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 6
	ML12138A081
Person / Time
Site:	Davis Besse
Issue date:	05/14/2012
From:	; FirstEnergy Nuclear Operating Co
To:	; NRC/RGN-III
References
	L-12-196
	Download: ML12138A081 (177)
	v • d • e

Exhibit 52

~ University of Colorado Dept. of Civil, Environmental & Architectural Engineering

~ Boulder College 01 Engineering and Applied Science t 303 492 8991 428 UCB Boulder, Colorado 80309-0428 1 3034927317 yunpmg.Xl@Colorado edu MEMORANDUM To:

Performance Improvement International 21 I I S EI Camino Real Suite 200 Oceanside, CA 92054 Attention: Dr-Chong Chiu From:

Prof. Yunping Xi

Subject:

Concrete Property Testing Results on Submitted Concrete Core Specimens

1. Introduction Concrete core samples were delivered to the University of Colorado at Boulder in Nov. 2011.

The concrete cores were cut into 13 samples for testing internal relative humidity, compressive strength, splitting tensile strength, coefficient of thermal expansion, accelerated creep, and freeze-thaw resistance. The identifications and dimensions of the samples will be described in the fo Ilowing sections together with testing resu It.

2. Internal Relative Humidity The level of internal moisture and the distribution of internal moisture in a concrete structure are important for evaluating shrinkage and freeze-thaw damage of the concrete. Internal relative humidity (RH) distribution of the concrete was measured by using thermal and moisture sensors SHT75 from Sensirion. The concrete core used for this test is identified as S6 11/8/11. Eight sensors were embedded in the concrete cylinder at different depths from the surface to measure continuously both internal temperature and RH. The distances of the sensors from the surface are 1.0 in, 1.5 in, 2.0 in, 2.5 in, 3.0 in, 5.0 in, 7.0 in, and 8.5 in. Distributions of RH in the concrete sample were obtained. Test results are shown in Table 1 and in Fig. I.

T ble I RH a a at I erent tunes (S.peclmen S6 1118/11) a test d t Depths from the surface (in)

Time 1

1.5 2

2.5 3

5 7

8.5 11 / 18/2011 01:53PM 60.07 63.57 64.26 59.07 60.44 67.01 67.92 69.41 11119/2011 01:53PM 59.7 63.08 63.07 58.32 59.5 67.07 68.01 69.14 11/20/2011 01 :53 PM 58.63 61.94 61.42 56.82 57.91 66.55 67.55 68.42 I 1121/20 I I 0 I :53 PM 58.14 61.32 60.44 56.02 57.26 66.52 67.48 68.24 Page 1 o!' 10

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9 Depth (inY Fig. 2 RH distributions at different times The RH values near the surface (from 1 in. to the range of 1.5 in. or 2 in.) are about 60%. The RH values at deeper locations approach to 70%, which is higher than the surface value. The values ofRH reflect the annual average RH value ofthe environment. The gradient of RH is not large in the time period of Oct. and Nov., 2011.

After the test for internal relative humidity, a ponding test was performed using the same cylinder. The concrete cylinder was placed upright with the outer surface facing up. A water column of 13 cm was placed on top of the cylinder. The purpose of the test was to examine the resistance of the concrete to water and moisture penetration. The sensor at 3.0 in was damaged during the first test. So, seven sensors were used in the ponding test with the distances to the top surface 1.0 in, 1.5 in, 2.0 in, 2.5 in, 5.0 in, 7.0 in, and 8.5 in.

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Exhibit 52 Fig. 3 shows the test data. The initial RH distribution is between 30% to 45%. This is because the test started on Feb. 7, 2012, more than two months after the fLrst test. Within two days, the water penetrated to the depth of2 inches, which indicated that the resistance of the surface layer concrete to water penetration is qu ite low. The concrete at 2.5 in. and deeper portion shows much higher resistance. The high moisture region with RH > 90% reached about 2.5 inches after four days of pond ing.

These results showed that the rate of moisture penetration into the concrete depends strongly on the quality of surface layer concrete, where microcracks may form due to various deterioration mechanisms such as drying shrinkage. In order to determine moisture resistance of the concrete at different locations of the structure, more samples need to be taken from the structure and tested.

3. Com pressive Strength The compressive strength of concrete was tested according to ASTM C39. It is for the unconfined compressive strength ofcylindrical concrete specimens. Four samples were used for the test. The identifications of the samples are shown in Table 2 and Fig. 4. Dimensions of the specimens and the test data are listed in Table 2.

Table 2 Compressive strength ofthe four specimens No Diameter (in.)

Length (in.)

fc' (psi)

Specimen description 1

2.65 5.76 5444 Hallway #J 2

2.65 5.76 6342 S9680-3 3

3.39 5.88 7990 S4 11/8111 4

3.39 6.38 10508 S4 11/8111 Fig. 4 Specimen # 1 (right) and #2 (left) after the compressive strength test

4. Splitting Tensile Strength The splitting tensile strength of cyl indrical concrete specimens was tested according to ASTM C496. The maximum load recorded, P, was used to calculate the splitting tensile strength of concrete samp les based on Eq. I.

Page 3 of 1 0

Exhibit 52 2P (Eq. I) j" =Trld in whichls, = splitting tensile strength; I and d are the length and diameter ofspecimen, respectively. Test results are shown in Table 3.

1 Table 3 S,prItfmg ensl e s ren gth t est d t aa Specimen Length (in)

Diameter (in)

Area (in!\\2)

Force (kips) lsi (psi)

Specimen description No. \\

4.2 3.68 10.63 67.110 957.43 S8 11/8/ 11 No.2 5.1 3.68 10.63 67.447 962.23 S8 11/3/11 No.3 5.4 3.68 10.63 58.585 835.8 S3 11/8111

5. Coefficient of Thermal Expansion (CTE)

Two cylindrical specimens (#S2 11/8/11, #S4 11/8/11) were used for the test. The diameters of the two samples are the same, 3.39 in. Thermal expansions of the two specimens were measured between two temperature ranges from noe to 400e and then from to 400e to 60oe, then linear coefficients of thermal expansion were calculated based on the test data.

The tests were conducted in an environmental chamber with temperature control. So, it is different from USBR 4910-92 conducted in u.S. Bureau of Reclamation, in which the specimens were submerged in water and the water temperature is varied to create the thermal expansion. The purpose of this test is to obtain eTE ofconcrete used in above ground structures.

Thermal sensors were installed inside of concrete samples to double check the internal temperature in concrete samples. When the internal temperature reaches the target temperature, deformation of concrete sample was measured after two hours of holding of the target temperature. This was to make sure that the internal temperature distribution in the cylinder reaches equilibrium (uniform distribution). The test data of the two specimens are shown in Table 4 and Table 5.

Table 4 eTE of Specimen #S2

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Lenfrth (in)

~L (ill) eTE (11°C)

Average eTE (11°C) 18 (from 22 to 40°C) 10.157 0.0013 0.00000711 0.00000875 20 (from 40 to 60°C) 10.157 0.002\\

0.00001034 38 (from 22 to 60°C) 10.157 0.0034 0.00000881 Table 5 eTE of Specimen #S4

~T (0C)

Length (in)

~L (in) eTE (11°C)

Average eTE (\\1°C) 18 (from 22 to 40°C) 10.284 0.0011 0.00000594 0.00000886 20 (from 40 to 60°C) 10.284 0.0024 0.00001167 38 (from 22 to 60°C) 10.284 0.0035 0.00000896 Page 4 of 10

Exhibit 52 The average value ofeTE of the two specimens = 8.8xl0-6 /0e. The eTE measured by USBR =

5.2x10-6 /oF = 9.4x I0-6 fOe. Our eTE value is slightly smaller than the eTE measured by USBR because when the concrete was heated in air, the measured thermal expansion is actually a combination of pure thermal expansion and drying shrinkage.

In add ition to the test of e TE in the temperature range above ooe, another series of tests was performed for the thermal strains in the temperature range below O°e. The purpose of this test was to examine the effect of ice formation on thermal expansion of the concrete. The testing sample was S2 11 /8/11. The diameter of the specimen is 3.46 in. and 11.5 in. long.

Two tests were performed. One is called dry test. The specimen was placed in a high temperature chamber for 14 days under 80 0 e to dry out the internal moisture, and then placed in a freezing chamber with programmable temperature control. The test started at 20oe, the temperature was reduced to 15°e in 30 minutes, stayed at 15°e for 3 hours3.472222e-5 days 8.333333e-4 hours 4.960317e-6 weeks 1.1415e-6 months , and the strain was measured. The process was repeated until the target temperature of -25°e. The test data are shown as the blue curve in Fig. 5.

Strain-Temp relations 10.00E+00

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After the test for the dry sample, the specimen was placed in a water tank for 68 hours7.87037e-4 days 0.0189 hours 1.124339e-4 weeks 2.5874e-5 months for saturation. The same testing procedure was used to obtain the thermal strain under low temperatures. The test was called wet test. The test data are shown as the red curve in Fig. 5. Jt is important to see from Fig. 5 that the concrete contracts upon cooling from 20 0 e to ooe, starts to expand from ooe to -15°e, and then starts to contracts again. The first reversal from contraction to expansion is due to the ice formation in the concrete, because the eTE of ice is about five times higher than the eTE of concrete. The second reversal is an indication of the completion of ice formation. Both ice and concrete contract upon a further cooling.

Page 5 of 10

Exhibit 52 The test results ind icated that with a high moisture content, the effect of ice formation on thermal strain of the concrete sample is significant, resulting in an expansion under the low temperature from O°C to -15°C. Because oflimited time, the internal moisture distribution in the sample may not be uniform, so the measured strains represent average values of the thermal strains. In order to determine the coupling effects among moisture content, low temperature, and ice formation, a more systematic experimental study with more samples is needed.

6. Accelerated Creep The accelerated creep tests were performed to obtain creep strain of the concrete used in Davis Besse Nuclear Power Station. The creep tests generally follow the procedure described in ASTM C-512 "Standard Test Method for Creep in Compression". Three accelerated creep tests were performed under 40°C (with and without humidity control) and 80 °C (with humidity control), respectively. Different relative humidity controls were used in the tests to find the effect of moisture level on the creep ofconcrete. Some ofthe basic terminologies used in this section are Basic creep - The long-term strain ofconcrete due to load ing without drying and heating.

Drying shrinkage - The long-term strain ofconcrete due to drying without load ing and heating.

Drying creep - The long-term strain ofconcrete due to loading and drying without heating.

Fig. 6 The MTS machine provides a stable compressive force at 16 kips 6.1 Testing method

One cylindrical specimens (#S2 11/8111) was cut into a cylinder of 11.5 in. long.

The diameter of the specimen is 3.46 in.

Two contact points were installed on the top and bottom portion of the specimen. The distance between the two contact points equals the gage length of the dial gauge to be used to measure the length change of the specimen.

The specimen was capped on top and bottom surfaces.

The specimen was loaded by a MTS machine mounded in an environmental chamber. The Page 6 of '0

Exhibit 52 loading level was kept as a constant 16 kips, which resulted in a compressive stress of 1702 psi. This stress level is less than 40% of the average compressive strength of the concrete (7,600 psi). Fig. 6 shows the loading setup.

The chamber was maintained at a constant temperature of40°C or 80°C.

For the first specimen, the temperature was kept at 40°C without humidity control. For the second specimen, the temperature was kept at 80°C, and the humidity was controlled in the range of 70% to 80%. For the third specimen, the temperature was kept at 40°C, and the humidity was controlled in the range of70% to 80%.

6.2 Test results The test results of the three accelerated creep tests are plotted in Figs. 7, 8, and 9. Fig. 7 shows the test results under 40°C without humidity control; Fig. 8 is for the test results under 80°C with humidity control; and Fig. 9 is for the test results under 40°C with humidity control.

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I Log Time (hours) 10 100 1000 Fig. 9 Test results obtained under 40°C with humidity control 6.3 Comparisons and discussions Comparison of Fig. 7 and Fig. 9 shows the effect of drying on concrete creep. The creep reading in Fig. 7 is about 0.0005 (500 microstrain) after 100 hours0.00116 days 0.0278 hours 1.653439e-4 weeks 3.805e-5 months , which is about 2.5 times the value in Fig. 9, which is about 0.0002 (200 microstrain). The difference between the two tests was the relative humidity in the chamber. In Fig 9, the relative humidity was controlled at the range of 70%-80%, so the test data represent basic creep ofthe concrete. In Fig. 7, there was no humidity control, so the test data represent a combination of basic creep and drying shrinkage, and that is why the measured strains are much higher. Basically, the effect of drying shrinkage is quite significant, and the creep of concrete under an arid environment (low humidity) is much higher than that in a humid environment.

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Page 8 of lO

Exhibit 52 Comparison of Fig. 8 and Fig. 9 shows the effect of temperature on concrete creep. The same humidity controls were used for the two tests, and thus there is no effect of drying shrinkage.

Combin ing Fig. 8 and Fig. 9 gives Fig. lOin log time scale and Fig. II in regular time scale.

Fig. II can be used to obtain creep compliance functions under two different temperatures, which can be used further to obtain the creep coefficient of concrete.

7. Freeze-thaw resistance The accelerated freeze-thaw tests were planned to obtain freeze-thaw ofthe concrete used in Davis-Besse Nuclear Power Station. The freeze-thaw tests generally follow the procedure described in ASTM C-666. The testing chest is shown in Fig.12. Each freeze-thaw cycle is approximately 2-3 hours. There will be 300 cycles. Two samples were used for the test: S9 680 3 and In Steam Room 602 #1 Fig. 12. The rapid freeze-thaw test chest After one day of testing, the temperature controller of the testing machine was broken. The test was stopped. After a new controller was installed, we did not re-start the test, because this is a long-term test and it cannot be done before the completion ofthe project.

A summary table is shown in the next page for identifications of all sample tested at University ofColorado.

Page 9 of lO

Exhibit 52 Tests at Univ. of Colorado Test Core identification Internal moisture and ponding test 56 11/8/11 Compression Hallway #1 59680-3 5411/8/11 5411/8/11 Splitting tension 5811/8/11 5811/8/11 5311/8/11 Coefficient of thermal expansion 52 54 Creep and CTE at low temperatures 5211/8/11 Freeze-thaw 59680-3 In Steam Room 602 #1 Page 10 of 10

Exhibit 53: C-0109 Roof Plans and Details Appendix VIII-54

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Exhibit 55: AC! 201.2R-08, Table 6.3

Exhibit 55 IGUIDE TO DURABLE CONCRETE Table 6.1-Effect of commonly used chemicals on concrete Moderate Phosphoric Organic acids Alkaline solutions Miscellaneous Acetic Formic Lactic Aluminum chlori(le I

Tannic Sodium hydroxide- > 20%

Ammonium nitrate Ammonium sulfate Sodium sulfate Magnesium sulfate Calcium sulfate Bromine (gas)

Sulfite liquor Slow Carbonic Sodium hydroxide' 10 to 20%

Sodium hypochlorite Ammonium chloride Magnesium chloride Sodium cyanide Chlorine (gas)

Seawater Soft water Negligible Oxalic Tartaric Sodium hydroxide' < 10%

Sodium hypochlorite Ammonium hydroxide Calcium chloride Sodium chloride Zinc nitrate Sodium dichromate Ammonia (liquid)

"The effect of potassium hydroxide is similar to that of sodium hydroxide, DePuy (1994), Taylor (1997), Skalny et al. (1998), Thomas and Skalny (2006), and Naik et al. (2006). Publications with particular emphasis on permeability and the ability of concrete to resist ingress and movement of water include Reinhardt (1997), Hearn et al. (1994), Hearn and Young (1999), Diamond (1998), and Diamond and Lee (1999).

6.2.3 Recommendations-Protection against sulfate attack is obtained by using concrete that retards the ingress and movement of water and concrete-making ingredients appro priate for producing concrete having the needed sulfate resistance. The ingress and movement of water are reduced by lowering the wlcm. Care should be taken to ensure that the concrete is designed and constructed to minimize shrinkage cracking. Air entrainment is beneficial if it is accompanied by a reduction in the wlcm (Verbeck 1968). Proper placement, compaction, finishing, and curing of concrete are essential to minimize the ingress and movement of water that is the carrier of the aggressive salts. Recommended procedures for these are found in ACI 304R, 302.1R, 308R, 305R, and 306R.

The sulfate resistance of portland cement generally decreases with an increase in its calculated tricalcium-alumi nate (C3A) content (Mather 1968). ASTM C150 permits the use of Type V sulfate-resisting cement and C3A with a maximum limit of 5%, and Type II moderately su1fate resisting cement and C3A limited to 8%. There is also some evidence that the alumina in the aluminoferrite phase of portland cement can participate in sulfate attack. Therefore, ASTM C 150 states that the C4AF + 2C3A in Type V cement should not exceed 25% unless the alternate requirement based on the use of the performance test (ASTM C452) is invoked. In the case of Type V cement, the sulfate-expansion test (ASTM C452) can be used instead of the chemical requirements (Mather 1978). The use of ASTM ClO12 is discussed by Patzias (1991).

Table 6.3 provides recommendations for various degrees of potential exposure. These recommendations are designed to protect against concrete distress from sulfate from sources external to the concrete, such as adjacent soil and groundwater.

Recommendations for the maximum wlcm and the type of cementitious material for concrete that will be exposed to sulfates in soil or groundwater are given in Table 6.3. Both Table 6.2-Factors influencing chemical attack on concrete Factors that accelerate or aggravate attack Factors that mitigate or delay attack I. High porosity due to:

I. Dense concrete achieved by:

i. High water absorption

i. Proper mixture proportioning' ii. Penneability ii. Reduced unit water content iii. Voids iii, Increased cementitious material content iv. Air entrainment

v. Adequate consolidation vi. Effective curing t

2. Cracks and separations due to:

2. Reduced tensile stress in concrete L Stress concentrations by:*

ii. Thermal shock L Using tensile reinforcement of adequate size, correctly located ii. Inclusion of pozzolan (to reduce temperature rise) iii. Provision of adequate contraction joints

3. Leaching and liquid penetration

3. Structural design:

due to:

L To minimize areas of contact

i. Flowing liquid§ and turbulence ii. Ponding iii, Hydraulic pressure ii. Provision of membranes and protective-barrier system(sJII to reduce penetration

'The mixture proportions and the initial mixing and processing of fresh concrete detennine its homogeneity and density.

tPoor curing procedures result in flaws and cracks.

Resistance to cracking depends on strength and strain capacity.

§Movement of water-carrying deleterious substances increases reactions that depend on both the quantity and velocity of flow.

"Concrete that will be frequently exposed to chemicals known to produce rapid deteriora tion should be protected with a chemically resistant protective-barrier system.

of these recommendations are important. Limiting only the type of cementitious material is not adequate for satisfactory resistance to sulfate attack (Ka10usek et al. 1976).

The field conditions of concrete exposed to sulfate are numerous and variable. The aggressiveness ofthe conditions depends on soil saturation, water movement, ambient temperature and humidity, concentration of sulfate, and type of sulfate or combination of sulfates involved. Depending on the aforementioned variables, solutions containing calcium sulfate are generally less aggressive than solutions of sodium sulfate, which is generally less aggressive than magnesium sulfate. Table 6.3 provides criteria that should maximize the service life of concrete subjected to the more aggressive exposure conditions.

Exhibit 55 Page 1 of 2

Exhibit 55 201.2R-24 ACI COMMIITEE REPORT Table 6.3-Requirements to protect against damage to concrete by sulfate attack from external sources of sulfate Severity of potential exposure Water-soluble sulfate (SO~ in soil, % by mass

Sulfate (S04)* in water,ppm wlcm by mass, max.tt Cementitious material requirements Class 0 exposure 0.00 to 0.10 oto 150 No special requirements for sulfate resistance No special requirements for sulfate resistance Class 1 exposure

>0.10 and <0.20

> 150 and < 1500 0.50i C 150 Type II or equivalent§ Class 2 exposure 0.20to<2.0 1500 to < 10,000 OA5i C150 Type V or equivalent§ Class 3 exposure 2.0 or greater 10,000 or greater 0040*

C150 Type V plus pozzolan or slag§ Seawater exposure See Section 604 See Section 6.4

Sulfate expressed as S04'S related to sulfate expressed as S03' as given In repOrts of chenucal analYSIS of portland cements as follows: S03% x 1.2 = S04%'

tACI 318, Chapter 4, includes requirements for special exposure conditions such as steel-reinforced concrete that may be exposed to chlorides. For concrete likely to be SUbjected to these exposure conditions, the maximum wlcm should be that specified in ACI 318, Chapter 4, ifi! is lower than that stated in Table 6.3 of201.2R.

tValues applicable to normalweight concrete. Tbey are also applicable to structural lightweight concrete except that the maximum wlcm ratios 0.50, 0.45, and 0040 should be replaced by specified 28-day compressive strengths of 26,29, and 33 MPa (3750. 4250, and 4750 psi), respectively.

§For Class I exposure, equivalents are described in Sections 6.2.5, 6.2.6, and 6.2.9. For Class 2 exposure, equivalents are described in Secrions 6.2.5, 6.2.7, and 6.2.9. For Class 3 ex sure, pozzolan and sla recommendations are described in Sections 6.2.5, 6.2.8, and 6.2.9.

Portland-cement concrete can be also be attacked by acidic solutions, such as sulfuric acid. Infonnation on acid attack is provided in Section 6.5.

6.2.4 Sampling and testing to determine potential sulfate exposure-To assess the severity of the potential exposure of concrete to detrimental amounts of sulfate, representative samples should be taken of water that might reach the concrete or of soil that might be leached by water moving to the concrete. A procedure for making a water extract of soil samples for sulfate analysis is given in Appendix A. The extract should be analyzed for sulfate by a method suitable to the concentration of sulfate in the extract solution, such as the photometer methods used in ASTM C1580. If the amount of sulfate detennined in the first analysis is outside of the optimum concentration range for the analytical procedure used, the extract solution should be either concentrated or diluted to bring the sulfate content within the range appropriate to the analytical method, and the analysis should be repeated on the modified extract solution.

6.2.5 Material qualification of pozzolans and slag for sulfate-resistance enhancement-Tests I year in duration are necessary to establish the ability of pozzolans and slag to enhance sulfate resistance. Once this material property has been established for specific materials, proposed mixtures using them can be evaluated for Class I and 2 exposures using the 6~month criteria in Sections 6.2.6 and 6.2.7.

Fly ashes, natural pozzolans, silica fumes, and slags may be qualified for sulfate resistance by demonstrating an expansion:s;; 0.10% in 1 year when tested individually with portland cement by ASTM ClO 12 in the following mixtures.

For fly ash or natural pozzolan, the portland cement portion of the test mixture should consist of cement with Bogue-calcu lated C3A of not less than 7%. The fly ash or natural pozzolan proportion should be between 25 and 35% by mass, calculated as percentage by mass of the total cementitious material.

For silica fume, the portland cement portion of the test mixture should consist of a cement with Bogue-calculated C3A of not less than 7%. The silica fume proportion should be between 7 and 15% by mass, calculated as percentage by mass of the total cementitious material.

For slag, the portland cement portion of the test mixture should consist of a cement with Bogue-calculated C3A of not less than 7%. The C3A should be calculated for the sum of the portland cement plus calcium sulfate in the cement. Some processing additions, if present in sufficient proportions, can distort the calculated Bogue values. Fonnulas for calculating Bogue compounds may be found in ASTM C 150.

The slag proportion should be between 40 and 70% by mass, calculated as a percentage by mass of the total cementitious material.

Material qualification tests should be based on passing results from two samples taken at times a few weeks apart.

The qualifying test data should be no older than I year from the date of test completion.

The reported calcium-oxide content analyzed in accordance with ASTM Cl14 of the fly ash used in the project should be no more than 2.0 percentage points greater than that of the fly ash used in qualifying test mixtures. The reported aluminum-oxide content analyzed in accordance with ASTM C 114 of the slag used in the project should be no more than 2.0 percentage points higher than that of the slag used in qualifying test mixtures.

6.2.6 Type II equivalent for Class 1 exposure ASTM C 150 Type III cement with the optional limit of 8% maximum C3A; ASTM C595 Type IS(MS), Type IP(MS), Type IS-ACMS), or Type IP-A(MS); ASTM Cl157 Type MS; or Any blend of portland cement of any type meeting ASTM Cl50 or Cl157 with fly ash or natural pozzolan meeting ASTM C618, silica fume meeting ASTM C 1240, or slag meeting ASTM C989 that meets the following requirement when tested in accordance with ASTMC1012:

Expansion S; 0.10% at 6 months Any fly ash, natural pozzolan, silica fume, or slag used should be previously qualified in accordance with Section 6.2.5.

6.2.7 Type V equivalent for Class 2 exposure ASTM Cl50 Type III cement with the optional limit of 5% maximum C3A or ASTM C150 cement of any type having expansion at 14 days no greater than 0.040%

when tested by ASTM C452 or ASTM el157 Type HS; or Any blend of portland cement of any type meeting Exhibit 55 Page 2 of2

Exhibit 56: Structural and Thermal Analysis Investigation

Exhibit 56 Redacted Thennal SIres, J.\\n~'IV~Il'<'

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Final Report Results & Comments Exhibit 56

- Summary Final Report -

Revision dated 24 Feb 2012 1.0 Duties and Responsibilities This report summarizes the activity of to the structural and thermal analysis investigation work performed at

-JLlv.,,,,,, nuclear power plant.

1.1 3DNastran_FEM Primary responsibility is the development of3D Nastran _Finite Element Models (FEM's) for use in computing thermal transient temperature distributions due to various environmental conditions. These 3D _

FEM's include pressure loading that result from wind due to Tornados and other Wind conditions during the winter and summer cases.

The 3D Nastran _

FEM is used in the thermal transient heat transfer analysis performed to compute solar heating/cooling for the following environmental conditions:

./ Summer Solstice (Hot w/o Wind, Hot w/34 mph Wind & Ave w/o Wind)

./ Autumn Equinox ( ""

"""" ")

./ Winter Solstice (Ave w/o Wind)

./ 1978 Blizzard

( w/105 mph Wind)

./ Vernal Equinox (Ave w/o Wind, Ave w/36 mph Wind)

Note: The initial series ofanalysis showed Vernal Equinox conditions were not critical.

The 3D Nastran..FEM's were also used to provide approximations for stresses &

deflections throughout the Davis-Besse Shield Build due to combined effects ofwind, thermal transients and 1.2 Nastran 2-D Plane-Strain _

Idealizations During the course ofthese and evaluated.

- - - - information. Performance Improvement international, LLC.

Page 1 of 39 2012

The total number of elernenlts and Exhibit 56 Redacted Thermal Stress Analysis:

Final Report Results & Comments 1.0 Nastran Finite Element Model Definitions The analysis code used for the transient thermal and structural analysis is MDlNastran 2010 v1.3.

MD Nastran is a general purpose fmite element program for performing linear, nonlinear structural analysis, vibration, dynamics and thermal analysis.

2.1 3D _

Finite Element Model Figure 2.1.1 shows an isometric view ofthe 3D Nastran _

idealization. Key point details of the development and definition for the 3D Nastran Thermal Transient and Structural FEM are the following:

./ The 3D Nastran Nastran 360 0 3D models idealize the entire Shield Building idealization was

./ Element size through the 30" concrete wall elements and node comprising the Nastran 3D

~odes. The total number ofdegrees offreedom

./ The reference drawings used to develop the Nastran 3D ~EM are Dwg. No. C-100, C-I04 & C-I09

./ The overall region idealized is from EL 567' 6" [base truncation level] to the top ofthe Dome EL 824' 3 Yz". The inside radius RIF 69'6" and outside radius ROF 72' [vertical wall thickness 2' 6"]. The Dome wall thickness = 2'.

./ Concrete Reinforcement

./

astran _

models idealize hoop and vertical reinforcement The 3D Nastran _FEM ""'....VB"'"'

results from other analysis u.v~", *...,.

_FEM is configured for use

& structural model is

.....,.,."Ull" a cross check validation of 3D Nastran

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Page 2 of 39 2012

Exhibit 56 Redacted 11191==

Th"ma' St,."

Fina' Report Res"" & Commen"

\\

i I

I

\\

,/ The ~D NastrAn _

models Idealize ~e entire'

\\

.J 'ld' contamment cpncrete oUl mg I

\\.

\\

I

\\

,/ The ~verall region ide~lized is ftom EL ~67' 6" [\\lase trunc~tion Ievt I] to the:top of th~ Dome ltL 82 4 ' 3 '/,",

\\

,/ The ipside ra~ius R,p ~! 69 ' 6":'S,d outsit radius 72' [1ertical "'1aU thickfess = 2' \\6"1. The Dome w thickness = 2'l AT II

./ Conorete Reir~forcemeI).t I

I

\\

\\ 3D_

Idealization: Key Enwlope Dlme)islons I

a.) Net hei~t of \\lIe ve1cal walls 4242' (EL 597' 6" to::L $09' 6")

\\

b.) Net height to the top \\ofDome,: 356' 9-112" (1\\:L 567' 6" to\\EL 834' 3-1/2")

,,),,,id, obm"",fV,",'" w"',,\\, !l9' (1';,\\69' 6")

d)y'",,",",oil Tru'1"' 2' 6" Ir,d, ' ' ~ Ro ' 72')

\\

y ~X millWfi~}tT.

-\\ Base EL 567' 6" Upper EL 824' 31/2" Truncation Level I

~e2.1.1 3D Nastra _

Model Idealization page 3 of 39

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2012

Exhibit 56 Redacted Thermal Stress Analysis Final Report Results & Comments "IEIJ!w

./

I Th~ 3D Nastran z

'(~..,

I I

Figure 2.1.2 3D Nastran _

Model: Steel Reinforcement

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2012 Page 4 of 39

Exhibit 56 Redacted Thermal Stress Analysis:

Final Report Results & Comments Concrete Figure 2.1.3 3D Nastran _Model: Typical Section Cut

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2012.

Page 5 of 39

Exhibit 56 Redacted Thennal Stress Analysis:

Final Report Results & Comments w:t'i'1'I!

A,

\\1:1I(>1"lal Pr:o]>l'I"tlt'<, for DaYh-Bp~,(' Oi'('rall_Fiultt> £1('lnPI!1 \\IoII('IIF£\\II hlpnllzrI I USSR T~Jt VoJuc-s,nr c-mall memorondum 11 January 2012 RI:vi SI}(( t) ~'nll 2012 I

Generic Steel jRebar & Inner Steel containmLt Eokut &!Oroh. "H~ot & Mo.. TroM/u" I

I TalJ\\t' I, ('ou(,I'('I(' & $1('('1 r"Ilr"""nloliye of \\peciiic re~ ioll' lIla~ reflee l loc:.lized lllc!.1 >lU'elllenI S. i I

TIt~rmtJl Pfopcrt/J:

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S_oI,Hi;J!.c,.

Em ~~Slvny, 'E j" f hEorm3 t f Xp.ioSlon Youn,', r-todulu'. E Po..nOrl*s RatJo. v

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atu/lb or'

,n/lntf (. 10")

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80 O. ~8 ~

1.610 86.074 I

O.IlO 0.15 680 2~,OO 0.30 Figure 2.1.4 Material Properties for Davis-Besse 3D Nastran _

Model

- - - - infonnation. Perfonnance Improvement international, LLC.

2012 Pag e 6 o f *39

Exhibit 56 Redacted Thermal Stress Analysis:

Final Report Results & Comments

-"'IJ!

2.2 Nastran 2D Plane-Strain Utility Idealization Figure 2.2.1 shows an overall isometric view of the complete 2D Nastran plane-strain_

idealization and a close-up view ofthe rebar and concrete mesh details. Key point details of the development and definition for the 20 Nastran plane-strain FEM are the following:

./ The 20 Nastran plane-strain models idealize a section cut through the Shield Building EL 683 ' 6".

at

./ The total number of elements and node comprising the Nastran 20 plane-strain model are

_elements and _nodes.

./ The total number ofdegrees of freedom ~OFs.

./ The reference drawings used to develop the Nastran 20 plane-strain No. C-110 ion idealized is at EL 683' 6" This elevation is defined as a reference elevations 683' 6" is along the vertical walls approximately half-way between EL 567' 6" [base truncation level] to just under the ring girder EL 80 I' 6-1 /2".

EL Owg.

. structural model is also used

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Page 7 of 39 2012

Exhibit 56 Redacted Final Report Results & Comments

,:Ii1@*

Thennal Stress Analysis:

I.

Close-U p VievJ of 2-D Plane-Strain fEM Rebar and Concrete De finition fulllsofTIetric View of 2-D Plane-Strain FEM Defin it ion Figure 2.2.1 20 Nastran Plane-Strain

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Page 8 of 39 2012

Exhibit 56 Redacted Thermal Stress Analysis: _

Final Report Results & Comments

.'i 3.0 Step-by-Step Analysis Process

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Exhibit 56 Redacted

-'ID">>-

Thennal Stress Analysis:

Final Report Results & Comments Figure 3.0.1 Schematic Flow Chart Representation of"Step-hy-Step" Analysis

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2012 Page 10 of 39

Transient thermal temperatures due to the Winter Solstice and 1978 Exhibit 56 Redacted Thermal Stress Analysis:

Final Report Results &Comments 3.1 Typical Output Results 2D Plane-Strain Utility Model The lane-strain idealization for the full 360 0 Shield Building wall was_

~ show examples ofoutput results These summary plots show the distribution of maximum principal and radial stresses for the peak summer solstice condition at 7:30 pm; respectively.

In these figures the S-W facing Flutes are showing the highest magnitude of maximum principal and radial stresses. The peak stress results occur at the outer rebar regions due to SCF effects where the overlapping rebar ends in the thick portion ofthe Flute.

Radial stresses are plotted for each ofthe selected time slices as the sun traverses the sky during the 24 hour2.777778e-4 days 0.00667 hours 3.968254e-5 weeks 9.132e-6 months period beginning.

One ofthe key aspects ofthis 20 plane-strain _

is that all ofthe action is in the thick portion of the Flutes with peripheral, secondary action along the rebar at the OF.

One ofthe key aspects ofthis study is that all ofthe action is in the thick portion ofthe Flutes with peripheral, secondary action along the rebar at the OF Figure 3.1.3 shows summary peak rad ial stresses during the 24 hour2.777778e-4 days 0.00667 hours 3.968254e-5 weeks 9.132e-6 months period for the environmental conditions listed above. From Fi re 3.1.3 the time slices ducing the highest rad iaI stresses

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Page 11 of 39 2012

Exhibit 56 Redacted Thermal Stress Analysis:

Final Report Results & Comments e tagruru~ of suess values sbown from mo :m plane stram

)ccur at pt"ak SCF ~{fects These stress contour plots are u!t~ [~q~!atlyely ~IeC[ concal ttme mten'3ls.

Malumum Principal Stress Highest In ThICk PorttOn of Flute OJetal( View @ EL 683' 6" Close-Up View @ EL 683' 6" I

I Figure 3.1.1 Summer Solstice Hot No Wind 7:30 pm, Constant Concrete CTE = 5.20 xlO-6 iniinJOF 2D Plane-Strain Maximum Principal Stress Distribution Note: The thick regions of the West and Southwest facing architectural flutes indicate the highest magnitude of maximum principal stress values resulting from the summer solstice conditions.

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2012 Page 12 of 39

A Exhibit 56 Redacted Thermal Stress Analysis :

Final Report Results & Comments mE.Ir_

The rn3g!l1tIldeSf SlUSS \\'3Iu&s shown from the ~D plane stram roce a! peak SCF effects. The~ stress contour plots are used to tlvely sekct cOl1c3lllme IDlerYais Maximum Radial St(e~s HIghest In Thick Portion of Flute Close-Up View @ EL 683' 6" Figure 3.1.2 Summer Solstice Hot No Wind 7:30 pm, Constant Concrete CTE = 5.20 x I 0-6 in/in/oF 2D Plane-Strain Radial Stress Distribution Note: The thick regions of the West and Southwest facing architectural flutes indicate the highest magnitude of maximum principal stress values resulting from the summer solstice conditions.

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2012 Page 13 o f 39

Exhibit 56 Redacted

-:'..-.'_l*

A >Thoemal St"" Analy,,,

Floal Report R"olts & Commeot,

.:miIU~ ~ ~I~ ~ ~ ~I~ II ~ ~ ~III ~ i 11111111111111111111111111 1111111111111111111111111 Summer Solstice Hot No Wind: 7:30 PM Summer Solstice Hot M mph Wind: 6:00 AM Summer Solstice Ave No Wind: 7:30 PM

...1_-____

11111111111111111111111111 1978 Blizzard Record Lov, 150 mph Wind: 5:00 AM Winter Solstice Ave No Wind: 7:30 AM

_r___ _

--.-..-.. -~

---1---.---...

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

_1... ---

t...

~

~ _

_~_......t~......

111111111111111111111111 111111111111111111111111 HHlUni Autumn Equinox Hot No Wind: 5:00 AM & 6:00 PM Autumn Equinox Hot 34 mph Wind: 5:00 AM Autumn Equinox Ave No Wind: 5:00 AM Figure 3.1.3 Survey Radial Stress Results: Nastran 2D Plane-Strain r EM; Heat Transfer Analysis; 24/1 Hour Time [192 - I hour Time Slices]

Note:

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2012 Page 14 of 39

Exhibit 56 Redacted Inal Report Results & Comments Thermal Stress Ana lysis:

.".IM 3.2 Typical Output Results 3D Figures 3.2.1 and 3.2.2 show examples of output results fcom the mapped thermal transient thermal stress analyses. These summary plots show the distribution of maximum principal and radial stresses for the peak summer solstice condition at 7:30 pm; respectively.

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2012 Page 15 of 39

Exhibit 56 Redacted

. ),.

Final Report Results & Comments

-"lM1W a.\\

Thermal Stress Analysis:

Overall View @ EL 683' 6" Close-Up View @ EL 683' 6" Figure 3.2.1 Summer Solstice Hot No Wind 7:30 pm, Constant Concrete eTC = 5.20 xl 0.6 inlin/oF 3D _

FEM Maximum Principal Stress Distribution Note: Non-Symmetric Thermal Stresses Due to Uneven/IIigher Heating Gradients on South Facing Panels.

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2012 Pag e 16 of 39

Exhibit 56 Redacted Final Report Results & Comments IImBIe-Thermal Stress Analysis

.a. '" >

1...... 1'~ ___

(,..,.

- ~

Radial SHess In Thick Portion of RUle o~ '" +76 pSI Overall View @ EL 683' 6" Close-Up View @ EL 683' 6" Figure 3.2.2 Summer Solstice Hot No Wind 7:30 pm, Constant Concrete CTE = 5.20 x l 0-6 irvin/oF 3D _

FEM Radial Stress D istribution Note: Non-Symmetric Thermal Stresses Due to Uneven/Higher Heating Gradients on South Facing Pane ls.

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2012 Page 1 7 of 39

Exhibit 56 Redacted Thermal Stress Analysis:

Final Report Results & Comments

.1:,.,'&.

4.0 Summary Results and Comm nts These results are from using the 3D FEA model for constant coefficient of thermal expansion (eTE) thermal stress analys is.

4.1 Summer Solstice Conditions Table 4.1 summarizes results from the Summer Solstice conditions. These results correlate with the hot daytime peak temperatures that occurred during the period from 1959 to 2004 in the Toledo, OH area. The "No Wind" condition removes heat by convection. The hot condition uses the high temperatures measured for June of 104°F during the day 84°F at night. The average conditions Llsed the average day temperature of 83°F and 63°F at night. For aU cases gravity is also include. The contribution due to pressure loading from wind has been demonstrated to have a negligible impact on overall stress results.

2D Nastran Plane-Strain Time Slice Peak Stress 3

Nastran _

Thick Flute POliioll FEM Peak Radial Stress Architectural Notch ID Case Description 1

Summer Solstice Hot No Wind 7:30 PM

+76 ps i

- 140 psi 2

Summer Solstice Hot 34 mph Wind 6:00 PM

+46 PSi

- 68 psi 3

Summer Solstice Ave No Wind 7:30 PM

+69 pSi

-126 PSi Table 4.1 Summer Solstice - Summary Resu lts for Radial Stress @ EL 683' 6" Figure 4.1.1 tlu'ough Figure 4. 1.3 show summary results listed in Table 4.1 for radial stress due to therma 1 transients.

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Page 18 of 39 2012

Exhibit 56 Redacted EE*Iv-A,\\".

Thermal Siress Analysis:

Final Report Results & Comments

...&a:>

'U' 0, =+ 76 psi Overall View @ EL 683' 6/1 Close-Up View @ EL 683' 6J/

Figure 4.1.1 Summer Solstice Hot No W ind 7:30 pm, Constant Concrete CTE = 5.20 xl 0.6 in/ inrF 3D _

FEM Radial Stress Distribution Note: Non-Symmetric Thermal Stresses Due to Uneven/ Higher IIeating Gradients on South Facing Pane is.

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201 2 Page 19 of 39

Exhibit 56 Redacted RliJ'ilIffiW Thermal Stress Analysi Final Report Results & Comments a~ " +46 pSI Overall View @ EL 683' 6/1 Close-U p View @ EL 683' 6" Figu re 4.1.2 Summer Solstice Hot 34 mph Wind 7:30 pm, Constant Concrete e TE = 5.20 xl 0-6 inlin!OF 3D _

FEM Radial Stress Distribution Note: No n-Symmetric Thermal Stresses Due to UnevenlHigher Heating Gradients on South Facing Pane ls.

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2012 Page 20 of 39

Exhibit 56 Redacted r.lIlilE1e Thermal Stress Analysis:

Final Report Results & Comments a&>

C1; =+ 68psl Overa ll View @ EL 683' 6" Summer Solstice Ave No Wind 7:30 pm, Constant Concrete CTE = 5.20 xl O-6 in/ in/oF 3D _

FEM Radial Stress Distribution Note: Non-Symmetric Thermal Stresses Due to Uneven/H igher Heating Gradients on South Facing Panels.

Close-Up View @ EL 683' 6" liigure 4.1.3

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2012 Page 21 of 39

Exhibit 56 Redacted Final Report Results & Comments

.IIQM_

Thermal Stress Analysis:

4.2 Winter Solstice & 1978 Blizzard Conditions Table 4.2 summarizes results fi'om the Winter Solstice and 1978 Blizzard conditions.

The 1978 Blizzard computed cold temperatures correlate with the coldest daytime peak temperature of -24°F that occurred during the 1978 blizzard, 105 mph southwest wind using low ambient temperatures, J05 mph wind present, 120°F steel secondary containment wall with gravity included. [Reference Exhibit 65]

2D Nastran Plnne-Strain Time Slice Pe.ak Stress 3D NastraI11IIII FEM _ eak Radial Stress Thick Flute Portion

.I\\rchitectural Notch ID Case Description 4

Winter 197813lizzard Record Low 5:00AM

79 pSi I

+190 psi 5

Winter Solstice Ave No Wind 7:30AM

- 20psi I

+53 psi Table 4.2 Winter Solstice & 1978 Blizzard-Summary esults for Radial Stress @ EL 683 ' 6" Figure 4.2. 1 through Figure 4.2.2 show summary results listed in Tallie 4.2 for radial stress due to thermal transients.

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Page 22 of 39 2012

Exhibit 56 Redacted R.fEfiW

..A.,;.,,) Therm al Stress Analysis:

Final Report Results & Comments Radial Stress In Thick Portion of Aute Or = 70 psi Overa ll View @ EL 683' 6" Figure 4.2.1 Close-Up View @ EL 683' 6" 1978 Blizzard Condition 5:00 am, Constant Concrete CIE = 5,20 X10-6 in/ill/oF Radial Stress Distribution

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2012 Page 23 of 39

Radial Stress In ThICk Ponlon of Flute Op = -20 PSi Exhibit 56 Redacted

..me>>

Thermal Stress Analysis: _

Final Report Results & Comments JiR" Overall View @ EL 683' 6/1 Close-Up View @ EL 683' 6/1 Figure 4.2.2 Winter Solstice Ave Temperatures 7:30 am, Constant Concrete eTE = 5.20 xl 0-0 inlin/oF Radial Stress Distribution

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2012 Page 24 of 39

Exhibit 56 Redacted Final Report Results & Comments

.';1+11.

Thermal Stress Analysis:

4.3 Autu mn Eq uinox Conditions Table 4.3 summarizes results from the Autumn Equinox conditions.

The Autumn Equinox conditions listed below correlate with the high September conditions during the 24 hour2.777778e-4 days 0.00667 hours 3.968254e-5 weeks 9.132e-6 months period when temperatures are at their lowest and at 3:30 pm when the temperatures on the Southwest facing panels are highest. The average and high September temperatures are computed with and without 34 mph wind condition present.

2D Nas tIan Plane-Strain 3D Nastrar FEM Peak Radial Stress ID Case Description Time Sli ce Peak Stress Thick Flute Porti on Architectural Notch

(;

Autumn Equinox Hot No Wind 6:00 PM

+49 psi

100 psi

~ A_~tumn Eguinox Hot 34mph Wind 5:00AM

30 psi

+ 79 psi 8

Autumn Equinox Ave No Wind 5:00 AM

- 20pSi

+ S8JlSi Table 4.3 Autumn Equinox - Summary Results fo r Radial Stress @ EL 683' 6" Figure 4.3.1 through Figure 4.3.3 show summary results listed in Tabl~ 4.3 for radial stress due to thermal transients.

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Page 25 of 39 2012

Exhibit 56 Redacted R{OPE*

Thermal Stress Analysis:

Final Report Results & Comments ik,

, *_Ilt

,--"m, Close-Up View @ EL 683 1 6" Autumn Equinox Hot No Wind 6: 00 pm, Constant Concrete CTE = 5.20 xl 0-6 in/inr F Radial Stress Distribu tion I ~..u"aa Radial Stress III Thick Porllon of FILJte 6

11 Overall View @ EL 683 1

Figure 4.3.1 information. Performance Improvement international, LLC.

2012 Page 26 o f 39

Exhibit 56 Redacted

-FETr-Thermal Stress Analysis:

Final Report Results & Comments A )

t........ l"11l~

"' I&.H1 Radial ~tressln Thick Portion of Flute o~ =-30 psi Overall View @ EL 683' 6/}

Close-Up View @ EL 683' 6/}

Figure 4_3.2 Autumn Equinox Hot 34 mph Wind 6:00 pm, Constalll Concrete CTE = 5.20 x 10-6 in/inr F Radial Stress Distribution

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2012 Page 27 of 39

Exhibit 56 Redacted WlOEJO-Thermal Stress Analysis:

Final Report Results & Comments

~

Radial Stres~ In Thick.

Portion of Aute

a. = *20 psi Overall View @ EL 683' 6/1 Close-Up View @ EL 683' 6/1 Figure 4.3.3 Autumn Equinox Ave No Wind 5:00 am, Constullf Concrete CTE = 5.20 xlO-6 in/inf' F Rad ial Stress Distribution

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2012 Page 28 of 39

Exhibit 56 Redacted Fi nal Report Results & Comments

.iG'g.

Thermal Stress Analysis:

4.4 Summary Results & Comment from 3D Nash"an Idealization The 3D models show that the region of highest maximum principal stress is at the outer most layers of concrete OF and inboard to the 1st rebar layer. From the OF layer ofrebar inboard, maximum principal stress levels drop off dramatically due to the high stress gradients.

These results should indicate lite regiol1s ofconcern at tile outer 2-3" ofconcrete 4.4.1 Summe!' Solstice Cases The S/W facing panels and architectural flutes indicate the highest magnitudes of maximum principal and radial stress.

> It is not believed the magnitude of radial stresses is sufficient to either initiate delamination cracks or propagate any cracks that may be present.

4.....2 Winter Solstice Cases

.. For the normally occurring winter cold temperatures radial stresses in the thick pOition of the architectural flutes are low or compressive.

,. For the low temperatures during the 1978 Blizzard event the magnitude ofradial stresses in the "notch" cut-out of the architectural flutes is approximately 190 psi.

> It is not believed the magnitude ofradial stresses is sufficient to either initiate delamination cracks or propagate any cracks that may be present.

4.4.3 Autumn Equinox Cases

)-

The S/W facing panels and architectural flutes indicate the highest radial stress.

It is not believed the magnitude of radial stresses is sufficient to either initiate delamination cracks or propagate any cracks that may be present.

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Page 29 of 39 2012

Exhibit 56 Redacted Final Report Results & Comments Therm al Stress Ana lysi s:

WI'EIiliW 4.5 Summary Results: 3D Nastran Idealization with Simulated 30'x30' "Crack" To investigate potential for extended crack growth in a pre-existing crack region, the 3D Nastran idealization was modified to simulate a 30' x 30' "Crack" The 30 ft x 30ft "failed" region It is desired to evaluate S/W facing flutes with and w ithout the simu lated "Crack" As shown in Table 4.5 the magnitude of maximum principal stresses increased a slight amoLlnt fro m C>MP= 162 psi (No crack) to 0MP= 184 psi (w/crack). there is only a marginal increase in the magnitude of rad ial stress, from 0R= 76 psi (No crack) to 0R= 92 psi (w/crack).

It is not believed that the increase magnitudes in either the radial or maximum pl'incipal stresses are suffic ient to propagate cracks that may have formed.

2D Nastran Plane-Strain Time Slice Peak Stress 3D Nastrar FEM Peak Stress Values at ' Crack" Radial Stress Max. Prine. Stress ID Case Description 9

Summer Solstice Hot No Wind; 7:30 PM

+ 76 ps i I

+162 ps i 10 Summer Solstice Hot No Wind; Crack 7:30 PM

+92 psi J

+ 184 psi Table 4.5 Summer Solstice with Simulated 30'x30' "Crack" Summary Results for Radial Stress @ EL 785' 10" Figure 4.5.1 through Figure 4.5.2 show views of the 3D Nastran.. FEM with the simulated "Crack" region.

Figure 4.5.3 shows summary stress results listed in Table 4.5 for maximum principal stress due to summer solstice thermal transients.

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Page 3 0 of 39 2012

Exhibit 56 Redacted A)

Thermal Stress Ana Report Results & Comments Figure 4.5.1 "Thin-Crack" region introduced as idealized the "Cracked" boundary at the OF Rebar

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2012 Page 31 of 39

Exhibit 56 Redacted

.*, ~.:.

Final Report Results & Comments A,

Thermal Stress Analysis:

Simulated -Crack" Region Bounds g R= OF Rebar [Rl\\~= 859.61S"Refj Azimuth & Elevation Top Ed~e = 19~5° to ~l--1. ~o at EL 800' 10" Bottom = 190.5"to 114.S0at EL 770' 10" Figure 4.5.2 "Thin-Crack" region introduced as idea lized the "Cracked" boundary at the Of Rebar

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2012 Page 32 of 39

Exhibit 56 Redacted Thermal Stress Analysis: _

Final Report Results & Comments aun..

Max. Ptll'lcipal stress with Na crack" awp== 162 pst Max. Pnt1clpal Stress at SHTllll.ted Crack o,..p:; 184 pSI Figu re 4.5.3 Summer Solstice Hot No Wind 7:30 pm, Constallt Concrete CTE = 5.20 x1 0-6 inlinf F Maximum Principal Stress Distribution

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2012 Page 33 of 39

Exhibit 56 Redacted Final Report Results & Comments Thermal Stress Analysis 4.6 Miscellaneous Plane-Strain Results - Overlapping Reinforcement Both the plane-strain and 3D models show regions of positive radial stress in the thick pOliion of the flutes. The average magnitudes of stress are about +350 psi. Sec f igure 4.6.1 The magnitude of these positive radial stresses are not believed high enough to cause cracks but the thick pOltion of the flutes is the only large region were radial stresses are positive.

It is known that there are regions where reinforcement overlaps in regions where tbe rebar either transitions to a different size or rebar ofthe same size is continued and the ovedap acts as a splice. The plane-strain models with overlapping rebar indicate that the effects of the localized stress concentration factor (SCF) around can be linked together to form a line of cracks. The overlapping rebar also makes it difficult to fill voids due to large aggrl;gatc blocking distribution of concrete paste.

Figure 4.6.1 shows results ofa parametric analysis to qualitatively view oiOtl1e effects when rebar is closely spaced or overlaps.

Localized cracks that may develop at the overlapping rebar (vertical & hoop) could link to the adjacent SCF point since the distance to next pair of overlapping rebar isn't ve ry far. The or layer is more susceptible to this crack propagation because it is the OF layer where maximum principal stress are highest. Overlapping rebar along the IF face doesn't have maximum principal stress available to propagate cracks.

In addition, it should be noted that with the exception of the localized SCF peak tension radial stresses, regions immediately adjacent to the high tension stresses show significantly lower stress... even negative (compression) values. It is believed any localized cracks around the rebar would not propagate due to these thermal stresses and the surrounding compressive stresses would arrest any localized cracks initiating due to the localized SCF points.

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Page 34 of 39 2012

4,

Exhibit 56 Redacted am@_

SCf effectS In Vertical Rebar Concre~e,nierface o~::: 1.290 psi 1Il:rn~nt 36Q70* PtANf STRAJN FCl mut.h' lIon.

I Propcrt.. 1. Con. f ~lc I.*, tcnoll. Ccnc.C'tc IPlbtt Tep A ~: o.m.1 S:I ~" ; 531.W1 l(An~c,"",~ lC' ("".dln4\\1!! Sy-tC l"tl 1 tJoddJ I3& =.96.21

, lc;d. U899

763.2745 Figu re 4.6.1 Typical Radial Stress Contour i1-om 2D Plane-Strain Nastran With SCF Effects due to Overlap of Yertical Rebar

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2012 Pa g e 35 of 39

Exhibit 56 Redacted Thermal Stress Analysis:

Final Report Results & Comments 4.7 Effects of Variable CTE=j{T}

Two (2) ofthe concrete core samples from Davis-Besse Shield Building were sent to the United States Bureau ofReclamation (USBR) for mechanical and thermal properties testing. The coefficient ofthermal expansion (CTE) was tested in accordance with the USBR test procedure 4910-92. The average value for CTE over the temperature range ofapproximately 33°p to 150 0p given is CTE = 5.20 x 10-6 in/inJ°P... constant value. [Reference Pigure 2.1.4]

The temperature range for Winter Solstice Average is indicated on Figure 4.7.1 (+27°P to

+63°P) and shows that CTE= ffT} remains within the linear range ofUSBR data. Therefore during average winter conditions a variable CTE = f{T} will produce the same results as constant CTE.

Figure 4.7.1 Qualitative Characteristics for nonlinear CTE= f{T}

Average Winter Temperature Range Shown Figure 4.7.12 shows the assumed CTE= f{T} with the computed temperature range from 1978 Blizzard cold temperatures. As shown on Figure 4.7.1 the temperature range from the 1978 Blizzard extends into the nonlinear range ofthe CTE = f{T} data suggested by Prof Xi.

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Thermal Stress Analysis:***** Final Report Results &Comments Exhibit 56 Redacted Figure 4.7.2 Qualitative Characteristics for nonlinear CTE= f{T}

Average Winter Temperature Range Shown For reference, recall Figure 4.2.1, "1978 Blizzard Condition 5:00 am, Constant Concrete CTE 5.20 xlO-Q iniinfF - Radial Stress Distribution", peak radial stress in the thick portion ofthe flutes are computed at GR= -70 psi.

Figure 4.7.3 shows radial stress contour from the 3D Nastran _

FEM for the 1978 Blizzard condition assuming a variable/temperature dependent concrete CTE=f{T} similar to Figure 4.7.2. Results for radial stress in the thick portion ofthe flutes are GR= +470 psi for the 1978 Blizzard condition compared to GR= -70 psi when CTE is constant.

The temperature range shown during the 1978 Blizzard cold conditions (-27°F to +32.4OP) does fall into the non-linear region when a variable CTE=f{T} is considered. Therefore, ifone views the 1978 Blizzard event as a catastrophic "once-in-a-lifetime" event then the concrete may have cracked way back then and the likelihood ofanother 1978 Blizzard is remote.

The simulated "crack" model described in paragraph 5.0 was addressed using the variable CTE=j{T}. Analysis results did not show and significant change in the state ofstress surrounding the simulated "crack" region when variable CTE=f{T} is used in place ofthe constant CTE.

It should be noted the variable CTE=1'{T} is based on "academic" predictions scaled to match the USBR test results and then extrapolated beyond the known test range. These results remain qualitative until conclusive data; precise material properties that is, are available. These qualitative results do suggest strong evidence to support the hypothesis that the 1978 Blizzard event could be one ofthe primary contributors to the cracks. Exhibit 61: explores variations on the variable CTE concept.

- - - - information. Performance Improvement international, LLC.

Page 37 of 39 2012

Exhibit 56 Redacted WffiE.~-

Thermal Stress Analysi Final Report Results & Comments A'

Radial Stress if! ThlcJc

~ortlOn of Flute oa =t 470 pS4 Overall View @ EL 683' 6" Close-Up View @ EL 683' 6" Figu re 4.7.3 Radia l Stress Contour from 3D Nastran 1_

FEM 1978 B lizzard Condition, Variable Concrete CTE = f{T} @ 93% Saturation

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2012 Page 38 of 39

Thermal Stress Analysi Exhibit 56 Redacted Final Report Results & Comments

./ The 3D Nastran models indicate the regions of interest for highest radial stresses are in the thick portions ofthe Flutes. The magnitudes of radial stresses from any ofthe thermal transient stress analysis are not sufficient to initiate or propagate cracks that may have formed \\vithout another mechanism for crack initiation & crack growth present. [Reference pages 19,20, 21,35 & 38]

./ The plane-strain and other sub-models show localized peaks in the radial stresses resulting from stress concentration factors (SCF) around discontinuities. These SCF effects can result

B:om (a.) overlap ofadjacent reinforcing bar, (b.) abrupt change in stiffness l2teel-Concrete]

and (c.) thermal gradients with abrupt change in coefficient ofthermal expansion (\\":TE).

[Reference page 35]

./ It is not unusual to have peak stresses at points where SCF's are known to exist and some localized dis-bonding of the concrete to the rebar may result. With the exception of these localized peak tension radial stresses, regions immediately adjacent to the high tension stresses show significantly lower stress... even negative (compression). Localized cracks that may develop around the rebar due to these SCF would not propagate due to thermal stresses alone and the surrounding compressive stresses would arrest any localized cracks initiating due to the localized SCF points.

./ The 3D Nastran _

models indicate stress gradients exist due to thermal transient conditions. The maximum principal stresses are largest at the outer most 2" - 3" of concrete at the outer rebar layer. Thermal stress gradients lead to significantly lower stress as one move inboard the radial direction fi"om the OF toward the IF ofthe Shield Building wall. At approximately 6" - 8" inboard of the outer most layers, radial stresses drop off to levels that would clearly not initiate cracks. [Reference Figures 4.1.1,4.1.2 & 4.1.3]

./ Some qualitative results suggest strong evidence to support the hypothesis that the 1978 B lizzard event could be one ofthe primary contributors to the cracks. These qualitative resu Its indicate the low temperatures during the 1978 Blizzard may be a catastrophic "once in-a-lifetime" that may have cracked concrete. At date of release of this report these results remain qualitative and academic until conclusive data in the form of precise material properties are available, i.e. CTE=j{T}, allowing for a quantitative re-assessment of the 1978 8 1izzard condition. [Reference Figure 4.7.3 ]

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Page 39 of 39 2012

Exhibit 57: Temperature Dependent eTE

Exhibit 57 Temperature dependent coefficient of thermal expansion (CTE)

Under low temperatures, concrete may expand (instead of contract) during a cooling period.

This possible expansion is due to ice formation in the concrete.

During a severe cooling process, the temperature of concrete in outer layer of the cylindrical wall is lower than that of inner layer. So, ice may form in the outer layer of the wall resulting in an expansion and ice may not form in the inner layer of the wall leading to continuous contraction. This special outer-expansion-and-inner-contraction deformation pattern can result in a tensile stress in the radial direction of the wall. Delamination cracking may occur in the case of excessively high radial tensile strength. The key issue here is the coefficient of thermal expansion (GTE) of concrete under low temperatures. Did the Davis Besse concrete in outer layer of the wall expand during the blizzard? If yes, how much did it expand? Was the expansion high enough to cause the cracking? These questions will be discussed in the following sections.

The effect oftemperature on CTE of concrete GTE of concrete depends on temperature. The reason is that the state of moisture inside of concrete depends on temperature. Under elevated temperatures (e.g. fire), liquid water turns into vapor which generate high vapor pressure. This case is not within the scope of this project, and thus will not be discussed further. Under low temperatures, liquid water or vapor turns into ice which is associated with a 9% volume expansion. After all moisture freezes, the effective GTE of concrete is a mixture of GTE of concrete and GTE of ice. The GTE of ice is about 5 times of the GTE of concrete without ice. Therefore, at a very low temperature, the GTE of concrete is not the same as the GTE of concrete at room temperature. Accurately speaking, the effective GTE of concrete depends both on temperature and on ice content, and thus on moisture content. The effect of temperature will be discussed first, and the effect of moisture will be discussed later.

40

.c:

'E g BO g

~ 120

<ii t;;

160 TIME {h)

Fig. 1 A typical strain and temperature chart (ASTM G671)

Literature review showed that the ice formation starts at QOG and completes at about -15°G or lower depending on the microstructure of concrete. This freezing process is due to the fact that the freezing point of water depends on pore size in concrete. The freezing point of water in large air voids is close to QOG, and thus the water in large air voids freezes first; and the freezing point in small pores could be well

Exhibit 57 below O°C, and the water in small pores freezes after the water in large voids. For a continuous cooling process, the strain of a concrete specimen is shown in Fig. 1, contracting first, expanding due to ice formation, and then contracting again after the completion of ice formation. In Fig. 1, the slope of the strain vs. tem perature curve is CTE. It is very clear that the slope of the curve, the CTE is not a constant.

Depending on concrete mix design and cooling condition (temperature range and freeze-thaw cycles), the curve could be significantly different, as shown in Fig. 2. The shape of the curve, Le. the temperature dependency of CTE is closely related to the internal structure of concrete.

4 LENGTH CHANGE MEASUREMENTS AFTER MULTIPLE FREEZE-THAW CYCLES (MORTAR BEAMS, w/c; 0.32. AGE TWO WEEKS)

~~O~~--1~O~O------O~O~~---_1~OO~~--_~200 TEMPERATURE (0C)

Fig. 2 Test data of concrete length changes under low temper~tureS(Bazant et aL 1988; see Refs. 18 and 19 for test results from T.C. Powers and R.A. Helmuth)

Experimental studies showed that properly air-entrained mortarf; contract upon freezing, while non-air entrained mortars or improperly air-entrained mortars expand. The expansion of the latter is attributed primarily to hydraulic pressure, owing to the rapid growth of ice,'which nucleates at low temperatures in laboratory samples (Sun and Scherer 2010).

Moisture in concrete Because of very small pores in concrete, water in the pores exists as a mixture of liquid and vapor in above ground concrete structures. Internal relative humidity (or pore relative humidity), RH of concrete are often used to represent internal moisture state of concrete. The internal RH can be correlated to the moisture content in concrete (the weight of moisture in concrete) by adsorption isotherms. Adsorption isotherm is a relationship among weight of internal mOisture, temperature, and RH.

Fig. 3 shows experimental results of adsorption isotherms in the literature in comparison with the predictions of a theoretical model developed by Xi et al. (1994). In the figure, the horizontal axis represents RH (where p is water vapor pressure and Ps is the saturation pressure at a given temperature); and the vertical axis stands for moisture content in gram of moisture in gram of concrete. So, with a given concrete mix Page 2

Exhibit 57 design, temperature, and RH, the moisture content in the concrete can be obtained from the adsorption isotherms.

It is important to note that RH = 100% does not mean all pores are filled up by water, it means the vapor pressure in the pore reaches the saturation pressure of vapor at the given temperature. At this stage, the concrete reaches its adsorption capacity, which is different from absorption capacity. There may be only limited layers of water molecular covering the surface of pore walls at RH ;:: 100%. When all pores are filled up by water, the concrete reaches its absorption capacity_

0.""....----------,

0..10 C).o.:a o

.5 s: 0.10

"-"" 0.0

'1110" 0582 Typ/!

c:ernet\\1 T.. 2$4'K t-26d1ya o:z I'OWIII'I lit aI. (1~1) 0.*

1.0 0.20

'1110- 0.45 T - 2911" K Typ.l~lint t-2555 day.

C.!

w/c";'O.ll T-~'K Type I CIlment 1-7day3 MUc,,*1 et aI. (tIil15) 0.".,-------------,

0..,...~---- -------,

Fig. 3 Adsorption test data and comparisons with predictions pf a theoretical model (Xi et al. 1994)

For above ground structures such as Davis Besse containment structure, it is more suitable to use RH and adsorption isotherms for estimating the internal moisture content. For under water structures, the absorption capacity is a better indicator.

The effect of moisture on CTE The general trend is that the higher the initial moisture content before ice formation, the larger amount of ice formed in the concrete, and possibly the larger dilation of concrete during the ice formation process.

The following figures show available test data in the literature.

Exhibit 57 4 ;

.r,,,....0'4';',,.

l (lflf!. 'Of 4}4Tv~.TI"1if o

Z

tIC.. 0' SO

-0\\

PlAIN Mj'X($

~O 15-

- 20 Fig. 4 Temperature vs. dilation of concretes at various levels of saturations (Grieve et al. 1987) a~

.~

.~...

Q

-100 300 200 100 0

~

I Ie.perature ("I::)

100 95

~;

~

IS "Iii j

Q

-100 l

0::.::.

1: 2 :0.65

-200_ 20 10 15 20 Te.,>er.tur. ("c) 100 95

r

~

gO IS 85 j Q

~

80 To.perature ("c)

Fig. 5 Dilatation of concrete and internal relative humidity (Zhou and Mihashi 2008)

Exhibit 57 In Fig. 4, there is almost no expansion when the saturation level in the concretes is below 90%.

However, in Fig. 5, there are significant expansions at low temperatures even the initial RHs in the concrete samples are below 90% (see the first and the second figures, the RHs are about 88%). Fig. 6 shows the dilations of concrete under dry and wet conditions.

One can see from Fig. 6(a) that the concrete at wet-freezing condition expands significantly (near 8 degree C), while the concrete at dry freezing condition does not expand at all. Therefore, the moisture content has a major effect on the expansion of concrete during ice formation process, and the extent of moisture effect depends on internal structure of the concrete. The internal structure of concrete depends on mix design, curing conditions, and age of the concrete. In Fig. 6(b), even if at wet-freezing condition, the concrete with proper air entrainment only expands slightly near 8 degree C).

o

~o

.(j

... *2(Wl.g

~on -

. *600 r-"'-~-'------------"----, Iro,

..'i:

~

C

700

-- w-ct - Ib:'l,.'zing (a) wlc = 0.5 non-air-entrained concrete (b) wlc =0.35 air-entrained concrete.

Fig. 6 Influence of saturation on dilation of concrete (Zuber and Marchand 2004)

In summary, at a sufficiently high RH level and under a continuous cooling process, there are three possible types of temperature-dependent thermal strains for concrete:

Type 1 - Contraction, significant expansion, and contraction, as shown in Fig. 1 and Fig. 6(a).

Type 2 - Contraction, slight expansion, and contraction, as shown by solid squares in Fig. 2 and Fig. 6(b).

Type 3 - Contraction, as shown by hollow circles in Fig. 2.

From Exhibit 52 Univ. of Colorado lab test report, the relation of thermal strain and temperature of Davis Besse concrete follows Type 3 when the sample is dry, and Type 1 when the sample is wet. So, Davis Besse concrete does expand under low temperatures.

/'

,i

-:tIJ\\)

.:§

Exhibit 57 The resulting temperature dependent CTE at the temperature ranges are:

T> 23°F (-5°C), CTE = 5.2 x 1Q-6rF (The same as USBR test data),

8.6°F (-13°C) < T < 23°F (_5°C), CTE = -4.94 x 10-6rF, 1.4°F (-1rC) < T < 8.6°F(-13°C), CTE =-43.1 x 1Q-6rF, Below 1.4°F (-1rC), CTE = 5.2 x 10-BrF Crhe same as t~e CTE under room temperature).

Conclusions CTE of concrete depends on temperature, internal moisture, arid internal structure of concrete. Some concretes may expand during the ice formation process if their internal moisture content is sufficiently high.

Davis Besse concrete showed expansive strains under low temperatures (Exhibit 52). Therefore, a tem re CTE was developed

Exhibit 57 References Bazant, Z.P., Chern, J.C., Rosenberg, A.M., and Gaidis, J.M. (1988) "Mathematical Model for Freeze Thaw Durability of Concrete", Journal ofAmerican Ceramic Society, 71(9), 776-783.

Grieve, R., Slater, W.M., and Rothenburg, L. (1987) "Deterioration and Repair of Above Ground Concrete Water Tanks in Ontario, Canada", Research Report to Ontario Ministry of the Environment, Golder Associates and W.M. Slater & Associates, Inc.

Sun, Z., and Scherer, GW. (2010) "Effect of Air Voids on Salt Scaling and Internal Freezing", Cement and Concrete Research, 40, 260-270.

Xi, Y., Bazant, Z.P., and Jennings, H.M. (1994) "Moisture Diffusion in Cementitious Materials: Adsorption Isotherm", Journal ofAdvanced Cement-Based Materials, 1,248-257.

Zhou, Z.Y., and Mihashi, H. (2008) "Micromechanics Model to Describe Strain Behavior of Concrete in Freezing Process", Journal of Materials in Civil Engineering, ASCE, 20(1), 46-53.

Zuber, B., and Marchand, J. (2004) "Predicting the Volume Instability of Hydrated Cement Systems upon Freezing using Poromechanics and Local Phase Equilibria", Materials and Stmctures, 37, 257-270.

Exhibit 58: Carbonation Lab Testing

Exhibit 58 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 Carbonation Based on my observation and examination of concrete-core samples received from the Davis Besse Shield Building, my findings for Carbonation are detailed in what follows.

Page 1 of 9

Exhibit 58 Carbonation in Concrete Carbonation in Concrete Laboratory tests and examinations were conducted on several concrete core samples to determine the extent of carbonation within the samples The cracked concrete samples, which are vulnerable to carbonation, were isolated and fractured in a plane perpendicular to the original cracked surface.

Figures A1 and A2 show examples of the carbonation depth as measured from the exterior surface. The exterior surface is the portion of the shield building that is exposed to the elements; it is the outer diameter surface.

Page 2 Page 2 of 9

Exhibit 58 figure A 1: Fracture Sample with Carbonation Laver (Core F2-790.0-4.S)

Figure A 2: Fracture Sample with Carbonation layer (Core F3-1)

Page 3 Page 3 of 9

Exhibit 58 The following table shows the nominal carbonation depth as measured from the exterior surface. The table lists the 16 samples used in determining the average nominal carbonation depth which, as previously stated, is 8.57 mm.

Table A 1; Nominal Carbonation-layer Depth from Exterior Surface (Carbonation Rate Determination)

Core Sample F3-1 5 11-1 511-2 S1 2-1 5 12-2 516-3 S5-1 55-2 57-1 I

Nominal Carbonation Depth From Exterior Surface (reference), mm

'11.7 9.33 10.00 8.59

8. 33

7. 73 7.90 7.87 7.75 57-2 57-3 S9-1 59-2 S-7-656.5-6.5 5-9-653-1 5 785-22.5 AVERAGE 9.07 7.56 10.05 7.65 8.84 8.65 6.06 8.57 Longitudinal Fracture Carbonation analysis was conducted for both longitudinal and transverse cracks.

Figure A3 shows a longitudinal crack for reference.. As can be seen, the longitudinal cross-section is defined as the plane that is parallel to the longer dimension of the core sample.

Several longitudinal cracks with no evidence of carbonation were evidenced. For example, Figure A4, Core F2-790.0-4.5, shows that a distance of 7 inches from the surface, no carbonation is detected.

Page 4 Page 4 of 9

Exhibit 58 Davis Besse Nuclear Plant F2*790.0-4.5

.. I II III

\\I t l

~

Iimiml Figure A 3: longitudinal Crack (Reference)

Figure A 4: Longitudinal Crack with no Evidence of Carbonation Page 5 Page 5 of 9

Exhibit 58 The Following table shows several samples with longitudinal cracks at various distances from the exterior surface. For the samples in the table listed below, ihere is no carbonation layers formed at any of the various distances within each respective sample.

Table A 2.: Carbonation Results from several Samples with longitudinal Cracks (No Carbonation)

Core Sample Crack Distance From Ex1erior Su rface Maximum Carbonation Depth, mm F2-790.0-4.S 17" Long Crack Longitudinal Cracl<

1 T'

0 F2-790.0-4.S 17" Long Crack Longitudinal Crack

1 1005" 0

F2-790.0-4.S 17" Long Crack Longitudinal Crack

1 13" 0

F2-790.0-4.S 17" Long Crack Longitudinal Crack

1 16" 0

F2-790.0-4.S 17" Long Crack Longitudinal Crack

1 20" 0

F2-790.0-4.S 5.5" Long Crack 45 Degree Crack

2 21"t022" 0

F2-790.0-4.S 5.5" Long Crack 45 Degree Crack

2 23" to 26" 0

F4-794.0-3.S 19" Long Crack Longitudinal Crack

1 9"

0 F4-794.0-3.S 19" Long Crack Longitudinal Crack

1 13" 0

F4-794.0-3. S 19" Long Crack Longitudinal Crack

1 17.5" 0

F4-794.0-3.S 19" Long Crack Longitudinal Crack I

1 20" 0

F4-794.0-3.S 19" Long Crack Longitudinal Crack

1 23" 0

F4-794. 0-3.S 19" Long Crack Longitudinal Crack

1 26':

0 FS-791.0-4 9" Long Crack Longitudinal Crack

1 7.5" 0

FS-791.0-4 9" Long Crack Longitudinal Crack

1 9'"

0 FS-791.0-4 9" Long Crack Longitudinal Crack

1 11 "

0 FS-791.0-4 9" Long Crack Longitudinal Crack

1 14.5" 0

8 2-798.S-4.S 5" Long Crack Longitudinal Crack

1 1"

0 52-798.5-4.S 5" Long Crack Longitudinal Crack

1 2.5" 0

5 2-798.5-4.5 5" Long Crack Longitudinal Crack

1 4"

0 Page 6 Page 6 o f 9

Exhibit 58 The following table shows the results from carbonation analysis on several core samples with longitudinal cracks.

Table A 3: Carbonation Analysis on Longitudinal Cracks Core 5ample Crack Longitudinal Crack Length Maximum Carbonation Depth, mm F3-1

1 7"

5.4 511-1 N/A N/A N/A 5 11-2

1 6 ~ "

0 5 11-2

2 6 ~"

2.09 512-1

1 5 ~"

0 5 12-1

2 1 Yz" 2.70 512-2

1° 1 %

2.99 5 16-3 N/A N/A N/A 55-1 N/A N/A N/A 55-2 N/A N/A N/A S7-1

1 2 ~/8 "

0 5 7-2 N/A N/A N/A 57-3 N/A N/A N/A 5 9-1 N/A N/A N/A 59-2 N/A N/A N/A

' N/A = No Longitudinal Crack

'Crack Fou nd Upon Sectioning Transverse Fracture Figure A6 shows an example of a transverse fracture sample, Core S5-2, which was determined to have a measured carbonation layer of 0.455 mm.

For comparison with Figure A6, Figure A7 shows an example of a transverse fracture sample, Core S7 -656.5-6.5, which was determined to have no carbonation layer.

Page 7 Page 7 of 9

Exhibit 58 Figure A 5: Carbonation Detected on a Transverse Fracture Sample (Core 55-2), 1.8 yrs.

8 7

Figure A 6: Transverse Fracture Sample with No Carbonation (Core 57-656.5*6.5)

Page 8 Page 8 of 9

Exhibit 58 The following table is a list of transverse fracture core samples along with their associated carbonation depth. Core samples with an asterisk in their identification number are those with transverse cracks as identified by the plant IR Inspection.

Table A 4: Transverse Fracture Carbonation Analysis Core Sample Fracture (Crack)

Distance From Surface, inches Maximum Carbonation Depth, mm F3-1

T1 7

0.582 S11-1*

T1 8

0 S11-1*

T2 5 1'2 0

5 11-2

T1 8

0 S12-1

T1 21 0

S1 2-2*

T1 51'2 0.264 S1 2-2*

T2 16 1'2 0.686 S1 6-3*

T1 14 1'2 0.343 S16-3*

T2 15 1'2 0

S16-3*

T3 21 1'2 0

S5-1 *

T1 9

0.500 S5-1*

T2 91'2 0.897 S5-1*

T3 14 1'2 0.604 S5-1*

T4 16 ~

0.893 S5-2

T1 12 "/8 0.445 S7-1*

T1 4 ~

0 S7-1*

T2 10 0

S7-1 *

T3 16 %

0 S7-2

T1 15 %

1.42 57*3*

T1 6 1'2 0.710 S9*1 *

T1 4 1'2 0.329 0

S9*1*

T2 12 ~

S9-2

T1 10 ~

0.388 "Core samples with transverse cracks as identified by the plant IR inspection Page 9 Page 9 of 9

Exhibit 59: Test Report from the United States Bureau of Reclamation (USBR)

Appendix VIII-60

United States Department of the Interior BUREAU OF RECLAMATION P.O. Box 25007 Denver, Colorado 80225-0007 IN REPLY REFER TO:

JAN 192012 86-68180 RES-3.40 MEMORANDUM To:

Performance Improvement International 2111 S EI Camino Real Suite 200 Oceanside, CA 92054 Attention: Dr. Chong Chiu Prom:

Katie Bartojay, P.E., Civil Engineer, Materials Engineering and Research Laboratory Group (MERL)

Subject:

Thermal Properties Testing Results 0 Specimens Materials Engineering and Laboratory Report No. MERL-2012-02 INTRODUCTION Six concrete core samples were delivered to the Bureau ofReclamation's Materials Engineering and Research Laboratory (MERL) on December 19, 2011. The specimens were identified as F4-791-2.5 #1 through #4 and S7-782.0-8.5 #5 and #6. AU six specimens were approximately 2.5-inches in diameter and 4-inches long.

The submitted samples were tested for thermal diffusivity, specific heat, and thermal coefficient oflinear expansion testing on concrete cores. Conductivity was calculated using the specific heat and diffusivity results.

CONCLUSIONS AND DISCUSSION Thermal Diffusivity Thermal diffusivity measures the rate at which temperature changes take place in concrete and is defined as an index ofthe facility with which a material will undergo temperature change [i]. Thermal diffusivity was tested in accordance with Reclamation's test procedure USBR 4909-92, "Thermal Diffusivity ofConcrete" (with modifications to account for upgraded equipment). Two concrete core specimens marked S7-782.0-8.5 #5 and S7-782.0-8.5 #6 were tested over three temperature ranges: 35°P to 75°P; 75°P to 115°P; and 115°P to 155°P. A small diameter hole was drilled from one end to accept a Exhibit 59 Page 1 of 5

Exhibit 59 thermocouple to be located at the approximate center of the specimen. The hole was filled with epoxy before testing.

Specific Heat Specific heat is the amount of heat required to raise the temperature of a unit mass of material one degree [i]. Specific heat was tested in accordance with Reclamation's t st procedure USBR 4907-92, "Specific Heat ofAggregates, Concrete, and Other Materials" (with modifications to account for upgraded equipment). Two c ncrete core specimens marked F4-791-2.5 #1 and F4-791-2.5 # 2 were tested over a te.mperature range of approximately 35°F to 150°F.

Thermal Conductivity Conductivity is the rate at which heat is transmitted through a unit thickness of material.

The coefficient of thermal conductivity (K) represents the uniform flow of heat though a thickness of material when subjected to a unit temperature difference between two faces

[i]. Thermal conductivity was calculated from the specific heat (c), diffusivity (0), and concrete density (P). The hardened density determined from this study was used in this calculation.

K =cpo

[ii]

Thermal diffusivity, specific heat and conductivity tests results are summarized in Table 1 and reported graphically in the Attachment.

Table 1 - Summary of tbermal properties of select cores

" Temperature....

. (oF)

. SQecific Heat (c}

Btu/(lbm*oF)

F4-791"2.5 #1&#2 Diffusivitv (d}

felhr

'. S7*782.0*8.5 #5&_~

Conduc.ivi!ll l!!ll Btu/(tf*hr*oF/ft) i--' Calculated 50 100 150 0.478 0.428 0.378 0.054 0.049 0.044 3.79 3.08 2.44 Typical ranges of these thermal properties for nonnal concrete[ii] are approximately:

0.02 to 0.06 ft2/hr for Diffusivity 0.20 to 0.28 Btullb per OF for Specific Heat 0.8 to 2.1 BtuJft2lhr °F/ft for Conductivity The specific heat values measured for the submitted specimens were not in the typical range for normal concrete. The calculated conductivity was also outside the range for normal concrete.

Coefficient of Linear Thermal Expansion Thermal coefficient of linear expansion is the change in a unit length per degree of temperature change of the concrete [iii]. Thelmal Coefficient of Expansion was tested in accordance with Reclamation's test procedure USBR 4910-92, "Coefficient a/Linear Thermal Expansion" (with modifications to account for upgraded equipment). Two Exhibit 59 Page 2 of 5

Exhibit 59 concrete core specimens marked F4-791-2.5 #3 and F4-791-2.5 #4 were tested over a temperature range of approximately 33°F to 150°F. Coefficient of linear thermal expansion tests results are summarized in Table 2 and reported in the Attachment.

Table 2 -Summary of coefficient of linear thermal expansion Specimen 10

.Average Coefficient Of*

.. LinearThermal.

Expansion.

(InchllnchrF)

F4-791-2.5 #3 5.2 x 10-6 F4-791-2.5 #4 5.1 x 10.6 Average 5.2 x 10.6 The coefficient of linear thermal expansion of concrete varies greatly with aggregate mineralogy and can be as low as 4-x 10.6 per degree F to as high as 13 x 10.6 per degree F[ii]. The values determined by this testing are in the range for normal concrete.

The test results derived from this work shall not be used to imply endorsement by the Bureau of Reclamation or the U.S. Government and cannot be used for advertising or commercial purposes.

Attachments cc: Dr. Yungpin Xi, University of Colorado, yungpin.xi@colorado.edu (electronic copy)

CJ Concrete, Mindess and Young, Prentice-Hall, lnc., 1981

[U] "Properties of Concrete, Fourth Edition" A.M. Neville, Pearson Education Limited, 2009.

(;;;] "Concrete Manual, Part I, Eighth Edition", A Water Resources Technical Publication, U.S. Department of Interior, Bureau of Reclamation, Denver, CO, 1988 Reprint.

Exhibit 59 Page 3 of 5

Exhibit 59 PII Core Specific Heat Test 0--E

..0

J t

Ci:J Q)

J:

~

'0

(\\J 0

r./)

0.055 OJ

~

0.050

~-

0.045

'Vi OJ 0.040

t:

is 0.035 0.65 0.60 0.55 0.50 0.45 0.40 0.35 0.30 0.25 0.20 I

Iy =-O.OOlx + 0.5277 1

I R2 =0.2978 1

I

~

I

./

~.

j I ~

u I

~ -

r.' *

I 40 60 80 100 120 140 160 Temgerature n Core 1 Core 2 Linear (Average) 0.060 PII Core Diffusivity --,-__-,____

+---+------11------\\-----+----1-----1 4-----I~~......ii+/-::=_---I-----+_---+---__l I-------------+-t-----+.-]f--.-I------l 0.030 40 60 80 100 120 140 160 Core 5 Core 6 Average Linear (Average)

PII Core Conductivit 4.50,-----,----,-

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

4.00 3.50 3.00 2.50 2.00 -f-----

/-----f-----r----+-------O-----I 40 60 80 100 120 140 160 Calculated Exhibit 59 Page 4 of 5

Exhibit 59 "JCIlJL....rwITtl 0.0007 I

0.0006 -\\,-

Specimen) =0.0000052 Specimen 4 =0.0000051 Dl ' *re "'I[fr:r:L.~.J 1l! If:nLlIlIH J,Jill' 0.0005

.JLdU.

-1 II, ~mLf1I:.: iJ I='

~

8 0.0004

.,.... 'j./

M

~

~ 0.0003 l h '/

-- Specimen 4

/~1"r Specimen 3 P

- Average

..J 0.0002 0.0001 -j

--l 0.0000 \\

" '/

o 20 40 60 80 100 120 140 160 180 cJLL rIT ILJI GJ ::T r HL T,- 1. '

Project: PI I Cores Test Date: 12/21/2011 Test Age: not known Bureau of Reclamation Materials Engineering and Research Laboratory Exhibit 59 Page 5 of 5

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 concrete pores does takes place at a lower temperature than 32°F due to surface tension which prevents the rearrangement to form ice. The end effect is that water in concrete freezes at varying temperatures depending on the pore size. This nonlinear dependence of the CTE with temperature is shown in Exhibit 57 and used as an input to the finite element analysis presented here.

Tests of moisture penetration were also performed at the University of Colorado at Boulder, which showed that a 1-0 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).

Page 4

Figure 1 below shows the location of ********************

I

~?~,I ~-- ------------ ------~~+~~=+/-~

~/

Vicinity of Flute 6 Studied in Detail Figure 1-Shield Building with Flute Numbers and Azimuth Locations _

Page 5

Exhibit 60: Test Report from the University of Colorado Appendix VIII-61

Exhibit 60 University of Colorado Dept. of Civil, Environmental. & Architectural Engineering Boulder College 01 Engineering and Apphed SCience t 303 492 8991 428UCB f 3034927317 Boulder, Colorado 80309-0428 yunping.xi@Colorado.edu Exhibit 60 Page 1 of 4

Exhibit61: Stress State during the 1977 and 1978 Blizzards Appendix VIII-62

Stress State during the 1978 and 1977 Blizzards Table of Contents Summary of Results.................................................................................................................................2 Modeling Summary.................................................................................................................................2 Overall Approach.................................................................................................................................2 Finite Element Software......................................................................................................................3 Temperature Conditions......................................................................................................................3

................................ :................................................................................... 3 Modeled Geometry.............................................................................................................................3 Material Properties..............................................................................................................................3 Coefficient of Thermal Expansion of High Moisture Concrete...............................................................4

........................................................................................................................ 5 Circumferential Temperature Distribution at O.F. Horizontal Rebar..................................................... 6

.................................................................................................7

...................................................................................................................................9 1978 Blizzard Condition..;.......................................................... ;....................................................... 10 1977 Blizzard Condition.....................................................................................................................14 Page 1

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

The blizzard of 1978 produced stresses above the tensile strength in the hoop direction, likely resulting in damage. The area exceeding the tensile strength is confined to a circumferential plane at the depth of the outer face main cylindrical wall under the raised shoulders.

The 1977 blizzard shows significantly lower stress compared to the blizzard of 1978. The hoop stress approached the tensile strength of the concrete and it is limited to a small area. For these reasons only minor damage, if any, is predicted.

Modeling Summary Overall Approach Page 2

Finite Element Software was used exclusively in the finite element analysis presented here.

Temperature Conditions The following two temperature conditions are presented in this report. The details of the temperature conditions and the selection of the time of day are summarized separately in the Root Cause Analysis Report:

I} Low temperature during the 1978 blizzard (105 mph wind, winter solstice, 5:00 AM)

2) Low temperature during the 1977 blizzard (76 mph wind, wintersolstice, 5:00 AM)

Expansion of concrete due to freezing of entrapped moisture was studied in the _

. This model is utilized to determine the stress state in a subsection of the structure spanning from the middle of one panel to the middle of the adjacent panel. The raised shoulders and the flute geometry are included in the model. Nominal steel reinforcement is included using a technique called The detailed stress concentration at the steel and concrete interface is not included in the model.

Modeled Geometry The drawings used as geometry input for this model are:

Drawing No: C-100 Rev. 5 "Shield Building Foundation Plan & Details SH. I" Drawing No: C-110 Rev. 6 "Shield 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 l.4!0") at 12" spacing. 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 2.1.4: Material Properties for Davis-Besse 3D 11I11IIII11III111 I Model Exhibit 56, Section 4.7: Effects of Variable CTE Exhibit 57: Temperature dependent coefficient of thermal expansion (CTE)

Page 3

ture Distribution at O.F. Horizontal Rebar (See Exhibit 65) The temperature profiles around the Shield Building at the outer face horizontal rebars are shown in Figure 2.

The figure shows 8 sets of double peaks for each temperature profile. The double peaks represent the warmer temperature under the shoulders. The temperature is warmer under the shoulders because there is a thicker layer of concrete at those locations which reduces the heat loss to the exterior during the blizzards.

Temperature (OF), Mid-Height, Outer Face Horizontal Rebar Depth 16 1917 Blizzard Temperature Calculation (Worst CClse) 1978 Blillard Temperature Calculali on (Worsl Case +20' F) 22.S 67 5 l ~ 15 2 17 5 Figure 2 - Circumferential Temperature Distribution at the O.F. Horizontal Rebar Depth Page 6

Figures 3 through 5 below depict the geometry and finite element mesh of the Figure 3 Geometry and Rebars Page 7

Figure 4 Detail of Flute Region Figure 5-

. Detail of Flute Region with Mesh Page 8

This section summarizes the results is used to make predictions about the delamination propensity due to the two blizzard conditions. This model does not attempt to make predictions of stress concentration effects around the included reinforcing bars due to lack of detail at the concrete/steel interface.

The tensile strength of the Davis-Besse concrete is in the range of 836 to 962 psi. The contours in the stress figures in this section are assigned an upper limit of 900 psi. A tensile stress exceeding 900 psi is indicated by light grey contours in the stress figures. The interpretation of any light grey area in the contour plots below is that damage may occur in that area. The damage that results from any tensile.

stress above the strength of the concrete depends on 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 more microcracking and eventually a structural crack.

The stress contour results shown in this section can be summarized as follows:

Higher tensile stress and larger stressed areas is predicted in the 1978 blizzard compared to the 1977 blizzard Blizzard of 1978:

o Tensile stresses high enough to damage the concrete is predicted o The high stresses are distributed over large areas in the observed crack locations under the thick sections of the shoulders and not in the thinner sections in the flute and panels Blizzard of 1977:

o Tensile stresses are lower or equal to the strength of the concrete o The highest tensile strength are confined to small areas under the thick sections of the shoulders Page 9

1978 Blizzard Condition The result ************** due to the 1978 bli zzard condition is shown in this section.

The temperature contours can be seen in Figure 6 and the stress results is shown in Figures 7 through

12.

rHll

+ S 210e+Ol

- +4.741e+01

. +4 276e+01

- +3,80ge +01

.. + 3.J42e.+Ol

- +2.87 5e +01

- +2.407e +01

. +L94 C~ +Ol

+ 1.4 /3~+Ol

+1. 006e + 0 1

- +S. 386e+OO

+ 7. t4 !le-Ol

-3.951e+OO ODS: m 122 J,od b AbaqusjSt;., nda,d 6.10-3 Tu~ Feb 21 18 :00 :59 P;,cific St:an r!"~ "d Ti me 20:12 Step: Step-l Incre ment 1: Step Time:::

1.000 Primary VlIr: NTH Deformed Va,: U Deformacon Scale Factor : +S.OOOe+02 Figure 6 - Temperature (OF) during the Blizzard of 1978; Deformation Scale Factor 500X Page 10

S, (*lu:*:. Prlnclpa t (Avg : 7S%)

+4.72.2e+03

+9.00Q'H02

+7. 5DOe+02

+6.0QOe+02

+4. SOOe+02

. +3.0006+02

+1.S0(}@t02

+0.000..+00

- - 1, SOOe -l 02

-3.000e +02

-4.S00e +Q2

-6.000'H 02

- 7. S00~ +0 2

- -9.0006+02

-1. 87Ge+03 008: m1 223.odb Abaqu:;:jStandard 6.10-)

Tue Feb 2118'00;S9 Pacific Standard 1lme 2012 y

Stl.\\p: Step-l Incre ment 1: Step Time =

1.000 L x Prlmary Var; S, '"'lax. Pr inct ~1 Defor med Var : U Deformation Scale FClctor: +S.ooOe+02.

Figure 7 - Max PrinCipal Stress (psi) during the 1978 Blizzard; Deformation Scale Factor 500X S, ~*lax. Pnncipal (Avg: 7S%)

+4.722e +03

- +1.200e+03

- +1.00De+03

- +8. 000e+02

- +6.0DOe+02

+4.000e+02

+2.000e+02

+O.OOOe+OO

-2.00Oc + 02

-4.000e+02

-G.OOOe+02

-8.000e+02

-i.000e+03

-1.200e+03

-1. B76e + 03 008: rn1223.odb AbaqusjStandard 6. 10-3 Tue Feb 2118;00;S9 PacifIC Standard lime 2012 y

Step: Step-l Increment 1: Step Tome =

1.000 L x Primary Va r: S, N ~x

. Prlnclp,,1 Deforrr.ed Var: U De forma~on Scale factor: +5.000e+02 Figure 8 - Max Principal Stress (psi) during the 1978 Blizzard; Deformation Scale Factor SOOX; Wider Contour Range (+/-1200 psi)

Page 11

5,511 (Cy1)

(Avg : 75%)

+3. 8S1e+03

+9.000e+02

+ 7.S00e+02

+6.00Qe+02

+4.S00e+O?.

+3.00De+02

+1. SOQe+02

+O.OOOtH *OO

- 1.500e +D2

-3.00OC + 02

-4.SCOc +02

-6.00~ + 0 2

-7. 500e+02

-9.000e +02:

-3.83Qe+03 008: m1223.odb Abaqus/Standard 6.10-3 Tue Feb2118 :00:59 Pacific Standard Time 2012 y

Step: Step-!

Increment 1: Step TIme =

1.000 L x Primary Var: 5, Sl1 (Cyl)

Deformed Var: U Oefor mi:ltion Sca le Factor: +S.OODe +02 Figure 9 - Radial Stress (psi) during the Blizzard of 1978; Deformation Scale Factor 50OX 5, 522 (Cy1)

(Avg: 750/0)

+1.932e +03

+9.000e+02

+7.S00e+02

+6.000e +02

+4.S0oe+02

+3.000e+OZ

+1.500e+ D2

+O.ooOe+OO

-1. SOOe+02

-3.000e+02

-4.500e+02

-6.000e+02

-7. S00e+02

-9.000e+02

. -3.47Ce+03

~

008: m1223. odb Abaqus/Standard 6. 10-)

Tue Feb 2113:00:59 Pacific Sttlndard Tl me 2012 Step: Step-l Increment l' Sta~ TIme =

1.000 L x Prim ~ry V,:,r: 5, 522 (Cyl)

Deformed Var: U Deformation Scale Factor: + S. 000e + 02 Figure 10 - Hoop Stress (psi) during the Blizzard of 1978; Deformat ion Scale Factor SOOX Page 12

5, 522 (Cyl)

(Avg: 75(:.'0)

+1.932e+03

+1.200e+ 03

+1. 0 0~ +03

+8.000e+02

+ S.OOOe +02

+4.00Ce *r*02

+2.000e +02

+O.OOO\\!+OO

-2.00Qe + 02

-4.000c +02

-S.OOOo::!of 02.

8. 000..+ 02

-1. 000Q+03

-1.20Qe+OJ

-3.47Ce+03

~

ODS: mI 2U.odb AbDqusjStandtl rd 6.10-3 Tue Feb 2118:00:59 Pacif,c Standard llme 2012 y

Step: SlOp-l Ir.crc ment 1: Step nme =

1.000 Pnmary Var: S, 522 (Cyl)

Deformed Var: U Deformation SCille Factor: + 5.000e+02 Figure 11-Hoop Stress (psi) during the Blizzard of 1978; Deformation Scale Factor 500X; Wider Contour Range (+/-1200 psi) 5, 5 33 (Cyl)

(Avg: 75%)

+1.396e+03

+9.000e+02

+7.500e+02

+6.00Ce+02

+4.S0Qe+02

+3.00Ce+02

+1 SOOe+02

+O.OOOe+OO

-1. 500e+02

-3.000e+02

-4.S00e+02

-G.OO Oe+02

-7.500e+02

-9.000e+02

-3.30ge+03 006: m1223.odo AbaqllsjStandard S. 10-3 Tue Feb 2. 1 18:00:59 PacIfic Standard li m 2.0n y

s t~ p ; Step-1 increment 1: Stl!P Time =

1.000 Primary Var": S, 533 (cyr)

Defo r rn~ V.::l r; U Deformation Scale Factor; +S.OOOe+02 Figure 12 - Vertical Stress (psi) during the Blizzard of 1978; Deformation Scale Factor 500X Page 13

The result 1977 Blizzard Condition due to the 1977 blizzard condition is shown in this section.

Figure 13 depicts the temperature distribution in the model. Figures 14 through 17 show the stress state in the max principal, radial, hoop, and vertical directions, respectively.

NTH

+4.676e+Ol

+4.225e+Ol

+3.774e+Ol

+3.323e +Ol

+2.B72e+Ol

+2.42.1e+ Ol

+1.970e + Ol

+I.S1Se+ Ol

+1. 069c -t-Ol

+6.177e+OO

+1.667e +OO

~

-2.8420-1*00

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Lx S~p : Step-l Increm ~n t 1: Step Ti me =

1.000 Prtm ~ ry Va.. : NT!l Deformed Var: U Deformation SC,;lle Factor : + $.OOOe + 02 Figure 13 - Temperature (OF) during the Blizzard of 1977; Deformation Scale Factor 500X Page 14

S, f'la x. Pnncipal (Avg: 750/0)

- + 1.632(H03

- +9,OOoe+02

. +7. S00e +02

. + 6. 000e+02

- + 1.S00e+D2

- +3.000e + 02

- + 1. 5QC.e+ 02

- +O.oOOe +OO

-LSOoe..*02

-3.000 ~ +02

-4. 500e+ 02

-6. QOOt! +02

- -7. S00e+02

-9.000e+ 02

-1. ? 14e+D3 008: m1222.od b Abaqus/Stdndard 6. 10-3 Tue Feb 2117:50:39 PiJciflC StllndDrd Time 2012.

Step : Step-1 Increrr ent 1: Step 'TIme :;

1.000 Pr1 m~ ry Var: 5, 1'>111x. Principal Deformed Var: U OeformaUon Scale FactOf-; +S.OOOe+02 Figure 14 - Max Principal Stress (psi) during the Blizzard of 1977; Deformation Scale Factor 500X 5, 511 (e yl)

(Avg: 75%)

+9.000e+02

+7.500e+02

+6.000e+02

+4.S0Ce+02

+3.000e+02

+1.S00e+D2

- +O.ooOe+OO

- -1. 500e+02

-3.000e+02

-4.500;:+02

-6.000e+02

~ -7.500e+02

-9.000e+02

-3.037e+03 008: m1222.odb Abaqus/Standard 6.10-3 Tue Feb.2 117:50:39 Padffc Stt:ndard T,me 2012 y

Step : Step-1 Increment 1: Step Time = 1.000 L x Pri ma ry Var: 5, Sl1 (eyl)

Deformed Var: U De ~o rm 3tjon 5cale Factor : +5.000e+02.

Figure 15 - Radial Stress (psi) during the Blizzard of 1977; Deformation Scale Factor 500X Page 15

S, S22 (Cyl )

(Avg : 75%)

+1.067e+03

.. +9,OCOO+02

+7.SQOe+02

+6.00~+02

+ 4.5001.: +01

+3. QOOe+02

+1. 500e-i 02

+O.OOOc+Ou

-1.S0Q(H 02

-3.QOOe+02

- -1, 500e+02

:~ ~gg;m Max 930

-9.000Q+02

- -3.601e+03 ODS: m 1222.odb Abaqus/Standard 6.10-3 Tue feb 211 7:50: 39 Pacific S~ndllrd Time 20 12 y

Step : Step-!

In crement 1: Step TIme =

1.000 pri lT\\.i'llJ'"lI Va r : 5, 5 22 (eyl)

Defor med Var: U Deformation Scale Factor : +S.DOOe+0 2.

Figure 16 - Hoop Stress (psi) during the Blizzard of 1977; Deformation Scale Factor 500X 5, 533 (cyl)

(Avg : 75%)

+1.002e + D3

+9.0aOe +02

+ 7.S00e+02

+6.QOOe +02

+ 4,500e+02

+3.000e+02

+1.5000 +02

+ D. QOQe*..OO

- -1.500e+02.

- -3.000 +02

- -4,SOOe+02

- -6.0QOe+02

-7.500e+02 Max 720

- *9.QOOe+02

- -S.S13e+03 ODS: m1222.odb Abaqus/Stundard 6.10-3 Tue Feb 2117:50:39 Pacific 5 ndard llme 201 2 y

Step: Step-1 Incremont 1: Step TIm

==

1.000 PnmlJry Var : 5, 533 (Cyl)

Deformed Var : U Ceformatlon Scale Factor: +S.OOCe+02.

Figure 17 - Vertical Stress (psi) during the Blizzard of 1977; Deformation Scale Factor 500X Page 16

Exhibit 62: Stress Analysis due to 105 MPH Wind load Appendix VllJ-63

Exhibit 62 Stress Analysis due to 105 mph Wind Load Summary of Results The results of the analysis presented in this report can be summarized as follows:

The wind pressu re does not produce stresses capable of delaminating the structure.

The 105 wind pressure load resu lts in a max principal stress of about 55 psi The 105 wind pressure load re sults in a radia l stress of less than 1 psi Modeling Summa ry was used exclusively in the finite element analysis presen ted here.

. II d Itli Il t 1"

The drawings used as geometry input for thi s model are:

Drawing No: (-100 Rev. 5 "Shield Building Foundation Plan & Details SH. 1" Drawing No: (-110 Rev. 6 "Shield Building Roof Plan Wall Section & Details" All vertical reinforcing bars in the containment shell section are modeled as rebar Ino (diameter 1.270")

at 12" spacing. The inner face horizonta l rebars are tl8 (diameter 1.000") at 12" spacing. The outer face horizontal rebars are 1/ 11 (dia meter 1.410") at 12" spacing. The vertical and horizontal rebars in the shoulder sections are #8 at 12" spacing.

Page 1 Page 1 of 6

Exhibit 62 1ateriCll Properties The material properties used for this analysis are summarized in Exhibit 56, Figure 2.1.4: Material Properties Figure 1 shows the radial displacement due to gravity and the 105 mph wind load. The maximum radial deflection is about 0.07 inch inward on the side of the structure facing the wind and about 0.07 inch outward on th e sides of the structure that are parallel to the wind direction.

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io ~ (; OC~' I '::-3 Direction of Wind Figure 1 - Radial Displacement (inches) due to Gravity and 105 mph Wind Load; Deformation Scale Factor =: 2000X Page 2 Page 2 of 6

Exhibit 62 Figure 2 depicts the max principal stress due to gravity and the 105 mph wind pressure load. The maximum stress due to the wind load is 55 psi although some larger stresses can be seen in the ring girder area. The stresses in the ring girder are a result of the dome weight and not due to the wind load.

$, {:tel '!:.

F~ n -' r;~1

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+J.jliA<l +Oi 2.... 1 l.e *' 01

+"1.05.72 + :0 1 2.!>&le !XI 55 psi yZ S~up : ii r lllr.Y*~1m.!*,."1in.:.:I *1 0 5m p '

L

>! I n ;; ro;orne~C 1 SC=[r 1'1 mp. '"

1. 1)(\\0

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l.::J rrttaber~ ~ U FEh:...l(-:

"I _ {tOC~*I* ;~3 Figure 2 - Max Principal Stress (psi) due to Gravity and 105 mph Wind Load; Deformation Sca le Factor =

2000X Page 3 Page 3 of 6

Exhibit 62 Figure 3 - Location Page L~

Page 4 of 6

Exhibit 62 Figures 4 and 5 below depict the max principal and radial stress The max principal stress shown in Figure 4 correlates well with shown in Figure 2. Both the location and magnitude of the max principal stress are in agreement. Figure 5 indicates that the radial stress is very low due to the combined gravity and wind load, The only location that experiences any significant radial stress is the corner of the flute. However, the corner of the flute location is a singularity due to the sharp angle between two elements in the finite element mesh. In the region of interest the rad ial stress is below 1 psi.

Page 5 Page 5 of 6

Exhibit 62

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Figure 4 - Max Principal Stress (psi) due to Gravity and 105 mph Wind S, S! l \\":~ IJ

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~.) ~ F,'n~jt. l * + ~ OJ I").oi+fC.I Figure 5 - Radial Stress (psi) due to Gravity and 105 mph Wind Page 6 Page 6 of 6

Exhibit 63: CFD Analysis of Shield Building

CFD ANALYSIS OF DAVIS-BESSE CONTAIN MENT TOWER ANALYSIS PERFORMED BY: __

JAN UARY 1, 2012

- - - inforrration. Perforrnanc:e Improvement international, LLC. ******

1

Davis-Besse Containment Tower Requirements The CFD analysis performed for this report includes:

Pages

No surrounding buildings

34mph from the Nort hwest (summer) 5-9

34m ph from the Southwest (winter) 10-14

72mph from the Southwest (winter) 15-22

With surrounding buildings

34mph from the Northwest (summer) 24-28

72mph from the Southwest (winter) 29-34

105mph from the Southwest (winter) 35-40

Tornado 41-44

Category F2

Traveled from t he Northwest to Southeast Boundary Conditions for the problem consisted of:

Winter

Ambient temperature of -13 °F.

Temperature of the containment tower remained at a constant 7°E

Summer

Ambient temperature of 104°F.

Te mperature of the containment tower rem ained at a constant 130°F.

Results extracted from the CFD:

Pressure distributions on t he surface.

Heat transfer coefficients.

Vorticity shedding calculated on t he 72mph case.

- - - - inform2tio;L Peric:rmJl1Ce Improvement intel*nation;;;i.. LLC ******

Model Creation

The CFD mesh consisted of 3.6 million cells to create the air volume.

Total size of the air volume was a 2,500 ft. diameter and a height of 670 ft.

Using a large air volume eliminates any wall effects.

_ sing a sma ll mesh size allows the vorticity shedding to be captured more accurately.

CONTAINMENT TOWER CFD MESH

Solution Method

The CFD program used for this analysis was Fluent version 13, an industry standard and proven analytical code.

-Incompressible ideal gas law was used, because the wind speeds are below Mach 0.55.

- The containment tower analysis without the buildings was done using a steady state solution.

- The containment tower analysis with the building was done using a transient analysis solution.

4 information. Performance I LLC._**_

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Davis-Besse Containment Tower CFD Results 34mph Northwest Summer Conditions

-1.33e-02

-1.86e-02

-2.40e-02

-2.94e-02

-3.47e-02 4.01e-02 4.54e-02

-5.08e-02

-5.62e-02

-6.15e-02

-6.6ge-02

-723e-02

~

DIRECTION

-1 ;33.~2

-1.86e-02

-2.40e-02

-2.94e-02

-3.47e -02 4.01e-02 4.54e-02

-5.08e-02

-5.62e-02

-6.15e-02

-6.6ge-02

-723e-02

-7.76e-02

-8.30e-02

-8.30e-02 FRONT BACK PR ESSURE CONTOU RS (psi)

- - - - in'i0l"1TI3tion, P erformal1c2Irnprovementim~mCltional, LLC. ******

6

HEAT TRANSFER COEFFICIENT HAND CALCULATIONS FOR ANALYTICAL COMPARISON TOWER = 130°F (54.4°C)

AIR TEMP = 10t;°F (40°C)

TEMP AVERAGE = llrF (42.22°C) v =0.1693 cmA2/s k =0.027 w/m*k Pr =0.71 U =15.20 m/s (34mph)

D=44.73m Re =U* D I v Re = 40,159,244 Nu =h* D/k l\\Ju = 0.3 + (0.62*Re AO.5*PrAO.33)/([1+(OA/Pr)AO.67)AO.25) " [1+(Re/282,OOO)AO.625)AO.8 Nu =38,092 h =(38,092

0.027 w/m

k) I 44.73m h = 22.99 w/mAYk

0.1761 BTU I hr"ftA2* of h =4.05 BTU I hr*ftA2* of ( This number compares to the front surface of the tower (slide 8). Region of comparison is the light blue and cyan)

This indicates the CFD model has predicted the correct surface heat transfer coefficients.

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I Davis-Besse Containment Tower CFD Results 34mph Northwest Summer Conditions 1.72e+o~

1.60e+01 1.4ge+01 1.37e+01 1.26e+01 1.15e+01 1.03e+01 9.16e+00 8.02e+00 6.87e+00 5.73e+00 1.15e+00 O.OOe+OO

~

DIREcnON FRONT 1.60..'*:01 1.498+01 1.378+01 1.26e+01 1.15e+01 1.03e+01 9.16e+00 8.02e+OO 6.87e+00 5.73e+00 4.58e+00 3.44e+00 2.2ge+00 1.15e+00 O.OOe+OO BACI(

Hx AREA OF COMPARISON Heat Transfer Wall Coefficients (Btu/ hr-ftJ\\2-0F)

- - - - inforTnatinn. Perform21lce impl"OVerne'1L internaticn;:;i LLC. ******

Davis-Besse Containment Tower CFD Results 34mph Northwest Summer Conditions 2.;"-02 1.8ge-02 1.35e-02 8.18.-03 2,82.-03

2~

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'33e-02

1.88H)2

2.400-02

294.-02

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~Jll0-02

~~

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8.30e-02 CROSS SECTION PRESSURE CONTOURS (psi)

CROSS SECTION VELOCITY CONTOURS (ft/s)

The cross section picture of the pressure contour shows a steady gradient pressure buildup in front of the tower.

At slow wind speeds the flow mainly stays attached except along the top front and aft edge.

The flow tries to stay attached, but flow separation happens at the bottom l1alf due to tile low pressure region.

The top dome has a profound effect on the flow separation.

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

- - - - lnformation. Performance Improver:,ent interllatondi, LLC. ******

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avis-Besse Containment Tower CFD Results 34mph Southwest Winter Conditions

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DIRECTION

-327e-02

-4.12e-02 4.98e-02

-5.83e-02

-6.68e-02

-7.54e-02

-8.3ge-02

-925e-02

-1.01e-01

-1.10e-01

-1.18e-01

-127e-01 L

-1.35e-01 FRO NT

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-5.83e-02

-6.68e-02

-7.54e-02

-8.3ge-02

-9.25e-02

-1.01e-01

-1.1'Oe-01

-1.18e-01

-1.27e-01 X

-1.35e-01 BACK PRESSURE CONTOURS (psi)

- - - - information. PerfolTnance Improvem ent inte l'llJt lOnal, LLC. ******

11

I HEAT TRANSFER COEFFICIENT HAND CALCULATIONS FOR ANALYTICAL COMPARISON TOWER =-13°F (-25°C)

AIR TEMP =TF (-13. 9°C)

TEMP AVERAGE =-3°F (-19.4°C) v = 0.1168 cm A 2/s k = 0.02248 w/m*k Pr=O.72 U =15.20 m/s (34mph)

D=44.73m Re=U *D/v Re =58,210,273 Nu =h*D/k Nu = 0.3 + (O.62*Re A O.5*Pr A O.33)/([1-:-(O.4/Pr)1I0.67]1I0.25) * [1+(Re/282,000)1I0.625]1I0.8 Nu =55,111 h = (55,111

0.02248 w/m*k) /44.73m h =27.7 w/mIl2*k '" 0.1761 BTU / hr*ftA2* OF h = 4.87 BTU I hr*ft" 2* OF ( This number compares to the front su rfa ce of the tower (slide 13). Region of comparison is the light blue and cyan)

Thi s indicates the CFD model has predicted the correct surface heat transfer coefficients.

- - - - l infofrnatio\\1, Performance Improvement intern2tion;;I, LLC. ******

12

I Davis-Besse Containment Tower CFD Results 34mph Southwest Winter Conditions 2.12e~1 1.96e+01 1.81e+01 1.66e+01 1.51e+01 1.36e+01 121e+01 1.06e+01 9.07e+00 1.51 e+OO O.OOe+OO Hx AREA OF COMPARISON 1.96e+01 1.818+01 1.66e+01 1.51e+01 1.36e+01 1.21 e+01 1.06e+01 9.07e+00 7.56e+00 6.05e+00 4.53e+00 3.02e+00 1.51 e+OO x~

O.OOe+OO BACI<

Heat Transfer Wall Coefficients (Btu/hr-ftA2-0F)

- - - - infnITnation. Performance Improvement internai.ional, l.LC. ******

13

I Davis-Besse Containment Tower CFD Results 34mph Southwest Winter Conditions

),56...02 2 70 *.o~

1.85e-02 8'-.0' lA3e-O)

'7.11.'()3

., 56e'()2

2A2...o2

-327...02

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-6.83...02

-6.668.02

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1'10.-01

1.18.-01

1 27e.Q1

1.35.-01

,l CROSS SECTION PRESSURE CONTOURS (psi)

CROSS SECTION VELOCITY CONTOU RS (ft/ s)

The cross section picture of the pressure contour shows a smaller pressure buildup in front of the 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 is vorticity shedding.

An effect of the flow separation at lower speeds will cause a cyclic pressure loads on the containment tower.

The top dome has increased the effect of flow separation.

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

1*-1 information. Performancf' Irnprovemeni: international, LLC ******

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-1.06e-02

-3.4Se-02

-S.83e-02

-3.4Se-02

-S.83e-02

~

BACK

HEAT TRANSFER COEFFICI ENT HAND CALCULATIONS FOR ANALYTICAL COM PARISON TOWER = -13°F (-25°C)

AIR TEMP = TF (-13.9°C)

TEMP AVERAGE = -3°F (-19.Ll°C) v = 0.1168 cm A 2/s k =0.02248 w/m*k Pr =0.72 U = 32.63 m/s (72mph)

D = 44.73m Re=U*D/v Re = 124,960,607 Nu = h*D/k Nu = 0.3 + (0.62*ReAO.S*PrAO.33)/([1+(0.4/Pr)AO.67]AO.2S) * [1+(Re/282,OOO)AO.62S]AO.8 Nu = 97,032 h = (97,032

0.02248 w/m*k) / 44.13m h = 48.76 w/mA2*k

0.1761 BTU / hr*ftA2* of h =8.587 BTU / hr*ftI\\2* of (This number compares to the front surface of thetower (slide 18). Region of comparison is the light blue and cyan)

This indicates the CFD model has predicted the correct surface heat transfer coefficients.

- - - - inhJl'm3cion. Perfol"fY1anc2 1mprovernent inLernaUonal, LLC. ******

17

Davis-Besse Containment Tower CFD Results 72mph Southwest Winter Conditions L

FRONT 327e+01 305e+01 2.83e+01 2.61e+01 2.40e+01 BACK 3.27e+01 3.05e+01 2.83e+01 Hx AREA OF COMPARISON Heat Transfer Wall Coefficients (Btu/hr-ftI\\2-0 F)

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18

Davis-Besse Containment Tower CFD Results 72mph Southwest Winter Condit ions

~

, 32e-02

-1 O6e-02

-345e-02

-S.83e-02

-8.22e-02

-1,Q6e-01 Large area of suction CROSS SECTION PRESSURE CONTOURS (psi)

CROSS SECTION VELOCITY CONTOURS (ft/s)

The pressure contour has stayed the same from the 34mph, but the pressure load and suction has increased.

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 72mph.

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 fl ow separation.

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

19

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I Davis-Besse Containment Tower CFD Results 72mph Southwest Winter Conditions Velocity path l.ines at 1/3 down from the top of the containment tower (ft/s)

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20

Davis-Besse Containment Tower CFD Results 72mph Southwest Winter Conditions Velocity path lines halfway from the top of the containment tower (ft/s)

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21

Davis-Besse Containment Tower CFD Results 72mph Southwest Winter Conditions Vorticity Shedding at 48ft from cylinder The flow tries to stay attached, but the flute causes separation.

Velocity contours 1/3 down from the top at 72mph Vorticity shedding frequency = (108 ft/s) / (48ft) = 2.25 hz 9.91e+01 9.01 e+01 811e+01 721e+01 6.3*le+01 5.41 e+O'I 4 51 e+01 360e+01

.70e+01 1 80e+01

- - - - 1information. Performanc21mprovemem International, ~L C. ******

Model Creation of Containment Tower with Buildings

- The CFD mesh consisted of 3.26 million cells to create the air volume.

- Total size of the air volume was a 2,500 ft. diameter and a height of 670 ft.

wall effects.

Using a small CONTAINMENT TOWER WITH BUILDINGS AIR VOLUME WITH BUILDINGS

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-Z94e-O'

.1948002 409e-Ol

-4 0ge.()~

-5 :!4e-01

.5.24e*01

.6 3ge-0?

-839e-01

-1S4e-O

. 754e*02

-8 70e-02

870e-D:!

-9 8fi~02

-98Se-02

-I IOe-O I

I IOe*OI

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-I ~6e-a l

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-1 7ge-OI

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~ With the addition of the surrounding buildings, the pressure has increased by O.027psi.

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25

I Davis-Besse Containment Tower CFD Results 34mph Northwest Summer Conditions I 78e+Ol 164e+01 r-~-

151e+Ol

, I 137e+OI I ~

',~:-,;"

123e+OI

\\ II 1 10e+Ol B 21e+00 II 6801e+00

~ 48e+00 4 11e+00 274e+00 137e+00 o OOe+OO 192e+OI I

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1370+00 0 000+00 FRONT BACK Heat Transfer Wall Coefficients (Btu/hr-ft"2-0F) 2G

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I Davis-Besse Containment Tower CFD Results 34mph Northwest Summer Conditions CROSS SECTION PRESSURE CONTOURS (psi)

CROSS SECTION VELOCITY CONTOURS (ft/s)

The pressure contours have dramatically changed with the addition of surrounding buildings.

There is a large low pressure region located above the building on the aft side of the containment tower.

The velocity vectors are disrupted from the buildings causing the flow to separate at lower wind speeds.

27

- - - - in'formation. Performance Improvement international, LLC ******

J Davis-Besse Containment Tower CFD Results 34mph Northwest Summer Conditions Large area of separated flow ll tl4e+Ol 636e+O I VELOCITY VECTORS (ft/s)

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-5 SSe-Ol Davis-Besse Containment Tower CFD Results 72mph Southwest Winter Conditions

3..13e-02

775&-02

-110e-01

-I 58e*OI

- 198e*OI

-238e-O I

-218e-01

-3.ISe-O!

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W ith the addition of the surrounding buildings, the pressure has increased by O.054psi.

BACK

- - - - irrformatior"l. Perfor, nanC2lrnprovemen1. inten-,atioi1.:;\\, U C. ******

30

I Davis-Besse Containment Tower CFD Results 72mph Southwest Winter Conditions 1*82e*OI

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- - - - information. Performance Improvemc=nt i nternati onClI, LLC ******

I Davis-Besse Containment Tower CFD Results 72mph Southwest Winter Conditions CROSS SECTION PRESSURE CONTOU RS (psi)

CROSS SECTION VELOCITY CONTOURS (ft/s)

The stagnation pressure region has shifted up towards the top of the containment tower. This is a result of the buildings being in front of containment tower.

The flow on the aft side of the tower is turbulent compared to the case with no buildings.

- - - - l lnfonn3tion. Perlorrnan(2 Improvement intemationai, LLC. ******

32

~I Davis-Besse Containment Tower CFD Results 72mph Southwest Winter Conditions 154e+02 1.42e+0..

1.30e+02 1 1ge+02 1.o7e+02 94ge+01 8 30e+01 711e+01 5.93e+01 4.74e+01 3'56e+0'1 237e;+ 01

1 1ge+01

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33

Davis-Besse Containment Tower CFD Results 72mph Southwest Winter Conditions 1 7ge+02 186e'*" 02 1 54e+O:!

142e+02 1..3tJe+02 1 18e+02 107e+(ts!

948e+01 8.30e+01 7 11e+Ol 5.93e+0 1 4.74e+0*\\

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I Davis-Besse Containment Tower CFD Results 10Smph Southwest Winter Conditions 5.23e*03

-844e-1r.1 I

., 74e*OI

264e-OI

353e-Ol

-443e-01

-5.33e-Ol

-622e-Ol

712e-Ol

-802e*Ol

-8.91 e-O I

-9.81e-Ol

-I07e+OO

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~I Davis-Besse Containment Tower CFD Results lOSmph Southwest Winter Conditions 270e*01 251~al

!32e+Ol

1 12e<-01

, 93e+Ol 1 74e+OI 155e+Ol 135e+Ol 2.51e*Ol e-Ol

1 121.'-01 1.93e* Ol 1 74~.Ol 155p-Ol I )5e.Ol 1 lBe ~OI 96&e.OO 713e+OD

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Davis-Besse Containment Tower CFD Results 10Smph Southwest Winter Conditions CROSS SECTION PRESSURE CONTOURS (psi)

CROSS SECTION VElOCITY CONTOURS (ft/ s) w The stagnation pressure region has shifted up towards the top of the containment tower. This is a result of the buildings being in front of containment tower.

The flow on the aft side of the tower is unsteady and turbulent.

The addition of the buildings has caused the flow to rise do to the pressure increase just before reaching the building.

This results in a higher pressure region at the midpoint causing a larger overturning moment.

38

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I Davis-Besse Containment Tower CFD Results 105mph Southwest Winter Conditions

'2 ~e+02 2 17e+02 200e+02 1.84e+02 I.B7e+02 1.S0e+02 1.34e+02 1.17e+02 I.OOe+02 8.35e+OI 334'6: +01

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- - - - information. Performanc21mproIJement international, LLC. ******

~I Davis-Besse Containment Tower CFD Results lOSmph Southwest Winter Conditions 250e+02

_ 34e+02

'2 t 7e+02 200e+02 184e+02 1.67e+02 1.50e+02 1.34 e+02 1,17e+02 100e+02 B,35e+O 1 VELOCITY VECTORS (ft/s)

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========1 Tornado Conditions

F2 Tornado with winds between 113 and 157m ph.

Tornado touched down just west of the Davis-Bessie power plant between 84Spm and 900pm on June 24, 1998.

.. Tornado was 100 yards wide an d t raveled southeast for 3 ~ miles.

Considerabie damage was noted along this path with some barns totally destroyed.

Slide 43 and 44 shows pressure contours on the buildings and co ntainment tower as the tornado passes.

4.62e-+02 4.33e+02 4,04E'-+02 375e-t02 3.46e+02 3 17e-t02 288e-+02 2.60e-t02 2 31 e-t02 2.02e-+D2 1.73e+02 1.44e+02 1.158+02 8.65e+01

~,.77 e-t{J 1 2.8%+01 6.988-03 VELOCITY VECTORS SIMULATING A TORNADO (ft/s)

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II

J Davis-Besse Containment Tower CFD Results CFD Analysis of Tornado Passing by Containment Tower

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I Davis-Besse Containment Tower CFD Results CFD Analysis of Tornado Passing by Containment Tower t

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Exhibit 64: Thermal Stress Analysis with Gravity and Wind Load

Exhibit 64 Thermal Stress Analysis with Gravity and Wind Load Table ofContents Summary of Results.................................................................................................................................2 Modeling Summary.................................................................................................................................2 Overall Approach.................................................................................................................................2 Finite Element Software......................................................................................................................3 Modeled Geometry.............................................................................................................................3 Finite Element Models.............................................................................................................................3 Descriptions......................................................................................................3

...................................................................................................................4

....................................................................................................................... 6 Thermal Stress Screening.........................................................................................................................7 Thermal Stress Screening Results.........................................................................................................8 Combination Load Cases......................................................................................................................9 Analysis Based on Measured Properties................................................................................................ 10 Circumferential Temperature Distribution at O.F. Horizontal Rebar................................................... 11

..........................................................................................................................12 Stress State during Hot Summer Condition............................................................................................ 13 Stress Analysis Results Summary.......................................................................................................13 Shoulder 10 location

.............................................................................. 14 Azimuth 225" Location

............................................................................... 17 Pagel Page 1 of 19

Exhibit 64 Summary of Results The results of the analysis presented in this report can be summarized as follows:

The temperature and wind conditions found to maximize the radial stress are not sufficient to delaminate the structure alone Thermally induced radial stresses is maximized at the hot summer temperatures At the location of the outer face horizontal rebar, the maximum radial stress due to temperature gradients, gravity, and wind is about 300 psi Modeling Summary Overall Approach

--~--

~~-~-

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

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

Page 2 Page 2 of 19

Exhibit 64 Finite Element Software analysis presented here.

was used exclusively in the finite element Modeled Geometf'y The drawings used as geometry input for this 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" All vertical reinforcing bars in the containment shell section are modeled as rebar #10 (diameter 1.2701J) 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 l.4lOIJ) at 12" spacing. The vertical and horizontal rebars in the shoulder sections are #8 at 12" spacing.

Finite Element Models focusing on two distinct different geometric Page 3 Page 3 of 19

Exhibit 64 Figures 1 through 3.

Figure 1

- Shoulder/Flute; All Mesh Shown Page 4 Page 4 of 19

Exhibit 64 Figure 2

-Shoulder/Flute; Mesh with Rebars Exposed Figure 3

- Shoulder/Flute; Rebar Mesh

~

~-

~

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

---~-----

PageS Page 5 of 19

Exhibit 64 Figures 4 through 6.

Figure 4 Shell Section; All Mesh z

~y x

Figure 5 Shell Section; Mesh with Rebars Exposed Page 6 Page 6 of 19

Exhibit 64 Figure 6 Shell Section; Rebar Mesh Thermal Stress Screening In order to understand the effect of the various thermal conditions that the containment structure may be subjected to, a screening analysis was performed. The screening analysis was performed using preliminary material properties before the official material properties were obtained.

The screening analysis considered a total of 32 thermal conditions. They included the summer and winter solstice, the spring and autumn equinox, windy and calm condition, as well as average and hot/cold ambient temperatures.

The six thermal conditions resulting in the highest radial stress in the screening analysis is analyzed with gravity and wind pressure loads in the next section.

Page 7 Page 7 of 19

Exhibit 64 Page 8 Page 8 of 19

Exhibit 64 Combination Load Cases The result of the screening analysis identified the thermal conditions most likely resulting in the highest These combination load cases were again solved with the preliminary material properties since the official values had not yet been obtained.

Page 9 Page 9 of 19

Exhibit 64 Analysis Based on Measured Properties The six cases predicted to result in the maximum radial stress is analyzed using measured material properties from samples taken from the Davis-Besse containment structure. The material properties used for the analysis are summarized in a separate section in the Root Cause Analysis Report (Exhibit 56, Figure 2.1.4: Material Properties The conditions analyzed using the measured material properties are the same six conditions presented in Table 2. They are listed below along with the time of day determined to produce the highest radial Page 10 Page 10 of 19

Exh ibit 64 Circumfcr ntiLll TCmpCI.IlLi e Dislributiun.1 O.F. HOI-izontal Rebar The temperature profiles for the six conditions resulting in the highest radial stress based on the screening analysis are shown in Figure 7. The temperature profile s are plotted in the circumferential direction around the shield building at the outer face horizontal rebar depth.

Temperature (OF), Mid-Height, Outer Face Horizontal Rebar Depth LL:'

Q III

l-L-

60 L

a;"'

c..

E III I

No Wind, Summer Solstice, Hot Temperature, 7:30 PM :I 40 No Wind, Summer Solstice, Average Temperature, 7:30 PM No Wi nd, Autumn Eq uinox, Hot Temperature, 6:00 PM 80 No Wind, Autumn Equinox, Average Temperature, 6:00 PM 34 mph Wind, Summer Solstice, Hot Temperature, 6:00 PM I 34 mph Wind, Autumn Equinox, Hot Temperature, 5:00 PM o

22.5 45 67.5 90 112.5 135 157.5 180 202.5 225 247.5 270 292.5 315 337.5 360 Azimuth n Figure 7 - Circumferential Temperature Distribution at O.F. Horizontal Rebar For each of the six temperature profiles shown in Figure 7 eight set of double valleys can be seen. The valleys represent the lower temperature under the thick sections of the shoulders. These areas are covered by thicker layer of concrete so it takes longer for them to heat up due to the hot exterior conditions. Figure 7 also shows that the azimuth 225° location corresponds to the hottest location around the structure. The condition resulting in the hottest temperature at the outer face horizontal rebar depth is labeled "No Wind, Summer Solstice, Hot Temperature, 7:30 PM." This is the temperature condition studied in the following sections.

Page 11 Page 11 o f 19

Exhibit 64 The south to south-west side has the highest thermal gradient do to the solar heating during the day.*

Figure 8 shows the location of the flutes, shoulders, and the azimuth convention for the Davis-Besse containment structure.

Figure 8 - Shield Building Flute Numbers and Azimuth Locations Pctge 12 Page 12 of 19

location Exhibit 64 Stress State during Hot Summer Condition The results shown in this section describes the detailed stress state in the hottest location around the structure for the hot summer condition (No Wind, Summer Solstice, Hot Temperature, 7:30 PM)

Figures 9 through 13 show the r'esults from the shoulder 10 location Figures 14 through 18 depict the same results from the azi muth 225 0

For each of the two locations the result is presented in five figures, The first figure shows the temperature distribution and the following four figures depict the stress state:

1, Temperature Distr'ibution 2,

Max Principal Stress 3,

Radial Stress 4,

Circumferential (Hoop) Stress 5, Vertical Stress The stress state is presented at the mid-height section

, The contour range is set to +/- 300 psi for all the stress figures so that they can be compared more easily, Stress 11.(1) ~is Rt;'~ults Smllmary The maximum stress is confined to the top and bottom of the outer face horizontal rebars, The maximum tensile stress is about 300 psi and not enough to crack the concrete, Page 13 Page 13 of 19

Exhibit 64 Shoulder 10 Location The temperature distribution, max principal stress, radial stress, hoop stress, and vertical stress in shoulder 10 ar-e depicted in Figures 9 through 13, I-espectively. Figure 9 shows that the shoulder surface is hotter than the flute surface. This is the result of more solar exposure on the shoulder su rface compare to the flute valley. Also, there is more surface area at the corner of the shoulder resulting in higher temperatu re during the hot ambient condition.

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Figures 10 through 13 depict the stress state using the max principal stress and the three stre ss components in a cylindrical coordinate system located at the containment structure center. The max principal and radial stresses are highest at the outer face horizontal rebar. The figures also indicate an in the Shoulder 10 Location area of high stress on the left edge of the model. This has been identified to be a si ngularity Comparing the stress in the three radial, hoop, and vertical directions (Figures 11 through 13 respectively) indicates th at the radia l component has the highest tensile stress. As shown in Figure 11, the radial tensile stress is below 300 psi which is less than the tensile strength of the concrete. It is concluded that the hot summer temperature condition is not capable of delaminating the structure in the flute/shoulder location.

Page 14 Page 14 o f 19

Exhibit 64 S, I"lax. Princip'"

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Figure 11-Radial Stress (psi) in the Flute/Shoulder Page 15 Page 15 of 19

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CefOfmed VM : U ecrormatlon So.:alc factor: +5 DOOe Figure 13 - Vertical Stress (psi) in the Flute/Shoulder PClge 16 Page 16 of 19

Exhibit 64 Azimuth 225 ' Location The temperature distribution, max principal stress, radial stress, hoop stress, and vertical stress in the shell area at azimuth 225° are shown in Figures 14 through lS, respectively. Figure 14 shows that the exterior surface is hotter than the interior. This is the result of the hot ambient daytime condition and the colder nighttime condition.

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Deformed Var: U Deform~tjon Scale F"ctor: +5.000e+02 Figure 14 - Temperature Distribution (OF) at the Azimuth 225 0

Location Figures 15 through 18 depict the stress state using the max principal stress and the three stress components in a cylindrical coordinate system located at the containment structure center. The max principal and radial stresses are highest at th e outer face horizontal rebar depth (see Figures 15 and 16).

Comparing the stress in the radial, hoop, and vertical directions (Figures 16 through lS, respectively) indicates that the radial component has the highest tensile stress. As shown in Figure 16, the radial stress is below 300 psi which is less than the strength of the concrete. It is concluded that the hot summer temperature condition is not capable of delaminating the structure in the shell section location (middle of a panel).

Furthermore, Figures 17 and 18 show that the hotter exterior surface temperature results in compression stresses in both the hoop and vertical directions due to expansion of the outer layer.

Pctge 17 Page 17 of 19

Exhibit 64

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X Figure 15 - Max Principal Stress (psi) at the Azimuth 225° Location S,S11(C),I)

(A'.*9,75%)

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Defo.-nleD VlIr : U CercI n1!.tlon Scale FlIctor: + 5 OOOe+02 Figure 16 - Radial Stress (psi) at the Azimuth 225° Location Page 18 Page 18 of 19

Exhibit 64 S, 522 (Cyl)

(A"g: 75%)

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DefQrmed Vl.Ir: U Defo rm~tl on Sc<lle F~c t:or: +5.000e+02 at th e Azimuth 225° Location Figure 17 - Hoop Stress (psi)

S, 533 (e\\,ll (A':g : 7 5%)

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Page 19 Page 19 o f 19

Exhibit 65: Thermal Analysis Appendix VIII-66

Exhibit 65 DAVIS-BESSE THERMAL ANALYSIS 1.0 ANALYSIS MODEL For the purposes of assessing the seasonal and daily variations in the temperature of the outer concrete shield building of the Davis-Besse reactor, a detailed 3D transient thermal analysis finite element based model was generated. This model was derived from the same 3D NASTRAI'J structural" model that was used for this effort and utilized the same node and element numbers. Additional surface flux thermal elements were added to the interior faces of the concrete in order to improve the capture of radiation heat transfer from the interior steel containment as well as convective heat transfer by the passage of the annular air. By preserving the same node numbers, this permitted directly mapping temperatures onto the NASTRAN structural model without having to interpolate temperatures between dissimilar meshes. This ensured that temperatures were accurately specified for all structural analyses performed with the NASTRAN model. Similarly, preserving the element ID numbers ensured proper specification of thermal properties for all of the materials present.

The thermal analysis model that was used for this effort and is shown in Figure 1. The majority of the concrete was modeled using linear 8 node brick elements. The steel rebar reinforcement was explicitly modeled using 1D bar elements that share the same nodes with the 3D solid elements used to represent the concrete.

Only in the dome region OF and IF rebar were membrane elements used to approximate the smeared thermal properties of the rebar and concrete based upon a volumetric averaging of their properties.

All pre-and post processing of the thermal analysis model was performed using MSC MD.PATRAN version 10.2. MSC MD.PATRAN is an open ended pre-and post - processor that facilitates the creation and post-processing of results for a number of different CAE solvers. This includes MD.NASTRAI\\J and ABAQUS, the two structural finite element analysis (FEA) solvers used for this effort. This enables models and results derived from one the analysis code to be converted into its equivalent in another code. In this way, the NASTf~AN thermal models and results files could be converted into an equivalent ABAQUS version.

Page 1 Page l of 23

Exhibit 65 By doing so, this obviated the need to generate a separate ABAQUS based thermal analysis model.

FIGURE 1. NASTRAN TRANSIENT THERMAL ANALYSIS MODEL Page 2 Page 2 of 23

Exhibit 65 2.0 THERMAL BOUNDARY CONDITIONS 2.1 RADIANT HEAT TRANSFER SOURCES The primary intent of thermal analysis was to ascertain the variation in temperature that occurs daily as well as seasonally. To accomplish this, it is essential that the variation in the position of the sun as it transits across the sky be properly modeled. This entails specifying the zenith angle, Z, or the angle of the sun relative to a normal pointing directly overhead. The zenith angle is a function of both the latitude as well as the time of year. It is derived from the following relationship (see http://edmall.gsfc.nasa.gov/inv99Project.Site/Pages/science briefs/ed-stickler/ed-irradiadiance.html ):

Z::: Zenith Angle::: The angle from the zenith (a point directly overhead) to the Sun's position in the sky. The zenith angle is dependent upon the latitude, solar declination angle and time of day.

Z::: cos-1 (sin tP sin 0+ cos tP cosH) q, =Latitude H

Hour Angle = 15° x (Time -12) (Angle of radiation due to time of day where time is given as the hour of the day from midnight)

~ = Solar Declination Angle Solar Declination Angles for the Northern Hemisphere Vernal Equinox March 21/226 0° Summer Solstice June 21/226 =+23.5° Autumn Equinox September 21/22 6 =0° Winter Solstice December 21/22 6

-23.5° The solar radiation that strikes the earth, also known as Insolation, is then simply given by 1= Scos (Z)

I::: Insolation or solar flux S::: solar flux -

1000 Watts/ m 2 -2.2 Btu/Hr - in2 (Clear day insolation perpendicular to the incident solar radiation)

Z::: Zenith angle Page 3 Page 3 of 23

Exhibit 65 The solar insolation that strikes the earth is strongly affected by the angle of incidence with the surface being radiated by the sun. The more oblique the angle, the lower the flux.

Consequently, in the latitudes farther from the equator, the solar insolation will be lower. In addition, seasonal variations will cause the solar declination to change by 47 degrees between the winter and summer solstices. Thus, the solar flux will be least during the winter and greatest during the summer in the northern hemisphere. The UV spectrum of sunlight is principally responsible for solar heating. It is strongly affected by the angle it passes through the atmosphere. This is shown in Figure 2 1.4

~---~ June solstice 1.2 1.0

-~

i

~

.5 0.8

.~

~ Sept """"

~

Yeal a:

~

Mar. equinox 0.6 0.4

"'"" ~

Dec. solstice 0.2 0

10 20 30 40 50 latitude (0)

FIGURE 2. SEASONAL AND ANNUAL VARIATIONS IN RELATIVE SOLAR UV-A RADIATION (340 nm) FOR DIFFERENT LATTITUDES (BASED ON JOHNSON ET AL 1976)

Page 4 Page 4 of 23

ML12138A081

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