide on the fracture surface ,

2. A potential iron oxide deposit removed and measured for thickness

3. Energy-dispersive X-Ray Spectroscopy (EDS) performed on the prospective iron oxide deposit to confirm iron oxide.

Energy-dispersive X-Ray Spectroscopy is an investigative technique that involves X-ray excitation of a sample. The identification of the samples chemical characterization stems from the interaction between the X-ray and the element's atomic structure .

According to the EDS tests performed , iron oxide was determined to exist on the concrete surfaces under examination.

Figure A 6: Iron OMlde Transfer onto Concrete (Core FZ-790.0-4.5J Page 6 Page 6 of 10

Exhibit 68 Figure A 7: Iron Oxide Deposit taken from Sample (Core F2-790.~.5)

Core F2-790p04p5, Iron Oxide Transfer Onto Concrete at Rebar Imprint Area, 20 kV 0 Fe Cn's 10K 500 Fe Ca IItn Si l~isi Iv! Fe j Jr.. .1 ,J

10. 20 Figure A 8: Energy-dispersive X-Ray Spectroscopy Results from Iran Oxide Deposit (Core FZ-790.0-4 5)

Page 7 Page 7 of 10

Exhibit 68 DavIs Besse Nuclear Plant Core 513-633.08 Figure A 9: Iron Oxide Transfer to Concrete (Core S13-633.08)

Figure A 10: Iron Oxide Transfer Into Macro Cracks of Concrete (Core 513-633.08)

Page 8 Page 8 of 10

Exhibit 68 Figure A 11: Iron Oxide Deposit taken from Sample (Core 513-633.08)

Figure A U: Iron Oxide Deposit taken from Sample (Core 513-633.08)

Page 9 Page 9 of 1 0

Exhibit 68 S13-633 08. Iron Oxide Deposit On Transverse Fracture T4. 20 kV Fe Cnts 1.0K 0

500 Fe Ca Mn AI Je

(

e I~!i~.l ....

~

M

10. 20.

Figure A 13' Energy-dispersive X-Ray Spectroscopy Results from Iron Oxide Oeposit (Core 513-633.08)

Page 10 Page 10 of 10

Exhibit 69: M-284a

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Exhibit 71: Comparison of Toledo Blizzards

4 Exhitiit 71 COMPARISON OF SNOWFALL, WIND SPEED AND TEMPERATURE FOR MAJOR BUZZARDS IN THE TOLEDO, OHIO AREA SNOWFALL SNOWFALL MAX WIND LOW TEMPERATURE 20 DAYS PRIOR DURING (MPH) (DEG F)

EVENT (INCHES) (INCHES)

BLIZZARD 84 (gust) -8, (850mb) 9.1 4 1977 78 (avg., est.) -2, (surface)

BLIZZARD l

-24 (850mb) 1978 21.5 12 105

-5, (surface) i-BLIZZARD 1994 11.9 r 3.5 45(avg., est.) -17, (surface)

Page 1 Page 1 of 1

Exhibit 72: Water and Moisture Transfer into Concrete

Exhibit 72 Water and Moisture Transport Processes in the Concrete Wall

1. Water and moisture transport mechanisms in concrete There are two types of moisture transport processes in Davis Besse containment structure that may possibly affect the ice-induced delamination. One may be called "top-down moisture transport", and the other may be called "Horizontal moisture penetration into the concrete" . The first one resulted in high moisture content near rebar region. When temperature decreases, the moisture near rebars freezes, lead ing to the sub-mode I delamination cracking as described in FM 3.6 Freezing near Rebar in Blizzard. The second one caused high moisture content in the outer layer of concrete, which leads to the sub-mode 2 delamination cracking as described in FM 3.6. The following sections describe the two types of moisture transport processes in Davis Besse containment structure.

Mechanisms 1: Top-down moisture transport The top-down moisture transport process assumes that the water comes from the top of the structure and slowly penetrates down within the concrete wall. The resulting cracking is delamination cracks in the wall caused by ice formation. Due to heterogeneous feature of concrete, the water comes down along random paths with the least resistance. That explains the sporadically distributed cracks in the wall. Patches of laminar cracks can distribute in a large area of the wall connected by narrow water paths behind the surface.

The identified areas of laminar cracking are detailed on Drawing C-III A. The document shows that the patches of cracks are generally located within the architectural shoulder areas of the Shield Building. The crack survey information also shows that in some shoulders the cracks are located a considerable distance (100+ feet below) from the top of the Shield Building wall. The cracks are also located in the top 20 feet of the cylinder wall.

Fig. 1 illustrates the top-down moisture transport process, which can be explained by using several evidences obtained during our investigation.

1. During the construction of the containment structure. The wall was built first and the dome was bu iIt two years and four months later. So, the jacking bars, dense rebars, and the top of the concrete wall were exposed to the environment. The two figures on the right side in Fig.

1 show the exposed wall, the dense rebars in the wall during construction, and dense rebars around the delaminated area. Many structures were built this way without any durability problem. In this structure, however, the initial defects may be generated by the jacking bars and dense rebars, together with the large aggregate used in the concrete. These factors resulted in high porosity concrete near the rebars and jacking bars, which formed a pathway for the water to penetrate down in the wall.

2. The cosmetic layer ofrebars near the surface of wall was evenly spaced at 12". The concrete near the cosmetic rebars has good quality, and thus there is no delamination around this layer ofrebars.

Exhibit 72

3. There is a flat concrete slab surrounding the dome (see the upper left figure in Fig. I) where snow can stay for a long time in the winter after a snow storm. Slow melting of snow provides long-term ponding of water on the top of the wall. During the inspection, a part of flat slab was covered by water, as shown in the upper left figure in Fig. I.

Jacking bars and dense rebars on the wall for more than two years before the dome was built Fig. 1 Failure mode of Type 1 freeze-thaw damage in the wall The concrete in the flat concrete slab exhibited typical sign of freeze-thaw damage (see the upper right figure in Fig. 1), such as scaling damage and cracking. Water can penetrate into the damaged concrete and slowly transport down to the concrete walJ below. Because of heterogeneous concrete properties, the water penetration did not occur uniformly in the wall.

Exhibit 72 The penetration could be faster at one location and stopped at another location, which explain the random distribution of the delamination cracking in the wall. There is no theoretical model that can be used to quantitatively estimate the rate of this type of mo isture penetration.

Mechanism 2: Horizontal moisture penetration into the concrete This moisture transport process assumes that the moisture comes from the side surface of concrete wall and slowly penetrates horizontally into the wall. The concrete with high moisture content may expand when temperature decreases (as explained in FM 3.6 Freezing near Rebar in Blizzard, and in Exhibit 57 for temperature dependent coefficient of thermal expansion). The resulting cracking is delamination cracks in the wall caused by a combined action due to the expansion of outer layer of concrete as well as the contraction of inner layer of concrete. Similar to the top-down mo isture transport process, the heterogeneous feature of concrete needs to be considered. More moisture will penetrate into the concrete wall in the locations where the resistance is low, and less moisture will penetrate into the concrete where the resistance is high.

This explains the sporad ically distributed cracks in the wall.

Water pressure Vapor pressure

,.......- -- 100% relative humidity Water pressure equivalent to 106 mph wind 90% relative humidity The interface between liquid pore water and water vapor

~

310 c )

ro ~

c ~

Concrete Wall Q)

> 310

' ~

-0 I )

-0 C ~

S ~

~

Liquid water Water vapor penetration diffusion Fig. 2 Liquid water and moisture penetration into concrete due to wind-driven rain

Exhibit 72 For above ground concrete structures like Davis Besse containment structure, there is hardly any liquid water in the porous system in concrete under an arid environment, the moisture content of concrete can be described by pore relative humidity, and the water vapor diffusion is a dominant process which is driven by the humidity gradient from the location of high humidity to the location of low humidity. During a blizzard, the effect of wind-driven rain is an important factor to consider. On the outer surface of concrete wall, the rain water is pressed by the wind pressure and penetrates into concrete. The liquid water penetration is a dominant process which is driven by the wind pressure on the surface. Fig. 2 shows the two transport processes in a concrete wall during a blizzard. The blue region represents the outer layer of concrete in which pores are saturated by liquid water, which may be called water region, and the yellow region stands for the inner layer where vapor water dominant, which may be called vapor region. The interface between the two regions is simplified as a line in this study, and in reality, the interface is a zone with a finite depth for the transition of the two regions.

The entire depth of the water region is shown in Fig. 2 as Lw (on the left) , which is important for our study. The partial depth of the vapor region, say higher than 90% pore relative humidity is shown as Lv (on the right) , which is also important for our study. The sum of the two depths is called Lm, representing the depth of concrete with high moisture content :

Lm = Lw + Lv (I)

In Section 2.1, Lw will be estimated by a theoretical model for liquid water transport in concrete taken into account the wind-driven rain effect on the surface of concrete water. In Section 2.2 ,

Lv will be estimated by a theoretical model for water vapor diffusion in concrete. The two estimations need several important transport parameters such as water permeability and moisture diffusivity, which are not available for Davis Besse concrete. So, the material parameters will be calculated by using available material models in the literature. Lm calculated from the estimated Lw and Lv will be compared to moisture profiles measured in a concrete sample from Davis Besse power station. The comparison also serves as a calibration of the material parameters used in the two theoretical models. The calibrated models can be used to better predict water and moisture transfer in the concrete during blizzards and other conditions.

2. Analytical model for liquid water transport in concrete The liquid water penetration process due to wind-driven rain is a very complicated process, which has been modeled by using coupled partial differential equations (Tariku et aL 2007) . In this study, because of limited time, we used a simplified approach which is described by the following governing equation similar to heat conduction:

oW oP _ K 02 p (2) oP 8t - f' ox 2

in which P = water pressure in concrete pores in psi ; W= water content in concrete in gram of water per gram of concrete; t = time in hour; and x = the depth into the concrete wall in inches; oW/oP = water capacity in l/psi (because W is a weight ratio), it is similar to the term of p eC in heat conduction where Pc is the density of concrete and c is heat capacity; and Kp = hydraulic

Exhibit 72 conductivity with a unit of in 2/(psi.hr). From Eq. (2), Kp is related to pressure gradient of water in pores of concrete. The water penetration equation, Eq. (2), can also be written as (3) in which Kw = water diffusivity, similar to the thermal diffusivity for heat conduction. The unit for Kw is in 2/hr.

K = Kp (4)

W aw/ap It should be noted that aWjaH already includes the density of concrete pc. Eq. (3) can be solved with given boundary condition and initial condition to obtain a water pressure distribution in concrete at any location and at any time. The boundary condition pet, 0) = Pb and the initial condition P(O, x) = Po must be prescribed first. Pb represents the water pressure on the surface of concrete wall due to wind-driven rain, and Po represents the initial water pressure in the concrete.

In order to find the internal water pressure distribution, the boundary condition and the material parameters must be determined first. Then, the depth of water region in Eq. (I), Lw, can be determined which is the distance from the surface to a location where the water pressure is considered to be low.

The boundary condition - Water pressure due to wind-driven rain Wind-driven rain (WDR) has a significant effect on moisture content in building envelops.

WDR intensity increases with wind speed (Hens 20 I 0). RlLEM tube method was developed to test the resistance ofa building envelop to WDR, as shown in Fig. 3. The tube is designed to be attached to a vertical wall. The water level in a tube is used to simu late the pressure of WDR at various wind speeds.

RELATIONSHIP BETWEEN ruBE WATER LEVEL AND WIND SPEED Grad<atlon ThllonlucaJ W ind

'" M ..rk(ml) Spa<!d (mph) 0 98. 1 0.5 94. 1 1.0 90.04 1.5 95.7 I 2.0 8L I I 2.5 75 .2 3.0 711 I 3.5 65.5 I 4.0 59.4 I 4.5 52-9 I 5 44.8 I Fig. 3 RILEM tube tester and relationship between water level and wind speed (Crissinger 2005)

Exhibit 72 When the water level reaches the top of tube at graduation mark 0 as shown in Fig. 3. the corresponding wind speed is about 98.1 mph. which is a high Category 2 hurricane (Crissinger 2005). At this water level, the height of water is 12 cm, and the corresponding pressure is 0.17 psi. This pressure level is lIsed in our analysis as the boundary condition.

The hydraulic conductivity Kp From Eq. (4). we need to first determine Kf'. the hydraulic conductivity. and fi WlAP. the water capacity. Kp can be expressed in two different ways as K pI' -~

- (Sa)

'7 K i, k (5b) or I' = - p"g 77 in which k = intrinsic permeability with a unit of in 2* which is the basic permeability of the material independent from the permeating media; and '7 = viscosity oftluid (water) in psi.s.

Viscosity of water '7 = 0.00 I Pa.s = 4.028 x 10. 11 psi.hr. It is important to note that Kp as defined in Eq. (Sa). K;. is related to pressure gradient. the superscript p is for pressure. and its unit is in2/psi/hr. This Kp is used as the field variable in our formulation of the governing equation. Eq.

(2). On the other hand. Kp is often defined as in Eq. (5b), K;: . It is related to the height of water column. the superscript" is for the height of water colunm. and its unit is inls.

Various test methods were developed for K: and K; :pressurized methods tor K; or water column methods for K::. The different testing methods resulted in ditIerent Kp. For the pressurized methods. the water flux J is proportiona I to the pressure gradient. J = K ~ (!::.PI L).

where L is the thickness of specimen. For the water colunm methods. the water nux J is also proportional to the pressure gradient. which is expressed by the height of water column.

J = K~(t:.h/ L). So, if a K; with the unit of inls is available for a concrete. the value must be converted to i,hpsilhr in order to be consistent with the definition of Eg. 5( a) for the analytical solution used in this study (will be described later). The conversion is K li K I' = _P (5c)

I' p ".g The value of K; varies in a very large range depending on the type of aggregate and water cement ratio of the concrete. Mehta and Monteiro (1993) indicated that with different type of

Exhibit 72 aggregate. K:: may vary more than [000 times for different concretes (see Table 5-2 of the reference).

In the case that the intrinsic permeability k is available, Eq. (Sa) can be used. However. the intrinsic permeability k tor Davis Besse concrete is not available. So, available test data in the literature will be used as reference in this study. As an example, the test data of k by Nehdi (1995) is k = 3.875xlO-IO in~ (25xl 0- 14 m~) for a concrete with wlc = 0.5 , entrained air = 2%. and fc'= 6400 psi (44 .2 Mpa) at 91 days, and thus K; = 9.62 in~/psi/hr. Another example is tTom the test results of Wang et a!. (1997) . They showed that K;: depends on the width of crack in 9 2 concret e and it is in the range of 1x 10- cmls to 1x 10- cm/s when the width of crack is between o to 500 /lm (see). As one can see_ the permeability values of distressed concrete vary in a very large range.

' .()DE.01 l00c..o2 .0 I...

1.OClE-03

'_.04 .6 1

U

' .OOE-O>

~ l.QQE.Qe J>

~ l, QI)E..()/

1I , 1IC£.oe

--+-cco ........ Ioedinv

l00E*OI

- . 0- * *am ....."OM!....

~ OCE *'O 1000 000 000 0 200 :>00 c..01< o~..,g D~,"'.nt ,m,eron.)

Fig. 4 Relationship between water permeability and crack widths (Wang et a!. 1997)

Exhibit 72 The depth o/water regio" Lw Using the boundary, condition, initial condition, and material parameters derived above, and a J D solution for Eq. (3) can be obtained P(f, x) = Po - (Po - Pr)crr[ x/

-J 4K" .f 1 (8) in which erf is the error function. Fig. 5 shows that after 4 days and 16 days continuous wind-driven rain at wind 106 hand 20 0 e, the internal water ure distributions in a 0 .20 -- ~

0.18

- After 4 days 0.16 VI

a. 0 .14

<l.I

- - After 16 days

..... 0.12

)

VI VI 0.10

<l.I

a. 0 .08

<l.I

+-'

0.06 ro 0 .04 S

0.02 0.00 o 1 2 3 4 5 6 7 8 9 10 11 12 13 14 Depth from the interface (in.)

During the 1978 blizzard, assuming average T = ooe and wind speed of98 mph. we obtained K/I~ 0.00193 in 2/hr. and the two curves shown in Fig. 6

Exhibit 72 for water pressure distributions after four days and 16 days continuous wind-driven rain. From Fig. 6, we can see that Lw is about two inches after four days and about three inches after 16 days wind-driven rain.

0.18 0 .16

- After ~ y_s Vl 0.14 a..

0.12 OJ - After 16 days

J Vl 0.10 Vl OJ 0.08 a..

" 0 .06 OJ oI-J rc 0.04 5

0 .02 0.00 o 1 2 3 4 5 6 7 8 9 10 11 12 13 14 Depth from the interface (in.)

Fig. 6 Water pressure distributions in the concrete with water diffusivity Kw = 0.00139 in 2/hr. under T = O°C and wind speed = 98 mph.

3. Analytical model for water vapor transport in concrete Water vapor transport is driven by mo isture concentration grad ient. The mo isture concentration is represented by pore relative humidity on concrete. The boundary (the starting point) of moisture diffusion is the interface between the water region and vapor region shown in Fig. 2, where the humidity boundary condition can be assumed to be 100%. This is actually a quite complicated problem, because the boundary (the interface) is not fixed, but moves to the right of the figure, depending on the depth of the water region. In this study, due to the tight schedule, the moving boundary problem is simplified as a one-dimensional moisture diffusion problem as described below.

The governing equation for the I D moisture diffusion in concrete is (Xi et al. 1994a, 1994b):

oW oR o2R

--=K- 2 (9) oR ot H ox in which oW/oR = moisture capacity in gram of moisture per gram of concrete, which is similar to the heat capacity p eC in heat conduction; and KH = coefficient of moisture diffusion, similar to the thermal conductivity. The unit of KH is in length 2/time, depending on the test data used for determining the parameter (will be explained later in detail). The subscript R is used here to represent humidity in concrete. The moisture diffusion equation can also be written as

Exhibit 72

( 10) in which D" = moisture diffusivity, similar to the thermal diffusivity. The unit for D" is the same as K/J because Wand H are weight ratio and pressure ratio, respectively, and thus have no unit.

D = Kf/ ( 11 )

/I rW?H It should be noted that ?W/2H already includes the density pc. Eq. (10) can be solved to obtain a moisture distribution in concrete at any location and at any time, which will be described in later sections. The boundary condition H(t. 0) = H, and the initial condition H(O,x) = H" must be prescribed. In this study. H, represents the relative humidity at the interface. 100%: and HII represents the initial relative humidity in the concrete. In order to find the internal moisture distribution fiom Eq. (10). the material parameters must be determined first.

Moisture capacity and coefficient ofmoisture diffusion Moisture capacity of concrete is not a constant but a function of H. Fig. 6 shows the effects of water-to-cement ratio. w/c. and H on the moisture capacity (Xi et a!. 1994a). The moisture capacity is highly nonlinear with respect to H. In fact. DH depends on H as welL which will be shown later.

1.6 w/c - (l.e T - lli6 OK 1.2 1- 28 daya Type I o<<neI'II

J

.r::

"0 ;7

~

O.B 0.0 .-t---r-----r~-__,_---r---j 0.0 0.2 0. 4 0 .6 03 1.0 Relative Humidity Fig. 6 The effects of water-to-cement ratio and H on the moisture capacity.

Exhibit 72 0.5 . , . . . - - - - - - - - - - - - - - - - ,

Relative humidity Fig. 7 The effect of relative humidity 011 the coefficient of moisture diffusion

Exhibit 72 As a summary, in the following moisture transport analysis

The density of concrete pc = 2400 kg/m 3 = 0.0868 Ib/in 3

aWjaH / pc = 0.35/0.0868 = 4.032 in 3/1b 2 2

The coefficient ofmoisture diffusion KH = 0.146 cm /day = 0.00094 in lhr.

2

The moisture diffusivity DH = 0.417 cm 2/day = 0.00269 in /hr.

The depth of vapor region Lv At the interface Hs = 100%, and at the deep part of the wall Ho = 70%, and the solution ofEq.

(10) is (IS) in which H(t,x) is the humidity distribution in the concrete at time t and depth x starting from the interface. The two curves for after 4 days and 16 days ofWDR are shown in Fig. 8.

1.00

- 0.95 "

>- 0.90 '

- After 4 days

0 0.85

- After 16 days E

l 0.80

~ 0.75

0 .70 ro Q) 0.65

~ 0 .60

~ 0.55 0 .50 o 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2 2.2 2 .4 2.6 2.8 Depth from the interface (in.)

Fig. 8 Internal humidity profiles in the vapor region after 4 days and 16 days ofWDR Hs = 100%, Ho = 70%, and the moisture diffusivity DH = 0.00269 in 2/hr.

Since only the high moisture region has effect on ice formation, we are interested in the depth from 100% to 90% relative humidity. The depth of high vapor density region is OJ in. after 4 days and 0.6 in. after 16 days ofWDR.

After a major blizzard, the effect ofWDR will disappear, but the moisture will stay in the wall in the water region if outside temperature is low. So, the depth of water region will remain more or less the same, while the depth of vapor region will increase because of continuous diffusion of moisture from the water region to deeper part of concrete. Neglecting the temperature effect,

Exhibit 72 Fig. 9 shows two curves of moisture distribution 20 and 80 days after a WDR. The depth of high vapor density region is 0.75 in. after 20 days and 1.5 in. after 80 days of the WDR.

1.00 0 .95

~ 0.90

> 0 .85

t=

'0 0.80 E 0 .75

.r:.

(l) 0.70

- - After 20 days

....>ro 0.65 aJ.... 0 .60

- After 80 days

....(l) 0.55 0

Q..

0 .50 o 0.250 5 0.75 1 125151.75 2 225252.75 3 3.2535 Depth from the interface (in.)

Fig. 9 Internal humidity profiles in the vapor region 20 and 80 days after a WDR Hs = 100%, Ho = 60%, and the moisture diffusivity DH = 0.00269 in 2/hr.

4. The depth of high moisture region Lm = Lw + Lv The material models used in the water and vapor transport analyses were developed not based on Davis Besse concrete, because the transport parameters for the concrete are not available.

Considering the large variation in the transport parameters, the variation of model predictions must be taken into account. One possible variation is the effect of aggregate and air-entraining agent used in Davis Besse concrete, which may increase the moisture diffusivity and thus increase the rate of moisture penetration (Xi 1995a, 1995b). An important variation is the distress in the concrete wall such as the surface cracking as shown in Fig. 4. The effect of internal damage on concrete diffusivity was analyzed by many researchers (see for example Xi and Nakhi 2005).

As a summary, the depth of water region Lw is much higher than that of vapor region Lv. Based on above preliminary and approximate analyses for solid concrete without major distress, the depth of high moisture region Lm is about 2 to 3 inches after a few days of WDR. This may be considered as a reference or guideline for determining the depth of high moisture region in the concrete wall. The present results are based on 10 analyses. The concrete in shoulder areas is subjected to 20 moisture penetration, and thus the high moisture region Lm in shoulder areas may be higher than that in the wall between shoulders.

Exhibit 72 References

1. Crissinger, J.L. (2005) "Measuring Moisture Resistance to Wind-Driven Rain Using a RlLEM Tube", Interface, Nov., 4-11.

2. Hens, H. (2010) "Wind-Driven Rain: From Theory to Reality", Buildings XI, ASHRAE, I 10.

3. Mehta, P.K., and Monteiro, P.J.M. (1993) "Concrete: Structure, Properties, and Materials",

Prentice Hall Inc., Englewood Cliff, NJ 07632.

4. Nehdi, M., Aitcin, P.c., and Perraton, D. (1995) "Investigation of the Performance of Mass Concrete Using in Situ Permeability Tests", Concrete Under Severe Conditions:

Environment and Loading, Vol. 1, Editors:P K. Sakai, N. Banthia, and O.E. Gjorv, E& FN Spon, 413-422.

5. Tariku, F., Cornick, S.M., and Lacasse, M.A. (2007) "Simulation of Wind-Driven Rain Penetration Effects on the Performance of a Stucco-Clad Wall", Buildings X, ASHRAE, 1 10.

6. Wang, K. , Jansen, D., Shah, S.P., and Karr, A. (1997) "Permeability Study of Cracked Concrete," Cement and Concrete Research, 27(3), 381-393.

7. Xi, Y., Bazant, Z.P., and Jennings, H.M. (1994a) "Moisture Diffusion in Cementitious Materials: Adsorption Isotherm", Journal ofAdvanced Cement-Based Materials, I , 248-257.

8. Xi, Y., Bazant, Z.P ., and Jennings, H.M. (1994b) "Moisture Diffusion in Cementitious Materials: Moisture Capacity and Diffusivity", Journal ofAdvanced Cement-Based Materials, I, 258-266.

9. Xi, Y. (1995a) "A Model for Moisture Capacities of Composite Materials - Formulation",

Computational Materials Science, 4, 65-77.

10. Xi, Y. (1995b) "A Model for Moisture Capacities of Composite Materials - Application to Concrete" , Computational Materials Science, 4, 78-92.

11. Xi, Y . and Nakhi, A. (2005) "Composite Damage Models for Diffusivity of Distressed Materials", J ofMaterials in Civil Engineering, ASCE, May/June, 17(3),286-295.

Exhibit 73: Laminar Cracking due to 1978 Blizzard

Exhibit 73 Laminar Cracking due to 1978 Blizzard Table of Contents Summary of Results ................................................................................................................................. 2 Modeling Summary ................................................................................................................................. 2 Overall Approach ................................................................................................................................. 2 Finite Element Software ...................................................................................................................... 2 Scenarios Modeled .............................................................................................................................. 3 Background ......................................................................................................................................... 3 Global Model Description .................................................................................................................... 3 Material Properties .............................................................................................................................. 3 Coefficient of Thermal Expansion of High Moisture Concrete ............................................................... 5

........................................................................................................................ 7 Circumferential Temperature Distribution at O.F. Horizontal Rebar ..................................................... 8

* * * *

Description ........................................................................................................... 9 Discussion ............................................................................................................................................. 11

_ _ _Results .................................................................................................................................. 11 1978 Blizzard Condition ..................................................................................................................... 12 1977 Blizzard Condition ..................................................................................................................... 14 Page 1 Page 1 of 15'

Exhibit 73 Summary of Results The results of the analysis presented in this report can be summarized as follows:

Cracking is predicted due to the 1978 blizzard given the following assumptions:

o Temperatures 20°F above calculated "worst case" temperatures (-24°F => -4°F) o Nonlinear CTE curve that is adapted from Exhibit 57 Figure 4 o Saturation depths of at least 3" o Reduced effective fracture toughness of the concrete (discussed below)

Cracking is not predicted due to the 1977 blizzard, even assuming worst case temperatures.

The locations of the cracking remain confined to the observed crack locations under the thick sections of the shoulders and to locations where the horizontal rebar is spaced on 6" centers.

Modeling Summary Overall Approach The analysis results presented here is a result of many different modeling efforts combined. The convective heat transfer due to wind around the shield building was calculated using a computational fluid dynamics (CFD) model taking into account the wind velocity and direction. The surface roughness, the raised shoulder geometry around the flutes as well as the auxiliary buildings was included in the CFD model. A separate heat transfer analysis consisting of the free-standing reinforced concrete shield building, the free-standing steel containment vessel, and the annulus/hollow space between the two are included. The transient heat transfer model included the duration of two or three days in order to reach accurate temperature predictions. Effects included are solar radiation (tracking of the azimuth and elevation of the sun), the ambient temperature variation from meteorological data, and the temperature on the inside of the steel containment vessel. The details of the wind and thermal modeling are reported separately in the Root Cause Analysis Report.

The calculated temperature distribution in the reinforced concrete shield building is mapped to the Abaqus Global Model. A coupled temperature-displacement analysis for each temperature condition is performed using the Abaqus Global Model.

Gravity is omitted* * * * * * * * * * * * * * * * *,. It is expected gravity would have relatively little effect on the results based on previous studies, which are summarized in Exhibit 64 Thermal Stress Analysis with Gravity and Wind Load.

Finite Element Software Abaqus Version 6.10 for Unux 64-bit was used exclUSively to solve the finite element analysis models presented here. Abaqus is generally considered the gold standard for nonlinear analysis.

Page 2 Page 2 of, 15

Exhibit 73 Scenarios Modeled The following two scenarios are presented in this report. The details of the scenarios and the selection of the time of day are summarized separately in the Root Cause Analysis Report.

1) low temperature and moisture intrusion during the 1978 blizzard

2) low temperature and moisture intrusion during the 1977 blizzard

Background

Expansion of concrete due to freezing of entrapped moisture was studied * * * * * * * * * . This model implements the* * * * * * * * * * * * * * * *

material model. This is the same material model that was used to successfully predict and reproduce the laminar cracking observed in the containment of the Crystal River 3 nuclear power station.

The horizontal rebars in the IF and OF rebar mats are modeled using solid elements, but as an approximation their cross-section is a 1" square, one element across. No other rebar is included in this model because doing so would be computationally prohibitive. This is an assumption made based on engineering judgment and is not expected to have a major impact on the results in this specific scenario because the laminar cracking near the OF rebar mat is driven by radial stresses and there are no radial rebars in the region where the cracks appear to initiate. Vertical stress components are also not expected to drive the damage, therefore the lack of vertical reinforcement is assumed to be a justifiable simplification. Due to these simplifications, which are necessary to make the model computationally tractable, the detailed stress concentrations at the steel and concrete interface are not resolved in the model and therefore these results should be interpreted as a qualitative indication of where cracking is likely and not an absolute quantitative assessment of crack growth.

Global Model Description The drawings used as geometry input for the global model are:

Drawing No: C-lOO Rev. 5 "Shield Building Foundation Plan & Details SH. 1"

Drawing No: C-110 Rev. 6 IIShield Building Roof Plan Wall Section & Details" All vertical reinforcing bars in the containment shell section are modeled as rebar #10 (diameter 1.270")

at 12" center to center spacing. The inner face horizontal rebars are #8 (diameter 1.000") at 12" spacing.

The outer face horizontal rebars are # 11 (diameter 1.410") at 12" spacing except in the top 20' of the walls, where the spacing is 6". The vertical and horizontal rebars in the shoulder sections are #8 at 12" spacing.

Material Properties The material properties used as input to the finite element analysis in this report are summarized in the following documents attached to the Root Cause Analysis Report:

-. Page 3 Page 3 of 15

Exhibit 73

Exhibit 56, Figure 2.1.4: Material Properties for Davis-Besse 3D Nastran Global Model

Exhibit 56, Section 4.7: Effects of Variable CTE

Exhibit 57: Temperature dependent coefficient of thermal expansion (CTE)

Material Properties specific to this analysis include the tensile strength (F T) and the fracture toughness or "strain energy" (G F). These are inputs to the Concrete Damaged Plasticity (CDP) material model, which was developed by Lee-Fenves and implemented natively in the Abaqus FEA software. The Lee Fenves material model and the implementation in Abaqus are recognized to accurately predict cracking in concrete by the concrete modeling community. This technology is new to the nuclear industry.

The tensile strength of the Davis-Besse concrete was tested and found to vary from roughly 500 psi to 1000 psi with an average strength close to 900 psi. The test methods used to measure the tensile capacity will generally produce varying results depending on sample size and test method (e.g. split tensile vs. direct tensile) . And, the samples used in these tests may have been too small for the standard test, thus over-predicted the strength. The strength of the concrete will also vary locally, with the concrete in the immediate vicinity of the rebars usually being weaker due to a higher void fraction in those areas. The samples were taken from regions away from the rebars (purposely to avoid the rebars) .

There is also some uncertainty in the modulus of elasticity. These models assumed E = 4.94Msi, when the stiffness may have actually been closer to S.9Msi, which is about 20% higher and would increase the stresses by up to 20%. In this case, we assume an effective strength (tensile capacity) of FT = 600 psi, which is toward the low end of the test results.

The fracture toughness ("strain energy", GF) of the Davis-Besse concrete was not measured . Various values of GF were tried and calibrated to the observed cracking. By means of comparison, the fracture 2

toughness of the Crystal River concrete was estimated to be about 0.4 in-lb/in

These values serve as starting points, FT =600 psi and GF =0.4 in-lb/in 2 Due to the known mesh size dependency in the Lee-Fenves material model, it is necessary to treat the measured strength parameters as starting points only. The mesh used for these models has very large elements relative to what the material model expects, and the stress concentrations are not fully developed. This inhibits cracking significantly. Therefore, it is necessary to use a lower set of strength parameters as an "effective strength" specific to the mesh. A sensitivity study is performed and a reasonable set of strength parameters is found to result in cracking of the concrete due to the conditions of the 1978 blizzard event. These strength parameters are FT =600 psi and G =0.18 in-lb/in F

2 These calibrated parameters are used in both the 1978 and 1977 cracking models to compare the relative severity of the two events. The parameters themselves are found to be well within the range of expected values for these parameters, which establishes this failure mode as having a high probability.

Page 4 Page 4 of 15

Exhib it 73 Coefficient of Thermal Expansion of High Moisture Concrete The coefficient of thermal expansion (CTE) of high moisture concrete is a highly nonlinear function of temperature. This is associated with the 9% volume expansion of the freezing of entrapped water. The freezing of water in small pores in concrete takes place at a lower temperature than 32°F due to surface tension , which prevents the formation of ice at 32°F. The water in the concrete freezes at varying temperatures depending on the pore size. This creates a nonlinear dependence of the CTE with temperature. This is shown in Figure 1 and is a key input to the finite element analysis presented here.

(Please see Exhibit 57 for a more detailed explanation of the nonlinear CTE.)

1.0 0E* 05 O.OOE+ OO

.... I

~

, tS -Lo aE- OS ~ -I 1

.~

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~

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~

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~

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50

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30

20 *10 10 20 30 40 SO 60 70 eo 90 100 11 0 120 l SO 14 0 Temp~rature rF)

Figure 1 - Nonlinear CTE for Davis-Besse Concrete with 93% Moisture Content (See Exhibit 57)

Without having enough information to perform a thorough 3D moisture diffusion analysis on the Davis Besse shield building model with measured properties taken from Davis-Besse concrete, it was necessary to make an assumption about the moisture gradient and proceed with that assumption in order to get a general qualitative response from the models. The most efficient way to do this was to assume that the moisture content gradients essentially match the temperature gradients. Rather than producing a constant depth of penetration, this produces a variable depth of penetration that takes into account the difference in the surface area to volume ratio in the shoulders.

The models assume that there are two regions. The two outermost contour regions are assumed to have a saturation of 93% (see Exhibit 57 Figure 4) . This region is aSSigned the nonlinear CTE plotted in Figure 1. The calculation of the saturated depth is discussed in detail in Exhibit 72 "Water and Moisture Transfer into Concrete".

The rest of the structure is assigned the linear CTE of 5.2e-6, as found and discussed in Exhibit 56, Figure 2.1.4 (Material Properties for Davis-Besse 3D Nastran Global Model) .

Page 5 Page 5 o f lS

Exhibit 73 Tests of moisture penetration were also performed at the University of Colorado at Boulder, which showed that a i-dimensional (10) depth of water penetration up to 3 or 4 inches is possible when there are winds in excess of 90mph (such as during the 1978 blizzard). The 1978 models are calibrated so that the depth of penetration is approximately 3 to 4 inches in locations subject to 10 moisture diffusion. In the flute valleys, where there is a very low surface area to volume ratio, the depth is about 3 to 4 inches.

In the "middle" of the walls between the shoulders, the depth is about 3.5 to 4 inches. In the corners of the shoulders, where there is a very high surface area to volume ratio, the depth of penetration transitions from 4 inches up to as high as about 14 inches in the corner. This is due to 20 moisture penetration in the shoulders, which appears to be highly significant. The 1977 models assume that the moisture depth of penetration is roughly half the 1978 case due to significantly less wind and precipitation in the 1977 case. Exhibits 66 and 71 summarize some key meteorological data during and prior to the blizzards of 1977 and 1978.

In the Results Section, Figure 7 shows the region that was assigned the nonlinear CTE as the two outermost thermal gradients (4°F to -4°F). Figure 10 shows the region that was assigned the nonlinear CTE for the 1977 case, the outermost thermal gradient (-4°F to -8*F). Depth of penetration is discussed in Exhibit 72.

Page 6 Page 6 of .15

Exhibit 73 Location of Figure 2 below shows the location of the Figure 2 - Shield Building with Flute Numbers * * * * * * * * * * * *

Page 7 Page 7 of 15

Exhibit 73 Circumferential Tempel'ature Distribution at O.F. Horizontal Rebar TT;h;e~t;e;m~p~e~r~a~tu~r~e~d~is~tr~i:b~u;ti~o~n~i;n~t;h~e~:::~p~resented here were calculated in* * * * * * * *

I . (See Exhibit 65) The temperature profiles around the Shield Building at the outer face horizontal rebars are shown in Figure 3. The models presented here use the worst case temperatures calculated for the 1978 blizzard with an offset of +20°F to simulate nominal temperatures. In this case, nominal temperatures will produce the most expansion and therefore the worst case stress condition for the building (see Figure 1). The +20°F offset brings the 1978 temperature gradients into rough equivalence with the lowest recorded ground temperatures during the 1978 blizzard (see Exhibit 66), which would be the expected low temperature condition assuming heavy cloud cover rather than a clear night sky (see Exhibit 65).

The 1977 blizzard model uses the worst case temperatures calculated for the 1977 blizzard with no temperature offset because the worst case 1977 temperatures are already in the range that will maximize expansion and cracking.

Temperature (OF), Mid-Height, Outer Face Horizontal Rebar Depth 16

- 1977 Blizzard Temperature Calculation (Worst Case)

- 1978 Blizzard Temperature calculation (Worst case +2Q* F) 12 10 22 .5 . 675 90 1125 135 157.5 180 Azimuth!")

202 5 225 2475 270 29ZS 31 5 3375 360 Figure 3 - Circumferential Temperature Distribution at the O.F. Horizontal Rebar Depth Page 8 Pag e 8 of 1 5

Exhibit 73 escriptiol1 Figures 4 through 6 depict the geometry * * * * * * * * * * * * * * * * * *

Figure 4 - * * * * * * *; Geometry and Rebars Page 9 Page 9 of 1 5

Exhibit 73 Figure 5 * * * * * *

Detail of Flute Region Figure 6- Section View of Flute Region with Mesh Page 10 Page 10 of 15

Exhibit 73 Discussion The result * * * * * * * * * * * * * * * * * * * * * * *

delamination propensity due to the two blizzard conditions, given the assumptions of the model.

The scale of damage ranges from 0 to 1 and represents the degree of tensile capacity lost from 0% to 100%. Any elements with red or magenta coloring are considered to have formed a structural crack, as they have less than 20% of their original strength left.

The damage that results from any tensile stress above the strength of the concrete depends on the 3D stress state as well as the strain energy available to open the crack. low strain energy results in microcracks and high strain energy results in a structural crack.

Its

* *Imodel results shown in this section can be summarized as follows:

The blizzard of 1978 scenario results in laminar cracking near the OF rebar mat.

The blizzard of 1977 shows some damage (microcracking) relatively close to the surface of the shoulders, and significantly less damage compared to the blizzard of 1978.

o It is important to point out that the distress in the concrete shoulder, predicted due to the conditions in 1977, could have also contributed to the damage in 1978 because distressed concrete can have a permeability that is 1,000 to 10,000 times higher than pristine concrete (see Exhibit 72). Thus, the assumptions being made in this analysis regarding depth of moisture penetration in the shoulders may actually be too shallow.

And as a result, the analysis presented here may be under-predicting the laminar cracking concentrated beneath the shoulders. Note that this is true regardless of the actual depth of penetration during 1977, as less depth of penetration will primarily make the damage more shallow (closer to the surface), which will still have the potential to greatly increase the permeability of the concrete near the surface.

laminar cracks developed most prominently at the OF rebar mat under the thick shoulder regions and not in the thinner sections in the flute and shell.

Page 11 Page 11 of lS

Exhibit 73 1978 Blizzard Condition The result due to the 1978 blizzard is shown in Figures 7 to 9. The temperature contours can be seen in Figure 7 and the cracking result is shown in Figures 8 and 9.

NTll 52

- 48 40 36 32

- 28

- 24 20 16

. 12 8

4 o

4 008: job.odb Abaqus/S~ndard 6 . 10-1 Tue Feb 14 17:46 :42 PO!Iclflc Smndard TIme 2012 y Step: Step expi!lnd-lce Increment 1: Step TIme = 1.000 Prlmary Var: NTll Figure 7 - Temperatures (OF) used in the 1978 Blizzard Analyses Page 12 Page 12 of 15

Exhibit 73 008 : Job.odb Ab"qus/Expllclt 6.10-1 Wed Feb 15 21:25:08 PlIcifJc Srnndllrd lime 2012 y Step: Step ' 2 -dT X Increment 20256: Step Tlme = 1.000

?rl mollfY VlJr: DAMAGET Deformed V"r: U Deformation Sc.!lle Factor: +1.Oe+OQ Figure 8 - Cracking Result due to 1978 Blizzard Conditions DAt"AGET 1.0 0 .9 0.8 0.7 0.6 0.5 0.'

0.3

. 0.2

. 0.1 0.0 ODB: Job.odb Abaqus/Explfclt 6.10-1 Wed Feb 1521:25 :08 Pacln, Standard Tlme 2012 Step : Step*2*dT Increment 20258: Step Time = 1.000 Primary Var: DAMAGET Deformed Var: U Deformtltfon Scale Factor : +l.Oe+OO Figure 9 - Cracking Result due to 1978 Blizzard Conditions showing regions with DAMAGE> 0.7 Page 13 Page 13 of 15

Exhibit 73 1977 Blizzard Condition The result from the* * * * * * *

due to the 1977 blizzard conditions is shown in this section .

Figure 10 depicts the temperature distribution in the _ Figures 11 and 12 show the cracking result .

NTll 47 40 36 32 28 24 20

- 16

- 12

- 8

4

0

*4

8 008: Job.odb Abaqus/StlIndard 6.10-1 Tue Feb 2115 :25 :31 Pacific S~ndard llme 2012 y Step: Stap-2-expand-lce Increment 1 : Step llme = 1.000 Pri mi:lry Var: NTll Deformed VM : U Deform~ti on Scale Factor: +1e+00 Figure 10 - Worst Case Temperatures (OF) calculated for the Blizzard of 1977 (see Exhibit 65)

Page 14 Pag e 14 of 15

Exhibit 73 Figure l1- * * * * * *,due to 1977 Blizzard Conditions DAMAGET 1.0

0. 9 0 .6 0 .7 0 .6 O. S 0.'

0.3 0.2

0. 1 0 .0

- L__ """',"'P. 2.dT Ia:; X Increment 20258 : Step l1me = 1.000 Primary Var : DAf'olAGET Figure 12 -Cracking Result due to 1977 Blizzard Conditions showing regions with DAMAGE> 0.7 Page 15 Page 15 of lS

Exhibit 74: Exhibit Not Used

Exhibit 75: Damage Propagation Analysis

Exhibit 75 Damage Propagation into Regions with High Rebar Density Table of Contents Summary of Results................................................................................................................................. 2 Modeling Summary ............................................ :.................................................................................... 2 Overa II Approach ................................................................................................................................. 2 Finite Element Software ...................................................................................................................... 3 Scenarios Modeled .............................................................................................................................. 3 Background ......................................................................................................................................... 3 Global Model Description .................................................................................................................... 3 Material Properties..............................................................................................................................4 Coefficient of Thermal Expansion of High Moisture Concrete ............................................................... 6 location of * * * * ........................................................................................................................ 8 Circumferential Temperature Distribution at O.F. Horizontal Rebar ..................................................... 9 Cracking * * * *IDescription ......................................................................................................... 10 Discussion ............................................................................................................................................. 12

* * *Results .................................................................................................................................. 12 Top 20' of the Wall location .............................................................................................................. 13 Steam line location ........................................................................................................................... 15 Page 1 Page 1 of 16

Exhibit 75 Summary of Results The results of the analysis presented in this report can be summarized as follows:

The purpose of this investigation is to study the sensitivity of the models response to location and high rebar density. This investigation identifies a plausible scenario in which damage under the shoulders propagates into the shell due to high rebar density, particularly along the top 20' of the walls and around the main steam line opening near to the Aux Building roof.

The environmental conditions and nonlinear CTE studied are the same as in Exhibit 74.

The locations of the cracking remain confined to the observed crack locations at the OF rebar mat, both under the thick sections of the shoulders and in locations where the horizontal rebar is spaced on 6" centers.

o The model at the top 20' of the walls shows some damage in the flute valley, which is in line with observation.

o The model near the Aux building roof shows less damage in the flute valley, which is also in line with observation.

Modeling Summary Overall Approach The analysis results presented here is a result of many different modeling efforts combined. The convective heat transfer due to wind around the shield building was calculated using a computational fluid dynamics (CFD) model taking into account the wind velocity and direction. The surface roughness, the raised shoulder geometry around the flutes as well as the auxiliary buildings was included in the CFD model. A separate heat transfer analysis consisting of the free-standing reinforced concrete shield building, the free-standing steel containment vessel, and the annulus/hollow space between the two are included. The transient heat transfer model included the duration of two or three days in order to reach accurate temperature predictions. Effects included are solar radiation (tracking of the azimuth and elevation of the sun), the ambient temperature variation from meteorological data, and the temperature on the inside of the steel containment vessel. The details of the wind and thermal modeling are reported separately in the Root Cause Analysis Report.

The calculated "worst case" temperature distribution in the reinforced concrete Shield Building is mapped onto the Abaqus Global Model with an applied offset of 20°F to simulate nominal conditions rather than worst case conditions. A coupled temperature-displacement analysis for each temperature condition is performed using the Abaqus Global Model. The computed displacements for the shield building and the mapped temperatures are subsequently used The usage of the technique is required

-. Page 2 Page 2 of 16

Exhibit 75 Gravity is omitted from _the global models * * * * * . It is expected gravity would have relatively little effect on the results based on previous studies, which are summarized in Exhibit 64 Thermal Stress Analysis with Gravity and Wind load.

Finite Element Software Abaqus version 6.10 for linux 64-bit was used exdusively to solve the finite element analysis models presented here. Abaqus is generally considered the gold standard for nonlinear analYSis.

Scenarios Modeled The following scenario is presented in this report. The details of the scenario and the selection of the time of day are summarized separately in the Root Cause Analysis Report.

o Temperatures 20 a F above calculated "worst case" temperatures (-24°F => -4°F) o Nonlinear CTE curve that is adapted from Exhibit 57 Figure 4 o Saturation depths of at least 3" o Reduced effective fracture toughness of the concrete (discussed below)

Background

Expansion of concrete due to freezing of entrapped moisture was studied in the Global Model Description The drawings used as geometry input for the global model are:

Drawing No: C-100 Rev. 5 ((Shield Building Foundation Plan & Details SH. 1#

Drawing No: C-110 Rev. 6 ((Shield Building Roof Plan Wall Section & Details" Page 3 Page 3 of 16

Exhibit 75 All vertical reinforcing bars in the containment shell section are modeled as rebar #10 (diameter 1.270")

at 12" center to center spacing. The inner face horizontal rebars are #8 (diameter 1.000") at 12" spacing.

The outer face horizontal rebars are # 11 (diameter 10410") at 12" spacing except in the top 20' of the walls where the spacing is 6". The vertical and horizontal rebars in the shoulder sections are #8 at 12" spacing.

Material Properties The material properties used as input to the finite element analysis in this report are summarized in the following documents attached to the Root Cause Analysis Report:

Exhibit 56, Figure 2.104: Material Properties for Davis-Besse 3D Nastran Global Model

Exhibit 56, Section 4.7: Effects of Variable CTE

Exhibit 57: Temperature dependent coefficient of thermal expansion (CTE)

Material Properties specific to this analysis include the tensile strength (FT) and the fracture toughness or "strain energy" (G F). These are inputs to the Concrete Damaged Plasticity (COP) material model, which was developed by Lee-Fenves and implemented natively in the Abaqus FEA software. The Lee Fenves material model and the implementation in Abaqus are recognized to accurately predict cracking in concrete by the concrete modeling community. This technology is new to the nuclear industry.

The tensile strength of the Davis-Besse concrete was tested and found to vary from roughly 500 psi to 1000 psi with an average strength close to 900 psi. The test methods used to measure the tensile capacity will generally produce varying results depending on sample size and test method (e.g. split tensile vs. direct tensile). And, the samples used in these tests may have been too small for the standard test, thus over-predicted the strength. The strength of the concrete will also vary locally, with the concrete in the immediate vicinity of the rebars usually being weaker due to a higher void fraction in those areas. The samples were taken from regions away from the rebars (purposely to avoid the rebars).

There is also some uncertainty in the modulus of elasticity. These models assumed E = 4.94Msi, when the stiffness may have actually been closer to 5.9Msi, which is about 20% higher and would increase the stresses by up to 20%. In this case, we assume an effective strength (tensile capacity) of FT =600 psi, which is toward the low end of the test results.

The fracture toughness ("strain energy", GF) of the Davis-Besse concrete was not measured. Various values of GF were tried and calibrated to the observed cracking. By means of comparison, the fracture toughness of the Crystal River concrete was estimated to be about 0.4 in-lb/in2

These values serve as starting points, Fr =600 psi and GF =004 in-lb/in 2 Due to the known mesh size dependency in the Lee-Fenves material model, it is necessary to treat the measured strength parameters as starting points only. The mesh used for these models has very large elements relative to what the material model expects, and the stress concentrations are not fully developed. This inhibits cracking significantly. Therefore, it is necessary to use a lower set of strength parameters as an "effective strength" specific to the mesh. A sensitivity study is performed and a

-. Page 4 Page 4 of 16

Exhibit 75 reasonable set of strength parameters is found to result in cracking of the concrete due to the conditions outlined above. These strength parameters are FT =600 psi and GF =0.09 in-lb/in 2

Page 5 Page 5 of l6

Exhibit 75 Coefficient of Thermal Expansion of High Moisture Concrete The coefficient of thermal expansion (CTE) of high moisture concrete is a highly nonlinear function of temperature. This is associated with the 9% volume expansion of the freezing of entrapped water. The freezing of water in small pores in concrete takes place at a lower temperature than 32°F due to surface tension , which prevents the formation of ice at 32T The water in the concrete freezes at varying temperatures depending on the pore size . This creates a nonlinear dependence of the CTE with temperature . This is shown in Figure 1 and is a key input to the finite element analysis presented here.

(Please see Exhibit 57 for a more detailed explanation of the nonlinear CTE.)

1.00E *OS a.OOEt 00

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50 3~ 10 10 20 30 40 50 60 70 80 90 1 00 11 0 1 20 13 0 14 0 Te:mpera1ure rF)

Figure 1- Nonlinear CTE for Davis-Besse Concrete with 93% Moisture Content (See Exhibit 57)

Without having enough information to perform a thorough 3D moisture diffusion analysis on the Davis Besse shield building model with measured properties taken from Davis-Besse concrete, it was necessary to make an assumption about the moisture gradient and proceed with that assumption in order to get a general qualitative response from the models. The most efficient way to do this was to assume that the moisture content gradients essentially match the temperature gradients. Rather than producing a constant depth of penetration, this produces a variable depth of penetration that takes into account the difference in the surface area to volume ratio in the shoulders.

The models assume that there are two regions. The outermost contour region is assumed to have a saturation of 93% (see Exhibit 57 Figure 4) . This region is assigned the nonlinear CTE plotted in Figure 1.

The calculation of the saturated depth is discussed in detail in Exhibit 72 "Water and Moisture Transfer into Concrete".

The rest of the structure is assigned the linear CTE of 5.2e-6, as found and discussed in Exhibit 56, Figure 2.1.4 (Material Properties for Davis-Besse 3D Nastran Global Model) .

Page 6 Page 6 o f 1 6

Exhibit 75 Tests of moisture penetration were also performed at the University of Colorado at Boulder, which showed that a 1-dimensional (lD) depth of water penetration up to 3 or 4 inches is possible when there are winds in excess of 90mph (such as during the 1978 blizzard). The shield building models are calibrated so that the depth of penetration is approximately 3 to 4 inches in locations subject to 1D moisture diffusion. In the flute valleys, where there is a very low surface area to volume ratio, the depth is about 3 to 4 inches. In the "middle" of the walls between the shoulders, the depth is about 3.5 to 4 inches. In the corners of the shoulders, where there is a very high surface area to volume ratio, the depth of penetration transitions from 4 inches up to as high as about 14 inches in the corner. This is due to 2D moisture penetration in the shoulders, which appears to be highly significant.

In the Results Section, Figures 7 and 9 show the region that was assigned the nonlinear CTE as the two outermost thermal gradients (4°F to -4°F). Exhibits 66 and 71 summarize some key meteorological data during and prior to the blizzards of 1977 and 1978. Depth of penetration is discussed in Exhibit 72.

Also note that the model being placed near the Aux Building roof assumes no contribution of heat from the Aux Building itself. However, it is worth noting that a similar stress state may exist in the Shield Building wall below the Aux building roof - because the elevated temperatures inside the Aux building will act to expand the concrete at the surface of the Shield Building wall, in a manner similar to the outer surface expansion due to ice above the Aux building roof.

Page 7 Page 7 of 16

Exhibit 75 Location Figure 2 below shows the location of the detailed * * *

in the containment structure.

Figure 2 - Shield Building with Flute Numbers and Locations * * * * *

-. Page 8 Page 8 of 16

Exhibit 75 Circumferential Temperature Distribution at O.F. Horizontal Rebar The temperature distribution in the presented here were calculated in separate heat transfer analysis and mapped to the Abaqus Global Model. (See Exhibit 65) The temperature profiles around the Shield Building at the outer face horizontal rebars are shown in Figure 3. The models presented here use the worst case temperatures calculated for the 1978 blizzard with an offset of +20°F to simulate nominal temperatures. In this case, nominal temperatures will produce the most expansion and therefore the worst case stress condition for the building (see Figure 1). The +20°F offset brings the 1978 temperature gradients into rough equivalence with the lowest recorded ground temperatures during the 1978 blizzard (see Exhibit 66), which would be the expected low temperature condition assuming heavy cloud cover rather than a clear night sky (see Exhibit 65).

Temperature (OF), Mid-Height, Outer Face Horizontal Rebar Depth 16

- 1977 Blizzard Temperature Calculation (Worst Case) 14 - 1978 Blizzard Temperature Calculation (Worst Case +20' F) 10

~ 8 "c.E

~

., 67S 90 11 15 135 157.5 160 :;:':'.2 5 215 147.5 292.5 3 15 337.5 360 Azimuth (' )

Figure 3 - Circumferential Temperature Distribution at the O.F. Horizontal Rebar Depth Page 9 Page 9 of 16

Exhibit 75 Cracking _ Description In the cracking both the inner face and outer face horizontal rebars are~::::::

in the circumferential direction, spanning one flute valley, one full shoulder, to the midpoint between the shoulders. The

. For this analysis, The temperatures and displacements

. The rebars and concrete are all modeled with solid continuum elements .

Figures 4 through 6 depict the geometry and finite element mesh of th Figure 4 - * * * * * * *

Geometry and Rebars Page 10 Page 10 of 16

Exhibit 75 Figure 5 Detail of Flute Region Figure 6 ection View of Flute Region with Mesh Page 11 Page 11 of 16

Exhibit 75 Discussion The result from the is used to make predictions about the delamination propensity due to blizzard conditions, given the assumptions of the model.

The scale of damage ranges from 0 to 1 and represents the degree of tensile capacity lost from 0% to 100%. Any elements with red or magenta coloring are considered to have formed a structural crack, as they have less than 20% of their original strength left.

The damage that results from any tensile stress above the strength of the concrete depends on the 3D stress state as well as the strain energy available to open the crack. Low strain energy results in microcracks and high strain energy results in a structural crack.

The cracking model results shown in this section can be summarized as follows:

The locations of the cracking remain confined to the observed crack locations at the OF rebar mat, both under the thick sections of the shoulders and in locations where the horizontal rebar is spaced on 6" centers.

o The model in the top 20' of the walls shows some damage in the flute valley, which is in line with observation.

o The model near the Aux building roof shows less damage in the flute valley, which is also in line with observation.

Overall, the results show good agreement with observed cracking in the areas studied.

-. Page 12 Page 12 of 16

Exhibit 75 Top 2 O' of the Wall Location The results at the top 20' location, perturbed by the 1978 blizzard are shown in Figures 7 and 8. The 20°F offset temperature contours can be seen in Figure 7 and the cracking results are shown in Figure 8.

NTll S2 48 44

<10 36 32 28 2'

20

- 16 12

- 8 o

008: job.odb A~qus/Stand&rd 6.10-1 Tue Feb 14 17 :46 :42 PlIclflc SlJ!lnd~rd Time 20 12 y Step: Step-2-explmd-lce Increment 1: Step Tlme = 1.000 Prim~ry V~r: NTll Figure 7 - Temperature Contours (OF) in the Top 20' of the Wall Page 13 Page 13 of l6

Exhibit 75 OAMAGET

1. 0 0.9

0. 8

0. 7 0 .6 0.5 0.4 0 .3 0.2 0 .1 0 .0 ODB : job ,odb A~qus/ E x p tlc lt 6.10-1 Wed Feb 15 2 1:25 :08 Pacirlc StlIndard TIme 2012 Step : Step dT Increment 20258 : Step Tlme = 1. 000 Primary Var : DAr-IAGET Deformed Var : U Deformation Scale Facb::lr : +1.Oe+ OO Figure 8 - Cracking Result in Top 20' of the Wall showing regions with DAMAGE> 0.6 Page 14 Page 14 o[ 16

Exhibit 75 Steam Line Location The results at the steam line location, perturbed by the 1978 blizzard are shown in Figures 9 and 10. The 20°F offset temperature contours are in Figure 9 and the cracking results are shown in Figure 10.

NTll S2

- 'IS 40 36 32 28 24 20 16 12 8

4 o

-4 008 : lob.odb A,b"qus/StlmdlH"d 6. 10-1 f'1on Feb 20 13:0$:$1 P",clflc Standllrd TIme 2012 Step: Step-2-expand-lce Increment 1: Step Tlme = 1.000 X PMmuy V"r: NTll Deformed Var : U Deformabon Scale Factor : +ie+OO Figure 9 - Temperature Contours (OF) at Steam Line Page 15 Page lS of l6

Exhibit 75 1.0

- 0.9

- 0.8 0 .7

- 0.6

- 0.5

- 0.4

- 0 .3 0.2 0.1 0.0 008: job.odb Abaqus/ExpUclt 6.10-1 Mon Feb 20 18 :46:37 Pacific St2Jndard TIme 2012 L Step : Step-2*dT X Increment 20258: Step TIme Primary Var: DAMAGET

= 1.000 Figure 10 - Steam Line Area Cracking Result showing regions with DAMAGE> 0.6 Page 16 Page 16 of 16

Exhibit 76: Sample List and Activity for Lab Testing

Exhibit 76 Performance Improvement International Providing a competitive advantage through research and applications To:

Fro Date: 02/27/2012

Subject:

Laminar Cracking of the Davis-Besse Shield Building - Sample List & Activity of Concrete-Core Samples You will find detailed in what follows a list of the concrete-core samples received from Davis Besse and the corresponding activities performed.

Page 1 of 6

Exhibit 76 Sam pie List & Activity Sample Handling Twenty Eight (28) total samples were received from Davis-Besse Nuclear Power Station for laboratory tests and examination .

Each sample is alpha-numerically identified and begins with either an 'F' or an'S' depending on the area of the shield building from which the sample was extracted, that is , from the Flute or from the Shoulder.

As the samples were received , they were photo documented in their 'As Received' condition, as can be seen in Figures A1 & A2 . Afterwards, the samples were marked from the "exterior surface" to the end of the sample in 1.0 inch increments, they were also photographed . The exterior surface is defined as the surface from which the core drill begins. Thus, it is possible for the exterior surface to be the inner or outer diameter surface of the shield building. The samples were also inspected and cracks identified .

Figure A3 & A4 show examples of these photo-marked samples.

Tables A1 & A2 show the list of samples received and the corresponding tasks performed on each respective sample . Note that the samples had varying degrees of testing and examination.

Table A3 documents the disposition of 6 out of the 28 samples which were not examined nor tested by methods described in tables A1 & A2 .

Page 2 Page 2 of 6

Exhibit 76 Table A 1: li~ of Samples and Tasks Performed on Samples ( Fracture was cut)

Examine Carbonation Isolate Aggregate for Photo Test Coarse Date Photo As Cracks Longitudinal Size and Sample Cracks as Crack(s) and Void Received Received and Cut Fraction, and Marked Fractures Size fractures (inches)

Mark (millimeters)

F3-1 10/29/2011 X X X X' x x x X S11-1 10/29/2011 X X X X X X 511-2 10/29/2011 X X X X X X 512-1 10/29/2011 X X X X X X 512-2 10/29/2011 X X X X X X 516-3 10/29/2011 X X X X X X 55-1 10/29/2011 X X X X X X 55-2 10/29/2011 X X X X X X 57-1 10/29/2011 X X X X X X X X 57-2 10/29/2011 X X X X' x X 57-3 10/29/2011 X X X X X X 59-1 10/29/2011 X X X X X X 59-2 10/29/2011 X X X X X X 5-9-653-11 11/10/2011 X X X X X X X X 5-9-785-22.5 11/10/2011 X X X X X X X X 5-7-656.5-6.5 11/10/2011 X X X X X X X X F2-790.0-4.5 12/2/2011 X X X X X F5-791.0-4 12/2/2011 X X X X X F4-794.0-3.5 12/2/2011 X X X X X 52-798.5-4.5 12/2/2011 X X X X X 53- 55' 11/2/2011 X 513-633.08 1/17/2012 X X X Page 3 Page 3 of 6

Exhibit 76 Table A Z: Ust of Samples and Tasks Performed on Samples (0 Fracture was cut)

Micro-Coarse Coarse Coarse Hardness Void Size Void EDSof Void Size Fine Coarse & Dry/Wet Date Micro Crack on One Size Iron Sample on Void Fine Carbonation Received Examination Transverse on Oxide Transverse Size Aggregate Tests Fracture Rebar Transfer Fractures and Only Imprint Cement F3-1 10/29/2011 X X X S11-1 10/29/2011 X S11-2 10/29/2011 X X S12-1 10/29/2011 X X S12-2 10/29/2011 X X S16-3 10/29/2011 X X S5-1 10/29/2011 X X S5-2 10/29/2011 X X S7~1 10/29/2011 X X S7-2 10/29/2011 X S7-3 10/29/2011 X X S9-1 10/29/2011 X X S9-2 10/29/2011 X X S-9-653-11 11/10/2011 X S-9-785-22.5 11/10/2011 X S-7-656.5-6.5 11/10/2011 X F2-.790.0-4.5 12/2/2011 X X X X F5-791 .0-4 12/2/2011 X F4-794.0-3.5 12/2/2011 X S2-798.5-4.5 12/2/2011 X S3- 55' 11/2/2011 X X S13-633.08 1/17/2012 X X Table A 3: Samples shipped out from Photometries for Mechanical and Physical Testing Sample Date Photo As Disposition Status 10 Received Received Shipped 10 Dr. Xi on 11nt2011 (14" piece only), Shipped 10 Dr. Xi on S9 11/1/2011 X Dr. Xi Results 11/14/2011 (7" piece only)

Main Steam 11/112011 X Shipped to Dr. Xi on 11/7/2011 Dr. Xi Results Room EOG 11/1/2011 X Shipped to Dr. Xi on 11/14/2011 Dr. Xi Results Hallway #1 Picked Up By Twinning Labs For Compressive Strength and MOE on S-1 11/10/2011 X Complete 11/14/2011 S-3 11/10/2011 X Picked Up By Twinning Labs For Split Tensile on 11/14/2011 Complete S-8 11/1012011 X Picked Up By Twinning Labs For Slep Loading Faligue Tesl on 11/14/2011 Shipped to Dr. Xi Page 4 Page 4 of 6

Exhibit 76 Figure A 1: Example of As Received Core Sample (Core F3-1)

Figure A 2: Example of As Received Core Sample with Transverse Crack (Sample 1# 59-1)

Page 5 Page 5 of 6

Exhibit 76 Figure A 3: Marked Sample at 1" Intervals, Cracks ldentif1f~d (Sample 1/ B -1)

Figure A 4: Marked Sample at 1" Intervals, Cracks Identified (Sample # 59-1)

Page 6 Page 6 of 6

Exhibit 77: Aggregate Size Distribution

& Void Fraction lab Testing

Exhibit 77 Performance Improvement International Providing a competitive advantage through research and applications To:

From:

Date: 02/27/2012

Subject:

Laminar Cracking of the Davis-Besse Shield Building - Concrete Sample Testing for Aggregate Size Distribution and Void Fraction Based on my observation and examination of concrete-core samples received from the Davis Besse Shield Building, my findings for Aggregate Size Distribution and Void Fraction are detailed in what follows.

Page 1 of 13

Exhibit 77 Aggregate Size Distribution and Void Fraction Analysis Aggregate Size Distribution ASTM C 856 Paragraph 8.1 allows for procedures to be chosen based on the purpose of the examination. The procedure used for determination of the Area Fraction of Aggregates in concrete will be discussed in what follows.

Of the three established methods for determining the area of a phase, the Delesse Method was chosen due to the limited sample sizes. In addition, Image J which is a government sponsored image analysis software program was used to facilitate this method.

The aggregate size is measured using Feret's diameter and coarse aggregates are considered those which have a major diameter of 0.3" or greater. Feret's diameter is used as defined in the in the Image J handbook:

Feret's diameter: The longest distance between any two pOints along the selection boundary, Improved in IJ 1.45m also known as maximum caliper.

The procedural outline along with an example is shown in what follows:

PageZ Page 2 of 13

Exhibit 77

1. The surface sample is polished then decorated with phenolphthalein which colors the concrete paste (purple) but not the aggregate. A micrograph is then taken .

Figure A 1: Micrograph of paste (red) I Aggregate mixture paste is decorated with phe.t'lolphthalem

2. The micrograph is then converted to grayscale via green channel for enhanced contrast.

Figure A 2: Micrograph of paste I Aggregate mixture converted to grayscale via green channel for enhanced contrast Page 3 Page 3 of 13

Exhibit 77

3. The micrograph is further converted to a Binary Selection of Aggregate Particles .

Figure A 3: Binary Selection of Aggregate Particles

4. The aggregate size is then represented by image analysis, using Image J .

.I ( ~ ) '" c ..... __.

t .- \

., . "j : (,

Figure A 4: Representation of Aggregate Sile by Image Analysis, (The minimum aggregate sile analyzed was set to 0,0004 square Inches)

Page 4 Page 4 of 13

Exhibit 77

5. A table is generated listing each aggregate and its corresponding Feret diameter and Area.

Area Feret °kArea Feretx Feret" FeretAn g le M.nFeret 246 o Ou2 u .O~tj 100 1 817 1 654 144 .3 4 5 0 .0 35 247 6 91 3 E-4 0 .044 100 0 .096 .684 33690 0 .021 248 7.064E-4 0.035 100 1 .644 . 663 143 . 130 0.030 249 6.189E-4 0.046 100 2 717 1 659 167 .005 0.019 250 5 .645E-4 0 034 100 2.446 1 .659 113.962 0 0'25 25 1 5 .645E-4 0 039 100 1.701 1 .665 138.576 0022 252 7457E-4 0 054 100 1892 1.671 159 228 0 .0 2 7 253 5 796E-4 0 .0-'.( 1 100 0 . 198 .6 82 146 .976 0 0 21 254 0.003 0080 100 0.412 .724 16."1 60 0.058 2 55 4 .076E-4 0.037 100 0 054 1 . 731 79.216 0 . 017 256 4 830E-4 0 037 100 0 .302 . 720 31.430 0027 257 0001 0042 100 , .62 8 . 71 1 163 .072 0 .037 258 0 033 0.352 100 2.4 41 1.797 15 751 0 135 259 0.001 0054 100 0 142 1.739 33 . 179 a 034 2 60 0 . 001 0054 100 0 073 1.757 47 603 0 0 33 261 0 . 002 a 060 100 1.512 1 .750 29 :539 0 050 2 62 6823 E - 4 0 043 100 1 821 1 .7 2 0 133 363 0 031 2 6:=: 0 005 a 138 100 o .:: 1 ~.:i4 1 .7 2 0 133 977 0089 2 64 0.002 0.075 100 1 682 1.724 111 .801 0044 2 65 0.001 0072 100 2419 . 790 67 . 166 0 .032 2 66 0.002 0.096 10Cl 0 268 727 '121 675 0.049 267 0 .001 0 04 8 100 1 :598 1 . 762 21 038 a 036 2 6Ei 0 .001 0 073 100 1 751 1 . 739 126.327 0027 269 5 .464E-4 0 061 100 0.078 1.795 42.709 0 . 018 270 0 . 002 0077 100 1 571 1.75 5 127.648 0 .046 271 0.001 0068 100 0 .021 1.765 129 806 0 029 272 0 .003 0.093 100 0 222 1 774 149 .589 0 055 273 0001 0.058 100 0 . 136 1 .812 30.379 0 .031 Figure A 5: Feret Diamtel and Area Table generated for each aggregate found In the analvsis

6. A summary report is produced for each sample under consideration, yielding the following information:

i. Aggregate Count ii. Total Area iii . Average Size iv. Area Fraction

v. Feret Diameter Slice jCount [Total Area j AVerage Size -,Area Fractlon J Feret Feree< Feret'C JFeretAng e t.1InFeret F3-1. Coarse Aggregate-2 JPG (green).IP9 273 2361 0 .009 40 7 0087 I A07 1 035 89362 (1049 Figure A 6: Output for a sample under consideration.

Page 5 Page 5 of 13

Exhibit n The following table shows the samples investigated and the subsequent results using The Delesse Method which was facilitated by Image J. For purposes of Aggregate Size Distribution , 5 samples were tested.

Table A 1: Aggregate Size Distribution Summary Number of Area Fraction Coarse Coarse Coarse Aggregate per Aggregate on Coarse Aggregate Number of Coarse Sample Aggregate Aggregate Square Inch Surface Aggregate Area Standard Aggregate/Square Identification Maximum Average Fraction, % Deviation, inch Size, inches Size, inches (Coarse+Fine) (Coarse+Fine) % inches F3-1 73.6 49 .3% 35 .5% 1.614 0598 0.327 2.4 S7-1 86 .6 46.8 % 31 .7% 1.364 0.653 0.275 2.1 S-7-656.5-6,5 64.6 44 .0% 33.2% 1.591 0.598 0.307 2.3 S-9-653-1 85 .3 44 .5% 28.3% 1.569 0568 0.297 2.4 S-9-785-22.5 101 .8 46 .8% 34 .8% 1.251 0.548 0.289 2.1 AVERAGE 82,4 46,0% 33% 1,48 0,59 0,3 2,26 Page 6 Page 6 of 1 3

Exhibit 77 AVoid Fraction Analysis on Core Samples A methodology that meets the intent of the ASTM C 457 Standard for Void Fraction Analysis will be detailed in what follows . Under the microscope, at 4X magnification , the areas of the coarse voids were summed and divided by the cross-sectional area of the sample under view. The micro-void fraction was measured on two samples (F-3 and S7-1) . The contribution of the micro voids to the total void fraction (micro + coarse voids) was verified to be less than 10% of the coarse measured void fraction.

Therefore , the micro voids were neglected in the results provided.

The under mentioned outlines the void fraction analysis method and provides some examples:

Coarse Voids on the Transverse Fracture Surface

1. Photograph the fracture surface or the rebar imprint as applicable using a magnification of 4 X (on screen).

2. Measure the voids diameter in the entire field of view and tabulate.

3. Calculate the total surface area of voids

4. Divide the total surface area of voids by the total surface area of the sample to obtain the void fraction .

5. Plot Void size distribution in histogram presentation Page 7

!'age 7 of 13

Exhibit 77 Coarse Void on the Longitudinal Fracture Surface

1. Cut the core sample longitudinally, along its centerline.

2. Photograph the cut surface using a magnification of 4 X (on screen).

3. Measure the voids diameter in the entire fieJd of view and tabulate.

4. Calculate the total surface area of voids

5. Divide the total surface area of voids by the total surface area of the sample cross section to obtain the void fraction.

6. Plot Void size distribution in histogram presentation.

Figure A 8: Void Fraction AnalysIs (Core 57-I, longitudinal)

Page 8 Page 8 of l3

Exhibit 77 A Micro-Void Size Longitudinal Cross-Sections:

I. Cut the core sample longitudinally, along its centerline.

2. Polish the cross-section and examine under optical microscope at SOX.

3. Photograph.

4. Measure the voids diameter in the entire field of view and tabulate.

5. Calculate the total surface area of voids.

6. Divide the total surface area of voids by the total surface area of the sample cross section to obtain the void fraction.

7. Plot Void size distribution in histogram presentation.

Figure A 9: Micro*Void Fraction Analysis (Core F3-1, Longitudinal)

"Note: The micro void fraction was measured on two samples (F-3 and S7-1) , and the contribution of the micro voids to the total void fraction (micro + coarse voids) was verified to be less than 10% of the coarse measured void fraction . This Methodology was used in lieu of ASTM C 457 since it meets the intent.

Reference micro void calculation analysis of samples F-3 and S7-1 .

Page 9 Page 9 of 13

Exhibit 77 The following tables show the void distribution summary for samples on which void fraction analysis was conducted. The samples with an asterisk in their identification number are core samples with transverse cracks as identified by plant IR inspection.

Table A 2: Coarse Void Size and Distribution Summary for Concrete not in Contact WIth Rebar Sample Total Number of Coarse Void Coarse Void Standard Coarse Void Identification Number of Coarse Maximum Average Deviation, Content by Coarse Voids per Size, Size, microns Area Voids/ Sq. Square inch microns microns Fraction, %

inch F3-1 31 voids / 3.6 6680 1917 1397 1.9%

8.51 Sq.

inches S12-2* 56 Voids / 25 .9 3610 1064 796 4.8%

2.16 Sq .

inches S16-3* 23 Voids / 10.6 5610 667 1120 1.2%

2.16 Sq.

inches S5-1* 31 Voids / 14.3 2420 987 521 1.9%

2.16 Sq.

inches S7-1

139 voids / 6.9 10800 1652 1579 3.2%

20.05 Sq.

inches S7-3* 63Voids / 29.1 2230 630 445 2.1%

2.16Sq.

inches S9-1

42 Voids / 19.4 3650 938 716 3.3%

2.16 Sq .

inches S-7-656.5-6.5 179 Voids / 19.1 4310 729 528 1.9%

9.37 Sq .

inches S-9-653-1 100 Voids / 12.6 10300 975 1125 2.6%

8 Sq. inches

$-9-785-22.5 90 Voids / 10.1 6710 1081 872 2.0%

8.95 Sq.

inches Page 10 P a g e 10 o f 13

Exhibit 77 AVoid Fraction Analysis at Areas near Rebar Contacts Void fraction analysis at areas near rebar contacts was conducted similarly to the aforementioned. Samples F2-790 and S13-633 are the only two samples received from Davis-Besse that show concrete/rebar interaction. Void fraction analysis was conducted on both samples.

Below are the micrographs that illustrate the areas under investigation and subsequent coarse void measurements. Again, the coarse voids (neglecting micro-voids) were summed and then divided by the cross-sectional area under examination .

Figure A 10; Cross-sectional view of rebar/concrete contact area (Core F2-790.0-4.5)

Page 11 Page 1 1 of 13

Exhibit 77 A

Figure A 11: Void Fraction Analysis (Core F2 790.0-4.5)

Figure A 12. Void Fraction Analysis (Core 513-633.08)

Page 12 Page 12 of 13

Exhibit 77 AThe following table shows several samples on which void fraction analysis was conducted . The samples with an asterisk in their identification number are core samples with transverse cracks as identified by plant IR inspection . The row highlighted in red shows void fraction analysis conducted on the concrete/rebar-interaction core sample, F2 790.0-4.5. The void fraction is roughly 3 times higher than for samples not in contact with rebar.

Table A 3: Coarse Void Size and Distribution Summary Sample Total Numberof Coarse Void Coarse Void Standard Coarse Void Identification Number of Coarse Voids Maximum Average Size, Deviation, Content by Coarse Voids per Square Size, microns microns microns Area

/ Sq . inch inch Fraction, %

F3-1 31 voids / 3.6 6680 1917 1397 1.9%

8.51 Sq.

inches 512-2* 56 Voids/ 25.9 3610 1064 796 4.8%

2.16 Sq.

inches 516-3* 23 Voids / 10.6 5610 667 1120 1.2%

2.16 Sq.

inches 55-1

31 Voids / 14.3 2420 987 521 1.9%

2.16 Sq .

inches 57-1* 139 voids / 6.9 10800 1652 1579 3.2%

20.05 Sq .

inches 57-3* 63Voids / 29.1 2230 630 445 2.1%

2.16 Sq .

inches 59-1

42 Voids/ 19.4 3650 938 716 3.3%

2.16 Sq.

inches 5-7-656.5-6.5 179 Voids / 19.1 4310 729 528 1.9%

9.37 Sq.

inches 5-9-653-1 100 Voids / 8 12.6 10300 975 1125 2.6%

Sq . inches 5-9-785-22.5 90 Voids/ 10.1 6710 1081 872 2.0%

8.95 Sq .

inches Core F2 790.0 56 Voids I 100 1988 526 505 6.4%

4.5 (Rebar 0.56 Sq.

Imprint Area) inches Page 13 Page 13 of 13

Exhibit 78: Microcrack lab Testing

Exhibit 78

Performance Improvement International Providing a competitive advantage through research and applications To:

From:

Date: 02/27/2012

Subject:

Laminar Cracking of Davis-Besse Shield Building - Concrete Sample Testing for M icrocracking Based on my observation and examination of concrete-core samples received from the Davis Besse Shield Building, my findings for Microcracking are detailed in what follows .

Page 1 of 3

Exhibit 78 Micro Cracks Micro Crack Examination Laboratory examinations were conducted to determine whether or not micro cracks existed on the samples received from Davis-Besse. Virtually all samples were examined. Cross-sections of samples were prepared, polished, and scanned . There was no evidence of micro cracking on any of the samples received. Cross-sections were inspected with magnifications of up to 500 times; recorded photos are at magnifications of 25 to 100 times.

There are numerous evidential photographs that show no micro-cracking, however, only a few will be shown below for illustration purposes.

Figure 1. Magnification at 100 Times Page 2 Page 2 of 3

Exhibit 78 Figure 2: Magnification at 15 Times (Core 57-1)

Figure 3: Magnification at 100 TImes (Core FS-791.o-4)

Page 3 Page 3 of 3

L-12-196, FENOC-Davis-Besse Nuclear Power Station, Unit 1 Docket No. 50-346, License No. NPF-3 Submittal of Contractor Root Cause Assessment Report-Section 7

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L-12-196, FENOC-Davis-Besse Nuclear Power Station, Unit 1 Docket No. 50-346, License No. NPF-3 Submittal of Contractor Root Cause Assessment Report-Section 7

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