Abstract
Failure of high level waste/spent fuel containers is most likely to occur as a result of mechanical overload following a period of corrosion that results in a reduction of the wall thickness and/or the degradation of the material properties. There can be significant interaction between the mechanical loads and the corrosion processes to which a disposal container is subjected which, in turn, can influence the mode and time of failure. Here, these interactions are illustrated for a single shell, carbon steel spent fuel container during its entire life cycle, from the time of manufacture through to the long term behaviour in a bentonite backfilled geological disposal facility. The evolution of the structural integrity of the container is illustrated through the use of failure assessment diagrams.
Keywords
Introduction
In 2008, the UK Government published the Managing Radioactive Waste Safely White Paper, 1 identifying the Nuclear Decommissioning Authority (NDA) as the implementer for the geological disposal for higher activity radioactive wastes in the country. The NDA has been delivering this commitment through its Radioactive Waste Management Directorate (RWMD). In April 2014, the RWMD has become a wholly owned subsidiary of the NDA, Radioactive Waste Management Limited (RWM). Building on work previously carried out by the RWMD, RWM continues a programme of research and development (R&D) to support the implementation of a geological disposal facility (GDF), which is described in their Technical Programme (Part C). 2 As part of this R&D programme, work considering candidate materials for containers for high level waste (HLW) and spent fuel (SF) is ongoing.
Prior to identifying a potential host site or defining a specific GDF design, generic studies are being performed to develop, among other things, an understanding of the likely corrosion behaviour and resulting durability of waste containers. The corrosion behaviour of candidate container materials during the post-closure period has been reviewed and the impacts of the operational phase considered on the basis of information available from other disposal programmes and the general scientific literature.3–5 Continuing work is considering the potential for interactions between specific UK wastes as well the need for developing tools to support the treatment of container failure in the safety case. The latter aspects, which are the focus of this paper, aim at taking into account the potential effect of both mechanical and corrosion degradation modes, and their interaction, on the integrity of the container during all stages from manufacture, interim storage, transportation, and ultimate disposal in a GDF.
Traditionally, the lifetimes of HLW/SF containers have been based largely on the corrosion behaviour of the container material, with the implicit assumption that failure due to mechanical failure mechanisms can be avoided through prudent design. Separate ‘corrosion allowances’ and ‘mechanical allowances’ are often defined. However, container failure will likely result from mechanical overload following degradation of the container by one or more corrosion processes resulting in a reduction in wall thickness and/or mechanical properties of the container material. For certain combinations of container material and exposure environment there can be significant interaction between the corrosion behaviour and the structural integrity. Therefore, it can be important to take into account both mechanical and corrosion degradation mechanisms, and their interaction, when predicting the lifetimes of HLW/SF containers.
The failure assessment diagram (FAD) is a useful method for illustrating the proximity of components containing defects to failure. 6 Typically, failure is assumed to result from either plastic collapse or brittle fracture, which are assessed separately. The proximity to failure by plastic collapse is determined from the ratio (L r) of the reference stress for the defect and a threshold stress, generally taken to be the flow stress of the material (i.e. the uniaxial true stress at the onset of plastic deformation). The proximity to brittle fracture (in the case of a linear elastic fracture mechanics assessment) is determined by the ratio (Kr) of the stress intensity factor for the defect and the fracture toughness of the material. In an FAD, a ‘failure curve’, separating regions on the FAD for which failure of the component is unlikely (below the curve) and for which failure is possible (above the curve), is identified on the basis of specific material related information derived from mechanical testing. Proximity to failure of a stressed component is then assessed on the basis of the proximity of the (L r, K r) ‘point’ to the failure curve. In the current application, the assessment point for assumed defects is calculated for various stages during the life cycle of the container, and a trajectory (or locus) of assessment points plotted on the FAD to illustrate the progression of the structural integrity with time.
Here, the consequences of interactions between corrosion and mechanical degradation mechanisms are illustrated for a single shell, carbon steel (C-steel) spent fuel (SF) container from the point of manufacture, through a period of interim storage, handling, and finally disposal in a bentonite backfilled GDF. The FAD approach is used to illustrate the progression in the structural integrity of the container for a hypothetical sequence of environmental and mechanical loading conditions, including a future hypothetical glaciation event. A summary of the main findings of this study is reported here but a more complete analysis can be found elsewhere. 7
Assessment scenario
Container and assumed defects
Figure 1 shows the advanced gas cooled reactor (AGR) spent fuel Variant 2 standardised container design considered in the current analysis. 8 The container is capable of holding fuel pins from 48 dismantled AGR fuel assemblies, consolidated in 16 stainless steel ‘slotted cans’ (3 assemblies per slotted can) and housed in four receptacle tubes. The void volume inside the container is approximately 1·7 m3.
Variant 2 (carbon steel) ‘standardised’ disposal container for AGR spent fuel 8
The container is fabricated from S275 grade C-steel with a nominal wall thickness of 120 mm and a minimum thickness for corrosion purposes of 75 mm (the minimum wall thickness, based on the design of the lid assembly). For the purposes of this study, the room temperature yield strength of the container material is taken to be 225 MPa and the Young's modulus 200 GPa. The fracture toughness of the parent material, heat-affected zone, and weld metal are assumed to be equal and are taken to be 220 MPa m1/2 (prior to effects due to hydrogen absorption)
Various defects are considered for the FAD analysis (Fig. 2). Internal and external semi-elliptical flaws with an initial depth of 10 mm are assumed to be present in the final closure weld (Fig. 2a and b). A similar sized defect is also assumed to be present on the inner surface of the base at a location previously shown to be highly stressed in the load configuration assumed (Fig. 2c). 8 Sensitivity of the results of this analysis to defect sizes up to 30 mm are presented in the more detailed analysis report 7 and are not shown here.
Summary of flaws assumed for structural integrity assessment (depth of all defects was assumed to be 10 mm in reference case)
Mechanical loading and environmental conditions
A hypothetical scenario involving a range of environmental and mechanical loading conditions has been defined to demonstrate the potential importance of considering combined corrosion and mechanical degradation mechanisms and to illustrate the application of the FAD approach. The scenario addresses all phases of the container lifecycle, including a period of interim storage immediately following loading and sealing of the container, transport of the container, and finally disposal in a GDF. The following conditions and features are considered in the assessment:
container sealed in 2025 with an air atmosphere (no inert gas used during sealing), with entrained water in some spent fuel elements due to perforation of the cladding during previous pond storage operations (assumed to amount to 1400 g H2O per disposal container) 9
consumption of O2 and water by internal corrosion, the latter resulting in H2 generation
a storage period of 50 years with institutional control for 90% of the time (assumed loss of control of temperature and relative humidity control for 1 year each decade)
loss of wall thickness due to internal and external corrosion during the storage period.
waste package emplacement in a GDF at 2075 with immediate emplacement of bentonite buffer and bentonite based tunnel backfill
vertical borehole emplacement, with 6·5 m borehole spacing and 25 m between disposal tunnels
maximum buffer swelling pressure of 7 MPa
GDF depth of approximately 650 m resulting in a hydrostatic load (in the absence of an ice sheet) of 6·5 MPa
saline ground water (approximately seawater salinity)
uneven saturation of the bentonite or variation in the buffer density, resulting in a period of asymmetric swelling pressure 8
peak container surface temperature of 90°C reached 20 y after disposal in the GDF, declining slowly to a temperature of 46°C after 10 000 y
continued loss of wall thickness after disposal, as a result of first aerobic corrosion and then anaerobic corrosion of the outer surface of the container
degradation of fracture properties of C-steel as a result of hydrogen absorption under anaerobic conditions
diffusion of hydrogen through the container wall as a result of the H2 concentration gradient between internal and external environments
assumed hypothetical glaciation event, with a peak ice sheet thickness of 1800 m at 55 000 years post-disposal
Based on this hypothetical scenario, the following timeline in the container lifecycle has been defined at which the structural integrity of the container has been assessed:
2025 – start of the storage period, at which point a portion of the entrained water has evaporated and internally pressurised the container, but no internal corrosion has occurred
2075 – immediately after emplacement of the waste package in the GDF, following the 50 year interim storage period and the associated external and internal corrosion and H2 generation
2095 – end of the aerobic corrosion phase in the GDF and the associated external wall loss, which is assumed to occur under largely unsaturated conditions (i.e. no external load at this time). Anaerobic corrosion, with the associated generation of H2, begins
2175 – 100 years post-closure and the time at which the GDF is assumed to be almost completely saturated, but with an asymmetric buffer swelling pressure applied to the container, in addition to the uniform hydrostatic pressure. The absorption of hydrogen by the container will lead to a reduction in the material fracture properties
2225 – 150 years post-closure, at which time it is assumed that the bentonite swelling pressure is uniform
3075 – 1000 years post-closure, after which time there will be additional wall thinning as a result of corrosion, an increase in the internal H2 partial pressure due to diffusion across the container wall, resulting in a decrease in the pressure differential across the container wall, and the container temperature will have dropped to about 70°C
12 075 – 10 000 years post closure, corresponding to additional wall loss and further increase in internal H2 partial pressure, and a continuing reduction in the container temperature to 46°C
37 075 – 35 000 years post closure, corresponding to the time at which approximately half of the original minimum wall thickness in the thinnest section of the container (75 mm thick) has been lost by corrosion
57 075 – 55 000 years post-closure, corresponding to the assumed time of the peak ice sheet thickness. Glaciation results in an additional external hydrostatic load and external H2 pressure, resulting in a significant increase in load and decrease in the fracture toughness of the container material.
The time dependence of the total external and internal pressures, the pressure difference across the container wall, and of the individual contributors to the external load are shown in Fig. 3.
Illustrative time dependent internal and external loading conditions during various stages of life cycle for carbon steel spent fuel container
Expected evolution of disposal system
Based on this set of environmental conditions, the container is considered to be susceptible to both general corrosion and, particularly during the post-closure period, hydrogen induced cracking.2,4,7 Because of the nature of the container material and assumed environmental conditions, pitting and stress corrosion cracking of the container are not expected to occur.2,4,7
During the operational period, the water entrained in the perforated fuel pins will lead to corrosion of the C-steel components inside the container as well as of the inner surface of the container wall. In this phase, external corrosion will occur due to the assumed periodic loss of environmental control during the period of interim storage. An atmospheric corrosion rate of 10 μm/year is assumed for the combined total of 5 year for which environmental control is lost. Internal corrosion, initially driven by aerobic processes (i.e. oxygen reduction) and, following the consumption of oxygen, by anaerobic processes (i.e. water reduction) will also lead to general corrosion (the potential for hydrogen embrittlement at this stage is considered low because of the low H2 partial pressure inside the container).
Following emplacement of the container in the GDF in contact with the bentonite buffer, general corrosion of the container will occur first under aerobic conditions (the initial inventory of O2 trapped in the pores of the buffer material is assumed to be equivalent to a uniform depth of corrosion of 200 μm which is consumed in a period of 20 year), and subsequently under anaerobic conditions (at an assumed rate of 1 μm/year). The H2 produced externally by anaerobic corrosion (and, to a smaller extent, internally) will result in the possibility of HIC. In particular, because of the presence of compacted bentonite, a gaseous H2 phase will develop at the external surface of the container, as the rate of diffusion of dissolved H2 through saturated bentonite is slower than the rate at which it is produced by anaerobic corrosion. The dependence of the fracture toughness on the H2 partial pressure (Fig. 4) 10 is taken into account in the assessment. Given the low yield strength of the selected container material, the susceptibility to HIC is low3,4,11 but, nevertheless, needs to be considered given the assumed presence of external and internal crack-like defects.
Reduction in relative fracture toughness of base material and welded carbon steels as result of exposure to hydrogen gas at elevated pressure, based on data from Ref. 10: best fit line shows trend used for assessment of effect of absorbed hydrogen on fracture toughness
There are various sources of mechanical load during the container lifecycle (Fig. 3). First, internal corrosion and/or radiolysis of the entrained water will result in the development of an internal H2 atmosphere during interim storage. (There will also be a component from any He released from fuel pins whose cladding may rupture during storage, which is not taken into account here).
Immediately following emplacement of the container in the GDF and during the period of aerobic corrosion there will be little change in the internal or external load, which is assumed to occur during the subsequent buffer saturation phase (left hand pane of Fig. 3). Starting in the year 2095, a hydrostatic and asymmetric buffer swelling pressure (minimum of 5·6 MPa and maximum of 8·4 MPa applied to one ‘side’ of the container, with all other surfaces experiencing a swelling pressure of 7 MPa) are generated outside of the container, reaching a peak external load of 14·9 MPa at 2175, after which the buffer swelling pressure becomes more uniform, resulting in an isotropic external load of 13·5 MPa in the year 2225. From the year 2095 onwards, anaerobic corrosion will produce diffusible hydrogen atoms, some entering the microstructure of the steel leading to its embrittlement and eventually combining as molecular hydrogen inside the void space in the container (leading to an increase in internal H2 partial pressure), others combining as molecular hydrogen directly outside the container, resulting in an external H2 atmosphere. The diffusion of hydrogen from the outside to the inside of the container (associated with the larger H2 partial pressures expected outside) will result in a slow increase in the internal H2 partial pressure up to the year 42 075 (right hand pane of Fig. 3). The time dependence of the internal H2 pressure depends on the diffusivity of hydrogen in the steel, the external H2 partial pressure, and the internal void volume of the container. As the internal H2 pressure increases, reaching a value of 5·2 MPa in the year 42 075, the pressure difference across the container wall decreases, thus mitigating some of the effect of the external load.
Starting in the year 42 075, an ice sheet is assumed to start forming during a hypothetical glaciation event lasting until the year 60 000, and during which the peak ice sheet thickness (assumed 1800 m) is achieved in the year 57 075. (A more recent assessment of the likely future glaciation events in the UK indicates that no ice sheet is expected to form for the next 170 000 y. 12 The earlier glaciation event assumed here is purely for illustrative purposes for the analysis). Because a continuous water column may exist from the surface of the ice sheet down to the depth of the GDF, the hydrostatic pressure increases as the ice sheet thickens. As well as contributing to an increase in the external load, this increased hydrostatic pressure also results in a higher H2 partial pressure at the surface of the container, a corresponding decrease in the fracture toughness of the container material (Fig. 4), and an increase in the flux of hydrogen through the wall [which partially offsets the increase in external load (Fig. 3)]. Starting in 57 075, the ice sheet is assumed to begin to recede resulting in a decrease from the maximum external load of 31·5 MPa.
Mechanical and corrosion degradation processes and their interaction
Figure 5 illustrates the interactions between the various mechanical loads and corrosion processes and how these jointly determine the container failure modes. In terms of the effect of mechanical loading on the corrosion processes (Fig. 5a), clearly any internal or external load that leads to an increase in tensile stress on the inner or outer surfaces of the container will increase the probability of HIC. The predominant role of internal and external corrosion, conversely, is to reduce the load bearing wall thickness and, for anaerobic corrosion, to reduce the fracture toughness of the material through the absorption of hydrogen. In terms of the impact on the container failure modes (Fig. 5b), all mechanical loads promote both plastic collapse (ductile overload) and brittle fracture through the increase in stress. Similarly, corrosion of the container wall promotes both failure modes by either reducing the wall thickness (and, hence, increasing the stress) or embrittling the container material (i.e. reducing the fracture toughness). The propensity to brittle fracture is directly related to the susceptibility to HIC.
a interactions between mechanical loads and corrosion processes; b effect of interactions on container failure modes
Results and discussion
Internal corrosion and pressurisation due to entrained water
Even though the amount of water assumed to be entrained inside the disposal container (1400 g) is considered to be a relatively conservative estimate (since it assumes no drying of spent fuel stored in water ponds before containerisation), the extent of internal corrosion and pressurisation of the container is expected to be small because of the large internal surface area and the large void volume. 9 A computational fluid dynamics (CFD) model was used to evaluate the distribution of internal temperatures, liquid water and water vapour. The temperature distributions at 2025 are shown in Fig. 6. Owing to natural convection, the maximum internal temperature (approximately 116°C) is predicted to be found at the top of the uppermost slotted cans (not shown). All other surfaces inside the container are expected to be below the saturation temperature and, hence, will be wetted and subject to corrosion. Immediately after sealing, the internal pressure is estimated to rise to 0·16 MPa, with about 290 g of the entrained water present as water vapour and the rest as liquid droplets or a pool of water in the bottom of the container. By the time the container is assumed to be emplaced in the GDF in 2075, the maximum temperature has decreased to 94°C.
Contours of temperature distribution for AGR spent fuel container immediately after sealing in 2025
Corrosion of internal C-steel components and of the inside of the container wall are predicted to consume the initially entrained water within a period of 1-10 year, depending upon the assumed corrosion rate. 9 Thus, the initial humid air atmosphere is replaced by a H2/N2 atmosphere (the atmospheric O2 initially present in the container having been consumed in the formation of iron corrosion products) with a maximum internal pressure at the end of the 50 year storage period of 0·23 MPa. The mean depth of corrosion is estimated to be 11 μm if uniformly distributed over all internal wetted surfaces and still well below any amount of concern to structural integrity even if assumed to localise in specific areas (e.g. at the bottom of the container). Even if corrosion is limited to a single receptacle tube containing the perforated fuel pins (i.e. where entrained water is expected to be present initially), the maximum depth of corrosion is predicted to be only 0·14 mm.
Therefore, neither the internal pressurisation nor the internal corrosion has a significant impact on the structural integrity of the container. 9
Effect of defect type and location on container integrity
Figure 7 shows the results of the assessment of the integrity of the three defects shown in Fig. 2 for 10 mm deep flaws plotted on a FAD. Each locus represents a series of assessment points for different times starting from the time of container sealing and up to the time of the peak ice sheet thickness. The trajectory of the locus of assessment points indicates both the proximity to failure as well as whether failure is more likely to occur by plastic collapse (as indicated by a horizontal trajectory), by brittle fracture (a vertical trajectory), or by some combination of the two.
Compilation of assessment points for 10 mm semi-elliptical defects on interior and exterior of closure weld and in base of C-steel container: 7 times shown for individual data points are number of years post-emplacement, ranging from −50 year at start of period of interim storage to 55 000 year corresponding to time of assumed glaciation event
The semi-elliptical surface breaking internal and externals flaws in the closure weld show a tendency towards failure by plastic collapse (Fig. 7). Prior to the onset of the assumed glaciation, the assessment points for these two defects are well within the FAD boundary curve. The depth of the external defect is assumed to decrease as the container corrodes in the GDF so that the locus of assessment points coincides with the L r axis at times greater than 10 000 year after emplacement corresponding to the point at which the defect is completely removed and there is no longer any likelihood of brittle fracture. For the internal semi-elliptical defect, the assessment point does not change very much over the first 10 000 year, as the increase in global stress resulting from the decrease in wall thickness is offset by the decrease in differential pressure due to the diffusion of hydrogen into the container and the development of an internal H2 pressure. Only after the wall thickness has been significantly reduced by corrosion does the assessment point move closer to plastic collapse. For this particular flaw location, the increased load and degraded fracture toughness do not result in a significant probability of brittle fracture, even in the presence of defects up to 30 mm in depth (not shown).
In contrast to the semi-elliptical weld defects, the internal flaw in the base of the container does exhibit some susceptibility to brittle fracture in the load configuration assumed. The locus of assessment points suggests possible failure of the container through a mixed brittle/plastic failure mode at the time of the peak glaciation (assumed here to occur approximately 55,000 y post-closure). Further analyses (not shown) suggest that, up to times of 35,000 y (i.e., after approximately 35 mm of the minimum 120 mm wall thickness has been lost by corrosion but before the high external loads due to glaciation) internal defects in the base of the container up to 30 mm present no threat to the integrity of the container.
Uncoupled analysis
For comparison, a more traditional analysis of the container integrity has been carried out in which mechanical- and corrosion-related failure modes are considered separately. 7 For the maximum external load of 31·5 MPa, a minimum wall thickness of ∼22 mm is required to prevent mechanical overload, leaving the remaining 53 mm (of the minimum 75 mm wall thickness) as the corrosion allowance. In this uncoupled approach, therefore, container failure is predicted to occur at approximately 50 000 years, as a result of plastic collapse of the container once the corrosion allowance has been consumed. Although, in the specific scenario considered in this study, the predicted container failure time is similar to that predicted by the coupled FAD approach for the various defects considered, the mode of failure is different, since a mixed failure mode, rather than plastic collapse, would be expected based on the FAD assessment for the internal defect in the base of the container. The coupled assessment also provides additional insight into the time dependence of container integrity.
Conclusions
The combined effects of mechanical and corrosion degradation mechanisms on the integrity of a spent fuel disposal container have been considered and illustrated using the failure assessment diagram methodology. The analysis has been performed for a single shell, C-steel spent fuel disposal container assumed to undergo a period of interim storage before emplacement in a geological disposal facility and backfilling with compacted bentonite.
For the material and design considered, in certain combinations of environmental and mechanical loading conditions, there can be significant interactions between corrosion and mechanical processes that need to be considered in evaluating likely failure modes and resulting durability of waste containers. Significant interactions include the reduction in wall thickness due to corrosion and the reduction in toughness of the container material due to absorbed hydrogen. A relatively long lifetime (>35 000 years) was estimated in the specific scenario analysed here for all defects with an assumed depth of 10 mm.
The results of such analyses are dependent on the detailed container design, the nature of the initial defects, the environmental and mechanical loading conditions, and the associated corrosion processes. The results for other container and GDF designs may well differ from those for the illustrative example presented here.
Footnotes
Acknowledgement
This work was funded by the UK Nuclear Decommissioning Authority under contract RWM005167.
*
The external H2 pressure will be governed by the gas breakthrough pressure for compacted bentonite, which is approximately equal to the sum of the buffer swelling pressure and the hydrostatic load.