Figure 4 summarizes the microstructural evolution during the breakdown process of NIFS-HEAT-2

Hot/cold roll Heat treatment

1373 K/RT 973 K 1273K 1373K 1573K

Ti-rich blocky precipitates (with N, O, C)

Elongation, band structure Dissolution

ingots.4 Bands of small grains aligned along the rolling direction were observed at the annealing temperature below 1223 K. The grains became homogeneous at 1223 K. The examination showed that optimization of size and distribution of Ti-CON precipitates are crucial for good mechanical properties of the V-4Cr — 4Ti products. Two types of precipitates were observed, that is, the blocky and the thin precipitates. The blocky precipitates formed during the initial fabrication pro­cess. The precipitates aligned along the working direc­tion during the forging and the rolling processes forming band structures, and were stable to 1373 K. Since clustered structures of the precipitates result in low impact properties, rolling to high reduction ratio is necessary for making a thin band structure or homo­genized distribution of the precipitates. The thin pre­cipitates were formed at ~-973 K and disappeared at 1273-1373 K. At 1373 K, new precipitates, which were composed of V and C, were observed at grain bound­aries. They seem to be formed as a result of redistri­bution of C induced by the dissolution of the thin precipitates. The impact ofthe inhomogeneous micro­structure can influence the fracture properties.14

Figure 5 shows the hardness as a function of final heat treatment temperature for three V-4Cr-4Ti materials: NIFS-HEAT-1, NIFS-HEAT-2, and US — DOE-832665 (US reference alloy).15 The hardness has a minimum at 1073-1273 K, which corresponds to the temperature range where formation of the thin precipitates is maximized. With the heat treatment higher than this temperature range, the hardness increases and the ductility decreases because the

Annealing temperature (K)

precipitates dissolve enhancing the level of C, N, and O in the matrix. Based on the evaluation of various properties in addition to the hardness as a function of heat treatment conditions, the optimum heat treat­ment temperature of 1173-1273 Kwas suggested.

Plates, sheets, rods, and wires were fabricated mini­mizing the impurity pickup and maintaining grain and precipitate sizes in Japanese, US, and Russian programs. Thin pipes, including those of pressurized creep tube specimens, were also successfully fabricated
in Japan maintaining the impurity level, fine grain size, and straight band precipitate distribution by maintain­ing a constant reduction ratio between the intermedi­ate heat treatments.16 The fine-scale electron beam welding technology was enhanced as a result of the efforts for fabricating the creep tubes, including plug­ging of end caps.17 In the United States, optimum vacuum level was found for eliminating the oxygen pick-up during intermediate annealing to fabricate thin-walled tubing of V-4Cr-4Ti.18 In Russia, fabrica­tion technology is in progress for construction of a Test Blanket Module (TBM) for ITER (International Ther­monuclear Experimental Reactor).19

Joining of V-4Cr-4Ti by gas tungsten arc (GTA) and laser welding methods was demonstrated. GTA

is a suitable technique for joining large structural components. GTA welding technology for vanadium alloys provided a significant progress by improving the atmospheric control. The results are summarized in Figure 6. Oxygen level in the weld metal was controlled by combined use of plates of NIFS — HEAT-1 (181wppm O) or US-8332665 (310wppm O) and filler wire ofNIFS-HEAT-1, US-8332665, ora high-purity model alloy (36 wppm O). As demonstrated in Figure 6, ductile-brittle transition temperature (DBTT) of the joint and the oxygen level in the weld metal had a clear positive relation. This motivated further purification of the alloys for improvement of the weld properties.20 Only limited data on irradia­tion effects on the weld joint are available at present.

The welding results in complete dissolution of Ti — CON precipitates and thus results in significant increase in the level of C, O, and N in the matrix. In such conditions, radiation could cause embrittlement. Some TEM observations showed enhanced defect clus­ter density at the weld metals. However, the overall evaluation of the radiation effects remains to be per­formed. Especially, elimination of radiation-induced degradation by applying appropriate conditions of post­weld heat treatment (PWHT) is the key issue.

For the use of vanadium alloys as the blanket of fusion reactors, the plasma-facing surfaces need to be protected by armor materials such as W layers. Limited efforts are, however, available for developing the coating technology. A low pressure plasma-spraying method was used for coating W on V-4Cr-4Ti for use at the plasma-facing surfaces. The major issue for the fabrication is the degradation of the vanadium alloy substrates by oxidation during the coating processes. Figure 7 shows the result of bending tests of the coated samples. The crack was initiated within the W layer propagating parallel to the interface and followed by cracking across the interface. Thus, in this case, the quality of W coating layer is the issue rather than the property of the V-4Cr-4Ti substrate or the interface. Hardening of substrate V-4Cr-4Ti by the coating occurred but was shown to be in acceptable range.21

Figure 8 is a collection of the products from NIFS-HEAT-2.

4.01.3.2.1 Irradiation creep: Macroscopic behavior

Under neutron irradiation, metals exhibit a high creep rate, much higher than the out-of-reactor ‘ther­mal’ creep rate, the creep rate increasing as the neu­tron flux increases. The behavior under irradiation of zirconium alloys, and particularly the creep behav­ior, has been studied extensively as pointed out by Franklin eta/.134 and Fidleris,150 because of the major importance of the prediction of the in-reactor defor­mation of the fuel assembly in the case of PWR and boiling-water reactor (BWR)169 or in-reactor struc­ture especially in the case of the CANDU

163,179

reactor.

It is usually assumed, for practical considerations, that the in-pile deformation consists of the sum of (i) the growth, (ii) the classical thermally activated out-of-pile creep, or so-called thermal creep, and (iii) the irradiation creep, strictly speaking.100,150,163,180 The ‘pure’ irradiation creep, subtracted from the two other components of the deformation, is the result of mechanisms which differ from the thermal creep and the growth. Nevertheless, these mechanisms are certainly coupled since they all imply disloca­tion loops, slip and climb of line dislocations, and point-defect bulk diffusion toward these defects. But very few authors have studied these potential

134,181

couplings.

The creep deformation under irradiation results, in fact, from two antagonistic phenomena. Indeed, while new deformation processes are activated, caus­ing the creep rate to increase, the thermal creep rate is strongly reduced by irradiation due to the irradiation-induced hardening. Indeed, it has been shown150 that a preirradiation reduces the thermal creep component of the deformation under irradia­tion. The effect of preirradiation on the reduction of the irradiation creep rate is particularly noticeable for RXA alloys. However, the hardening effect saturates at fluence of about 4 x 1024nm—2 and is followed by a steady-state creep rate. Concerning cold-worked materials, the effect of the preirradia­tion is much lower, according to Fidleris.150

As reported by several authors,134,150,153,182 the metallurgical state of the zirconium alloy has a sig­nificant effect on the in-reactor creep resistance. Indeed, while cold working may improve the thermal creep resistance of Zircaloy in certain test directions and stress range, it increases the in-reactor creep rate appreciably.150,153 Nevertheless, the creep sensitivity to the initial dislocation density is significantly lower than the growth sensitivity to the initial dislocation density.171 On the other hand, the grain size does not seem to have a significant effect on the creep strength in the range from 1 to 70 pm.

The in-reactor creep rate is very sensitive to irra­diation as well as loading conditions. The effects of flux, as well as the effect of stress, are usually described by a power correlation. The effect of tem­perature is usually described by an Arrhenius equa­tion.134 However, since it is in general very complex to distinguish between the ‘pure’ irradiation creep and the thermal creep, the authors usually use an overall creep constitutive law (eqns[1] and [2])163,180 and only growth is taken into account as a separate defor­mation component.

e ethermal—creep T ^irradiation—creep T egrowth

ecreep T egrowth Ц

with

ecreep = K s”ffexp(jT^

where e is the strain rate in s—1; s is the effective stress for thermal creep in MPa; n is the stress exponent; Tis the temperature in K; Qis the activation energy inJ; R is the gas constant, 8.31 J K—1 mol—1; ф is the fast neutron flux in n m—2 s— (E > 1 MeV); p is the flux exponent; and K is a constant for thermal creep in s—1 (MPa)—n(nm—2s—1)—p. According to various authors,134,150 the flux exponent (p) has been assigned values ranging from 0.25 to 1. A flux exponent ofp = 1 is commonly obtained for CANDU pressure tube deformation.163,183 For uniaxial creep tests per­formed at 280 °C on cold-worked Zy-2, Tinti184 has obtained a flux exponent increasing from 0.6 to 1.0 with increasing instant flux.

A stress exponent of n = 1 is obtained at 300 °C for low applied stress (s < 100 MPa). As the stress increases, the stress exponent increases, reaching values up to n= 25 for 450 MPa applied stress for cold-worked Zr-2.5% Nb.183

The effect of temperature on the creep rate can be rationalized by plotting the creep rate in an Arrhenius plot (logarithm of the creep rate vs. inverse temperature). The activation energy is then the slope obtained in this plot. It can be seen in Figure 23 that for low temperatures, the creep activa­tion energy Q/R is very low, between 2000 and 5000 K.150,163 The irradiation creep at low temperature is therefore nearly athermal. At higher temperatures, the dependence increases rapidly toward values of Q/R of 25 000-30 000 K. These last values are close to the activation energy measured for thermal creep. These observations tend to prove that for low-temperature regime, mainly ‘pure’ irradiation creep mechanisms are activated. As the temperature increases, the ther­mal creep mechanisms become activated, yielding to activation energy close to the thermal creep values.

It has also been shown by several authors that while the thermal creep of zirconium alloys is aniso­tropic, the irradiation creep remains strongly anisotropic.150 According to Holt,171 the anisotropy of irradiation creep is nevertheless slightly lower than that of thermal creep.

4.01.3.2.2 Irradiation creep: Mechanisms

Various mechanisms for irradiation creep have been proposed in the literature as reviewed by Franklin et a/.,134 Holt,163,171 Matthews and Finnis,181 and Was.9 A nice history of the proposed mechanisms for both zirconium alloys and stainless steels is given by Franklin et a/.134 These mechanisms can fall mainly into two large categories:

1. The mechanisms based on stress-induced prefer­ential absorption (SIPA) of point defects by line dislocations arising from different fundamental phenomenon. These mechanisms lead to the climb of edge dislocations under applied stress, yielding a creep deformation.

2. The mechanisms based on climb-enhanced disloca­tion glide mechanisms, which are essentially a com­bination of climb of dislocations due the absorption of point defects under irradiation and glide resulting in a creep deformation. For this category of mechan­isms, the strain is essentially produced by glide but the strain rate is controlled by the climb.

Other mechanisms involving irradiation-induced loops have also to be added to these two categories of deformation mechanisms involving line disloca­tions. Indeed, the stress-induced preferential nucle — ation (SIPN) of loops or the stress-induced preferential growth of loops due to SIPA can lead to an additional creep strain.

The SIPA mechanism is based on the fact that under an applied stress, the bias of the dislocation becomes dependent on the orientation of the Burgers vector with respect to the direction of the stress.105,134,181 Indeed, as described previously, due to a higher relaxation volume, the sink strength of an edge dislocation toward SIAs is higher than toward vacancies. This difference in sink strength is the bias of the edge dislocation. It can be shown that a dislo­cation with a Burgers vector parallel to the applied stress exhibits a higher bias toward SIAs than a dislo­cation with a Burgers vector perpendicular to the applied stress. Therefore, under irradiation, the net flux of SIAs (SIA flux minus vacancy flux) toward dislocations, with Burgers vector parallel to the applied stress, is higher than the net flux of SIAs toward dislocations with Burgers vector perpendicular to the applied stress. This difference in the absorption of point defects by different types of dislocations leads to dislocation climb, resulting in a creep strain. The SIPA creep rate is insensitive to the grain size but is sensitive to the dislocation network.

However, it has been seen that for growth, the anisotropic diffusion of SIAs is believed to play an important role in the deformation mechanism. There­fore, any irradiation creep model proposed for zir­conium should also include anisotropic diffusion. The SIPA model that includes anisotropic diffusion is called the SIPA-AD model and has been reviewed by Matthews and Finnis.181

In the case of RXA zirconium alloys, the irradia­tion creep mechanisms are not clearly identified yet. Indeed, since the initial dislocation density is very low, another deformation mechanism has to be acti­vated. The creep strain could be partly due to the preferred nucleation and/or growth of the (a) type loops in the prismatic planes. Indeed, according to the SIPN or SIPA mechanism, the nucleation or growth of interstitial (a) loops can be favored in the prismatic planes perpendicular to the applied stress. For the same reason, the nucleation or growth of vacancy (a) loops can be favored in the prismatic planes parallel to the applied stress, leading to a resulting creep strain. According to Faulkner and McElroy,185 an applied stress increases the mean diameter of (a) loops without affecting the density, proving that the SIPA mechanism is efficient in their experiment. However, the growth of (a) loops under an applied stress can explain the measured creep strain only for low strain levels. Indeed, this creep strain should remain limited since the (a) loop den­sity and mean loop diameter saturate at relatively low doses. Since the initial dislocation density is very low in RXA zirconium alloys, creep mechanisms involv­ing climb of dislocations due to the SIPA mechanism or climb-plus-glide of dislocations require the gener­ation of a dislocation network. It is possible that (a) loops coalescence occurs, resulting in the creation of a dislocation network that is able to climb and glide under stress.181,186 However, this network is clearly observed only at 400 °C.67 Other types of dislocation sources, such as Frank-Read or Bardeen-Herring sources,147 can also be activated under both irradia­tion and applied stress, leading to the creation of a dislocation network that undergoes a SIPA or climb — enhanced glide mechanism.

It should also be pointed out that in order to explain the observed creep rate, some mechanisms must be activated that allow the dislocations to over­come the high density of dislocation loops during their climb and glide motion, even for low applied stress. It is possible, as pointed out by MacEwen and Fidleris187 in the case of Zr single crystal, that the gliding dislocations are able to clear the loops during in-pile deformation, leading to the dislocation chan­neling mechanism. All these mechanisms probably occur in series, as proposed by Nichols,188 explaining the evolution of the stress dependency as the stress increases. Indeed, according to this author, for zero applied stress, growth of zirconium occurs, and then as the stress increases, (a) loop alignment occurs (SIPA on loops). For higher stress, the climb of line dislocations via SIPA takes place, and then the dislo­cation climb and glide processes occur at even higher stress. For very high stress, close to the YS, disloca­tion channeling occurs.

For cold-worked zirconium alloys, such as SRA Zircaloy or cold-worked Zr-2.5Nb alloy,163 the SIPA mechanism on the initial dislocations is a likely mech­anism for irradiation creep. However, according to Holt,171 the creep anisotropy of cold-worked zirco­nium alloys computed from the SIPA mechanism assuming only (a) type dislocations is not in agree­ment with the experimental anisotropy. The anisot­ropy computed from the climb-plus-glide mechanism assuming 80% prism slip and 20% basal slip is in good agreement with the experimental anisotropy, demon­strating that climb-plus-glide mechanism is probably the effective mechanism. It should also be pointed out that, since dislocations climb toward grain boundaries or toward other dislocations, recovery of the initial dislocation network occurs. In order to maintain a steady-state creep rate, multiplication of dislocations should also occur either via loop coalescence or via dislocation sources, as discussed previously.

It should also be pointed out that, as there is a coupling between swelling and irradiation creep in stainless steel,181 we could assume a coupling between growth and irradiation creep to occur in zirconium alloys due to the effect of the stress on the partitioning of point defects.134,162 Nevertheless, the simple assumption of two separable deformation components has proved to hold correctly for the results given in the literature.163,180

4.01.3.2 Outlook

Concerning damage creation and point-defect clus­ter formation, improvement in the knowledge of anisotropic diffusion of SIAs as well as better under­standing of the microstructure of vacancy and inter­stitial (a) loops and basal (c) vacancy loops (origin of the loop alignment, origin of the corduroy contrast for instance) has to be aimed at. Multiscale modeling approaches coupled with fine experimental analyses of the irradiation microstructure (high-resolution TEM, synchrotron radiation analyses, tomography atom probe, etc.) should bring new insight concerning the previous points mentioned but also elements in order to propose modeling of the microstructure evolution during irradiation: for instance, origin of the alignments of Nb precipitates, stability of p-Nb pre­cipitates, etc.

Concerning the mechanical behavior of Zr alloys after irradiation, multiscale modeling of the postirra­diation deformation with a better understanding of the dislocation channeling mechanism and under­standing of its effects on the postirradiation mechan­ical behavior are needed.

Moreover, better understanding of the postirradi­ation creep deformation mechanisms is also needed using multiscale modeling.

The last point concerns the deformation mechan­isms under irradiation. In that field, the basic questions are still without answers: What are the irradiation creep deformation mechanisms? What are the coupling between the deformation under irradiation and the thermal creep and growth? Progress has to be made especially using in situ deformation devices under irradiation, coupled with modeling approaches. (See also Chapter 1.01, Fundamental Properties of Defects in Metals; Chapter 2.07, Zirconium Alloys: Properties and Characteristics and Chapter 5.03, Corrosion of Zirconium Alloys).

GBE is an emerging field, which promises methods to improve the performance of materials, whose deg­radation in service is caused by the presence of high angle boundaries. The concept, first proposed59 by Prof. T Watanabe in the early 1980s, envisages improvement ofproperties ofmaterials by controlling the grain boundary character distribution (GBCD). Many processes like diffusion, precipitation, segrega­tion, sliding, cavitation, and corrosion are kinetically faster along high angle grain boundaries. Hence, it is possible to decelerate these detrimental pro­cesses by replacing the random boundaries with low energy ones, coincident site lattice (CSL) boundaries (denoted by the ‘sigma number,’ S, which is defined as the reciprocal of the fraction of lattice points in the boundaries that coincide between the two adjoining grains on the basis of CSL model). Another prerequi­site for GBE is to completely destroy the interconnec­tivity of random grain boundary network. The insight in the field of GBE was achieved with the advent of computer assisted EBSD (electron back scatter diffraction) technique developed during the 1980s.

The embrittlement in ferritic steels is known to be caused by segregation phenomena. The kinetics of segregation can be controlled by suitable selection of the nature of grain boundaries. GBE has been applied60-63 to combat embrittlement problems in ferritic steels. The task of carrying out GBE using experimental methods is time consuming. Hence, it is prudent to resort to computational methods, which need to be validated using selected experiments. A 3D Poisson-Voronoi grain structure, simulated using MC technique was employed to study60 (Figure 12(a)) intergranular crack percolation using percolation theory. The percolation threshold was estimated to be 80%. To apply this model to specific alloys like ferritic steel, system specific characteris­tics need to be incorporated61 in the model. One such attempt is to define the propensity of the grain boundaries for propagation of cracks based on rela­tive values of the grain boundary energy and the

0.185 0.180 0.175 0.170 0.165 0.160 0.155 0.150 0.145 0.140 0.135 0.130 0.125

0 5 10 15 20 25 30 35 40 45

(d) Charpy impact energy (J)

energy required for propagation of cracks. These calculations were carried out (Figure 12(b)) for two different grain sizes. The prediction of finer grain size being favorable to reduce embrittlement was con­firmed (Figure 12(c)) experimentally. The GBCD, that is, the distribution of various grain boundary types has been evaluated62 in modified 9Cr-1Mo fer­ritic steel using EBSD technique. The experimental observations confirmed the reduction of DBTT by 20 K with reduction in grain size. The fractal analysis of the fracture surface demonstrated (Figure 12(d)) that the tortuous path which cracks need to follow in fine grain sample is responsible63 for the observed reduction in the propensity for embrittlement.

It is shown clearly that the low energy boundaries can be introduced in engineering materials in three different methods: preferential nucleation of low angle boundaries around twins or controlled recov­ery or orientation relations during phase transforma­tion, if some of the variants happen to result in CSL boundaries. Significant improvements in properties using GBE have been achieved64 in many austenitic stainless steels, in contrast to ferritic steels. The major challenges in the application of GBE to Cr-Mo fer­ritic steels arise from the following factors: lower twinning probability, higher stacking fault energy, and limited variants with CSL boundaries during g! a transformation during cooling.

The push for higher operating temperatures in tur­bine engines, as well as in reactor designs for both terrestrial and space applications, has frequently

governed the periodic scientific examinations of refractory alloys. The historical examination of Nb and its alloys is typical of this, with early studies of the irradiation properties exploring the potential uses of these alloys in fusion energy and fission type space reactor. While these alloys have favorable properties, such as elevated temperature capability and compati­bility with liquid alkali metals for energy applications, and attractive physical properties such as thermal conductivity, much of the work on Nb and Nb-base alloys has examined the nonirradiated properties.

It is worth putting into perspective the relatively small commercial market for niobium-base alloys. Approximately, 75% of all niobium metal is used as minor alloying additions in steel, and only 1-2% is produced in the form of niobium-base alloys. The total market for niobium-base alloys in the mid-1990s was <105kgyear — (100metric tonnes year-1).6 By comparison, 60 tons of niobium was used in 1961 for the SNAP-50 reactor program alone, and substantial additional quantities were used for other research projects such as the NERVA (Nuclear Experiment for Rocket Vehicle Applica­tions) program.7 Throughout the history of the vari­ous space reactor programs, dozens of alloys were examined, with several brought to near-commercial production. However, today only the Nb-1Zr and C-103 (Nb-10Hf-1Ti) alloys remain commercially available for use in the sodium vapor lamp and rocket or turbine engine exhaust nozzles.

Nb-1Zr has historically been considered the only niobium-base alloy with a sufficiently mature database (mechanical properties including thermal creep, chemical compatibility, fabrication, and weld­ing knowledge) to be considered a near-term candi­date for radiation environments.7-11 Developed for high ductility and good weld characteristics, the alloy shows less-than-desirable thermal creep strength at elevated temperatures compared to other refractory alloys. Though the C-103 alloy has greater short-term elevated temperature strength than that of Nb-1Zr, its long-term properties show no improvement over Nb-1Zr.12 Nonetheless, Nb-1Zr is the only Nb-base alloy with a significant radiation effects database.

Despite the periodic programmatic interest in the use of Nb and Nb-base alloys, no clear fundamental study of the irradiated properties for a specific appli­cation has been performed or completed. Much of the data available on the irradiated properties is scattered and easily spans a time frame of several decades, which can lead to misinterpretations of results on the basis of either the limited scientific knowledge of the time, lack of understanding of the sensitivity of properties on impurity concentrations, or aging effects. Radiation effects data are limited to the examination of swelling and tensile properties, with no information regarding fracture toughness or irra­diation creep performance.

The following sections deal with radiation effects on the properties of Nb and Nb-1Zr specifically. While some initial scoping examinations have been performed on other Nb-base alloys, these are relatively inconsequential and based on the less-than — desirable ductility, thermal stability, or welding cap­abilities of these alloys.

4.08.5.1 Application and Technical Issues

Generation IV advanced nuclear power systems are proposed; the temperature and dose regimes for their operation are shown in Figure 23.42 Among them, the supercritical water-cooled reactor (SCWR) and the lead fast reactor (LFR) require a higher neutron dose

at an operating temperature of 600 °C. It is known that 9Cr-ODS steels have superior compatibility with sodium, but their corrosion resistance is not adequate for SCPW and LBE at a temperature >600 °C. Thus, the most critical issue for the application of 9Cr-ODS steels to SCWR and LFR is to improve their resistance to corrosion.

It has been reported that the addition of chromium (>13 wt%) and aluminum (4 wt%) to ODS steels quite effectively suppresses corrosion in an SCPW and LBE environment. In general, however, an increase in the Cr content often results in increased susceptibility to ther­mal aging embrittlement. Furthermore, the addition of Al significantly reduces steel strength at high tempera­tures. Recent progress in R&D of high Cr-Al-added ODS ferritic steels is summarized in the proceedings of the International Conference of Advanced Power Plants (ICAPP) 2009. The oxidation and corrosion performance of Al-added 16Cr-ODS steels in SCPW and LBE environments is described in Section 4.08.7.

A mechanism for the irradiation-induced creep of graphite was proposed by Kelly and Foreman53 which involves irradiation-induced basal plane dis­location pinning/unpinning in the graphite crystals. Pinning sites are created and destroyed by neutron irradiation (radiation annealing). Under neutron irradiation, dislocation lines in the basal planes may be completely or partially pinned depending upon the dose and temperature of irradiation. The pinning points were speculated to be interstitial atom clus­ters 4 ± 2 atoms in size,54,55 that is, the same defects clusters assumed to contribute to the reduction in thermal conductivity. The interstitial clusters are temporary barriers as they are annealed (destroyed) by further irradiation. Thus, irradiation can release dislocation lines from their original pinning site and the crystal can flow as a result of basal plane slip at a rate determined by the rate of pinning and unpinning of dislocations. Kelly and Foreman’s the­ory assumes that polycrystalline graphite consists of a single phase of true density p0 and apparent density p. The material may be divided into elementary regions in which the stress may be considered uniform and which may be identified as monocrystalline graphite. Significantly, the model excludes porosity. It is further assumed that the only deformation mode is basal plane slip for which the strain rate is determined by

exz k(sxz)f [9]

and

£yZ = k(syz)f [10]

where f is the fast neutron flux; k, the steady-state creep coefficient, and a is the stress in the given direc­tion. The microscopic deformation assumes the usual relationship between the basal plane shear strain rate (є) and the mobile dislocation density (O), and is given by

e = Obn = kaf [11]

where b is the Burger’s vector and n is the dislocation velocity as a function of the pinning point concentra­tion in the basal plane as the pins are created and destroyed by neutron flux. The dislocation line flow model used the flexible line approach where the dislo­cation line is pinned/unpinned and the dislocation line bowing is a function of the line tension and pin spacing. The concentration of pinning sites increases under irradiation from the initial value (from intrinsic defects) to a steady-state concentration. The initial creep rate is high and decreases to a steady-state value as the pinning concentration saturates at a level controlled by the neutron flux and temperature. This saturation would be expected to occur over the same dose scale as the reduction of thermal conductivity to its saturation limit (see Section 4.10.5.2).

Thus, a two stage model can be envisioned where the primary creep rate is initially high and falls to a secondary or ‘steady-state’ creep rate. The steady-state creep term should be the dominant term when the dose has reached values at which physical property changes due to dislocation pinning have saturated (see Section 4.10.5.2). Kelly and Foreman state that at higher tem­peratures the steady-state (secondary) creep rate (k)

would be expected to increase because of (1) incom­patibility of crystal strains increasing the internal stress and thus enhancing the creep rate, and (2) additional effects due to the destruction of interstitial pins by thermal diffusion of vacancies (thermal annealing as well as irradiation annealing). Kelly and Foreman53 further speculate that the nonlinearity of creep strain with stress, which is expected at higher stress levels, may also be related to the high-dose dimensional change behavior of polycrystalline graphite.56 The possibility of other dislocation and crystal deformation mechanisms being involved in irradia­tion creep must also be considered. For example, prismatic dislocations may play an enhanced role at high temperatures (>250 °C) when the graphite lat­tice is under stress, as suggested by others.57 Are there mechanisms of dislocation climb and glide that need to be explored? Can dislocation lines climb/glide past the assumed interstitial cluster barriers via a mechanism that is active only when structural rear­rangements occur during irradiation? This behavior is analogous to carbons and graphites undergoing thermal creep when they undergo structural reorga­nization, that is, during carbonization and graphitiza — tion Read More 22.08.2015   Comprehensive nuclear materials Dimensional Change As discussed previously (see Section 4.11.11.2), when irradiated, graphite crystallites, as simulated using HOPG, expand significantly in the V direction and shrink in the ‘a direction. These dimensional changes are reflected in the behavior of polycrystal­line graphite, but the volumetric changes, although relatively large are much smaller than those seen in HOPG. The reason for this is attributed to the many microcracks, which range in size from the nano — to microscale; see Figure 32. While these cracks can accommodate the crystal growth in the ‘c,’ the shrink­age in the V direction will directly be reflected in the polycrystalline behavior. However, as previously discussed, the crystallite dimensional change rate is much greater below ^300 °C than above that temperature. Read More 22.08.2015   Comprehensive nuclear materials Boiling water reactors Of the BWR plants that have been licensed for com­mercial operation in the United States, ^30% utilize either reinforced or prestressed concrete primary containments. Leak tightness of each of these con­tainments is provided by a steel liner attached to the containment inside surface by studs (e. g., Nelson studs) or by structural steel members. Exposed sur­faces of the carbon steel liner are typically painted to protect against corrosion and to facilitate decontami­nation should it be required. A portion of the liner toward the bottom of the containment and over the basemat is typically embedded in concrete to protect it from damage, abrasion, etc. due to corrosive fluids and impact. A seal to prevent the ingress of fluids is provided at the interface around the circumference of the containment where the vertical portion of the liner becomes embedded in the concrete. BWR con­tainments, because of provisions for pressure sup­pression, typically have ‘normally dry’ sections (drywell) and ‘flooded’ sections (wetwell) that are interconnected via piping or vents. Requirements for BWR containments include the following: 1. Provide an ‘essentially’ leak-tight barrier against the uncontrolled release of radioactivity to the environment for all postulated design basis acci­dent conditions. 2. Accommodate the calculated pressure and tem­perature conditions resulting from a loss-of — coolant accident. 3. Withstand periodic integrated leak-rate testing at the peak-calculated accident pressure that may be at levels up to and including the containment design pressure. 4. Permit appropriate periodic inspection of all impor­tant components and surfaces, and the periodic test­ing ofthe leak tightness ofcontainment penetrations. The containment vessel can also provide structural support for the NSSS and other internal equipment. The containment foundation, typically a basemat, provides the primary support and transfer of load to the earth below. Figure 2 presents a cross-section of a BWR Mark I reinforced concrete containment. Each of the three BWR primary plant types (Mark I, Mark II, and Mark III) incorporates a number of reinforced concrete containment internal structures. These structures may perform singular or several functions, including the following: 1. Radiation shielding; 2. human accessibility provisions; 3. NSSS and other equipment anchorage/support/ protection; 4. resistance to jet, pipe whip, and other loadings produced by emergency conditions; 5. boundary of wetwells and pool structures, and allow communication between drywell and wetwell (Mark II and III); 6. lateral stability for containment; 7. transfer of containment loads to underlying foun­dation; and 8. transfer of fuel to reactor (Mark III). As many of these functions are interrelated with the required containment functions, these structures are considered to be safety-related. Of the BWR plants that utilize steel primary containments, all but the pre-Mark plant type have reinforced concrete structures that serve as second­ary containments or reactor buildings and provide support and shielding functions for the primary con­tainment. Although the design parameters for the secondary containments of the Mark I and Mark II plants vary somewhat, the secondary containments are typically composed of beam, floor, and wall struc­tural elements. These structures typically are safety — related because they provide additional radiation shielding; provide resistance to environmental/opera — tional loadings; and house safety-related mechanical equipment, spent fuel, and the primary metal con­tainment. Although these structures may be massive in cross-section to meet shielding or load-bearing requirements, they generally have smaller elemental thicknesses than primary containments because of reduced exposure under postulated accident loadings. These structures may be maintained at a slight nega­tive pressure for collection and treatment of any air­borne radioactive material that might escape during operating conditions. Other structures include such things as founda­tions, walls, slabs, and fuel/equipment storage pools. The spent — and new-fuel storage pools, and the pools for reactor internals storage, typically have a four- wall-with-bottom-slab configuration. The walls and slab are composed of reinforced concrete members lined on the interior surface with stainless steel. Cross-sections of these members are generally large because they must support a large pool of water and heavy fuel/component loads produced by high-density fuel storage considerations. The fuel storage pool in Mark III plants is located within the primary containment. Read More 22.08.2015   Comprehensive nuclear materials Void Swelling and Bubble Swelling The progressive accumulation of high ‘cavity’ densi­ties (1012-1017cm-3) leads to a macroscopic increase in volume of the steel. The concentration of these cavities tends to increase with decreasing tempera­ture or with increasing He, H, and residual gases such as O and N. ‘Cavity’ is a generic distinction for a hole in the matrix. Identifying a specific cavity as being either a bubble or void is not as simple as might be imagined, however. In general, bubbles are rela­tively small, gas-pressurized, existing sometimes at equilibrium pressures, although not necessarily at lower temperatures where they can be significantly over-pressurized. One defining feature is that bubbles tend to grow slowly by gas accumulation while voids are either totally or partially vacuum-filled, but which are free to grow rapidly via vacancy accumu­lation without further gas addition. It is well known that bubbles can serve as nuclei for voids, accounting for the known tendency of helium especially to accelerate the onset of void swelling and to increase the cavity density. In some strongly helium-generating environments, there can also develop a late-term surge of tiny bubbles forming at very high densities in the interstices between earlier nucleated voids at much lower densities. This is a consequence of the 59Ni two-step transmutation sequence that accelerates helium production after voids are already nucleated and growing.114 As dis­cussed earlier in Section 4.02.6 these ‘helium-filled’ bubbles are probably pressurized with stored hydrogen as well as helium. Interestingly, the onset of this late — term bubble evolution does not change the steady-state swelling rate even though the cavity density increased by several orders ofmagnitude once helium generation accelerated strongly with the 59Ni sequence. For most engineering applications in nuclear sys­tems it is void swelling that is the most important contributor to dimensional instability. In the absence of physical restraint or applied stress field void swell­ing distributes its strains isotropically with the most famous published example shown in Figure 49.115 When restrained in any direction, however, the swelling-induced stresses activate irradiation creep (to be discussed later), which then redistributes the strain in the unrestrained directions, as shown earlier in Figure 16 where fuel pins locally restrained by a spirally wrapped wire evolved into spiral fuel pins. At any given altitude on the fuel pin the interaction between wire and cladding the cross-section becomes oval in shape and the resulting deformation is called ‘ovality.’ It is important to note that, contrary to popular opinion, swelling and irradiation creep are not sepa­rate processes, but are ‘two sides of the same coin.’ These phenomena are two manifestations of the radiation-enhanced dislocation motion required to move the material previously located at the void positions to the outer boundaries of the grains. This process is operating even in the absence of stress to produce swelling, but responds selectively to shear stresses generated either by externally applied or internally generated forces. While swelling attempts to be isotropic, irradiation creep redirects mass flow anisotropically. As will be shown later irradiation creep can operate before the onset of swelling but is accelerated when swelling begins. 20% CW 316 Fluences beyond FFTF goal Figure 49 Macroscopic swelling (~10% linear as measured by length change, ~33% volumetric, as measured by density change) observed in unfueled 20% cold-worked AISI 316 open cladding tube at 1.5 x 1023ncm~2 (E > 0.1 MeV) or ~75 dpa at 510°C in EBR-II. Note that in the absence of physical restraints all relative proportions were preserved. Reproduced from Straalsund, J. L.; Powell, R. W.; Chin, B. A. J. Nucl. Mater. 1982, 108-109, 299-305. Void swelling is probably the most heavily researched and published radiation-induced phe­nomenon, although pressure vessel embrittlement has also received a similar amount of attention. A comprehensive review on void swelling and irradi­ation creep was written in 1994 1 and is now being revised116 not only to incorporate new insights devel­oped over the past decade and a half, but also to revise some earlier perceptions that have not sur­vived more recent examination. A brief summary of current knowledge relevant to the purpose of this review is provided in the following sections. In some crystal systems, especially simple body — centered cubic (bcc) metals, the void swelling process is inherently self-limiting, usually saturating at some value below 5%.9 Such saturation is usually accompa­nied by a process referred to as ‘self-organization’ whereby voids arrange themselves in three­dimensional arrays that exhibit the same crystalline orientation as that of the crystal structure. Unfortu­nately, for most face-centered cubic (fcc) metals, espe­cially stainless steels, self-organization and saturation of void swelling do not operate under most reactor­relevant conditions, and as a result swelling in austen­itic stainless steels is an inherently unsaturable process. Void swelling normally exhibits a transient or incu­bation regime where either no swelling or very slow swelling occurs before swelling moves to a steady-state rate. Tens of percent swelling have been reached in a number of reactor-relevant irradiation histories, and values of 80-90% swelling without hint of impending saturation have been attained in both model and com­mercial alloys during neutron irradiation.1,117,118 Swelling in excess of 200% was reached during proton irradiation of 316 stainless steel and saturation was eventually observed at ^260% swelling.119 An example of apparently nonsaturable void swelling in AISI 316 is presented in Figure 50.117 Note that the onset of rapid swelling, defined by termination of a ‘transient’ regime, is dependent on both irradiation temperature and dpa rate. The dpa rate dependence of the transient is not easily dis­cerned in Figure 50 where each irradiation temper­ature in this experiment is coupled with a specific dpa rate, with the range of dpa rates increasing ^65% from lowest to highest. It will be shown later that dpa rate is a very strong determinant of void swelling. The transient regime is terminated when the condi­tions for both void nucleation and especially rapid void growth have been attained. The conditions for void nucleation must be favor­able to end the transient. This usually requires Figure 50 Swelling determined by density change as a function of irradiation temperature and dose, as observed in 20% cold-worked AISI 316 irradiated in the EBR-II fast reactor. Reproduced from Garner, F. A.; Gelles, D. S. In Proceedings of Symposium on Effects of Radiation on Materials: 14th International Symposium; ASTM STP 1046; 1990; Vol. II, pp 673-683. All measurements at a given temperature were made on the same specimen after multiple exposures with subsequent reinsertion into the reactor. This procedure minimized specimen-to-specimen data scatter and assisted in a clear visualization of the posttransient swelling rate. attainment of a dislocation network to the quasi­equilibrium value of ^3 x 10cm~ , either by reduction of higher cold-worked densities or build up from lower densities characteristic of annealed alloys.1 It also requires that the temperatures be low enough to guarantee sufficient supersaturation of vacancies or that elements (P, Si, Ni) that strongly increase the effective vacancy diffusion coefficient, and thereby depress void nucleation, be low enough or have been reduced via precipitation. Helium and other gases influence void nucleation and under some situations where nucleation is difficult can serve to shorten the transient duration. Rapid void growth after sufficient nucleation of voids requires not only the attainment of the quasi-equilibrium dislocation density, but also that dislocation network be a ‘glissle’ network capable of moving mass quickly. Voids previously nucleated but still embedded in a ‘sessile’ microstructure composed primarily of Frank loops can grow but not quickly. Therefore, significant unfaulting of Frank loops is a prerequisite for termination of the transient and the onset of the high swelling rate. As also shown in Figure 50, the terminal post­transient swelling rate of AISI 316 is typical of all austenitic stainless steel at ~1% per dpa, essentially independent of all irradiation or material vari — ables.1,120 This terminal rate also appears to be char­acteristic of Fe-Cr-Mn, Fe-Cr-Mn-Ni, and simple Ni-base alloys, although for the nickel-base %’/%» stabilized alloys the transients are generally much longer and insufficient amounts of swelling were attained in most studies to allow confirmation of the full generality of this statement of a universal steady — state swelling rate for all fcc alloys.1,121,122 In Fe-Cr-Mn and Fe-Cr-Mn-Ni alloys removal of highly diffusing Mn from voids and grain bound­aries via the inverse Kirkendall effect leads to these sinks becoming coated with lower-swelling ferrite phase, thereby producing a late-term decrease in the average swelling rate.1 , Read More Календарь Ноябрь 2024 Пн Вт Ср Чт Пт Сб Вс   1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30   Рубрики 1 BIOFUELS A Look Back at the U. S. Department of Energy’s Aquatic Species Program: Biodiesel from Algae A. Worrall ABOUT THE AUTHOR ACCELERATOR DRIVEN SUBCRITICAL REACTORS ADVANCED BIOFUELS Advanced Biofuels and Bioproducts Advanced separation techniques for nuclear fuel reprocessing and radioactive waste treatment ADVANCES IN Advances in Biochemical Engineering/Biotechnology Advances in Biofuels Advances in Biorefineries Alcoholic Fuels Algae Energy All alternative energy Alternative transportation An Indispensable Truth An Introduction to Nuclear Materials AN INTRODUCTION. TO THE ENGINEERING. OF FAST NUCLEAR REACTORS ANNEX III. AHWR. Bhabha Atomic Research Centre, India ATOMIC ACCIDENTS BACKGROUND Bioenergy Bioenergy — Realizing the Potential Bioenergy from Wood BIOENERGY. RESEARCH:. ADVANCES AND. APPLICATIONS BIOETHANOL BIOFUEL'S ENGINEERING PROCESS TECHNOLOGY BIOFUELS BIOFUELS 1 Biofuels and Bioenergy Biofuels for Road Transport Biofuels from Agricultural Wastes and Byproducts BIOFUELS FROM ALGAE Biofuels Refining and Performance BIOGAS BIOGAS 1 BIOMASS — DETECTION, PRODUCTION AND USAGE Biomass and Biofuels from Microalgae Biomass Conversion Biomass Gasification and Pyrolysis BIOMASS NOW — CULTIVATION AND UTILIZATION BIOMASS NOW — SUSTAINABLE GROWTH AND USE Biomass Recalcitrance Biomasses identities Biotechnological Applications of Hemicellulosic Derived Sugars: State-of-the-Art Biotechnological Applications of Microalgae Biotechnology. for Fuels and Chemicals Cellulosic Energy Cropping Systems Comprehensive nuclear materials DESIGN FEATURES TO ACHIEVE. DEFENCE IN DEPTH IN SMALL AND. MEDIUM SIZED REACTORS Design of Reactor Containment Systems for Nuclear Power Plants ENERGY Energy Efficiency ENERGY SERVICES IN THE HINDU KUSH HIMALAYAN REGION Energy Storage Estimating Loss-of-Coolant Accident (LOCA) Frequencies Through the Elicitation Process EUROSUN 2008 EuroSun2004 EuroSun2008-10 EuroSun2008-11 EuroSun2008-13 EuroSun2008-2 EuroSun2008-3 EuroSun2008-4 EuroSun2008-5 EuroSun2008-7 EuroSun2008-9 EXAMPLES OF REACTIVITY-CONTROL SYSTEMS Exploitation des creurs REP Fast Breeder Reactor Programs: History and Status Fast Reactor Safety. (Nuclear science. and technology) Fertilization Frequency Response Testing. in Nuclear Reactors Fuels and Chemicals. from Biomass Gebaudetypologie Bayern Geothermal Energy GREEN BIORENEWABLE. BIOCOMPOSITES Handbook Nuclear Terms Handbook of biofuels production Handbook of Small Modular Nuclear Reactors Hydrogen — Fuel Cells Hydropower IAEA RADIATION TECHNOLOGY SERIES No Infrastructure and methodologies for the. justification of nuclear power programmes Innovative Biofibers from Renewable Resources Integral design concepts of advanced water cooled reactors Introduction to Nuclear Power Liquid Biofuels: Emergence, Development and LIQUID, GASEOUS AND SOLID BIOFUELS — CONVERSION TECHNIQUES Management von Biogas-Projekten Materials' ageing and degradation in. light water reactors Microbes and biochemistry of gas fermentation Modern Power Station Practice Natural circulation data and methods for advanced water cooled nuclear power plant designs Natural gas Neutron Scattering Applications and Techniques Nuclear and Radiochemistry Nuclear Back-end and Transmutation Technology for Waste Disposal NUCLEAR CHEMICAL ENGINEERING NUCLEAR ELECTRIC POWER Nuclear energy Nuclear Power and the Environment Nuclear power plant life management processes: Guidelines and practices for heavy water reactors NUCLEAR POWER PLANTS NUCLEAR POWER the PUBLIC Nuclear Reactor Design NUCLEAR REACTOR ENGINEERING Nuclear Reactors 1 NUCLEAR REACTORS 2 Others Particle Image Velocimetry (PIV) Passive Safety Systems and Natural Circulation in Water Cooled Nuclear Power Plants PHYSICS OF. HIGH-TEMPERATURE. REACTORS POWER Pretreatment Techniques for Biofuels and Biorefineries Principles of Fusion Energy PROCESS SYNTHESIS. FOR FUEL ETHANOL. PRODUCTION Production of Biofuels and Chemicals with Ionic Liquids Progress, Challenges, and Opportunities for. Converting U. S. and Russian Research Reactors Pumping Pyrolysis Radioactive waste management and contaminated site clean-up Recycling of Biomass Ashes Renewable Energy Products Rudiger Meiswinkel, Julian Meyer, Jurgen Schnell Second Generation Biofuels and Biomass Small modular reactors (SMRs) the case of Russia solar energy Solar Thermal and. Biomass Energy Sonar-Collecttors SonSolar Springer Series in Materials Science Study on Neutron Spectrum of Pulsed Neutron Reactor Sustainable Biotechnology Switchgrass Technologies for Converting Biomass to Useful Energy The Asian Biomass Handbook The Experimental Analyze Of The Solar Energy Collector The Future of Nuclear Power Thermal biomass. conversion and utilization —. Biomass information. system Thoria-based Nuclear Fuels Type of Membrane Contactor Suitable for CO2-O2 Exchanges in Culture Broth Vereinfachtes Verfahren WAS INTENTIONALLY LEFT BLANK Why We Need Nuclear Power wind power WORKSHOP ON NUCLEAR REACTION DATA AND. NUCLEAR REACTORS:. PHYSICS, DESIGN AND SAFETY Автономное энергообеспечение Альтерантивная энергия Альтернативные источники энергии и энергосбережение Без рубрики Биотопливо Ветровая энергия Ветрогенераторы, солнечные батареи и другие полезные конструк­ции Возобновляемые источники энергии ВОПРОСЫ ТЕОРИИ. И ИННОВАЦИОННЫХ РЕШЕНИЙ. ПРИ ИСПОЛЬЗОВАНИИ. ГЕЛИОЭНЕРГЕТИЧЕСКИХ СИСТЕМ ГЕЛИОЭНЕРГЕТИКА Другая энергия Зарядные станции Источники питания Источники энергии ИСУ Концентраторы солнечного излучения КРЕМНИЕВЫЕ СОЛНЕЧНЫЕ БАТАРЕИ Получение кремниевых пластин для солнечной энергетики Применение солнечной энергии Разное Солнечная электроэнергия Солнечная энергия Солнечная энергия — использование Солнечные батареи Солнечные коллекторы Солнечный ДОМ ТОНКОПЛЕНОЧНЫЕ. СОЛНЕЧНЫЕ ЭЛЕМЕНТЫ. НА ОСНОВЕ КРЕМНИЯ ФОТОЭЛЕКТРИЧЕСКОЕ. ПРЕОБРАЗОВАНИЕ. СОЛНЕЧНОЙ ЭНЕРГИИ Экологическая техника Электромобили Электроэнергия Элементов Энергетика Энергия Последние записи Как выбрать гостиницу для кошек 14 декабря, 2021 Сегодня каждый, кто собирается в отпуск и не знает, с кем оставить своего котика или кошку, может во[...] Грузы для тентованных грузовиков в Челябинске 1 декабря, 2021 Оказавшись в час пик вблизи любой оживленной автомагистрали, вы увидите большое количество тентованн[...] Нихромовая проволока: пару слов об отличительных особенностях 21 сентября, 2021 На современном рынке предложено много видов проволоки и одним из самых востребованных считается пр[...] Заявки на доставку груза фургонами и термобудками по всей России 17 сентября, 2021 Рядом с супермаркетами, торговыми центрами и небольшими магазинами шаговой доступности можно увидеть[...] Информация о характеристиках газовых резаков 7 сентября, 2021 Основное предназначение устройств газовой резки металломатериалов - смешение горючих элементов с кис[...] Свежие записи Как выбрать гостиницу для кошек Грузы для тентованных грузовиков в Челябинске Нихромовая проволока: пару слов об отличительных особенностях Заявки на доставку груза фургонами и термобудками по всей России Информация о характеристиках газовых резаков © 2024 Солнечная и другая альтернативная энергия Powered by WordPress | Designed by: Wordpress Themes | Thanks to Solar Energy