thermal shock fracture of silicon carbide and its

THERMAL SHOCK FRACTURE OF SILICON CARBIDE ANDITS APPLICATION TO LWR FUEL CLADDINGPERFORMANCE DURING REFLOODYOUHO LEE*, THOMAS J. MCKRELL, and MUJID S. KAZIMIDepartment of Nuclear Science and Engineering, Massachusetts Institute of Technology (MIT)77 Massachusetts Avenue, Cambridge, MA 02139*Corresponding author. E-mail : [email protected]

Invited September 12, 2013Received September 17, 2013Accepted for Publication September 24, 2013

1. SiC CLADDING AS A REPLACEMENT FOR THECURRENT ZIRCALOY CLADDING

Zircaloy cladding prevents fission-products’ releaseinto the coolant while imposing major limits on nuclearreactor designs, and safety. These limits are mainly dueto zirconium based alloy embrittlement through chemicaland radiation damage, early pellet-cladding mechanicalinteraction (PCMI), and restricted mechanical performanceand chemical stability at high temperature. Today, thereis a demand for higher burn-up and enhanced safety forlight water reactors. Therefore, the limitations of zirconiumbased alloy cladding are being viewed more criticallygiven recent events. Hence, Light Water Reactor (LWR)performance and safety would be considerably improvedby finding a replacement for its cladding that demonstratesbetter ability to withstand the more challenging LWRconditions [1].

A cladding made of silicon carbide (SiC) has beenproposed as a replacement for the current cladding, madeof zirconium (Zr) based alloys. SiC is already widely usedin many applications involving harsh environments, suchas combustion engines. It is also attractive for nuclearreactor applications, especially as a cladding material. SiCcaptures less neutrons than Zr, demonstrates higher strength

at high temperatures, has good chemical stability, andresistance to radiation damage. In short, many of the SiCproperties fit well with cladding requirements. However,SiC is a SiC is a brittle material brittle material and has alower thermal conductivity than zirconium based alloy,thus its introduction into reactors should be subjected tocareful evaluation. As such, feasibility of a fuel rod withSiC cladding in LWRs should be subjected to a high levelof scrutiny to ensure improved performance in operationand under accident conditions [1].

2. CURRENT STATUS OF SIC CLADDINGRESEARCH AND TECHNICAL ISSUES OF SiC CLADDING FAILURE [1]

A fuel rod cladding made of silicon carbide has beenstudied as a replacement for the current zircaloy claddingin several places around the world [1-9]. ManufacturingSiC cladding made of triple SiC layers - monolith/fibercomposite/and environmental barrier coating (EBC) hasbeen developed to dimensions approaching the currentLWR fuel rod design [5,7,8]. Radiation performance ofSiC/SiC ceramic matrix composite (CMC) cladding has

SiC has been under investigation as a potential cladding for LWR fuel, due to its high melting point and drasticallyreduced chemical reactivity with liquid water, and steam at high temperatures. As SiC is a brittle material its behavior duringthe reflood phase of a Loss of Coolant Accident (LOCA) is another important aspect of SiC that must be examined as part ofthe feasibility assessment for its application to LWR fuel rods. In this study, an experimental assessment of thermal shockperformance of a monolithic alpha phase SiC tube was conducted by quenching the material from high temperature (up to1200ºC) into room temperature water. Post-quenching assessment was carried out by a Scanning Electron Microscopy (SEM)image analysis to characterize fractures in the material. This paper assesses the effects of pre-existing pores on SiC claddingbrittle fracture and crack development/propagation during the reflood phase. Proper extension of these guidelines to anSiC/SiC ceramic matrix composite (CMC) cladding design is discussed.KEYWORDS : Fuel, Cladding, Silicon Carbide, Quenching, Safety

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http://dx.doi.org/10.5516/NET.02.2013.528

been proven to be promising but heavily dependent uponmanufacturing processes [2,10]. Efforts have been madeto evaluate SiC oxidation performance under Loss ofCoolant Accidents (LOCAs) during its service as LWRcladding [1,6,7,11]. These SiC oxidation studies havedemonstrated orders of magnitudes slower reaction rate thanthat of Zr. This is an indicator that major failure modes forSiC cladding would be principally different from failuremodes of the Zr cladding stated in the U.S Code of FederalRegulation, Title 10, Part 50.46, “Acceptance Criteria forEmergency Core Cooling Systems (ECCS) for Light-WaterNuclear Power Reactors” (10 CFR 50.46) [1,12,13]: 10 CFR50.46 states:

1. Peak cladding temperature. The calculated maximumfuel element cladding temperature shall not exceed(1204°C).

2. Maximum cladding oxidation. The calculated totaloxidation of the cladding shall nowhere exceed 0.17times the total cladding thickness before oxidation(Equivalent Cladding Reacted, ECR).

3. Maximum hydrogen generation. The calculated totalamount of hydrogen generated from the chemicalreaction of the cladding with water or steam shallnot exceed 0.01 times the hypothetical amount thatwould be generated if all the metal in the claddingcylinders surrounding the fuel, excluding the claddingsurrounding the plenum volume, were to react.

4. Coolable geometry. Calculated changes in coregeometry shall be such that the core remains ame-nable to cooling.

5. Long-term cooling. After any calculated successfulinitial operation of the emergency core cooling system(ECCS), the calculated core temperature shall bemaintained at an acceptably low value and decayheat shall be removed for the extended period of timerequired by the long-lived radioactivity remainingin the core.

It is worth noting how pervasive the effects of claddingoxidation are in the establishment of the current U.S NRCLOCA criteria. Indeed, the fundamental mechanism ofcladding embrittlement of zircaloy during LOCA is dueto micro-structural changes of the cladding with oxidation[14,15]. That is, the oxidized cladding cross section exhibitsan oxide layer, an oxygen stabilized alpha-phase layer, anda region of prior beta-phase. Importantly, oxidation ofzircaloy above the alpha-to-beta transformation temperatureresults in inherently brittle phases for the regions affectedby oxygen. Hence, ductility of zircaloy cladding is signifi-cantly impaired with oxidation, and embrittlement can leadto cladding fragmentation during the quenching phase ina LOCA. The ability of the cladding to withstand the thermalshock stresses during the reflood phase of LOCA is closelyrelated to the degree of oxidation reaction [13,16]. Thecurrent allowable peak cladding temperature (1204°C)and the maximum oxidation (17% ECR) criteria were

chosen in such a context – these limits are adequate toensure the survival of the cladding under the thermal shockduring the reflood phase of LOCA. The maximum hydrogengeneration limit is also affected by cladding oxidation. Inaddition, the coolable geometry criterion concerns thechange in coolant channel geometry due to potentialblockages through brittle cladding failure. The claddingbrittle failure is predominantly caused by lower ductilityas a result of oxidation during LOCA. The Long-termcooling criterion is also affected by cladding oxidation,as the oxide layer formed on the cladding surface lowersthe cladding thermal conductance. Today, the U.S NRCis modifying 10 CFR 50.46 to reflect the fact that thecurrent limits of maximum cladding temperature andmaximum oxidation are not conservative for high burnupcladding [17,18]. In particular, hydrogen embrittlementof zircaloy is significantly exacerbated with burnup. Thenew rule is focusing on maintaining appropriate ductilityas the unit of measure to determine survivability duringthe quench process and any other unforeseeable event.

Hence, research conducted prior to this point assures thatfailure mechanisms, hence safety criteria, of SiC claddingwould principally depart from the current practice estab-lished based on Zr based cladding. Indeed, the attempt touse SiC definitely is a radical departure from the presentexperience, with different material (ceramic) from stainlesssteel and metal based alloys. At this early stage of ourassessment of SiC cladding behavior in LWRs, the focusshould be on understanding failure modes because theywill reveal the feasibility, performance, and appropriatedesign and safety criteria [1].

A key failure mode of SiC cladding can be expectedto arise from its brittleness, when fast fracture occurs underexcessive tensile stresses. In this study, thermal stressescaused by rewetting of fuel rods during the reflood phaseof a LOCA are investigated as a probable mode for excessivetensile stresses in fuel rod cladding.

3. BACKGROUND ON SiC CLADDING THERMALSHOCK FRACTURE

Failures of a load bearing structure can be either of theyielding-dominant or fracture-dominant (fast fracture)types. Fast-fracture dominant failures are fractures thatoccur before general yielding. For such failures, the sizescale of defects, which is of major significance, is essentiallymacroscopic, since general plasticity is not involved butonly highly localized plasticity is involved with flaws ordefects [19]. Fast fracture of ceramics due to lack of ductilityto accommodate defects undergoing plastic deformationhas been a relatively well understood subject. MonolithicSiC undergoes a fast fracture failure mode if tensile stressesare excessive [20,21]. SiC/SiC composites exhibit rathercomplex modes of failures that show some degree ofductility [23,24], which is sometimes called brittle-like

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LEE et al., Thermal Shock Fracture of Silicon Carbide and its Application to LWR Fuel Cladding Performance During Reflood



(or quasi-ductile) failure. For the SiC cladding design forLWR application, the CMC design is being considered[1,2,7]. The CMC design employs monolith SiC as theinner most layer of the cladding, mainly to keep the fissiongas inside the rod, and provide strength to the cladding. Theouter monolith layer protects against the corrosive actionof water, while the middle SiC/SiC composite, characterizedby a higher fracture toughness than the monolith, is used toprotect the inner monolith and to make the cladding failurehappen in a less drastic manner. A thin CVD preparedenvironment barrier coating (EBC) is employed at theouter most protective layer. Figure.1 illustrates a typicaldesign for CMC layers.

Under accident conditions, fuel cladding may experiencesignificant tensile stresses due to both thermal shock, andrapid depressurization of the core operating pressure. Forceramics, cracks in compression tend to get closed upand propagate stably, and may twist out of their originalorientation to propagate parallel to the compression axis[25,26]. Fractures are not caused by rapid unstable propa-gation of one crack, but by the slow growth of many cracksto form a crushed zone. It is not the size of the largestcrack that counts but that of the average crack size [25].In contrast, cracks in tension tend to open and propagateunstably perpendicular to the applied stress direction. Insuch a case it is often the largest crack that governs failure.Hence, the projection of stress distributions around flawlocations inside the cladding should be determined toanalyze thermal shock performance.

A triplex cladding is regarded as a more robust structurefor thermal stresses induced by quenching than the claddingstructure made of a sole monolith; the CMC structure hasa composite layer of high fracture toughness with additionalcrack arresting capabilities. The outer most layer of thecladding experiences the greatest tensile stress as it sees thesharpest temperature gradient in reflood cases of LOCA.EBC is a monolithic SiC, which exhibits lower fracturetoughness compared to composite materials. In case ofthe failure of EBC upon quenching, propagating crackswould run into the neighboring composite layer, which isunfavorable for crack growth. Hence, understanding ofthe CMC fracture upon quenching requires a detailed

description of stress fields in each layer, flaw distributions,and crack propagation mechanisms between the layers.This study explores monolithic SiC performance uponquenching as a preliminary attempt to envision materialperformance of the EBC layer and the innermost monolith.Thus, it provides a building block for understanding theCMC cladding behavior.

4. EXPERIMENT

An experiment facility was built to bring SiC specimensup to 1500ºC and drop them into a pool of water, as illus-trated in Fig.2. The tubular SiC specimens are suspendedin the air inside a quartz tube located at the center of thefurnace. By employing bottom-flooding with tubularsamples, this experiment was designed to demonstratesimilar experimental designs/conditions that were used toestablish the current Zr cladding safety criteria [13,34]written in 10 CFR 50.46. A B-type thermocouple readsthe temperature adjacent to the outer surface of the quartztube, where the SiC specimen is located. The temperaturereading is recorded by the data acquisition system (DAS).Temperature calibrations were made between this B-typethermocouple reading and the temperature obtained by athermocouple attached to the sample’s surface. Comparingthese two temperatures, an empirical relation between thefurnace temperature and the true sample surface temperaturewas established and used to report SiC specimen tempera-tures. SiC specimens were suspended inside the furnaceuntil it reached constant temperature. Then, specimenswere quickly dropped into the room temperature water pool(~22ºC) by an air-pressure driven rod. A high speed videocamera was used to record the quenching of the specimens.Recorded quenching videos were used to analyze transientboiling states for later application in modeling.

Fig. 1. CMC SiC Cladding Layers(Figure in Courtesy of Stempien.et.al. [10]) Fig. 2. Quenching Experiment Facility

The material used in this experimental study wasmonolithic tubular Hexoloy -SiC with a density of3.05g/cc, obtained from Saint-Gobain. The specimens haddimensions of 14mm OD, a thickness of 1.55mm, and aheight of 13mm. The sample was received as a long tube,and was cut into the specimen size. Cut end surfaces werepolished by a grinding wheel, and then ultrasonicallycleaned with detergent added to water, deionized (DI) water,acetone, and methanol prior to furnace exposures. Post-quenching examinations were conducted with scanningelectron microscope (SEM) analysis for all tested specimens.

5. RESULTS

39 SiC specimens at temperatures ranging from 350ºCto 1174ºC were quenched into deionized water at 22ºC.A minimum of three independent tests were performedfor each temperature, except for 1033ºC. Experimentalresults are summarized in Table 1.

Shattered specimens are ones that immediately brokeinto multiple pieces upon quenching. Cracked specimensare ones that were observed to have crack growth by eithervisual examination or SEM analysis. Thus, all shatteredspecimens are regarded as cracked specimens. The experi-mental results show that SiC specimen temperatures above350ºC result in crack formation for the tested temperatureresolution. For those crack inducing quenching temperatures,SiC specimens are expected to undergo strength degradationafter the thermal shock. Past thermal shock studies conductedwith SiC found threshold material temperatures for strengthdegradation [27]. In this study, we used survival probabilityas a measure of thermal shock tolerance, which is definedas the ratio of the number of crack-free samples / number

of total samples tested after quenching. Visual observationsof cracks in SEM analysis are limited to only surfacecracks. Hence this may underestimate thermal shock damageon the material. Nevertheless, it can still reveal a strengthdegradation trend with quenching temperature as shownin Figure.3.

Tested SiC materials are pressureless sintered -SiC,which is characterized by considerable porosity. Represen-tative SEM images of pre-quenched specimens are shownin Fig. 4 and Fig. 5. Pores can essentially play as a pre-existing flaw where stress is concentrated. The SEM analysisin Fig.4 shows average pore diameters 20-50μm that aredistributed uniformly. Neither pores nor pre-existing flaws



SiC T(±5o)

350ºC

400ºC

450ºC

500ºC

550ºC

600ºC

700ºC

795ºC

1033ºC

1174ºC

4

4

3

3

3

3

4

4

2

9

0

0

0

0

0

0

0

1

2

9

0

2

3

3

3

3

4

4

2

9

0.0

0.0

0.0

0.0

0.0

0.0

0.0

25.0

100.0

100.0

0.0

50.0

100.0

100.0

100.0

100.0

100.0

100.0

100.0

100.0

SamplesTested

Shattered CrackedShattered

(%)Cracked

(%)

Obtained Data

Table 1. Thermal Shock Experiment Results

Fig. 3. SiC Specimen Survival Probability Based on CrackGrowth

Fig. 4. SEM Image of Cross Sectional Ends of an As-ReceivedTubular SiC Specimen

Fig. 5. SEM Image of Side Surface Microstructure of an As-Received Tubular SiC Specimen

were observed on the side surface of as-received tubularspecimens, Fig. 5.

Fractured surfaces of shattered samples showed thatcracks propagate through grains (transgranular fracture),Fig.6. Transgranular crack propagation was commonlyobserved for all cracked samples at lower temperaturesas illustrated in Fig. 7. Transgranular fractures explainfairly straight looking cracks with smooth edges, whichare different from the faceted fracture surfaces resultingfrom intergranular fractures. To enhance fracture toughness,intergranular rather than transgranular fractures need tobe promoted. Transgranular fracture cracks take a straightpath through grains, whereas intergranular cracks do notenter grains, instead traveling along the grain boundaries.This allows the branching of the crack through interlockinggrains, enhancing overall toughness [28].

Once formed, cracks tend to run axially and radiallyacross the entire thickness of tubular specimens. Thisindicates that hoop stress is the most dominant stressdirection for crack propagation. Crack growth exhibiteddifferent behavior for different quenching temperatures.Higher quenching temperatures caused a wider crackwidth as shown in Fig. 8 and Fig. 9.

Cracks that were formed at 1174ºC are about 25 μmwide while those at 450ºC were less than 5 μm. Temperaturegradients inside a quenched material are steeper whenquenching from higher temperatures due to the initialtemperature difference. Steeper temperature gradients leadto a greater thermal expansion mismatch inside a material,

causing higher stress levels. Higher stress levels are energet-ically more favorable for crack propagation and a materialaccommodates a higher strain energy release rate by creatinglarger cracks. Cracks exhibited a tendency to be linked atpores as shown in Fig. 10. This explains that pores (sites ofvoid where no elastic potential energy can be accommodated)are energetically favorable for crack propagation.

6. THERMAL STRESS MODELING OF TRANSIENTLWR FUEL RODS

A rigorous analysis of the observed experimental resultswould come from understanding the fracture mechanismunder certain stress fields. There have been many studieson transient stress field calculations for quenched materials.



Fig. 6. SEM Image of Intragranular Fractured Surface of aQuenched SiC Specimen (T=1033ºC)

Fig. 7. SEM Image of Transgranular Crack Propagation of aQuenched SiC Specimen (T=500ºC)

Fig. 8. Cracked SiC Specimen after Quenching at T=1174ºC





Generally, transient energy equations are solved for afixed heat transfer coefficient and thermal diffusivity ofthe material to obtain transient temperature fields. Thistemperature field is then inputted into elastic stress-strainequations to obtain transient stress fields as illustrated inthe schematic diagram in Fig. 11.

Although the convention of one sided coupling of energyand stress-strain equations is a first principle approachthat can be applied universally for various thermal shockcases, care should be taken in using it. Heat transfer modesbetween a quenched material and coolant should be carefullyaddressed. Conduction may be the dominant heat transfermode for the early portion of the transient before anyappreciable convective heat transfer mode such as filmboiling occurs [29,30]. Maximum stresses may be createdduring this early portion of the transient. At the instant whena SiC specimen meets quenching water, the instantaneousSiC surface temperature can be found by assuming thesituation as a contact of a semi-infinite solid. This approxi-mation is reasonable for the early portion of the transient,during which temperatures in the interior are essentiallyuninfluenced by the change in surface conditions andconduction is the dominant heat transfer mode [31].

The instantaneous surface temperature, Ts, can be foundby the energy balance at the interface [31]

where k is conductivity, is density, Cp is heat capacity, andTi is the initial temperature (22ºC) of the water. Assumingthat the surface temperature remains constant for a brief

instant of the early portion of the transient, material temper-ature distribution T(x, t)SiC can be calculated as follows [31]

where Ti,SiC is the initially uniform SiC temperature, Ts isthe instantaneous surface temperature found by Eq.1, x isthe position inside the sample, d is thermal diffusivityand t is time. Obtained T(x,t) is inputted into the followingequations to yield transient stress fields.

where r is radial stress, is hoop stress, z is axial stress,E is Young’s modulus, is the poisson ratio, is thermalexpansion coefficient, r is the radial location, a is the radial

Fig. 11. Schematic Diagram for Thermal Stress Analysis

Fig. 12. Schematic Illustration of SiC Surface Temperature during quenching with Water

(1)

(2)

(3)

(4)

(5)

Elastic Modulus, E

Thermal Conductivity, k

Density,

Heat Capacity, Cp

Thermal Expansion Coefficient,

400 GPa

29.35 W/m-k

3100 kg/m3

1298 J/kg-k

5.1x10-6

Table 2. Input for -Hexoloy SiC Properties for Thermal StressCalculation

position of the innermost surface, b is the radial positionof the outermost surface, and T is the difference betweenthe sample temperature T(x,t) and a constant temperatureat which material is stress-free. Note that the shell of thetested tubular specimens was treated as a semi-infinite planein the Cartesian coordinate in Eq.(1) and Eq.(2) for thetemperature field calculation over the very short time aftercontact. Effects of the curvature of the tubular specimenin temperature fields are negligible as the shell thicknessis much smaller than the outer diameter (thickness 0.11of OD). The following material properties evaluated attemperatures of 1033ºC were used [33].

Stresses are tensile at the outer surfaces while com-pressive in the middle of the quenched specimen as shownin Fig.13. Results shown in Fig.13 exhibit sharp temperatureand stress gradients during the earlier portion of the transientwhile the interior temperatures are unaffected. The stressesare projected over a finite thickness of a critical flaw andincreases stress intensity. If thermal stress intensity is largerthan the material’s fracture toughness, fractures are initiatedPores were uniformly populated over the entire thicknessof the tested specimens. Pores/flaws near the surfaces aremore significant in contributing to fracture than internalpores/flaws, which is a primary reason why surface qualitycontrol of SiC cladding is important.

Note that the calculated stresses based on the conductionmodel in the very early portion of a transient are the ceilingfor true maximum stresses. They assume water as a neigh-boring continuum, where heat can flow without an interfacethermal resistance. In reality, a heat transfer mechanismwould be somewhat mixed between conduction and convec-tion. Thermally induced agitations of water molecules dueto rapid heat conduction and water movement next to thesample surface with dropping the specimen would result

in a certain macroscopic movement of water molecules.This macroscopic movement of water would provide anadditional mechanism for heat transfer. Also, contactresistance in heat transfer would exist between the quenchedmaterial and the neighboring water during transient con-duction. The lowest limit for the true maximum stress wouldbe the immediate boiling heat transfer at time zero. Thiswould impose thermal resistance at the beginning andneglect the conduction dominant phase of the transient.Thus, the true maximum stresses would be bracketed bythese two limiting cases

Instant Boiling Model (Boiling at time 0) < True Maximum Stress < (6)Pure Conduction Model (Conduction at time 0)

Even for a convective heat transfer after an appreciableconvection starts, the heat transfer coefficient rapidlychanges with time depending on the sample temperature.Hence, using a single heat transfer coefficient may leadto an unrealistic interpretation of the experimental results.Consequently, current thermal shock models set a limiton the use of commercially available software for thecladding temperature analysis upon quenching during areflood phase. Current understanding of transient energyanalysis for fuel rods during accident scenarios does nottake into account the discussed technical details, becauseit uses the instant boiling model at time 0.

A large break loss of coolant accident (LBLOCA)analysis in a typical PWR was run by RELAP-5 as a testcase (input of the U.S. NRC) [32]. Cladding thermalconductivities and heat capacities were modified to SiCproperties found in Carpenter’s work [2]. Fig.14 showspeak fuel rod cladding temperature with time at the axiallocation of 1.811m.



Fig. 13. Transient Temperature and Hoop Stress Distributions in the Specimen Thickness at Time = 0.1ms, Initial SiC Temperature = 1033ºC, Water Temperature = 22ºC

Fuel rewetting starts at 100~125 seconds with the rapidtemperature drop. During the rewetting process, thermalstresses induced by the temperature gradient inside thecladding are expected to be small with similar claddingouter and inner surface temperatures. This may be anunrealistic snapshot of cladding temperatures for the veryearly phase of quenching where the inner most claddingsurface is not affected by the outermost temperatureperturbation. Such an instant moment would be the timewhen the cladding experiences the greatest stress. Thecurrent RELAP-5 models and numerical schemes are notintended to handle such rapid transient thermal hydraulic-stress coupling. The same can be said for the NRC oxidefuel transient analysis code, FRAPTRAN 1.4. Such arigorous thermal hydraulic-stress coupled model is notrequired by the current criteria set for the zircaloy cladding.These criteria are indeed empirical judgment supportedby many redundant experimental data. Since brittle fractureof zircaloy upon quenching is a conditional failure thattakes place after considerable oxidation, rigorous modelinghas been made primarily for the modeling of oxidationand hydriding. Quenching performance of zircaloy oxidelayer and underlying brittle phase has been predominantlyaddressed by means of experimentation with qualitativeexplanation in terms of the degree of oxidation damageand retention of ductility. For SiC cladding, this approachmay not be acceptable because brittle fracture upon quench-ing is not a conditional failure mode. Rigorous investigationsof structural failure in terms of imposed excessive stressesshould be pursued. Efforts are being made to explain pre-sented experimental results with modeling. Experimentalresults in Table.1 and survival probability shown in Fig.1do not necessarily imply the same behavior for CMCcladding. Stress fields that tested specimens are differentfrom the reality of an actual CMC cladding. The maindifference comes from (1) two sided quenching for testedspecimens while only the outer most surface sees coldwater in case of an actual fuel rod, and (2) additional

tensile stresses would be imposed for an actual fuel roddue to depressurization of the reactor.

7. CONCLUDING REMARKS

In thermal shock experiments, pressureless sintered-SiC exhibited vulnerability to transgranular fracture

with temperature dependence of survival probability. Poresacted essentially as pre-existing flaws and transgranularcracking with pore bridging was observed. Well preparedCVD -SiC with minimal porosity is worth testing as acomparison in terms of thermal shock performance.Fracture models being developed for -SiC can readilybe applied to -SiC with a correction on pre-existingflaw size and geometry. The heat transfer mechanism hasa dominant role on temperature gradient inside the materialand therefore stress fields. Rigorous treatment of materialbehavior, such as thermal hydraulic coupling during theearly portion of transient is absent in current codes. Throughthis study, advancements in thermal shock models that arereadily applicable to LWR fuel rods of ceramic cladding,including SiC, are being developed.

ACKNOWLEDGMENTSFinancial support from the INL Academic Center of

Excellence at MIT and AREVA Fellowship in NuclearEnergy Technology at MIT are appreciated. Visiting Frenchstudents Aline Montecot and Yann Song are acknowledgedfor their assistance, particularly on SEM analysis. Theauthors appreciate samples provided by Saint Gobain.

NOMENCLATUREk – Thermal Conductivity [W/m-k]

– Density [kg/m3]T – Temperature [ºC or K]

– Stress [MPa]E – Young’s Modulus [GPa]



Fig. 14. SiC Cladding Temperature History During a Design Based LBLOCA

– Poisson Ratio [-] – Thermal Expansion Coefficient [K-1]

d – Thermal Diffusivity [m2/s]

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