Localized corrosion continues to be one of the most difficult forms of corrosion to predict and monitor, making implementing mitigation strategies limited. This reality leads corrosion engineers to use the blunt instrument of alloy selection as the only means of (hopefully) ensuring that such an attack does not occur on an engineering asset. In the late 1960s and early 1970s, the advent of electrochemical techniques in corrosion science was heralded as a way to address the problem by determining the critical conditions needed for pits to nucleate and grow and also those required for pit growth to stop (i.e., the pits to repassivate).

The conceptual framework first suggested by Pourbaix, et al.,1  that a wetted surface needed to exceed some potential (usually referred to as the pitting potential, Epit) for pits to nucleate, and those pits would grow until the potential was below a lower potential (usually referred to as the repassivation potential, Erp). The position of these potentials relative to the corrosion potential (Ecorr) was also critical. If Ecorr was above Erp but below Epit, then the surface would be passive unless there was a disruption in the passive film (e.g., a mechanical breach). Slow cyclic potentiodynamic polarization (CPP) curves were heralded as a means of determining both the pitting potential (Epit) and Erp in a single measurement. The increasing availability of commercial electrochemical workstations promised a revolution in material selection against localized corrosion.

The excitement came to a crashing halt when Brian Wilde, then at the Ohio State University, published papers2-3  that seemed to show that the measured values of the repassivation potential were highly sensitive to experimental variables. The work showed that the value of Erp was a strong function of scan rate and vertex current density, with results decreasing by hundreds of mV with increasing amount of corrosion damage. It was concluded that Erp was not a property of a material/solution combination and thus not of use for engineering purposes. For almost two decades, the interest in the engineering application of Erp waned. Rosenfeld, et al.,4  provided additional supporting data. Although it was still considered a valid concept, there seemed to be no way to operationalize it.

In the early 1990s, work on a proposed repository for high-level nuclear waste (HLNW) was proceeding apace. Yucca Mountain (YM), Nevada had been selected as the location for the repository. The containment concept was defense in depth: the HLNW would be contained in stainless steel containers each of which would be placed inside an additional corrosion-resistant outer container. These would be buried under YM which is in the desert west of Las Vegas. The dry climate would limit the amount of water available; the mountain would provide a barrier to that water reaching the canisters, and then the corrosion-resistant outer canister would protect the stainless steel container which would protect the HLNW. At early stages after emplacement (several hundred years), the canister surfaces would be hot enough that any seepage water would evaporate upon contact, leaving behind the dissolved species it had collected during its transport through the mountain. Eventually, the canister surfaces would cool sufficiently that the salt on the surface would deliquesce, forming an electrolyte that would initially be saturated, but then dilute as the temperature continued to fall. Aqueous corrosion of any kind would not start until the deliquescence of the salts on the surface. That said, under conditions of elevated temperature and concentrated chloride solutions, there was concern that the canisters would eventually be breached by corrosion, allowing radionuclides to be carried to groundwater several hundred feet below the emplacement zone. EPA regulations required DOE to provide a technical basis for the safety of any design such that the biosphere near YM would be exposed to only acceptable levels of radiation over the design life. This design life increased from initial values of approximately 103 y to 106 y. Prediction of corrosion performance for those time scales was unprecedented and daunting.

Projects aimed at developing such a technical basis were created and executed under the separate auspices of the U.S. DOE and the U.S. NRC. The latter work was being performed at the Center for Nuclear Waste Regulatory Analyses (CNWRA) at the Southwest Research Institute (SwRI) as part of their development of a Total System Performance Assessment. The U.S. DOE work focused on establishing the rates of both general and localized corrosion that would develop in repository environments, primarily through exposure testing of a wide range of alloys in solutions whose compositions aimed to replicate those expected to develop on the canisters due to the seepage of water through the fractures in YM.

The team as CNWRA took a different tact, particularly regarding localized corrosion. They made the compelling argument that the uncertainty associated with the application of laboratory-measured corrosion rates to field environments rendered a corrosion-allowance approach of limited utility with such long service lives. The propagation of uncertainty in localized corrosion rates would render any prediction highly suspect in terms of reliability. Rather than using a rate-based approach, Dunn, Cragnolino, and Sridhar considered what conditions could exist in a repository that would preclude localized corrosion from propagating, even if initiated. Thus, the concept of Erp was particularly attractive in that once established, a go-no-go decision could be made on any combination of alloy and environment. In addition, one could in principle monitor the Ecorr of the material, and as long as it was below Erp, localized corrosion could not propagate, even if the passive film was breached.

The challenge involved how to measure Erp. In a series of papers, the CNWRA team explored several methods, all of which involved initiating and growing localized attack (i.e., pitting/crevice corrosion) at a highly positive potential for some time, and then lowering the potential until the current density fell to low levels, indicative of repassivation. Others had used a similar approach to establish Erp for different alloy/environment systems. The key differences are that Dunn, Cragnolino, and Sridhar decided that (a) the anodic charge density passed (Q) would be the critical independent variable, and (b) they would extend the range of Q to values much, much higher than in most previous studies in an attempt to establish a lower limit to Erp. They supplemented their work by extracting Erp vs. Q data from the literature to demonstrate the existence of a plateau in Erp for sufficient levels of Q for a range of alloy/environment combinations, as shown in their Figure 3.

The most powerful part of the work of Dunn, Cragnolino, and Sridhar was their demonstration of the validity of the concept via a large matrix of long-term (nearly 3 y) potentiostatic tests that demonstrated that when the potential was above the properly measured Erp, localized corrosion eventually occurred, but when the potential was at or below Erp no form of localized corrosion (either pitting or crevice corrosion) was observed. The higher the potential above Erp, the shorter the time until localized corrosion is initiated.

The importance of the work of Dunn, Cragnolino, and Sridhar is its provision of a basis for engineering implementation of the concept of repassivation potential. The approach has been the basis for further experimental work, modeling, and monitoring. Experimental work has applied the concept to a wide range of alloys.5-,9  Other work has sought to connect this work to other aspects of localized corrosion.10-,14  Characterization of the effects of experimental variables on Erp has also been the focus of work.15-,17  The use of Erp in corrosion modeling has allowed predictive frameworks to be developed18-,21  and applied for corrosion monitoring.22 

The impact of the Dunn, Cragnolino, and Sridhar papers in the engineering application of corrosion science is on par, in my view, with the work of Stern and Geary23  in establishing polarization resistance as an effective means of evaluating corrosion rate as well as Pourbaix’s development of E-pH diagrams.24  In all three cases, the publication provided a tool of great use to corrosion engineering that was based on a combination of deep insight and strong fundamental science.

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Lifetime predictions of containers used for the disposal of high-level radioactive waste (HLW) are necessary to determine the overall performance of a HLW repository. Failure of the containers by localized corrosion and release of radionuclides into the accessible environment may occur if chloride-containing groundwater infiltrates the repository and interacts with the containers. To assess the repository performance, it is necessary to develop a methodology for predicting the long-term susceptibility of container materials to localized corrosion. Laboratory tests have demonstrated that pitting and crevice corrosion are initiated when the corrosion potential of the material exceeds the repassivation potential for localized corrosion. Once corrosion is initiated, the propagation of the localized attack will result in rapid penetration of the waste packages. However, no localized corrosion can be initiated, and all active corrosion sites repassivate when the corrosion potential of the material is lower than the repassivation potential. Results of this investigation demonstrated that the corrosion potential and repassivation potential are useful parameters to predict the long-term localized corrosion susceptibility of corrosion- resistant candidate container materials such as type 316L (UNS S31603) stainless steel and alloy 825 (UNS N08825).

Editor’s Note: Originally published in Corrosion 56, 1 (2000): p. 90-104. When citing “An Electrochemical Approach to Predicting Long-Term Localized Corrosion of Corrosion-Resistant High-Level Waste Container Materials,” please cite the original version.

KEY WORDS: alloy 825, chloride, high-level nuclear waste containers, localized corrosion, nickel-based alloys, radionuclides, repassivation potential, stainless steel

INTRODUCTION

The primary goals of the U.S. Department of Energy (DOE) updated waste containment and isolation strategy for the proposed repository at the Yucca Mountain (YM) site are the near-complete containment of radionuclides within waste packages (WP) for several thousand years and acceptably low annual doses to a member of the public living near the site.1  The WP consists of containers, fillers, basket materials, and waste forms (spent fuel [SF] or vitrified waste). One of the most important system attributes recognized in this strategy is container lifetime. To predict the container lifetime confidently, the adequacy of methodologies for extrapolating short-term laboratory data to long-term performance needs to be demonstrated.

Several WP designs have been evaluated by the DOE over the history of the YM repository program. The canistered fuel design is considered most likely to be used for high-level radioactive waste (HLW) disposal.2  In this design, either 21 pressurized water reactors (PWR) or 40 boiling water reactors (BWR) assemblies are contained in a type 316L (UNS S31603)(1) stainless steel (SS) canister (3.5 cm wall thickness) surrounded by an inner overpack made of a corrosion-resistant alloy (2.0 cm wall thickness), which is contained in an outer overpack made of a corrosion allowance material (10 cm wall thickness). The total surface area of such a WP is ∼ 37 m2 and the loaded weight is ∼ 65,000 kg. Alternate designs include WP for uncanistered fuel, small-canistered fuel (12 PWR or 24 BWR SF assemblies), and vitrified waste. Nickel-based alloys, such as alloys 825, 625, and C-22 (UNS N08825, N06625, N06022, respectively) presently are being evaluated for the inner corrosion-resistant overpack, and a carbon steel (such as A516 grade 55) is proposed for the outer corrosion allowance overpack.3  If groundwater contacts the container, it is expected by the DOE that uniform corrosion of the carbon steel overpack will proceed at a slow and predictable rate. Evaporative effects and interactions with concrete in the repository may result in an alkaline groundwater that, in the presence of chloride, can lead to localized corrosion of the carbon steel surface.4  If the carbon steel overpack is perforated, then the inner overpack will be exposed to the repository environment. Corrosion of the inner overpack may be mitigated by the use of a corrosion-resistant alloy and galvanic coupling of the inner overpack to the carbon steel overpack.5-6 

The most common approach to long-term corrosion prediction involves obtaining the corrosion rates of candidate materials exposed to environmental conditions that are similar to, or preferably more severe than, the anticipated environmental conditions of the particular application. These rates then are integrated over time to predict service life. The corrosion rates are obtained through the use of controlled laboratory tests, coupon exposures in the field, or performance evaluation of similar materials in analogous applications. An example of this approach is found in the DOE Total System Performance Assessment (TSPA) conducted in 1995.7  This analysis uses atmospheric and natural water corrosion data to estimate the corrosion rate of outer overpacks in the repository. For the corrosion-resistant inner overpack, degradation was assumed to occur only by localized corrosion that was described by a pit growth rate that varies with temperature. The general corrosion of the corrosion-resistant container is determined by the passive dissolution rate of the material. This approach, referred to here as the empirical corrosion rate approach, has many drawbacks when applied to corrosion estimates for the repository time scales (eg., 103 years to 104 years). These limitations include:

First, the relationship between the laboratory or field service environments in which corrosion rates are obtained and the repository environment seldom can be established clearly. Inferring analogous behavior between short-term experiments and the repository setting is especially problematic in the case of the hydrologically unsaturated zone at the YM site. In the proposed repository, the fluid in the environment surrounding the container (typically referred to as the near-field or very near-field environment) may evolve in time and may vary in ionic strength from that of a dilute groundwater to that of a saturated electrolyte because of a complex combination of processes including evaporation, interaction with minerals and man-made materials, and radiolysis.

Second, several corrosion phenomena, especially localized corrosion and stress corrosion cracking (SCC), have relatively long initiation times that may be longer than the time elapsed in short-term tests, which are typically < 7 days.8  Once corrosion initiates, however, the rates can be quite high. The initiation time may be a function of the material and test environment. Thus, short-term corrosion rate data may be more useful in eliminating materials that are not suitable for a given environment than for predicting the long-term performance of materials.

Third, it is difficult to evaluate effects of design features in corrosion processes such as galvanic coupling. Hence, a large number of long-term tests (i.e., several weeks to years), using various possible design configurations, have to be conducted.

Finally, proof of the performance of a material in a repository cannot be obtained easily because of the lack of long-term exposure data for the candidate corrosion-resistant container materials. If feasible, models developed to simulate the long-term corrosion behavior of a material should be constructed from a mechanistic understanding of uniform and localized corrosion processes. The predictive capability of the models can be verified using a diverse set of corrosion data. However, the empirical corrosion rate approach is not based on a mechanistic understanding of the processes involved. As a result, long-term estimations made using this approach are more uncertain. The usefulness of the empirical corrosion rate approach is constrained further when most of the field and laboratory measurements of corrosion rates do not include measurement of the corrosion potential of the metal in the test environment. As a result, data do not provide a suitable means to predict the performance of the material when the duration of the exposure is increased. Because the corrosion potential of the test specimens is not always measured, valuable information necessary to predict the onset of localized corrosion in long-term exposures often is not available.

An electrochemical approach can overcome many of the limitations of the empirical corrosion rate approach by explicitly calculating corrosion rates in terms of constituent electrochemical reactions9-10  or by using critical electrochemical parameters for determining the onset of various corrosion processes.5,11  An electrochemical approach has several advantages.

First, the occurrence of various corrosion modes such as pitting, crevice corrosion, and SCC can be accelerated in short-term tests using a range of environments and applied potentials. The critical potentials for the occurrence of various corrosion modes then can be determined. The extrapolation from these short-term data is made then by calculating the evolution of the natural potential of the metal (corrosion potential) in the repository environment and comparing these to calculations derived from the measured critical potentials. The calculations of corrosion potentials are supported by fundamental electrochemical models based on the kinetics of the corrosion reactions. Thus, a clearer and more defensible relationship can be established between the test environments and the anticipated repository environment.

Second, the initiation time for localized corrosion and SCC explicitly can be considered in terms of its relationship to the corrosion potential. In some cases, kinetics of the corrosion modes can be measured independently and supported by electrochemical models.

Finally, the corrosion performance of a given material in diverse applications can be assembled in a single map of critical electrochemical parameters. This approach, similar to the one used by Staehle,12  could provide greater confidence in the performance prediction.

Electrochemical models do not dispense with the need to measure corrosion rates; rather, they assign different corrosion rates to different regimes of corrosion, with the delineation of the regimes of corrosion and the rates of corrosion within each regime supportable by fundamental, mechanistic reasoning. It also must be emphasized that electrochemical models, just as other models, are only as good as the quality of the underlying data. Thus, electrochemical models offer an additional advantage over empirical corrosion rate models because they can be used to evaluate critical data needs.

A number of investigators have proposed the use of the repassivation potential (Erp) as a parameter to predict the long-term initiation of localized corro- sion13-16  and SCC.11,17  If the corrosion potential (Ecorr) of the metal exceeds Erp, then localized corrosion can be initiated. Prior to initiation of localized corrosion, the metal is expected to corrode at a slow rate corresponding to the passive current density. After the initiation of localized corrosion, the pit would propagate through the metal container at a rate that is orders of magnitude greater than would be expected during passive dissolution. If the Ecorr of the inner barrier decreases below Erp, active localized corrosion would cease and the metal will repassivate. The concept of repassivation is not new. Pourbaix, et al., proposed the use of the Erp (also called the protection potential) below which localized corrosion would not be initiated.18  However, later investigations by Wilde,19  and Rosenfeld, et al.,20  indicated that the Erp decreased with increasing extent of prior corrosion, leading to the conclusion that this parameter cannot be used as a conservative criterion. Hence, to be a conservative parameter for predicting long-term localized corrosion, it must be demonstrated that a lower bound value for Erp can be established.

In the present study, data on type 316L SS and alloy 825 generated by the Center for Nuclear Waste Regulatory Analyses (CNWRA) program during the past few years was reviewed along with data from the literature to demonstrate the conservatism of Erp. Additionally, the applicability of Erp for prediction of the long-term occurrence of localized corrosion was examined using results of ongoing long-term tests. Finally, the current mechanistic understanding of localized corrosion initiation and repassivation processes was assessed.

EXPERIMENTAL PROCEDURES

Materials

A variety of corrosion-resistant and corrosion allowance materials are being tested in the CNWRA program. However, this study focused mainly on alloy 825 and type 316L SS because its objective was to establish the predictive methodology for corrosion- resistant materials. The compositions of type 316L SS and alloy 825 are given in Table 1.

Table 1.

Composition of Alloy 825 and Type 316L Stainless Steel (wt%)

Composition of Alloy 825 and Type 316L Stainless Steel (wt%)
Composition of Alloy 825 and Type 316L Stainless Steel (wt%)

Environments

Tests were conducted at 95°C±2°C in an ASTM G-3121 -type electrochemical cell with ≈ 1,500 mL solution containing 1,000 ppm Cl, 85 ppm , 20 ppm , 10 ppm , and 2 ppm F, all as sodium salts. The solution composition reflected the expected composition of groundwater that may be expected in a HLW repository with an increased chloride concentration. The effect of chloride concentration was studied by adding chloride as sodium chloride (NaCl) to the above solution. However, because of the solubility limitation, solutions with chloride concentration > ∼ 5.5 M were achieved by the use of lithium chloride (LiCl) or magnesium chloride (MgCl2) solutions, without the addition of other anionic and cationic species. The pH of the concentrated MgCl2 solutions were measured to be in the range from 2.0 to 2.7. In tests conducted under air- saturated conditions, a high-purity mixture of 79% N2 + 21% O2 (carbon dioxide [CO2] free air or zero air) gas was used to purge the solution.

Localized Corrosion Tests
Five configurations of the various types of specimens used in pitting and crevice corrosion tests in this investigation are shown in Figure 1.
FIGURE 1.

Schematic of various specimen configurations used in the study of localized corrosion: (a) CPP, (b) lead-in-pencil, (c) longterm test, and (d) crevice.

FIGURE 1.

Schematic of various specimen configurations used in the study of localized corrosion: (a) CPP, (b) lead-in-pencil, (c) longterm test, and (d) crevice.

Close modal

Cylindrical specimens of alloy 825 and type 316L SS measuring 6.2 mm in diameter and 48.6 mm long with a 600-grit finish (Figure 1[a]) were used in cyclic potentiodynamic polarization (CPP) tests. These tests were conducted in an ASTM G-522 -type five-neck flask equipped with a calibrated thermometer, a platinum counter electrode, and a salt bridge with a porous silica tip filled with test solution. The salt bridge was connected to a saturated calomel electrode (SCE) that was maintained at room temperature and served as the reference electrode. Test solutions were deaerated thoroughly with high- purity N2 (99.999 %) through a glass frit bubbler. Prior to the start of each test, the Ecorr of the specimen was measured vs the reference electrode using a high-impedance (100 MΩ) electrometer. The potential of the test specimens were increased at a rate of 0.167 mV/s using a computer-controlled potentiostat. Potential scans were started at the Ecorr. The current density of the specimen was monitored continuously. At a current density of 5 mA/cm2, the direction of the potential scan was reversed and the potential of the specimen was decreased using the same scan rate. The potential scan was stopped upon reaching the initial value of Ecorr. Following each scan, the specimen was examined for localized attack. The purpose of the CPP tests was to establish initiation potentials for pitting (Ep) and crevice corrosion (Ecrev), as well as Erp over a wide range of environmental conditions (10−4 M Cl to 9 M Cl, 25° C to 95°C). An SCE maintained at room temperature was used as a reference in these tests.

Simulated, single-pit propagation tests were conducted on alloy 825 using lead-in-pencil specimens (Figure 1[b]). The advantage of this geometry is that the location of pitting is known and the geometry is amenable to theoretical analyses using relatively simple analytical models. In addition, the repassivation of very deep simulated pits can be investigated with this specimen geometry. The disadvantages of the single-pit specimen are that the morphologies and growth rates of the simulated pit specimens are different than those of naturally formed pits. As a result, the growth rates obtained in single-pit tests cannot be used to predict the penetration of naturally formed pits. Prior to the start of the tests, specimens were weighed to ±0.03 mg. The specimens then were covered with a polyolefin tubing that allowed only one end of the specimen to be exposed to the solution. Electrical contact to the specimen was made outside the test cell. The specimen was oriented so that the opening of the unidirectional pit faced upward. The pit had an area of 0.041 cm2. Corrosion was initiated at potentials of 500 mVSCE to 600 mVSCE. After ≈ 5 h, the anodic current density was constant, indicating that stabilization of the actively growing pit was achieved and the potential was decreased to values from −300 mVSCE to 100 mVSCE at a rate of 5 mV/s. The current and potential were recorded allowing real time calculation of pit depth and propagation rate. This calculation was made using the theoretical weight loss of 2.64 × 10−4 g/coul assuming congruent dissolution of Fe, Cr, and Ni in alloy 825 as Fe2+, Cr3+, and Ni2+, respectively.23  At the conclusion of the tests, specimens were weighed again to confirm the accuracy of the propagation rate calculations. The purpose of these tests was to measure the growth rate and repassivation of single pits and compare them to measurements made on samples where an uncontrollable number of multiple pits formed.

Potentiostatic polarization tests were conducted using cubic specimens (Figure 1[c]). These specimens had a total surface area of 15 cm2, including polished and Cr-depleted mill-finished surfaces, as described in previous publications.15,23  Potentiostatic and open- circuit corrosion tests also were conducted using specimens with a controlled crevice (Figure 1[d]).

Potentiostatic pitting and crevice corrosion initiation times were measured by polarizing cylindrical and creviced specimens to potentials above and below the Erp measured in CPP tests. Specimens were connected to a computer-controlled multichannel potentiostat. A constant potential was applied for test intervals of 28 days. Between test periods, specimens were examined for visible signs of corrosion and weight loss using an analytical balance with ±0.00005 g reproducibility (standard deviation) while the test solutions were changed. No further exposures were conducted on specimens after localized corrosion occurred. Open-circuit tests were conducted with a setup and solutions similar to those used for the potentiostatic polarization tests. However, instead of controlling the potential of the specimen with a potentiostat, the corrosion potential of alloy 825 was monitored continuously in air-saturated, 1,000 ppm Cl solutions at 95°C. Periodically, the test was interrupted and the specimen was examined for signs of localized corrosion. These methods allowed the validity of using Erp from short-term CPP tests as a parameter for determining long-term performance to be tested.

Since at open circuit, no net current flows in the system, the corrosion kinetics cannot be measured using electrochemical techniques on a single specimen. Thus, a series of open-circuit corrosion tests was conducted with a creviced alloy 825 specimen connected to a large noncreviced alloy 825 plate that served as a cathode through a zero-resistance ammeter (ZRA). If crevice corrosion occurred, the potential of the crevice specimen decreased, resulting in an anodic current in the external circuit. Tests were conducted in solutions containing 1,000 ppm Cl at 95°C±2°C. The redox potential of the test solution was controlled by using a redox couple such as Cu+/Cu2+ added as chloride salts with hydrochloric acid (HCl) or NaCl additions to achieve the desired total chloride concentration. The area ratio of the alloy 825 cathode plate to the creviced area of the test specimen was at least 10:1. During the tests, the current was measured with a resolution of 1 nA. The potential of the specimens also was recorded with a resolution of 0.1 mVSCE maintained at room temperature. This test method allowed the measurement of corrosion rates under natural exposure conditions at various redox potentials.

Pit repassivation tests also were conducted on alloy 825 specimens (Figure 1[a]). Tests were conducted by initiating pits at 600 mVSCE. After initiation, pits were grown at 400 mVSCE. Following the accumulation of a specified charge density, the applied potential of the specimens was decreased to the desired value. The repassivation time was measured as the time required for the current density to decrease < 50 μA/cm2 after the potential of the specimen was decreased rapidly from 400 mVSCE to the desired potential. Details of this technique have been published previously.11,16  These tests allowed the validity of Erp to be tested on specimens with multiple pits.

RESULTS

Effects of chloride concentration on Ep and Erp of type 316L SS and alloy 825, as measured in CPP tests, are shown in Figures 2 and 3, respectively. Values of Ep and Erp plotted in these figures were measured in deaerated environments so that a large potential range could be investigated without the interference of oxygen reduction reaction. In addition to Ep and Erp, the Ecorr for these alloys measured in air-saturated and deaerated solutions are shown in Figures 2 and 3. It was apparent that Ep and Erp were strongly dependent upon the concentration of chloride > 10−3 M. At chloride concentrations < 10−3 M, Erp values increased substantially and approached the measured Ep, indicating that pitting did not occur in all of the specimens tested in low chloride concentrations. Where pitting did not occur, the Erp had no physical meaning and Ep corresponded to the onset of the oxygen evolution reaction on the alloy. The net result was a large scatter in the data and rather large 95% confidence intervals (±200 mV) on the repassivation potential for alloy 825 (Figure 3) at chloride concentrations < 10−2 M. At chloride concentrations in the range from 10−2 to 1 M, a large difference between Ep and Erp was observed. The critical potentials for alloy 825 were higher than those for type 316L SS, especially at low chloride concentrations. However, the difference between these potentials was reduced significantly at very high chloride concentrations.
FIGURE 2.

Effect of chloride concentration on critical potentials for pitting of type 316L SS at 95° C. The specimens were immersed completely exposing the crevice between the specimen and washer. The scan rate for Ep and Erp measurements was 0.167 mV/s.

FIGURE 2.

Effect of chloride concentration on critical potentials for pitting of type 316L SS at 95° C. The specimens were immersed completely exposing the crevice between the specimen and washer. The scan rate for Ep and Erp measurements was 0.167 mV/s.

Close modal
FIGURE 3.

Effect of chloride concentration on critical potentials for pitting of alloy 825 at 95°C. The specimens were immersed completely exposing the crevice between the specimen and washer. Values of Ep and Ep were measured using a scan rate of 0.167 mV/s in deaerated solutions.

FIGURE 3.

Effect of chloride concentration on critical potentials for pitting of alloy 825 at 95°C. The specimens were immersed completely exposing the crevice between the specimen and washer. Values of Ep and Ep were measured using a scan rate of 0.167 mV/s in deaerated solutions.

Close modal

Another notable observation (Figure 3) was the relative independence of Erp on pH between pH 1 and pH 8.2. This observation was consistent with the observations in the literature, as well as the current understanding of pitting process. The pH inside the pit was dictated by the hydrolysis of cations and became independent of the external pH.

An expression for Erp in terms of temperature and Cl concentration over the range of 10−3 M to 9 M (pH 1 to 8.2) was obtained by using the lowest measured value of Erp for each Cl concentration as a conservative criterion to predict the initiation and repassivation of localized corrosion. The general expression for Erp is given in Equation (1), with specific parameters for alloy 825 over the Cl concentration range from 10−3 to 9 M in Equation (2):5 
formula
formula
where Erp (mVSCE) is the repassivation potential for localized corrosion (pitting and crevice corrosion), and (T, mVSCE) is the repassivation potential at 1 M chloride concentration (Equation [1]).

Previous work has established the dependence of Erp on other environmental factors such as the concentrations of , , and F.24  As shown in Figures 2 and 3, thiosulfate () decreased the Erp significantly for both alloys. The effect of observed in this study was similar to that found by Nakayama, et al.25 

Although the presence of oxygen was not expected to alter the values of Ep and Erp, the occurrence of the oxygen reduction reaction determined the corrosion potential of the container materials in a repository environment. The corrosion potential data measured in air-saturated solutions also are plotted in Figures 2 and 3. In addition, the corrosion potential of alloy 825 in Cl- solutions with the addition of 5 mM hydrogen peroxide (H2O2) is shown in Figure 3. As mentioned before, the occurrence of localized corrosion in a given environment can be gauged by comparing the Ecorr under anticipated redox conditions and Erp in the same environment, but without the redox species. While redox species such as oxygen and H2O2 were not expected to alter the Erp significantly, it generally is convenient to obtain Erp values in the absence of these redox species to expand the range of potentials studied. From Figure 2, it was anticipated that type 316L SS would exhibit localized corrosion in air- saturated solution with chloride concentrations >10-3 M, as a result of the Ecorr being more noble than the Erp. In contrast, the data plotted in Figure 3 suggests that alloy 825 did not suffer localized corrosion in air-saturated solutions even at chloride concentrations as high as 4 M since the Ecorr was less noble than the Erp. However, predictions based on the data shown in Figures 2 and 3 are valid only if the values of Erp and Ecorr are not time-dependent. Long-term tests, conducted as part of this investigation to determine the effect of exposure time on the value of these parameters, indicated the Ecorr increased with time while Erp was not time-dependent.

Long-term Ecorr data for alloy 825 in a 1,000-ppm Cl- solution is shown in Figure 4. Despite the fluctuations, a trend of an increasing Ecorr was readily apparent. Over the course of almost 600 days of testing, Ecorr increased from an initial value of -270 mVSCE to a high of 160 mVSCE. Erp values for alloy 825 in ppm Cl-, ranging from 50 mVSCE to 100 mVSCE, are indicated in this figure as reported previously by Dunn, et al.26  After ≈ 350 days, the Ecorr exceeded the upper value of Erp (100 mVSCE) and crevice corrosion was observed. Although the depth of attack in each case was shallow and the weight loss minimal, crevice corrosion was observed visually each time Ecorr exceeded 100 mVSCE.

FIGURE 4.

The evolution of corrosion potential of alloy 825 exposed to a 1,000-ppm Cl solution at 95°C. Crevice corrosion was observed when Ecorr exceeded the Erp measured in CPP tests.

FIGURE 4.

The evolution of corrosion potential of alloy 825 exposed to a 1,000-ppm Cl solution at 95°C. Crevice corrosion was observed when Ecorr exceeded the Erp measured in CPP tests.

Close modal

The effect of penetration depth on the Erp for pitting and crevice corrosion is shown in Figure 5. These tests were conducted by initiating pitting or crevice corrosion at potentials of 500 mVSCE to 600 mVSCE. After initiation, the localized corrosion was allowed to propagate at 400 mVSCE.11  Specimens then were repassivated by reducing the potential at a rate of 5 mV/s until the current density was continuously < 50 μA/cm2. After repassivation, the depth of the pits and crevice-corroded regions were measured. The relationship between depth of attack and Erp can be determined by varying the time during which localized corrosion was allowed to propagate. From the results plotted in Figure 5, there was a strong dependence of Erp on pit depth for shallow penetrations. For deep pits, the Erp appeared to approach a lower limit. For creviced specimens with shallow penetrations, there was considerable scatter in the data. This could have been related to the location of the localized attack within the crevice region. Because of the stochastic nature of localized corrosion, attack did not always occur at the same location inside the crevice. When depths of attack were shallow (i.e., low charge density), the diffusion distance for attack initiated near the center of the crevice area was greater than for attack initiated near the edge of the crevice- forming washer. As a result, the Erp for shallow attack initiated near the center of the crevice was expected to be lower than the Erp for attack located near the edge of the crevice. Variation in the measured Erp values for shallow penetrations ranging from −80 mVSCE to 20 mVSCE was attributed to variation of the initiation site location within the crevice area. The Erp ranged from −80 mVSCE to 20 mVSCE for a scan rate of 5 mV/s. When crevice attack was allowed to propagate to depths > 1.0 mm, however, the Erp approached a lower bound value of −80 mVSCE.

FIGURE 5.

Effect of prior pitting and crevice corrosion depth on Erp for pitting and crevice corrosion for alloy 825 in 1,000 ppm Cl solution at 95°C.

FIGURE 5.

Effect of prior pitting and crevice corrosion depth on Erp for pitting and crevice corrosion for alloy 825 in 1,000 ppm Cl solution at 95°C.

Close modal

The effect of applied potential as well as corrosion potentials under various redox conditions on the time to initiate localized corrosion is shown in Figure 6. Results showed short-term tests and long-term tests conducted by holding test specimens at constant applied potentials using a potentiostat. The corrosion current density of the specimens were recorded throughout the duration of the tests. When localized corrosion was initiated, the current density increased from a passive current density of ≈ 10−7 A/cm2 to values > 10−4 A/cm2. In addition, corrosion products were observed on these specimens and localized corrosion was verified by optical examination at the conclusion of the tests. At potentials near the Ep measured in CPP tests (e.g., 600 mVSCE), the initiation time was on the order of 200 s. Decreasing the potential to 500 mVSCE resulted in initiation times of 200,000 s. In contrast, relatively short initiation times for crevice corrosion were observed even at applied potentials several hundred millivolts below the Ep. Figure 6 also shows pitting and crevice corrosion initiation time for specimens at open-circuit potentials under various redox conditions (1,000-ppm Cl solutions with either 1:1 or 1:7 Cu2+/Cu+ added as a redox couple). The error bands in these data points denote the variation in the corrosion potentials for the same redox couple used. No pitting corrosion was observed on the boldly exposed surfaces (i.e., the surface of the specimen not covered with a crevice forming device) of any of the creviced specimens. In addition, no pitting corrosion was observed on a boldly exposed specimen in which the open-circuit potential was in the range from 300 mVSCE to 440 mVSCE. It was clear that the initiation time for crevice corrosion under natural exposure conditions was consistent with that under applied potential conditions.

FIGURE 6.

Effect of applied and corrosion potentials under various redox conditions on pitting and crevice corrosion initiation time for alloy 825 in 1,000 ppm Cl solution at 95°C.

FIGURE 6.

Effect of applied and corrosion potentials under various redox conditions on pitting and crevice corrosion initiation time for alloy 825 in 1,000 ppm Cl solution at 95°C.

Close modal

Repassivation times are not shown in Figure 6. However, these experiments demonstrated that the repassivation time increased with an increase in po- tential.16  The Erp for alloy 825 measured in CPP tests with a 1,000-ppm Cl solution also is shown in Figure 6. It was evident that the initiation time for localized corrosion increased as the potential was decreased and approached Erp. At potentials < Erp, no localized corrosion was initiated. This was indicated by the results of several creviced specimens continuously maintained at applied potentials < Erp (Figure 6). The observation that no crevice corrosion had been initiated on long-term test specimens maintained below the Erp measured in short-term tests suggests that the Erp was a stable parameter, independent of time. Many of the long-term tests shown in Figure 6 were still ongoing at various applied and natural potentials on alloy 825.

Figure 7 shows the current density vs time data for an alloy 825 specimen tested in 1,000 ppm chloride solution at an applied potential of 0 VSCE, which was 100 mV below the maximum Erp value. Test data for 1,000 days showed that the specimen initially had a current density of 8 × 10−5 A/cm2, as well as some current spikes associated with the dissolution of the Cr-depleted surface layer. In the later stages of the test, the current density was much lower, on the order of 1.0 × 10−7 A/cm2, reflecting the absence of any current increase associated with the initiation of localized corrosion. The weight of the specimen remained relatively constant and no localized corrosion was observed by metallographic examination. In addition, no localized corrosion was observed on specimens tested at less noble potentials.
FIGURE 7.

The current density for alloy 825 in 1,000 ppm Cl solution at 95°C held at 0 mVSCE, which is 100 mV below the Erp.

FIGURE 7.

The current density for alloy 825 in 1,000 ppm Cl solution at 95°C held at 0 mVSCE, which is 100 mV below the Erp.

Close modal

Figure 8 shows current density vs time for creviced alloy 825 specimens maintained at potentials between 300 mVSCE and 500 mVSCE. At the higher potentials, the current density quickly increased, indicating that crevice corrosion had been initiated rapidly. In contrast, at lower potentials, significant increases in the current density were observed only after some incubation time. At the conclusion of the tests, all specimens were examined. For specimens tested at 400 mVSCE and above, crevice corrosion was initiated on at least 14 of the 24 crevices formed by the serrated polytetrafluoroethylene (PTFE) crevice washer. For specimens tested at potentials of 350 mVSCE or less, only three or four crevices actively corroded. For all specimens, the current density was calculated using only the area of the corroded regions. Each individual crevice had a surface area of 0.06 cm2 yielding a total creviced area of 1.44 cm2 using two serrated crevice washers with 12 teeth per washer.

FIGURE 8.

Effect of applied potential and time on crevice corrosion growth under multiple crevice washers. Alloy 825 in 1,000 ppm Cl solution at 95°C.

FIGURE 8.

Effect of applied potential and time on crevice corrosion growth under multiple crevice washers. Alloy 825 in 1,000 ppm Cl solution at 95°C.

Close modal
To better understand the pit growth and repassivation process, pit propagation tests were conducted on alloy 825 in a 1,000-ppm chloride solution at 95°C, using lead-in-pencil pit specimens with initial pit depths of 5 mm. Pits were activated at 500 mVSCE to 600 mVSCE before the potential was decreased to values from −300 mVSCE to 100 mVSCE. The results for several tests conducted with an initial pit depth of 5 mm are shown in Figure 9. When the potential was decreased from 600 mVSCE to −300 mVSCE, the current density rapidly decreased and then returned to values of 10−4 A/cm2. After several hours at this value, the current slowly decreased to values < 10−7 A/cm2, characteristic of a passive behavior. After repassivation of the pit, the calculated propagation rate was < 10−10 cm/s. As in the case of boldly exposed specimens, repassivation time increased with an increase in potential. While some crevice corrosion occurred on all test specimens, the degree of crevice corrosion was found to be much greater at 0 mVSCE. With time, complete repassivation of the pits was observed at potentials ≤ to 0 mVSCE.
FIGURE 9.

Effect of applied potential and time on single-pit growth rate using a lead-in-pencil-type electrode. Alloy 825 in 1,000 ppm Cl solution at 95°C.

FIGURE 9.

Effect of applied potential and time on single-pit growth rate using a lead-in-pencil-type electrode. Alloy 825 in 1,000 ppm Cl solution at 95°C.

Close modal

DISCUSSION

Significance of Repassivation Potential

These results have been used in the Engineered Barrier System Performance Assessment Code (EBSPAC) to predict HLW container lifetimes.5  In this code, the Erp is used as a threshold parameter and the Ecorr is used as an enabling parameter that triggers localized corrosion and SCC when it exceeds Erp. For deep pits and crevices (i.e., >1 mm), Erp is mainly dependent on the material, Cl concentration of the bulk environment, and temperature. Hence, for a given material/environment combination, Erp is a stable parameter that is independent of time. Certain anionic species, such as (Figures 2 and 3) may decrease Erp while others such as NO3 may act as inhibitors and increase Erp. In the specific case of a HLW repository, the environment surrounding the WP will vary with time. Modeling calculations have predicted that the oxygen concentration,27  pH, temperature, and Cl concentration will change significantly after emplacement of the HLW.28  As shown in Figures 2 and 3, Erp can be measured rather rapidly using the CPP technique and can be expressed in terms of a well-established logarithmic relationship to the Cl concentration.29  The reduction in Erp as the Cl concentration and temperature are increased indicates that, in environments with low Cl concentrations (and/or low temperatures), a high Econ. is needed to initiate localized corrosion. However, as the aggressiveness of the environment is increased, localized corrosion can be initiated at lower values of Ecorr.

One of the important findings of this study is the constancy of Erp with depth of pitting or crevice corrosion beyond a certain depth. As mentioned previously, the use of Erp fell into disfavor because some initial investigations that were limited to very shallow pit depths showed that there was no lower bound value for Erp.19  The effect of pit depth, expressed as charge density (Q), on Erp for a variety of alloy-environment combinations, as compiled from the literature data, is shown in Figure 10. Some of the papers referenced report only current density measurements, while others report pit depth and current density.19,23,30-34  Without knowledge of the pit depth distributions, calculating maximum pit depths is not possible. Pit depth, however, is proportional to Q. For hemispherical pits, the penetration depth is proportional to Q1/3. The absolute value of Erp, of course, varies between systems, but it is clear that, for each system, (i.e., low values of Q) if sufficient pit growth occurs, the Erp attains a stable, lower-bound value. For shallow pit depths, the Erp is a strong function of depth for all of the systems investigated. At a more microscopic scale, experiments on thin metallic films of different thicknesses produced by sputtering have shown that Erp initially decreases with pit growth, but then increases slightly and attains a stable value.35  The summary of literature investigations plotted in Figure 10 clearly indicate that only a few investigators have explored sufficiently the relationship between localized corrosion penetration depth and Erp. From this figure, the early conclusion of Wilde is shown,19  that Erp is not a bounding parameter was based on experiments where pit growth was insufficient to attain stable Erp values. Studies conducted at the CNWRA30  and else- where32  have shown that Erp is a lower bound value independent of pit depth. The determination of Erp is dependent on test parameters such as the backward scan rate from the pit growth potential.29  Dunn, et al., have shown that the time required to repassivate a pit of a given depth increased with the potential.16  Thus, a slower backward scan rate would be expected to result in higher Erp. This accounts for slightly lower Erp values for deep pits presented in Figure 5 (backward scan rate of ∼ 5 mV/s), where Erp was in the range of −100 mVSCE to −50 mVSCE compared to those in Figure 3 (scan rate of 0.167 mV/s) where the Erp was in the range of −50 mVSCE to 100 mVSCE for a chloride concentration of 1,000 ppm (0.028 M). The Erp measured using the lead-in-pencil- type electrode (Figure 9) was similar to (0 mVSCE) than that of the CPP tests shown in Figure 3, despite a much deeper pit (5 mm). These experiments had the slowest scan rate because they were conducted potentiostatically. Since the potential is expected to change slowly over time, Erp values obtained using slow scan rates should be the most appropriate for predicting long-term container performance.
FIGURE 10.

Effect of pit/crevice depth on Erp on a number of alloy-environment combinations as found by a variety of investigators.

FIGURE 10.

Effect of pit/crevice depth on Erp on a number of alloy-environment combinations as found by a variety of investigators.

Close modal

Erp values shown in Figure 6 span a potential range of 150 mV. This span is a result of the test technique. The 150-mV range of Erp values would likely be reduced if the material was tested using either a much slower scan rate or a potentiostatic test method. The scan rate used in the present study (0.167 mV/s) may result in some variation in maximum penetration depth that, in turn, alters the measured Erp values (Figure 5) for shallow penetrations. While these test methods may reduce the range of the measured Erp values, it is expected that the value of the Erp of a material in a given environment will be distributed over a range of potential values. The key difference between the distributions of Ep and Erp is that the distribution of Ep values decreases with time, whereas the distribution of Erp values are not time-dependent. This is readily apparent from the results shown in Figure 6.

Potentiostatic testing carried out for times ranging from a few days to > 1,000 days (Figure 6) clearly indicates that localized corrosion can be initiated at potentials far less than the Ep measured in CPP tests (Figure 3). Thus, the Ep measured in CPP tests is not a useful parameter to predict the long-term initiation of localized corrosion. At potentials close to Ep, multiple crevice corrosion sites were initiated on the specimens. As the potential was reduced, the number of crevice corrosion sites decreased. For specimens tested at potentials near the Erp, localized attack usually occurred at only one site. From the results shown in Figure 6, it is apparent that the initiation time for the creviced specimens under applied potentiostatic control compared well with the initiation times measured in open-circuit tests with Cu2+/Cu+ redox couple added to the solutions to increase the redox potential of the environment. These data provide increased confidence that the concept of repassivation potential can be applied to materials exposed to natural redox conditions. No localized corrosion has been initiated to date at potentials below the Erp even after 1,000 days of testing, indicating that this parameter, as measured in short-term tests, can be used as the basis for long-term predictions of container performance.

Pit Propagation Rate

Results obtained with alloy 825 single-pit specimens (Figure 9) indicated that, at potentials below the Erp, the dissolution rate is approximately the same as the passive dissolution of alloy 825 (< 2 × 10−7 A/cm2 or < 2 × 10−3 mm/y). At potentials above the Erp, the simulated pits dissolved at a very rapid rate (1 × 10−2 A/cm2 or 102 mm/y). However, these pit growth rates were obtained under potentiostatic conditions where there was no cathodic limitation. Additionally, the geometry of the single-pit specimens used inert walls that did not allow the lateral propagation of the localized corrosion front. Thus, the propagation rate cylindrical pits formed using these specimens do not represent pit propagation rates under natural conditions.

The measured crevice corrosion propagation rate for type 316L SS in 1 M chloride solution is shown in Figure 11. These measurements of the propagation rates of pits in creviced specimens from open-circuit tests indicated that the propagation rates at open circuit, while still unacceptably high, decreased with time.26  The rate-controlling process for crevice corrosion growth under natural potential conditions was not clear at this time. The exponent of −0.62 in the rate equation lies between −0.5 for transport or internal ohmic-controlled pit growth and −0.67 for external ohmic-controlled pit growth mechanisms.36  In either case, these results clearly indicated that the pitting multiplication factor approach used in some performance assessment models, such as that in DOE TSPA, was not valid for corrosion-resistantmaterials.7  This is because the pitting multiplication factor assumes that the pit growth rate is proportional to the uniform corrosion rate. For the corrosion-resistant materials, pit growth rates are generally much higher than the uniform corrosion rates and have a very different time-dependance as a result of differences in the rate-controlling processes.
FIGURE 11.

Crevice corrosion growth rate of type 316L SS under open-circuit conditions in a 1 M chloride solution.

FIGURE 11.

Crevice corrosion growth rate of type 316L SS under open-circuit conditions in a 1 M chloride solution.

Close modal
Long-Term Prediction of Localized Corrosion

For localized corrosion to be initiated in a repository setting, the Ecorr of the corrosion-resistant barriers must be more noble than the Erp. Initial measurements of the Ecorr for type 316L SS shown in Figure 2 indicated that localized corrosion in deaerated environments was not likely. However, the Ecorr was greater than Erp in air-saturated environments and the possibility of localized corrosion existed. Although clearly more resistant, especially at low chloride concentrations, alloy 825 also was susceptible to localized corrosion when Ecorr exceeded Erp. Long-term corrosion potential data for alloy 825 in a 1,000-ppm Cl-based solution are shown in Figure 4. The increase in the corrosion potential of this specimen was most likely the result of the dissolution of the mill-finished surface, which previously had been shown to be deficient in Cr compared to the bulk material. As a result, the initial surface layer corroded at a faster rate than the polished surfaces. This was supported by the initial weight loss observed in the early stages of exposure. The Cr-depleted layer yielded a lower Ecorr because the alloy had a higher passive current density. After the Cr-depleted layer was removed, the exposed surface contained an increased amount of Cr. Passive films formed on surfaces that were not depleted in Cr had improved corrosion resistance and lower passive dissolution rates.

Since the Ecorr was determined by the potential at which the sum of all anodic currents and all cathodic currents equaled zero, a decrease in the passive current density resulted in an increase in the Ecorr. Other candidate corrosion-resistant WP materials such as alloys 625 and C-22 also will have a mill-finished surface that is depleted in Cr as a result of the production process.37  A similar increase in Ecorr was observed with thermally oxidized specimens. The corrosion potential of alloy 825 in air-saturated 1,000-ppm chloride solutions increased by as much as 200 mV after oxidation in dry air at 100°C for 30 days prior to the exposure to aqueous environ- ments.23  Thickening and aging of the oxide layer in air-saturated solutions or during thermal exposures to dry air can reduce the anodic dissolution kinetics resulting in an increase in the Ecorr.38 

The Ecorr measurement shown in Figure 4 emphasizes the need to monitor the Ecorr in long-term laboratory and field immersion tests. Because the parameters used in calculating the Ecorr from electrochemical models are uncertain, especially for highly irreversible reactions such as oxygen reduction,6  calibration of models against measured open-circuit potentials is necessary when these electrochemical parameters are used to predict the long-term evolution of corrosion potentials. The Ecorr can decrease after significant localized corrosion has occurred because of the influence of active pit growth region on the overall Ecorr. The extent of decrease would depend on the relative area ratio of active localized corrosion area to the passive corrosion area. The lowering of Ecorr after significant localized corrosion growth occurs was not addressed in the present study.

Several field experiences can be analyzed to show the applicability of Erp in predicting the occurrence of localized corrosion. Bernhardsson and Mellstrom reported failure caused by pitting of alloy 825 compressor coolers on a gas lift platform.39  The environment was chlorinated seawater at 42°C, and failure took place after 4 months (minimum penetration rate was estimated to be ∼ 5 mm/y). The Erp for alloy 825 in 0.5 M chloride (the approximate concentration of chloride in seawater is 0.55 M) at 20°C is 60 mVSCE.11  Since the Erp was expected to decrease with an increase in temperature and the Ecorr in chlorinated seawater had been measured to be ∼ 200 mVSCE,40  this in-service failure was consistent with the use of Erp as a determining factor for the initiation of localized corrosion. In stagnant, aerated 3% NaCl solution at 60°C, Bernhardsson and Mellstrom also found that type 316L SS underwent crevice corrosion after 2 months, while alloy 825 did not.39  Although these are still relatively short-term tests, the results are consistent with even shorter-term Erp values. For example, Okayama, et al., reported a value of −360 mVSCE for Erp of type 316 SS and −100 mV for alloy 825 in 3% NaCl at 80°C.41  The Ecorr of these alloys in aerated chloride solution at 95°C has been reported to be —280 mVSCE and is anticipated to increase slightly with a decrease in temperature as a result of higher solubility of oxy- gen.11  This value of Ecorr is higher than the Erp of type 316L SS and lower than that of alloy 825, thus explaining the localized corrosion of the former.

Protection against corrosion in highly oxidizing pulp and paper bleach plant washer fluids using controlled potential has been reported.42  In this case, the material is typically a type 317L SS (Fe-19% Cr-12% Ni-3.5% Mo [UNS S31703]) and the Ecorr in chlorine- containing waters is in the range from 100 mVSCE to 500 mVSCE, depending upon residual chlorine. This is far higher than the potential at which crevice corrosion was observed in potentiostatic tests (considered to be roughly equal to the Erp). For type 317L SS in this environment (typically 1,000 ppm Cl to 5,000 ppm Cl and pH 2), Erp was measured to be ∼ −30 mVSCE. In these cases, cathodic protection by maintaining the potential at −500 mVSCE resulted insuccessful performance of the washer drums for > 10 years.

In filtered, natural seawater (mean temperature: 25.2°C), Kain observed crevice corrosion on type 316L SS after 60 days, but no crevice corrosion on alloy 59 (61% Ni-22% Cr-15% Mo [UNS N06059]).43  Slight crevice corrosion was observed on alloy C-276 (Ni-16% Cr-5% Fe-16% Mo-4% W [UNS N10276]). Kain measured a Ecorr in the range of 200 mVSCE to 300 mVSCE. In natural seawater, microbially enhanced cathodic reduction of oxygen is believed to be responsible for the high Ecorr values. Again, the observed crevice corrosion of type 316L SS is consistent with the measured Erp of 0 mVSCE at that tempera- ture.41  Many of Fe-Ni-Cr-Mo alloys were tested in natural, filtered seawater at 30°C for 30 days by Hack.44  Specimens were fitted with crevice washers of PTFE. In this limited duration test, alloys C-276 and 625 (Ni-22% Cr-5% Fe-9% Mo-4% Nb) suffered no localized corrosion, while type 316 SS and alloy 825 suffered significant crevice corrosion. The Erp for alloy 825 in 0.5 M Cl was −60 mVSCE, which was lower than the Ecorr values reported for these alloys in natural seawater. An evaluation of all the published data on the Erp values of alloy 825 and type 316/316L SS indicates that, if the chloride concentration becomes higher than = 0.5 M, alloy 825 may not be significantly better than type 316L SS in localized corrosion resistance. The relatively superior localized corrosion resistance of alloys 625 and C-276 compared to alloy 825 is consistent with experimental evidence that the Erp of alloys 625 and C-276 is higher than that of alloy 825.41  The ongoing experimental studies at CNWRA on alloys 625 and C-22 also support these field observations.

Mechanistic Considerations of Initiation and Repassivation

Pit and crevice corrosion initiation have been thought of as originating from different processes. Pit initiation is well established as a stochastic process, with many metastable pits nucleating and repassivating dynamically on the metal surface below the critical potential for stable pit initiation.29,45  Crevice corrosion initiation traditionally has been thought of as a deterministic process dictated by the development of a critical solution composition with a high chloride concentration and a low pH deep in the crevice such that the passive film is dissolved46  or a critical potential drop caused by ohmic resistance of the crevice electrolyte such that active corrosion is initiated.47  However, recent evidence has shown that crevice corrosion in highly corrosion-resistant alloys occurs essentially by the nucleation and growth of pits. Sridhar and Dunn showed that critical solution composition and potential drop did not occur until an increase in crevice current was detected, indicating that pits nucleated inside the crevice first (measured by an increase in current density) followed by growth of these pits to a sufficient spatial extent to affect crevice chemistry.48  Shinohara, et al., used the Moire fringe technique to show that pits nucleated inside the crevice and then spread in depth and lateral extent.49  Hence, at least for the corrosion- resistant materials, a view of crevice corrosion is emerging where the crevice is thought to provide a long enough diffusion path to stabilize metastable pits.50  The observation of the similarity of Erp for deep pits and crevices in this study reinforced this point of view.

Pit initiation theories can be classified as those that focus on the properties of the passive film and factors leading to the destabilization of the film and those that focus on processes in embryonic pits and factors leading to the stabilization of these pits.29,45  The present study indicated that the latter theories are more relevant to predicting the long-term initiation of localized corrosion. However, this does not imply that passive film properties are unimportant for the long-term initiation of localized corrosion of a material. Clearly, the passive film properties determine the evolution of corrosion potential that dictates the triggering of localized corrosion. Equally important, if the repassivation process is conceived as a competition between maintaining a salt film and a concentrated chloride solution in nascent pits vs the formation of a passive oxide film, then properties of passive films will play a role.

Burstein and Mattin proposed that the embryonic pits covered by the remains of the original oxide passive film acquire a critical solution chemistry because of metal dissolution inside them and consequent chloride ion migration.51  This solution may reach saturation with respect to metal chloride depending upon the dissolution rate (current) and volume of pit embryo. The presence of increased chloride concentration and lower pH as a result of hydrolysis results in a critical pit solution that maintains active dissolution of the metal inside the pit and an increase in the metastable pit size. From this point on, the metastable pit can adopt two paths: the oxide film cover can rupture (perhaps caused by osmotic pressures created by the chloride concentration difference between the inside and outside of the pit embryo) and the in-rushing external solution can dilute the pit solution and repassivate the pit; or the pit embryo can attain such a size that the oxide ruptures and in-rushing solution does not have a significant dilution effect so that the pit embryo continues to grow and becomes a stable pit. The path taken will be determined by many factors, including the dissolution current (applied potential), geometry, and external solution concentration. This concept is essentially similar to the pit initiation model originally proposed by Galvele.52  If the applied potential is sufficiently high, continued dissolution of the metal helps to maintain the critical concentration and no repassivation occurs. If a crevice is present or the surface roughness is high, the transport path is lengthened and dilution by external solution is minimized. Thus, the pit is stabilized. Scully, et al., expressed a similar concept in terms of a critical ratio of pit current to pit radius (assuming hemispherical pits).53  The presence of a critical pit chemistry for maintaining active pit growth recently has been verified experimentally by Sridhar and Dunn.54 

If the model proposed by Galvele is valid,52  the Erp would be expected to decrease continually with the depth of the pit. This decrease of Erp with pit depth would result from the lengthening of the transport path that would require less current (or potential) to maintain the critical pit chemistry. However, Erp is an externally measured potential. The potential at the bottom of the pit (the pitting front) is decreased by ohmic resistance in the pit solution, which increases with pit depth.55  Hence, these competing factors would ensure that, in deep pits or crevices, the Erp would become constant above a certain pit depth. This was observed in the present case.

The detrimental effect of on lowering Erp also can be explained by the previously discussed mechanism. It has been shown that, in the active condition, adsorbs as atomic sulfur on the metal surface.45  Further, it has been shown that adsorbed sulfur increases from 5 to 12 times the active dissolution current, depending upon the alloy system.45  In such a case, adsorbed sulfur will stabilize the pit embryo and prevent repassivation. Lower metal dissolution rates (lower applied potentials) then would be required to promote repassivation in the presence of . The detrimental effect of reduced sulfur species may be important because of the possibility of the presence of sulfate-reducing bacteria (SRB) near the HLW containers. Although the end products of SRB activity are sulfides, thiosul- fate can be produced as a transient species in the reduction of sulfate56-58  have shown that the effect of thiosulfate in accelerating pit initiation is similar to that of sulfide in aqueous chloride solutions. In addition, the dissolution of sulfide inclusions in the metal also can result in the formation of thiosulfate in solution.59 

CONCLUSIONS

  • Results of this investigation showed that the repassivation potential for alloy 825 and type 316L is dependent upon temperature and composition of the environment, in particular, the oncentration of Cl anions. Numerous test techniques designed to test the adequacy of the repassivation potential have shown that pitting and crevice corrosion cannot be initiated at potentials lower than the repassivation potential. In addition, after initiation, all localized corrosion ceased to propagate if the potential of the material was reduced below the repassivation potential. After repassivation, the corrosion rate of the material was determined by the passive current density.

  • Long-term testing of alloy 825 indicated that the corrosion potential under air-saturated conditions increased over a period of 600 days to values higher than repassivation potential measured in short-term experiments. Crevice corrosion was observed when the corrosion potential was a few millivolts greater than the repassivation potential. Under these conditions, the propagation rate for localized corrosion will determine the lifetime of containers made of these materials. The long-term increase in corrosion potential is believed to be caused by the growth and aging of the passive film. Similar increases in potential were observed on specimens covered with a thermally aged oxide film over shorter periods of time. Many of the parameters that are necessary to calculate the corrosion potential in an air-saturated environment, such as exchange current densities for the anodic and cathodic reactions and transfer coefficients for the reactions, are not known for the substrates of interest. However, the measured values of corrosion potential can be used to calibrate the modeling of the evolution of the corrosion potential under repository conditions.

  • To judge the performance of a given alloy or container design over the expected performance period of a HLW container, it is necessary to characterize the environment contacting the container. Some of the key parameters include the time-dependent composition of the aqueous environment, especially chloride concentration, oxygen concentration, and temperature profiles of the WP. Other important parameters are related to galvanic effects between various overpack materials. From these results, it may be concluded that, for a given environment, the repassivation potential is conservative and an effective parameter to predict the long-term performance of corrosion-resistant materials.

ACKNOWLEDGMENTS

The authors acknowledge the helpful reviews and suggestions of V. Jain, B. Sagar, and B. Leslie.

(1)

UNS numbers are listed in Metals and Alloys in the Unified Numbering System, published by the Society of Automotive Engineers (SAE) and cosponsored by ASTM.

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