The effect of salt loading on chloride-induced stress corrosion cracking of 304L austenitic stainless steel under atmospheric conditions

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The effect of salt loading on chloride-induced stress corrosion cracking of 304L austenitic stainless steel under atmospheric conditions G.G. Scatigno , P. Dong , M.P. Ryan , M.R. Wenman PII: DOI: Reference:

S2589-1529(19)30305-9 https://doi.org/10.1016/j.mtla.2019.100509 MTLA 100509

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Materialia

Received date: Accepted date:

13 August 2019 14 October 2019

Please cite this article as: G.G. Scatigno , P. Dong , M.P. Ryan , M.R. Wenman , The effect of salt loading on chloride-induced stress corrosion cracking of 304L austenitic stainless steel under atmospheric conditions, Materialia (2019), doi: https://doi.org/10.1016/j.mtla.2019.100509

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The effect of salt loading on chloride-induced stress corrosion cracking of 304L austenitic stainless steel under atmospheric conditions G.G. Scatignoab, P. Donga, M.P. Ryanc, M.R. Wenmana* a

Department of Materials and Centre for Nuclear Engineering, Imperial College London, London SW7 2AZ, UK EDF Energy, Barnwood, Gloucester, GL4 3RS, UK c Department of Materials and London Centre for Nanotechnology, Imperial College London, London SW7 2AZ, UK b

* Corresponding author. Tel.: +44 (0) 207 594 6763. E-mail address: [email protected] (M.R. Wenman)

Abstract The effect of salt loading on chloride-induced stress corrosion cracking in 304L was studied at atmospheric pressure. Stress relieved samples were uniaxially pre-strained to 5% and were loaded with nine levels of MgCl2, investigating Cl- deposition levels from 1.7 x 10-4 to 3.1 x 10-2 g.cm-2. Samples were subject to 60 MPa stress, 90°C at 70% relative humidity, for 480 hours. A direct correlation between chloride deposition and the density of cracking and corrosion was observed between 5.7 x 10-4 and 1.96 x 10-2 g.cm-2. Crack propagation rates were constant between salt loadings of 5.7 x 10-4 and 2.4 x 10-2 g.cm-2 at 1-2 μm.hr-1. Keywords Austenitic stainless steels; Corrosion; Stress-corrosion cracking; Chlorides 1. Introduction The properties of austenitic stainless steels (ASS) such as corrosion resistance, manufacturability, ductility, abundant industrial experience, and low cost, make ASS a favourable alloy choice for many industries. However, ASS can fail if conditions exist that lead to stress corrosion cracking (SCC) [1,2]. SCC takes place when a combination of tensile stress (residual or applied), susceptible material composition and a specific environmental chemistry, are present [2–5]. Different alloy systems are susceptible to different chemical species [2,4,6,7] that are not predictable a priori. SCC cracks are particularly detrimental as they produce a pseudo-brittle failure of otherwise extremely ductile alloys [8–10]. 304L ASS, one of the most widely used grades of steel, is especially susceptible to SCC in the presence of salts containing Cl- ions [10–13]. Within the nuclear industry, 304L is used as piping material, for the cooling circuit in the nuclear power plant (NPP), and as a material for canisters used for interim storage of spent nuclear fuel (SNF). NPP and interim storage facilities are often located in coastal areas and thus exposed to sea-salt deposition under atmospheric conditions [2,5,10,11,14–17]. Sea-salts contain a mixture of chloride and sulfate salts [18]. Although the main component of sea-salt is NaCl which has a deliquescence relative humidity (RH) of around 75%, the presence of highly hygroscopic salts such as MgCl2 (deliquescence RH of 33%) means that sea-salt particles start wetting much earlier [19–22]. Currently only a third of SNF is reprocessed and with a worldwide lack of permanent disposal facilities, interim storage of SNF using dry casks has become the primary option in the near term [14,23,24]; their use is expected to increase. The cooling ponds at many reactor sites are either at, or near capacity, thus dry casks will have to be implemented in the very near future [25]. The dry casks will have to survive in interim storage for at least 40-60 years (current UK forecast are 100 years) [5,14,15]. As any cracking takes place on the outer surface of the component, this form of SCC is termed atmospheric [5,10]. Because of the relatively long-term use of the SNF storage, it is necessary to address all the issues that may affect the integrity of the casks. Microstructure, surface finish, residual stresses, chemistry and salt loading are all key factors that may affect the durability of the dry cask [3,10,14,26–32]. Dry casks containing SNF are often placed within a building, which uses natural air circulation to passively cool them. In the USA, outside storage are also currently in use 1

[14,15,24]. This will lead to salt deposition on the casks. It is predicted that the initial temperature of the casks will be too high (around 150˚C, but may be higher, depending on the heat load of the SNF loaded) to reach the salt deliquescence point, avoiding the wetting of the canister surface and inducing SCC [14]. However, as the SNF’s radioactivity decays, the heat production will decrease exponentially with time and temperatures ideal for wetting and SCC propagation will be reached. The canisters will plausibly reach different levels of salt loading on their outer surface, depending on their location within the storage building and varying wind patterns. The effect of salt loading has so far rarely been assessed, as the standard laboratory experiments deal with solution concentration rather than weight of salt per unit area [20–22,33]. Salt deposition is not the only concern. SCC will only take place if the right material, aggressive chemicals and tensile stress are present [15].The role of applied and, to an extent, residual tensile stress has been widely investigated in the literature [10,12,28,34,35] with respect to Cl-SCC of ASS. The majority of cask designs have welds, which will likely produce significant residual stresses, as well as a machined surface finish, which also produces additional surface residual stresses [14,28]. These residual tensile stresses can be enough to trigger SCC if the aggressive species is present on the component, in this case Cl- ions. In order to recreate the microstructural residual stresses that are inevitably present in an engineering component, the samples were uniaxially strained to a level of 5% cold work prior to testing. The level of cold work was chosen based on a study performed by Scatigno et al. [35]. The study indicated that the highest crack density was observed on samples with plastic strains of 0.5-2%. Therefore, 5% cold work was chosen to represent a medium level of cracking [10,35], obtained when combined with an applied stress of 60 MPa, in the hope that this would give more differentiation to the chemical aspects i.e. the level of aggressive Cl- species rather than the mechanical effects of stress. The real engineering application of the casks require an understanding of the effect of salt loading on the likelihood of SCC to occur. For the tests described here MgCl 2 was chosen rather than real or model sea-salt. This was for several reasons: 1) acceleration of the tests, as real sea-salt may be somewhat less aggressive than MgCl2 since the deliquescence point of MgCl2 is much lower, at 33% RH compared to 70-80% RH for other common Cl- salts) [20–22]; 2) there is a wide range of existing literature based on MgCl2 to compare with directly; 3) a study with real sea-salt would be complicated by the presence of organic components and would have to be site specific [36,37]. This would therefore need to be studied in future when there is a baseline to compare against. In this work MgCl2 salt was deposited on to specimens prior to them being subjected to beam bending that produced a uniform upper surface tensile stress of 60 MPa. Samples were tested for 480 hours at 70% RH and 90°C. From correspondence with the nuclear utility, a chloride deposition rate of 0.87 mg m-2day-1 would be expected for a coastal indoor dry store facility. This would convert to a yearly chloride deposition of 3.1755 x 10-5 g.cm-2. It is expected from modelling estimates by the utility that the initial surface canister would be 150oC and would take approximately 25 years to reach temperatures below 100oC suitable for SCC. Canisters stored in equivalent outdoor facilities would realistically be expected to have higher salt deposition rates [38]. Nine different levels of salt loading were tested as described below (ranging from 1.7 x 10-4 to 3.1 x 10-2 g.cm-2) and measurements of corroded area, crack density and crack growth rates were carefully assessed.

2. Experimental The material tested was standard 304L ASS, complying with BS EN 10008-2-2005 requirements for grade 1.4307 (304L) [10]. The composition supplied by the manufacturer is shown in Table 1.

2

Table 1: 304L steel composition (wt.%).

Grade 304L

C 0.022

Cr 18.19

Mn 1.23

Ni 8.35

P 0.034

S 0.004

Si 0.25

Co 0.152

Cu 0.078

Fe Bal.

Rectangular section dog-bone samples were machined (see Fig. 1b) out of a 304L plate using electrodischarge machining (EDM). The samples were heat-treated at 900 ºC for 30 minutes under an argon atmosphere and quenched in water to allow for recovery and to fully stress relieve the samples, while avoiding recrystallisation and grain boundary sensitisation [10]. The samples were subsequently prestrained by placing the specimens in uniaxial tension. A fixed displacement of 5% plastic strain was applied using a Strauss servo-hydraulic tensile machine to each sample. An extensometer was used to accurately measure the elongation. A fixed displacement rate of 0.1 mm.s-1 was used, equivalent to a strain rate of 2 x 10-3 s-1. The actual strain measured for the 24 samples ranged from 5.00% (specimen 4b) to 5.10% (specimen 3b) with a mean of 5.06% and a standard deviation of 0.02%. (a)

(b)

Fig. 1: (a) Stainless steel jigs used to produce beam bending. (b) geometry of bend test specimen with dimensions (mm) [10,39].

The sample surfaces were ground using standard metallographic techniques, using SiC paper from 800 grit up to 4000 grit, in order to remove Cu residues left by the EDM cutting, and to provide a known finish for all samples. The samples were thoroughly cleaned and degreased using distilled water and propan-2-ol. The samples were then sprayed with a saturated solution of MgCl 2 and isopropanol using an airbrush. The solvent was then evaporated, and the process was repeated until the required level of salt loading was achieved. Nine different levels of chloride depositions were tested with 2 samples per loading, as reported in Table 2 (an extra 2 samples were repeated for the two lowest deposition levels due to the unusual crack densities and crack morphologies observed initially). Samples were weighed before and after MgCl2 deposition using a Fisher Scientific MH-214 3

analytic scale (accuracy 0.1 mg) to calculate the salt deposition. The Fisher scientific scale was operated in a fume cupboard and was insulated from the environment by a case. Limitations with the precision of the digital scale meant that it was not possible to perform any tests below 0.000171 g.cm2 with any degree of accuracy.

Table 2: Salt loadings given as chloride deposition levels for the 9 levels investigated.

Sample

1a 1b 2a 2b 3a 3b 4a 4b 5a 5b 6a 6b 7a 7b 8a 8b 8c 8d 9a 9b 9c 9d 10a 10b

Chloride Deposition (g.cm-2) 0.0319 0.0309 0.0231 0.0244 0.0212 0.0166 0.0133 0.0101 0.00440 0.00400 0.00226 0.00287 0.000931 0.00109 0.000619 0.000573 0.000532 0.000553 0.000165 0.000148 0.000174 0.000199 0 0

Average Chloride Deposition (g.cm-2) 0.0314 0.0237 0.0188 0.0117 0.00420 0.00257 0.00101

0.000569

0.000171

0

The samples were mounted on 304L ASS jigs (see Fig. 1a) to produce beam bending, applying a 60 MPa tensile stress on the upper surface (where the salt was deposited) of the specimens [39]. The load was applied and maintained using a compression spring. The deflection necessary to achieve the required applied tensile stress was calculated for each specimen and measured using a Mitutoyo ABSOLUTE Digimatic Indicator (accuracy 0.001 mm). The loaded jigs were then placed in an airtight, non-metallic enclosure. A beaker containing a saturated ionic solution of NaCl was placed in the enclosure along with the samples to ensure that a target humidity of 70% RH was maintained [40]. The container with the samples mounted on the jigs and the beaker were placed in an oven to achieve a temperature of 90 °C (± 1°C). The total duration of the test was 480 hrs (20 days). A data logger was placed within the container, inside the oven to monitor temperature and RH. The samples were inspected every three to four days and photographed using a digital camera to study the macroscopic 4

corrosion product formation and the pitting/cracking of the samples with different salt loadings. After testing, the samples were washed thoroughly using distilled water and propan-2-ol to remove the remaining salt and any loose corrosion products from the surface. To observe fine cracks, the samples were cut into six parts and mounted in Bakelite. A low deformation circular saw with an aluminium oxide blade disc was used to section the samples along the transverse direction (TD), shown in Fig. 2, at the lowest feed rate of 0.005 mm.s -1 to avoid the introduction of further residual stresses. Half of all the mounted samples were subjected to standard metallographic preparation using SiC paper from 800 grit up to 4000 grit and then polished using a solution of OPS (oxide polishing suspension). Optical characterisation was carried out using the Olympus-BXS1 optical microscope. The lowest magnification set-up was used to record and reconstruct the morphology of the cracks. For crack counting, rules were imposed on what was considered a crack and what was ignored: (a) cracks that had propagated all the way across the sample and/or through thickness were counted; (b) only cracks longer than a 50 µm threshold were considered and (c) branched cracks were only consider when they exceeded 50 µm in length and deviated at the beginning of the crack propagation (i.e. bifurcation see Fig 2).

TD

RD Fig. 2: Schematic of cracks that were or were not considered for the crack counting, along with a schematic of sample geometry: rolling direction (RD) and transverse direction (TD).

The other half of the samples were cross-sectioned twice (at an approximate interval of 2 mm with cuts parallel to the rolling direction, RD), to measure the crack depths and provide an alternative crack count for comparison. This was again performed using the circular saw at the lowest feed rate. The cross-sections were mounted in Bakelite and ground/polished using the same method as described previously. An optical microscope with stitching software was used to reconstruct the cracks at the lowest magnification for analysis. Scanning electron microscopy (SEM) was also carried out using a Jeol 6400 microscope, in secondary electron mode, at an accelerating voltage of 20 kV and working distance of 15 mm to look at the morphology of the cracks. The pictures recorded were analysed with an open source imaging software, ImageJ. A threshold contrast was used to isolate corroded/tarnished areas against unspoiled metal and the corroded areas was measured using ImageJ software. The evolution of the corroded area was recorded against time to a maximum of 100%, i.e. corrosion of the entire gauge length. The data obtained resembled a classic first order response to a step input, shown in eq. 1. (1) where A and k are fitting constants. A gives the final plateau value of the system (in this case total corroded area) and t is time. Origin Pro 2017 software was used to fit the corroded area data using the exponential decay function equivalent to eq. 1. The time constant, τ (the value of t where t = 1/k), was calculated for each data set from the fitted values of k and plotted against salt loading. τ is commonly used as a standard to define the time response of the system. At time τ, y = 0.63A, which indicates the time taken for 63% coverage of the maximum corroded area (plateau height).

5

3. Results & Discussion 3.1. Corroded area The start and end states of the corroded 304L samples are shown in Fig. 3. The salt loading has had a clear influence on the surface corrosion of the 304L samples. At low salt loadings samples (samples 5-9, below 4.27 x 10-3 gcm-2 chloride deposition) the corroded surfaces are consistent with quite localised corrosion, whereas at high salt loadings, corresponding to samples 1-4 (above 1.20 x 10-2 g.cm-2), the corrosion is more homogenous forming dark stains of corrosion products all over the surface of the specimens. Sample set 10, surprisingly, had also developed some surface corrosion despite no chloride salt being deliberately deposited on them. It is possible that there may be some exposure to chloride ions from an aerosol formed by the ionic bath in the test chamber. However, the corrosion products formed were superficial and completely removed after rinsing in water. Upon optical microscopy examination, the samples tested with no chloride salt deposited (10a and 10b) displayed no evidence of pitting or cracking.

Fig. 3: Comparison of samples before and after the 20-day corrosion testing.

Fig. 4 shows an example of a crack recorded on sample 6b after grinding of the corroded surface. Corrosion products are still present on the surface, but a shallow crack is already visible. An example of pitting observed in the samples that can potentially lead to SCC is shown in Fig. 5. Fig. 6 shows a close up of a pit, on sample 6a, prior to grinding. The characteristic lacy cover is visible, which helps to shield the pit and create a localised aggressive environment where a lower pH than the bulk system can be established more easily, helping pit growth. [3,26].

6

Stress Corrosion Crack

Fig. 4: Example of a stress corrosion crack on sample 6b after initial grinding to remove surface corrosion products.

Fig. 5: An example of pitting observed on sample 4a.

7

Lacy Cover

Fig. 6: Close up of a typical corrosion pit observed on sample 6a.

The coverage of the corroded areas as measured for the nine different salt loadings are compared in Fig. 7. After 20 days, differences in the total corroded areas achieved are apparent; generally the lower salt loading had less corrosion coverage than higher salt loadings – as expected according to literature [41,42]. The fitted graphs resemble the increasing form of an exponential decay curve. The decreasing rate of visible surface corrosion is due to the decreasing amount of untarnished area (i.e. remaining available area for corrosion) with time and due to the decrease in the concentrations of salt at the surface with time (due to chloride diffusion into pits and cracks). The reference samples corresponding to no salt deposition showed a final coverage of around 30%, which is in contrast with literature that reports the presence of a threshold of Cl - below which corrosion does not take place [41,42]. This difference, however, is likely to be due to salt aerosol in the chamber depositing onto the surface (the threshold for corrosion in 304L has been reported to be as low as 7 μg.cm-2 [42]) rather than an indication that 304L ASS can corrode in the absence of a salt.

8

100 100 -2 -2 Chloride Deposition Salt loading (gcm ) (gcm ) 0.042 0.031 0.032 0.024

Corroded Area (%)

Corroded Area (%)

80 80

0.025 0.019 0.016 0.012

60 60

40

20

0.0056 0.0042 0.0035

0.0026

0.0014

0.0010

40

0.00076

0.00057

0.00023

0.00017 0.00

0.00

20

0 0 0

5

0

10

5

15

10

Time (days)

15

20

20

Time (days) Fig. 7: The evolution of corroded areas with time for the different salt loadings used. Error bars show standard deviation.

For the two lowest chloride loadings (1.7 x 10-4 g.cm-2 and 5.7 x 10-4 g.cm-2), the fitted exponential curve seems to have plateaued well before the entire surface has corroded. This is possibly due to the availability of salt being a limiting factor; i.e. there is not enough salt to induce corrosion across the entire gauge area of the specimens or potentially some degree of cathodic activity protecting the surface. As the salt loading increases it appears that the maximum corroded area also increases. This resembles the studies by Zhang et al. where they have explored changes in crack densities with time, for various surface residual stresses. In a similar fashion to Fig. 7, they have observed what appears to be plateaus for different levels of residual stresses, suggesting that residual stresses were the limiting factor on corrosion rate in their study, whereas here it appears to be the salt concentration at the sample surface [34]. It is also interesting to point out the differences in the corroded areas for the four highest salt loading samples. Samples corresponding to 3.1 x 10-2, 1.9 x 10-2 g.cm-2 and 1.2 g.cm-2 have all reached >90% corroded area. However, although the curve fit for 2.4 x 10-2 chloride loading plateaued at around 80% surface corrosion, the fitting of the curve has no weighting to the upper plateau points (the fit here is not so good) and the actual measured final corroded area was closer in value to the other three high salt loading points. The fitted curve equations in Fig. 7 were used to obtain the time constant values, , which represents the length of time taken for the corroded area to reach 63% of the plateau value.  is plotted against chloride deposition levels in Fig. 8. There appears to be a negative linear trend between 2.6 x 10-3 and 2.4 x 10-2 g.cm-2 where increases in salt loading decreases the time response of the system, indicating higher corrosion rates at higher salt concentrations. At salt loadings greater than 2.4 x 10-2 g.cm-2, there is a significant increase in  suggesting that the corrosion rate had decreased with further increases to salt loading. This is likely to be due to the concentration of chlorides already being 9

saturated so the corrosion rate limiting step is the cathodic reaction, i.e. the availability of oxygen. With very high salt loadings, a sufficiently thick liquid film may be formed, which along with the reduced solubility of oxygen in concentrated salt solutions would result in slower oxygen diffusion to the cathodic sites [43]. For salt loading values below 2.6 x 10-3 g.cm-2, the fitted τ values may not follow the linear trend shown in Fig. 8, however, it is not possible to determine if this is a true trend.

6

Time constant (days)

5

4

3

2

1

0 0.00

0.01

0.02

0.03

0.04

Chloride Deposition (gcm-2) Fig. 8: The time constant, , versus salt loading. Time constant represents the time at which 63% of the maximum corroded area had been achieved. Error bars represent standard deviation. A linear fit has been plotted between 0.0026 and 0.024 -2 g.cm chloride deposition with an R-squared value of 0.971.

3.2. Crack densities The number of cracks observed per unit area is reported in Fig. 9. The graph appears to display three distinct regions: a very low salt loading region (from 0 g.cm-2 to 1.0 x 10-3 g.cm-2) that exhibits an unexpectedly high crack density, but with a very high level of scatter; a second region that seems to show a direct correlation between crack density and salt loading (from 1.0 x 10 -3 g.cm-2 to 1.9 x 10-2 g.cm-2), and a third very high salt loading region (above 1.9 x 10-2 g.cm-2) where the density of cracks falls with increasing salt loading. Fig. 9 shows a linear fit that was applied for the middle region between 1.0 x 10-3 g.cm-2 and 1.9 x 10-2 g.cm-2 chloride deposition, where a linear increase in crack density was observed with increasing salt load. Due to concerns over removing shallow cracks in highly corroded samples, an alternative crack count is provided in Fig. 10 made from the cross-sections of the samples i.e. the surface layer was not removed. Two cross-section cuts (along the RD) were made across each sample, at 1/3 and 2/3 widths, and the crack numbers were averaged between them. The trend observed here mimics the trend seen previously in Fig. 9: a highly cracked low salt loading region; a linearly increasing region and a decreasing crack number region associated with the high salt loading regime. The peak at very 10

low salt loadings is still present but it is much less prominent than before. However, this is likely due to the reduced length of these micro-cracks (less than 1 mm), which means many of them would not have been accounted for as they would not have been bisected by the cross-section cuts.

0.30

Very Low Salt Loading Region

Linear Region

High Salt Loading Region

Crack density (mm-2)

0.25

0.20

0.15

𝑦

5.89𝑥 + 0.00394

0.10

0.05

0.00 0.00

0.01

0.02

0.03

0.04

Chloride Deposition (gcm-2) Fig. 9: Crack number densities recorded for the different salt loadings. Error bars show standard deviation. A linear fit was -2 applied between 0.0010 and 0.019 g.cm chloride deposition with an R-squared of 0.796.

11

Crack number (mm-1)

0.35 Very Low Salt Loading Region 0.30

Linear Region

High Salt Loading Region

0.25 0.20 0.15 𝑦

.5𝑥 + 0.0749

0.10 0.05 0.00 0.00

0.01

0.02

0.03

0.04

Chloride Deposition (gcm-2) Fig. 10: Crack counting performed via the cross-sectioned samples. Error bars show standard deviation. A linear fit was -2 applied between 0.0010 and 0.019 g.cm chloride deposition with an R-squared of 0.943.

It is worth noting that despite similar corroded areas between samples of the same salt loading, the standard deviation in the number of cracks observed is significantly higher for samples at lower salt loadings (between 0 and 1 x 10-3 g.cm-2). This is due to a large difference in the observed crack densities between equivalently tested samples. The difference in the crack densities may be a result of other factors, such as material heterogeneity between samples, which may be exacerbated at low salt loading levels. Second phase particles, especially MnS, can play a very important role in promoting pitting and thus SCC, and these may differ significantly from sample to sample, [3,4,26]. Strain heterogeneity is always present and, as previously suggested by Scatigno et al. [35], can potentially promote micro-residual tensile stresses through incompatibilities between soft grains and hard lessdeformable grains, promoting in turn SCC itself [44–46]. An upper limit for the observed crack numbers has been reached at a chloride deposition of 1.9 x 10 -2 g.cm-2. Further increases in salt loading appeared to have an inhibiting effect on the number of cracks formed, with crack densities falling to almost 0 mm-2 at the highest tested salt loading. However, it is worth noting that despite the reduced number of cracks, one of the samples exposed to 3.1 x 10-2 g.cm-2 salt loading still displayed a through-thickness crack after 12 days of testing, suggesting that propagation rates have not decreased. Although the decrease in recorded surface crack numbers at the higher salt loadings could be due to crack coalescence, where several cracks may combine to form a longer crack, it is unlikely to be the sole reason for the fall in crack densities, as a similar decline is seen in the cross-sectional crack count. It is possible that the decrease in the observed crack densities at such high levels of salt loading is due to an increasing wet film thickness that could limit oxygen diffusion to the corrosion pits, reducing the likelihood of SCC initiation.

12

A relatively high number of cracks, compared to moderate salt loadings, were observed in some of the very low salt loading samples (below 1 x 10-3 g.cm-2). Samples 9a and 9b (1.7 x 10-4 g.cm-2 salt loading) exhibited regions with very densely packed cracks. However, despite the numbers of cracks observed were relatively higher, the cracks themselves appeared to be much finer and shorter in morphology than cracks observed in the higher salt loading samples; a comparison is shown in Fig. 11. On the other hand, samples 9c and 9d did not show the same densely packed cracking behaviour but the cracks that were observed were still very fine and short. This resulted in the very large standard deviation across sample set 9. In comparison, the cracking in higher salt loading samples was more defined in both length and width. This suggests that even very small chloride depositions are sufficient for SCC initiation, however, the availability of chloride at these concentrations may limit the rate of propagation. A comparison of the types of cracks observed in sample 2a is made with samples 9a and 9d in Fig. 11. In all three images, evidence of pitting associated with cracks can be observed and there appears to be more pits formed on sample 9a, which has led to the increased number of cracks. However, the pits are noticeably larger in sample 2a than sample 9a, consistent with literature observations that increased chloride deposition density results in larger pit diameters [42]. Fig. 12 shows SEM images of cracks from sample 2a and 9a; both samples exhibit a transgranular zig-zag pattern characteristic of chloride-induced SCC. (b)

(a)

(c)

Fig. 11: Examples of the differences in cracking found in the (a) very high salt loading sample 2a; (b) very low salt loading sample 9a; (c) very low salt loading sample 9d.

13

(b)

(a)

Fig. 12: Secondary electron SEM images showing morphology of SCC cracks for (a) high salt loading sample 2a; (b) very low salt loading sample 9b.

The presence of zig-zag “micro” cracks have been previously reported by Zhang et al. in which they have observed such cracks forming perpendicularly to the grinding direction. The study showed that small increases in residual stresses from grinding processes can greatly exacerbate the amount of micro-cracks observed [34]. The samples used in this study have all been ground to 4000 grit to remove surface contamination, however, it is difficult to ensure that the residual stresses introduced are identical through the manual grinding procedure. Although these were heat treated afterwards, some of these surface residual stresses may have remained and lead to the very different cracking patterns observed in sample set 9. It is possible that the effect of small changes in residual stresses on crack density is exacerbated at a low salt loading levels due to these micro-cracks only being visible in less corrosive environments; at higher salt loadings these shallow cracks would be obscured by the heavy surface corrosion. Furthermore, in order to expose cracks, the corroded sample surface was ground; for higher salt loading samples, more effort was required to remove the corrosion products and the fine surface cracks may have been removed as well. The crack densities were compared to the final corroded areas and the time response in Fig.13. The time constant, which can be used as a measure of the initial rate of corrosion, does seem to show a trend with SCC susceptibility according to Fig. 13a. As the time constant decreases (indication of a faster initial corrosion rate) the number of cracks observed increases seemingly linearly. This suggests that the initial surface corrosion seen is proportional to the number of SCC cracks. However, from Fig. 13b, the final corroded areas do not appear to have any correlation with the extent of SCC. This is likely due to a combination of effects: number of new cracks is likely to fall over time due to release of residual tensile stress from existing cracks; final corroded area plateauing at different values depending on chloride deposition levels and other contributions to the final corroded surface area such as pitting corrosion and depassivation of the steel.

14

(b)

0.15

0.15

Crack Density (mm-2)

Crack Density (mm-2)

(a)

0.10

0.05

0.10

0.05

0.00

0.00 2

4

6

30

40

50

60

70

80

90

100

Corroded Area (%)

Time Constant (Days)

Figure 13: Comparison of crack density against: (a) time consant, (b) final corroded area. Error bars show standard deviation.

3.3. Crack depths The crack depths were also measured from the cross-sections and converted to average crack velocities, shown in Fig. 14. It can be seen that despite the chloride loading varying by a factor of 40 (between 5.7 x 10-4 g.cm-2 and 2.4 x 10-2 g.cm-2), the crack velocities, in general, remain relatively consistent between 1 and 2 μm.hr-1, possibly suggesting that, unlike corroded surface area, salt loading has little impact on average crack propagation rates in this range. The values of average crack velocities calculated are of the same order of magnitude as the maximum crack velocities from a study by Örnek et al.who have reported crack velocities of between 5.4 and 6.8 μm.hr-1 for chloride depositions between 5.07 x 10-5 and 1.45 x 10-3 g.cm-2 [41]. This suggests that crack propagation rate is likely to be dependent on Cl- diffusion rate rather than Cl- deposition density, as the Clconcentration may already be saturated at crack tips, even at very low bulk concentrations, as suggested by Spencer et al. [10]. The crack velocity is much higher for the highest salt loadings samples (3.1 x 10-2 g.cm-2) due to the data representing a single through-thickness crack. Despite the significantly decreased cracking densities at such high salt loadings, the single crack that had developed had still propagated very rapidly through the entire thickness of the sample. It is possible that this isolated instance of a crack may be due to a sample defect, but even still it would not be safe to rely on a thick salt crust to prevent SCC occurring, as when cracking does occur it can still cause component failure.

15

7

Crack Depth Velocity (μm.hr-1)

6

5

Sample Set 10 Sample Set 9 Sample Set 8 Sample Set 7 Sample Set 6 Sample Set 5 Sample Set 4 Sample Set 3 Sample Set 2 Sample Set 1

4

3

2

1

0 0.00

1.71x10-4 5.69x10-4 1.01x10-3 2.57x10-3 4.20x10-3 1.17x10-2 1.88x10-2 2.37x10-2 3.14x10-2

Chloride Deposition (gcm-2)

Fig. 14: Average crack velocities for the samples. Error bars show standard deviation. Note that Sample Set 1 lacks error bars due to the crack velocity being derived from a single observed crack.

Despite the average crack velocities being relatively consistent, high inherent variation in crack velocities was observed. This variation could possibly be due to heterogeneous distribution of microstructural residual stresses from the imposed cold work, resulting in potentially high localised stresses. This could in turn result in the formation of dominant cracks early on and propagation at an accelerated rate, i.e. the single through-thickness crack observed in the highest salt loading sample that appeared to have a much higher crack velocity. By the end of the experiment multiple samples corresponding to medium and high salt loadings (above 2.6 x 10-3 g.cm-2) had developed thoughthickness cracks. Samples 1a, 2b, 3a, 3b, 4b, 6a, 6b all exhibited such cracks, as shown in Fig. 15. Samples 5a and 5b had not developed through thickness cracks by the end of the experiment, however, the samples failed when removed from the jigs suggesting that they were already very close to failing. Even the lowest salt loading samples, despite not cracking through, displayed dominant cracks that propagated further into the material than the average cracks (maximum crack velocity of 1.6 μm.hr-1). Dominant cracks will release much of the applied tensile stress and lower the residual stresses remaining in the sample, potentially resulting in reduced development of further cracks. Higher salt loading, within the linear region, could increase pit formation and allow for more pit-tocrack transitions to occur before the retardation of cracking from the release of stress by dominant cracks.

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3a

3b

4a

4b

6a

6b

Fig. 15: The end state of some of the corroded samples. The samples were ground to remove the salt and corrosion products in an attempt to expose surface cracks.

The lowest salt loading samples appear to have slower cracking velocity compared to all other salt loadings and much smaller standard deviations due to the numerous shallow micro-cracks. It is possible that these cracks behave differently due to the differences in their origin (micro-cracks could be forming due to residual stresses from surface preparation techniques). Zhang et al. have claimed that the micro-cracks could reach a dormant layer where cracks are arrested when they are sufficiently away from the ground surface. They have estimated that the mechanically affected layer, in their case, is around 90-110 μm and that the micro-cracks would arrest shortly afterwards [34]. Here, it is unlikely that the cracks will arrest in the dormant region due to additional microstructural residual stresses from cold work and the majority of the micro-cracks observed are around 100-200 μm deep. Dominant cracking has also been observed in samples 9a, 9b and 9c where a few cracks have progressed much deeper into the specimen, exceeding 600 μm in length. From this study it is not possible to determine if there is a lower threshold level of Cl - ions where the chloride deposition is so small that cracking, aside from surface micro-cracking, does not occur during the lifetime of a component. Even if such a threshold did exist, it would be hard to deem the tested component safe from SCC since any chloride containing salt would release chloride ions in aqueous solution when the deliquescence point is reached. Dissolved in water, even a very dilute chloride solution could reach crevices and other confined geometrical areas and become concentrated enough to cause SCC. A situation similar to the one proposed by Jivkov et al. could arise [47]. At very low concentrations of Cl-, like at very low levels of applied tensile stress, the kinetics of the phenomenon may be slow enough that SCC may not be observed in the lifetime of the components. However, at low concentrations of chloride, surface residual stress effects appear to be exacerbated and maintaining a good surface finish becomes critical to prevent formation of micro-cracks. Allowing very high levels of salt to deposit on surfaces to achieve a salt crust could be dangerous as despite lower crack densities observed, the cracks formed can still propagate through the depth of the material quickly.

4. Conclusions Beam bending applying a uniform stress of 60 MPa was carried out for 480 hours on 24 different samples, pre-strained to 5% cold work. Nine different levels of MgCl2 salt were deposited on sample surfaces, between 0 and 3.1 x 10-2 g.cm-2. Pitting and crack evolution were observed on all samples with salt deposited. Conclusions drawn on the effect of salt loading level for the nine levels of salt tested were:

17

1. There appears to be a direct positive linear correlation between the level of salt deposited onto a sample, the visibly corroded area and the crack number density for salt loadings between 5.7x 10-4 g.cm-2 and 1.9 x 10-2 g.cm-2. 2. Above 1.9 x 10-2 g.cm-2, crack density decreases, possibly due to the formation of a salt crust that could impede access to oxygen/water and reduce pitting. However, when a crack does form, propagation appears to be unimpeded. 3. At very low salt loadings between 0 and 5.7 x 10-4 g.cm-2 an unusually high number of microcracks were observed for some samples. These cracks were shorter and finer than cracks seen at higher salt loadings and have been suggested to be due to the effects of surface residual stresses present. 4. Crack depth velocities, calculated from sample cross-sections, showed a similar rate of around 1 - 2 μm.hr-1 for samples between 5.7 x 10-4 and 2.4 x 10-2 g.cm-2 suggesting that bulk chloride concentration has little effect on crack propagation rates, if above 1.7 x 10-4 g.cm-2. All samples that had cracked showed the presence of dominant cracks that had propagated much faster than the mean crack velocity. These dominant cracks have resulted in sample failure in 9 out of the 12 samples tested between 2.6 x 10-3 and 3.1 x 10-2 g.cm-2 salt loading. 5. The final corroded area for samples at different salt loadings appeared to change. Higher salt loadings generally reached a higher plateau of percentage corroded area than lower salt loadings suggesting some dependence of the maximum extent of corrosion with the amount of chloride deposited at the surface. 6. The time response for each of the different salt loadings were fitted and it was found that there appears to be a direct negative correlation between salt loading and the system time constant between 2.6 x 10-3 and 2.4 x 10-2 g.cm-2. This suggests that rates of corrosion are enhanced with salt loading. At higher salt loadings the time constant increases, which implies that corrosion rate has reduced, possibly due to the excess salt becoming a barrier to oxygen and water. Dry storage canisters in the UK are currently planned to be able to remain in interim storage for at least 100 years, allowing significant deposition of salt to occur in their lifetime. Storage facilities generally rely on passive measures meaning that the accumulated salt will not be removed via washing. From the results of this study, it seems that even low salt depositions of 1.7 x 10-4 g.cm-2 (corresponding to a canister in service for 5 years in an enclosed dry storage facility) have the potential for initiating SCC if deliquescence can occur. Although initially the canister surface temperatures would exceed 100 oC due to decay heat, which would be too high for the deposited salt to wet, this would likely fall to sub 100 oC within 25 years, allowing salt deliquescence to occur. As the length of service of the canisters increases, the number of SCC cracks would increase due to the increased salt loading and longer exposure times. It is unlikely for the canisters to remain in interim storage long enough to reach the high salt loadings that appeared to limit crack initiation (1.9 x 10-2 g.cm-2) and even if such an event did occur, propagation of already formed cracks would likely be unimpeded. Although this study was performed using MgCl2, sea-salt would likely give a similar trend albeit with lower crack densities due to being less aggressive, however the influence of sitespecific organic components will also need to be considered. Furthermore, another consideration for future work would be how the varying level of gamma radiation at the surface of the canisters as the spent nuclear fuel decays would affect the SCC behaviour.

18

Declaration of interests The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.

Acknowledgments The authors would like to thank EPSRC Grant No. (EP/I003088/1) and the PROMINENT group for their financial support for Giuseppe Scatigno, Mark Wenman and Mary Ryan. Mary Ryan currently holds the RAEng/Shell Research Chair. All authors also thank EDF Energy and the EPSRC Centre for Doctoral Training in Nuclear Energy (EP/L015900/1) for financial support. References [1]

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Graphical abstract

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