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Mariano Kappes obtained his Bachelor’s degree in Materials Science and Engineering at the Instituto Sabato in Argentina in 2006 and then obtained his PhD in 2011 at the Ohio State University. His PhD thesis was distinguished with the Morris Cohen Award, awarded annually by The Electrochemical Society to outstanding graduate research in the field of Corrosion. Since 2014, he is a research scientist at the National Agency of Atomic Energy in Argentina and a Professor at the National University of General San Martin. He holds a position at the National Scientific and Technical Research Council as scientist since 2015.
Fluorides, bromides, and iodides, despite being less common than chlorides, are present in various environments of industrial relevance. Stainless steels suffer pitting corrosion in solutions of all halides except fluorides, which can be understood considering that fluoride is the anion of a weak acid. The aggressiveness of the rest of the halides for pitting corrosion is on the order Cl − > Br − > I − for stainless steels with Mo content below 3 wt.%. Mo is not as effective in inhibiting Br − pitting corrosion as it is for inhibiting Cl − pitting corrosion. Most of those observations were rationalized based on the effect of anions on pit growth kinetics. Sensitized austenitic stainless steel suffers stress corrosion cracking (SCC) in solutions of all halides, albeit chlorides seem to be the most aggressive. Fluoride SCC is relevant for SCC under insulation of stainless steels, and standards and regulations developed to mitigate this problem consider this ion as aggressive as chloride. For the solubilized stainless steels, aggressiveness toward SCC is in the order Cl − > Br − . The SCC of solubilized stainless steels was not observed in solutions of F − and I − , and the possible reasons for this fact are discussed.
Chlorides are present in a wide variety of industrial environments ( Kolotyrkin, 1963 ; Brown, 1977 ), including fossil and nuclear power plants, food, paper, pulp, and chemical industry and petroleum refineries. Therefore, pitting, crevice corrosion, and stress corrosion cracking (SCC) of stainless steels were extensively studied in chloride solutions ( Brown, 1977 ; Cragnolino et al., 1981 ; Cragnolino & Macdonald, 1982 ; Haruyama, 1982 ; Sedriks, 1996 ; Frankel, 1998 ; Szklarska-Smialowska, 2005 ). Given those considerations, this review will be focused on SCC and pitting corrosion of stainless steels in solutions of halides other than chlorides, in particular, at a temperature below 100°C.
The SCC of stainless steel was observed in solutions of all halide ions. The SCC in F − and I − solutions was only reported for alloys in the sensitized condition ( Table 1 ). Sensitization occurs by exposure of stainless steels to elevated temperatures during welding, heat treatment, or service conditions, causing chromium carbide precipitation at grain boundaries and depletion in chromium close to grain boundaries, thus, favoring intergranular (IG) attack and intergranular stress corrosion cracking (IGSCC). A remarkable observation is that IGSCC was observed in laboratory tests even when F − concentration in the solutions was as low as 1 ppm ( Ward et al., 1969 ; Theus & Cels, 1974 ). Pitting corrosion of stainless steels, on the other hand, occurs in solutions of all halides except F − ( Table 1 ). Moreover, F − can act as an inhibitor of pitting and crevice corrosion in chloride solutions ( Yamazaki, 1994 , 1996 ), and this was explained based on the fact that it is the anion of a weak acid.
Selected properties of the halides and overview of observed modes of corrosion of stainless steel in halide solutions.
Halide ion, X | F | Cl | Br | I | References |
---|---|---|---|---|---|
Radius of X (nm) | 0.133 | 0.181 | 0.196 | 0.219 | |
(kJ/mol) | −472 | −351 | −326 | −294 | |
pKa of HX | 3.17 | −7 | −9 | −10 | ; |
Concentration in seawater (ppm) | 1.3 | 19×10 | 65 | 0.05 (present as I and ) | ; ; ; |
Pitting of stainless steel in X solutions | No | Yes | Yes | Yes | ; ; |
SCC of stainless steel in X solutions | Only in sensitized condition | Yes | Yes | Only in sensitized condition | Rhodes, 1969; ; ; Griess et al., 1985; Takemoto et al., 1985; ; ; ; ; ; ; |
Fluorine (F), chlorine (Cl), bromine (Br), iodine (I), and astatine (At) are the elements of the halogen family that occur naturally on Earth. Astatine is the scarcest of the naturally occurring elements ( Greenwood & Earnshaw, 1997 ), and it is radioactive with a half-life of 8.1 h for its most stable isotope ( Lide, 2005 ). Therefore, At is not further discussed in this review. The rest of the elements in the halogen family achieve a stable configuration by forming diatomic molecules: F 2 and Cl 2 are gases, Br 2 is a liquid, and I 2 is solid at ambient temperature and pressure. However, due to their reactivity, halogens occur in nature as halides (X − ), and iodine also occurs as iodate ( IO 3 − ) ( Greenwood & Earnshaw, 1997 ). The abundance of elements in crustal rocks decreases with an increase in the atomic mass of the halogen element ( Greenwood & Earnshaw, 1997 ). While fluoride is more abundant than chloride in crustal rocks, fluoride is mainly present in minerals that are scarcely soluble in water, like fluorite or fluorspar (CaF 2 ), cryolite (Na 3 AlF 6 ), and fluorapatite (CaF 2 ·3Ca 3 (PO 4 ) 2 ), whereas chloride is mainly present in water-soluble rock salt (NaCl) ( Chambers & Holliday, 1975 ; Greenwood & Earnshaw, 1997 ). Chloride is the dominant anion in ocean water, where its concentration is 1.9 wt.% ( Greenwood & Earnshaw, 1997 ), about 15,000 times higher than F − ( Warner, 1971 ) and 290 times higher than Br − ( Stine, 1929 ). Iodine is less abundant than lighter halogens both in the earth’s crust and in ocean water. Iodine is a micronutrient necessary for various organisms ( Ito, 1988 ). In seawater, it can be present as I − and IO 3 − , with total inorganic iodine in ocean water at around 0.05 ppm ( Ito, 1988 ). The concentration of halide ions in seawater is summarized in Table 1 . Chloride dominates in aerosols in the atmosphere, and they can contact metals by direct deposition or after dissolution in rainwater. The main sources of chlorides in aerosols are the ocean, road deicing salts, or the products of the combustion of fossil fuels and residues ( Willison et al., 1989 ).
The stable oxidation number of fluorine and chlorine throughout the range of stability of water is −1 ( Pourbaix, 1966 ). In dilute solutions of hydrofluoric acid (HF), F − dominate at pH >3.17 ( Mccoubrey, 1955 ), which is the pk a of HF ( Table 1 ). Below this pH, HF dominates ( Pourbaix, 1966 ). In more concentrated solutions (greater than 3.8 g F/l), the bifluoride ion, HF 2 − , predominates in the pH region close to pk a ( Pourbaix, 1966 ). Hydrochloric acid, HCl, is a strong acid; hence, it is completely dissociated in water, and the stable form of chlorine in water solutions is as Cl − . Bromine is stable as Br − in almost the entire range of water stability, except for very low pH and oxidizing conditions, where it can react according to ( Pourbaix, 1966 ):
Iodine has the lowest standard reduction potential of the halogen group ( E 0 =0.62 V SHE ), and aqueous I − solutions are thermodynamically unstable in the presence of dissolved oxygen, reacting to give IO 3 − ( Pourbaix, 1966 ):
IO 3 − is the thermodynamically stable form of iodine in seawater, but I − is produced by biologically mediated reduction of IO 3 − or under reducing conditions ( Ito, 1988 ). In acid media, the intermediate formation of I 2 or I 3 − (triiodide) occurs according to ( Pourbaix, 1966 ):
4.1 fluorides.
In industrial applications of stainless steels, some relevant sources of F − are thermal insulation ( Takemoto et al., 1985 ; Whorlow et al., 1997 ) and electrode coatings and fluxes used during welding operation ( Ward et al., 1969 ; Takemoto et al., 1985 ). Fluorinated hydrocarbons used in lubricants and gaskets can liberate F − at the temperature of operation of pressurized or boiling water reactors ( Brown, 1977 ). Fluorides are present in pickling solutions of stainless steels. Pickling is a surface process that removes metallic contamination, besides welding and heat treatment scales. This process is typically performed by immersion in nitric and hydrofluoric acid solutions, but it is explicitly not recommended for sensitized stainless steels by ASTM A380 ( ASTM A380-17, 2017 ), due to possible intergranular attack and SCC even under ambient conditions ( Berry et al., 1973 ). Fluoridation is the addition of F − to a public water supply to prevent tooth decay, and it is targeted to maintain a concentration of about 1 ppm in water ( Greenwood & Earnshaw, 1997 ). Cryolite is used in the electrolysis of alumina for metallic aluminum production ( Greenwood & Earnshaw, 1997 ). Fluorine has a key role in the nuclear fuel cycle ( Crouse, 2015 ), where it is used to produce gaseous UF 6 from which fissionable isotopes can be separated by different technologies.
Industrial applications of bromine compounds used to be dominated by 1,2-dibromoethane or ethylene dibromide, a compound added to gasoline as a lead scavenger, until environmental legislation limited the use of lead-based anti-knock additives ( Greenwood & Earnshaw, 1997 ; Thomas et al., 1997 ). Ethylene dibromide and methyl bromide have applications as pesticides as well ( Greenwood & Earnshaw, 1997 ). Other major applications of bromine compounds are as flame retardants, in plastics, fibers, rugs, and carpets. Silver bromide, AgBr, is an active compound in photographic films. Some catalysts for chemical industries contain bromides, and this caused pitting of a type 316L stainless steel component handling organic acids ( Ohtsu & Miyazawa, 2012 ). Bromides are present in high-density brine completion fluids, applied to deep oil and gas wells to balance the high pressure of the well and maintain the borehole stability ( Liu et al., 2014 ). Those fluids typically contain bromides or chlorides of Ca, Zn, and Na ( Sridhar et al., 2017 ). Calcium bromide is used for controlling mercury emissions of coal-fired power plants ( Ladwig & Blythe, 2017 ). The salt decomposes in the furnace to yield bromine or hydrogen bromide, and they react with elemental mercury. Oxidized mercury compounds are then more easily captured in downstream wet scrubbers for flue gas desulfurization. However, bromides can cause localized corrosion problems of materials used for wet scrubbers ( Ozturk & Grubb, 2012 ).
Concentrated LiBr brines have applications in absorption refrigeration systems ( Itzhak & Elias, 1994 ; Srikhirin et al., 2001 ). The main advantage of those systems is that a compressor is not used. Water is used as the refrigerant, and when it evaporates at low pressure, heat is absorbed from the environment. This water vapor is absorbed by concentrated LiBr brines. The brine, now diluted, is heated in a generator. The brine becomes concentrated, and evaporated water from it is condensed in a heat exchanger. The condensed water and concentrated brine are now ready to start a new cycle. Stainless steels are frequently used in metallic parts of those systems, which motivated corrosion studies in concentrated (above 50%) LiBr brines at temperature in the range from 80°C to 160°C ( Griess et al., 1985 ; Guiñon et al. 1994 ; Itzhak & Elias, 1994 ; Itzhak et al., 1996 ).
Iodine solutions are commonly used in redox titrations. Iodine is scarcely soluble in water, but solubility increases in iodide solutions by the following complexation reaction ( Harris, 2007 ):
I 3 − solutions can be used to titrate solutions of reducing agents, using starch as an indicator. Oxidizing solutions are treated with I − to produce an excess of I 3 − , which is then titrated with thiosulfate ( Harris, 2007 ).
Triiodide solutions have applications as antiseptic for cuts and wounds ( Chambers & Holliday 1975 ). Stainless steels 304L and 316L exhibit pitting in iodine solutions or solutions exposed to iodine vapor, as reported by Tsukaue et al. (1993 , 1994a , b ) at temperatures in the range from 50°C to 80°C. Triiodide is involved in the cathodic reaction (Eq. (3). Iodides produced in this reaction favor the solubility of iodine so that the reactant of the cathodic reaction is regenerated by Eq. (5). This causes the accumulation of triiodide in stainless steels, resulting in stable pit growth ( Tsukaue et al., 1994a ), b ). Iodine is a dangerous fission product that could be released in accidents of nuclear power plants ( Wren et al., 1999 ). The amount of iodine released to the environment is monitored with sampling systems that have stainless steel lines that transport gases through filters and absorbers where the presence and concentration of radionuclides are analyzed ( Evans & Nugraha, 2002 ). It was shown that with water vapor content above 75% and above 10 −9 m I 2 in the gas, aqueous pitting corrosion by iodine is possible ( Evans & Nugraha, 2002 ). Besides the integrity problem in the material of the sampling line, the amount of iodine transmitted through the line is reduced, causing errors in the monitoring process ( Evans & Nugraha, 2002 ).
Pitting corrosion of stainless steels involves breakdown of the passive film, metastable pitting, and stable growth of pits ( Frankel et al., 2017 ). Breakdown of the passive film was explained by halide adsorption and thinning of the passive film, halide penetration, and the film-breaking mechanism, as reviewed in-depthly elsewhere ( Frankel, 1998 ; Szklarska-Smialowska, 2005 ; Soltis, 2015 ). Halide ions affect properties of the passive film, for example, XPS studies show that the thickness of the passive film of iron in buffer solution decreases as the concentration of Cl − , Br − , and I − increases ( Khalil et al., 1985 ), and Cl − is the most aggressive for a given concentration. The point defect model was also applied to analyze passivity breakdown in solutions of the different halide ions ( Macdonald & Lei, 2016 ). The model successfully predicts that the ability of halide ions to cause passivity breakdown is in the order F − ≪Cl − >Br − >I − . According to this model, passivity breakdown first requires absorption of the halide ion into the passive film. In short, halide ions absorb by occupying surface oxygen anion vacancies, and this process requires expansion of the oxygen vacancy to accommodate the halide ion with a larger diameter, dehydration of the halide ion, and insertion of the halide ion into the expanded oxygen vacancy. By considering the Gibbs free energy required for each of those processes, a minimum in breakdown potential was predicted for Cl − solutions. The exceptionally high stability of hydrated F − ions is elucidated in Table 1 by the Gibbs free energy values of halide anion hydration, Δ G 0 hyd , ( Macdonald & Lei, 2016 ). The high energy required for F − dehydration results in the absence of passivity breakdown in F − solutions.
Despite those effects of halide ions on properties and breakdown of the passive film, it is known that breakdown events of the passive film can occur at a very high rate, as evidenced during metastable pitting at potentials below the stable pitting potential, E pit ( Frankel et al., 2017 ). Those frequent events of passive film breakdown and metastable growth are not a problem for structural integrity if repassivation of the passive film is rapid ( Frankel et al., 2017 ). Furthermore, stainless steels for industrial applications usually have appreciable fractions of inclusions or second phase particles, which provide favorable sites for pit initiation and possibly sulfur-containing aggressive species to the local environment ( Frankel, 1998 ). Therefore, the main attention is often targeted to understanding the conditions required for stable growth, or in other words, those that make a metastable pit grow stably ( Frankel, 1998 ; Newman, 2001 ; Li et al., 2018 ). Like many other variables that affect pitting corrosion resistance ( Newman, 2001 ), the aggressiveness of the different halide ions can be understood in terms of their effect on stable pit growth rather than on pit nucleation.
The aggressiveness of the different halide anions can be ranked by comparison of pitting potential ( E pit ), the repassivation potential ( E rp ) or the critical pitting temperature (CPT) at a given molar concentration of the anion, or by evaluating the minimum amount required for stable pit growth ( Szklarska-Smialowska, 2005 ).
Hydrofluoric acid is a weak acid, with a pK a of 3.17 ( Mccoubrey, 1955 ) ( Table 1 ). The rest of the acids are strong, meaning that they completely dissociate in water. The localized acidification theory of pitting corrosion ( Galvele, 1976 ) predicts that the anions of weak acids act as inhibitors. In accord with those considerations, pitting corrosion of stainless steels was not observed in fluoride solutions at room temperature ( Streicher, 1956 ; Koch, 1993 ). Studied materials where this was verified include quenched and tempered martensitic 403 stainless steel ( Pahlavan et al., 2016 ), supermartensitic 13 Cr stainless steel ( Macdonald & Lei, 2016 ), sensitized 304 stainless steel ( Zucchi et al., 1988 ), 17 wt.% Cr, 12 wt.% Mn and 0.61 wt.% N stainless steel, and 18 wt.% Cr with 9 wt.% Ni stainless steels ( Tzaneva et al., 2006 ). The absence of pitting was confirmed by either observation of specimens after the test or absence of hysteresis in the return potential scanning curve.
While type 304 sensitized stainless steel suffer intergranular attack and SCC in the presence of F − ions ( Ward et al., 1969 ; Theus & Cels, 1974 ), breakdown of the passive layer was not observed in the absence of stress at room temperature ( Trabanelli et al., 1988 ; Zucchi et al., 1988 ). The Cr content near the grain boundary of sensitized stainless steel is controlled by the bulk chemical composition of the stainless steel, time, and temperature in the carbide precipitation region, and grain size. For type 304 stainless steel, scanning transmission electron microscope (STEM) studies coupled with X-ray energy-dispersive spectroscopy (EDS) revealed that the Cr content near the grain boundary can decrease to 13 wt.% or less ( Rao, 1979 ; Ford & Silverman, 1980 ; Thorvaldsson & Salwén, 1984 ). Despite this depletion in Cr at the grain boundary, as the potential was increased in the anodic direction, polarization curves of sensitized type 304 stainless steel in F − solutions at room temperature exhibited a passive behavior, followed by a current increase in the potential range corresponding to oxygen evolution ( Zucchi et al., 1988 ). No hysteresis was observed on the potential scan in the noble direction, confirming that localized corrosion did not occur even at the most positive potentials ( Zucchi et al., 1988 ). In accord with those studies, anodic polarization in F − solutions did not cause breakdown of the passive layer in martensitic stainless steels ( Pahlavan et al. 2016 ), where the bulk Cr content was close to the Cr content near the grain boundary of sensitized 304 stainless steel.
Despite stable pitting is not observed in fluoride solutions, Macdonald and Lei (2016 ) reported metastable pitting for a 13 Cr martensitic stainless steel in 0.1 m NaF solution. The rate of metastable pitting events and the average current density increased in the order F − <I − <Br − <Cl − . The average current density of metastable pits in F − solutions was one order of magnitude lower than in Cl − solutions ( Macdonald & Lei 2016 ). The low frequency of those events and their low current might explain why other authors working in similar systems reported the absence of metastable pitting in F − solutions ( Pahlavan et al., 2016 ).
It was shown that F − increases the pitting initiation ( Yamazaki, 1994 ) and crevice corrosion initiation and repassivation ( Yamazaki, 1996 , 1997 ) potential of type 304 stainless steel in Cl − solutions. This increase in the pitting potential when F − is added to Cl − solutions was also reported by Pahlavan et al. (2016 ) for type 403 martensitic stainless steel. Interestingly, when F − was added to Cl − solutions, the rate of nucleation of metastable pits increased ( Pahlavan et al., 2016 ). Therefore, F − appear to inhibit the stable growth of pits, and this was explained on the basis of Galvele’s localized acidification theory ( Galvele, 1976 ). The high pK a of HF results in F − bonding to H + at pit bottoms, therefore, inhibiting pit growth ( Pahlavan et al., 2016 ).
Stable pitting of stainless steels was observed in solutions of the rest of the halides. The strength of the acid increases from HCl to HI ( Table 1 ) ( Chambers & Holliday, 1975 ), but because they are all strong and completely dissociated in water, this difference in strength can only be observed in solvents more acidic than water, for example, acetic acid. In aqueous solution, this difference in strength is irrelevant because they are all leveled out to H 3 O + ( Chambers & Holliday, 1975 ), and in fact, the aggressiveness of the different anions for pitting corrosion of stainless steels and passive iron in aqueous solutions generally increases from I − <Br − <Cl − ( Tousek, 1975 ; Janik-Czachor, 1979 ; Szklarska-Smialowska, 2005 ; Tzaneva et al., 2006 ; Pahlavan et al., 2016 ). A notable exception to this rule is observed in Mo alloyed stainless steels, where it was found that Br − ions are more aggressive than Cl − ( Guo & Ives, 1990 ; Kaneko & Isaacs, 2002 ) above a certain content of Mo, as discussed in a subsequent section.
Pitting potentials ( E pit ) decrease with an increase in the concentration of the aggressive ion, [X − ], according to ( Galvele, 1976 ; Szklarska-Smialowska, 2005 ):
where a and B are constants. The B value is controlled by the charge of the aggressive anion and complexation reactions between metal cations and anions inside the pit ( Nguyen et al., 2019 ). Without complexation, a B value of 59 mV is predicted for all halide ions at room temperature ( Galvele, 1976 ; Nguyen et al., 2019 ). For pure iron in borate buffer solutions containing Cl − , Br − , or I − , Janik-Czachor (1979 ) reported a B value of 100 mV, irrespective of the nature of the halide ion. For pure iron, this is an unusually high value, considering other studies ( Galvele, 1976 ; Nguyen et al., 2019 ) in Cl − solutions, where it was reported to be close to 59 mV. On the other hand, for stainless steels that suffer pitting at room temperature, the reported B values in Cl − solutions are close to 90 mV ( Galvele, 1976 ; Laycock & Newman, 1997 ; Nguyen et al., 2019 ). The B value increases with the stability constant of the metal-chloride complex, and stainless steels contain chromium, which forms complexes more stable than iron ( Nguyen et al., 2019 ). Measurements in martensitic stainless steels ( Pahlavan et al., 2016 ) revealed that the B value for Cl − , Br − , and I − was 130, 80, and 76 mV, respectively. For type 316 austenitic stainless steel, a slope of 91 mV was reported for Cl − solutions and of 75 mV in Br − solutions ( Pahlavan et al., 2019a ). In other words, an equal increase in halide ion concentration will cause a higher decrease in pitting potential for Cl − than for the rest of the halides.
The halide ion radius and the strength of the adsorbed halide-metal bond increase with atomic mass ( Table 1 ) ( de Castro & Wilde, 1979 ; Szklarska-Smialowska, 2005 ) so that larger iodide ion adsorption is thermodynamically more favored than fluoride adsorption. Therefore, the anions that adsorb more strongly to the metal are less aggressive for pitting ( Szklarska-Smialowska, 2005 ). A plausible explanation for this fact could be that adsorption of halide ions to the bare metal surface exposed to an acid solution can inhibit active dissolution of the metal, under certain conditions. Cl − , Br − , and I − additions to a H 2 SO 4 solution result in lower rates of active metal dissolution, as it is observed for mild steel ( Jesionek & Szklarska-Smialowska, 1983 ), pure nickel ( Abd El Rehim et al., 1986 ), and 18 Cr-8 Ni stainless steel ( Asawa, 1971 ). An inhibiting effect on carbon steel corrosion is also observed when KF is added to 0.01 m H 2 SO 4 solutions ( Sekine et al., 1994 ), but this is probably related to a buffering effect of F − ions.
The strength of adsorption of halides to active metal, the surface coverage of halides and the inhibition efficiency increase with increasing halide ion size ( Jesionek & Szklarska-Smialowska, 1983 ; Abd El Rehim et al., 1986 ). However, this inhibiting effect of halides on active dissolution of metals in acid solutions is observed below a concentration that is specific for each halide ion and material ( Asawa, 1971 ; Jesionek & Szklarska-Smialowska, 1983 ; Abd El Rehim et al., 1986 ) and around 10 −2 m ( Abd El Rehim et al., 1986 ). For Ni in sulfuric acid solutions ( Abd El Rehim et al., 1986 ), when halide ions are present at a concentration above 10 −2 m , they accelerate anodic dissolution and shift the active to passive transition to higher potential, and the aggressiveness increases in the order I − <Br − <Cl − . A similar effect is observed in 18 Cr-8 Ni stainless steels in 4 N H 2 SO 4 ( Asawa, 1971 ), for concentrations of halide ions up to 1 m . At the bottom of a pit under sustained growth, the concentration of halide ions can be much higher, on the order of 8 m or more ( Mankowski & Szklarska-Smialowska, 1975 ). The effect of halide ion on metal dissolution at those concentrations was studied with the artificial pit electrode, as will be discussed below.
Pitting corrosion studies with wire electrodes showed an “anomalous” higher aggressiveness of Br − vs. Cl − ( Carroll & Lynskey, 1994 ), under a wide range of experimental variables. The pitting potential was lower in Br − than in Cl − solutions for types 316, 304L, and 302 stainless steel wires, measured in solutions of pH ranging from 3 to 9 and halide concentration ranging from 0.1 m to 1 m . This anomalous effect was associated to a refining effect of the wire-drawing process on inclusions like sulfides ( Pistorius & Burstein, 1992 ), which act as favorable sites for pit initiation of stainless steels in Cl − solutions ( Szklarska-Smialowska, 2005 ). On the other hand, pitting potentials in bromide solutions were similar whether they were measured with wire electrodes or electrodes cut from bars, suggesting that pitting initiation in Br − solutions relies more on the adsorption of Br − ions than on interaction with inclusions. According to the authors ( Carroll & Lynskey, 1994 ), Br − has a greater tendency to adsorb on the metal surface than Cl − , and this has a greater effect on pit initiation when there is a low population of sulfides. Measurements of pitting potential of 316L stainless steels electrodes cut from bars, with a regular population of sulfides, showed the higher aggressiveness of Cl − vs. Br − usually observed for low Mo stainless steels.
Pit growth can be studied independently of pit initiation with the “lead-in-pencil” or artificial pit electrode technique ( Laycock & Newman, 1997 ; Kaneko & Isaacs, 2000 ; Ernst & Newman, 2008 ). A survey of the literature published up to date reveals that this technique was used to study pit growth kinetics in Br − vs. Cl − solutions ( Kaneko & Isaacs, 2000 ; Pahlavan et al., 2019a , b ), but no studies were reported in I − solutions. In this technique, a wire of the metal or alloy to be studied is embedded in epoxy, and after dissolving back the metal at an anodic potential, a one-dimensional artificial pit is eventually formed after the conglomeration of smaller pits ( Laycock & Newman, 1997 ). The electrochemical potential is then scanned in the active direction, while the current is measured. Figure 1 shows the polarization curves of artificial pit electrodes obtained in Br − vs. Cl − solutions, for an 18 wt.% Cr-12 wt.% Ni austenitic stainless steel ( Kaneko & Isaacs, 2000 ). Similar results were obtained by Pahlavan et al.(2019a) for a type 316 austenitic stainless steel. A region where current is independent of potential is observed in Figure 1 . In this region, a salt film is precipitated on the metal surface, and the current density of the unidimensional pit is diffusion limited ( i lim ) ( Laycock & Newman, 1997 ):
Anodic polarization curves of 18 wt.% Cr–12 wt.% Ni stainless steel, obtained for a 0.44-mm deep artificial pit electrode in 1 m LiCl and 1 m LiBr bulk solutions. The electrochemical potential was scanned at a rate of 5 mV/s in the active direction. E T is the transition potential, where anodic dissolution shifts from diffusion control to activation/ohmic control. Adapted from Corrosion Science, 42(1), Kaneko and Isaacs, Pitting of stainless steel in bromide, chloride and bromide/chloride solutions, 67–78, Copyright (2000), with permission from Elsevier.
where n is the average charge of metal ions, F is Faraday’s constant, D is the effective or average diffusion coefficient of metal cations, C sat is the molar concentration of metal ions in the saturated halide salt solution, and x is the depth of the unidimensional pit. Notice that Eq. (8) is based on Fick’s first law of diffusion and, therefore, is strictly valid during steady state. Because of metal dissolution, x increases with time. However, relative changes of x during the characteristic time of diffusion, x 2 / D , are small enough for the approximation to be valid ( Laycock & Newman, 1997 ). As a consequence, despite x increases as potential is scanned in the active direction, i lim is independent of potential for scan rates in the range of 1 mV/s to 10 mV/s, typically used in this experiment ( Laycock & Newman 1997 ; Li et al. 2019 ).
It is expected that the most prominent element in the alloy will precipitate first ( Bocher et al., 2010 ). For stainless steel, ferrous halide should dominate in the salt film. However, how chromium, nickel, and the rest of the alloying and impurity elements affect the precipitation of iron halides is not known in depth. Kaneko and Isaacs (2000 ) argue that the concentration of saturated FeBr 2 should be similar to saturated FeCl 2 and around 5 m . In that case, the reason for the higher diffusion-limited current density reported in the literature ( Kaneko & Isaacs, 2000 ; Pahlavan et al., 2019a ) and schematized in Figure 1 resides in a higher diffusion coefficient of metallic cations in Br − vs. Cl − solutions, which is roughly estimated to be within a factor of 1.5. As potential is decreased, eventually, the salt film is dissolved, which is indicated by a small hump in the i vs. E curve ( Ernst & Newman, 2008 ) ( Figure 1 ). Current then decreases sharply with potential, due to ohmic drop and activation control.
The transition potential, E T , is the minimum applied potential required for salt film precipitation at the pit bottom. From Figure 1 , this potential can be estimated as the potential, where E=i lim , on the left of the hump associated with salt film dissolution ( Laycock & Newman, 1997 ; Ernst & Newman, 2008 ). The actual potential at the pit bottom might be lower due to an ohmic drop in potential, and it can be estimated by different procedures detailed in the literature ( Gaudet et al., 1986 ; Laycock & Newman 1997 ; Li et al., 2019 ). An increase in E T measured with pencil electrodes correlated with an increase in E pit measured on flat electrodes, as the Br − /Cl − concentration ratio increased in solutions with a total halide concentration of 0.2 m ( Pahlavan et al. 2019b ). Kaneko and Isaacs (2000 ) reported similar results for 1 m Br − and 1 m Cl − solutions. According to Laycock and Newman (1997 ), stable pitting requires the precipitation of a salt film at the pit bottom, thus, explaining the correlation between E T and E pit . While a solution saturated in metallic cations is the most aggressive solution attainable at the pit bottom at a given temperature, recent studies confirm that stable pitting can be sustained in less concentrated solutions ( Srinivasan & Kelly, 2017 ; Li et al., 2019 ). For example, according to Srinivasan and Kelly (2017 ), stable pitting of a 316L stainless steel at room temperature requires a concentration around 50% of the saturated solution at the pit bottom ( Srinivasan & Kelly, 2017 ), otherwise the pit repassivates.
A deeper understanding of how measurements with the artificial pit electrode can explain E pit differences in Br − vs. Cl − solutions can be achieved with a suitable model for pit growth ( Li et al., 2018 ; Pahlavan et al., 2019b ). During pitting corrosion, metallic cations are produced by anodic dissolution at the pit bottom and diffuse out of the pit down a concentration gradient. Mathematically, a stable growth of pits requires that ( Li et al., 2018 )
where i diss,max is the maximum anodic dissolution current at the pit bottom for a given temperature, pit solution concentration, and potential, and i diff,crit is the critical current density for diffusion of metal cations out of the pit. In other words, the rate of production of metallic ions at the pit bottom must compensate its loss by diffusion out of the pit, otherwise, dilution and repassivation occur. The expression for i diff,crit for unidimensional pits is similar to Eq. (8), but C crit is a fraction of C sat ( Li et al., 2018 ):
From artificial pit experiments and a diffusion model ( Gaudet et al., 1986 ), Kaneko and Isaacs (2000 ) argue that to prevent pit repassivation, a higher concentration of metallic ions is required in Br − vs. Cl − solutions, i.e. C crit in Br − will be higher than C crit in Cl − in Eq. (10). As previously discussed, the diffusion coefficient of metallic cations is higher in Br − vs. Cl − solutions, according to experiments conducted by Kaneko and Isaacs (2000 ). Therefore, from Eqs. (9) and (10) and for a given pit depth x , pit stability in Br − solutions requires a higher dissolution current i diss,max at the pit bottom. Finally, considering the curves presented in Figure 1 , it is inferred that for a given current density, a higher potential is required to attain it in Br − vs. Cl − solutions. Notice that this analysis was made considering unidimensional pits, but it can be extended because similar expressions hold for hemispherical pits ( Li et al., 2018 ). All those contributions explain the higher E pit observed in Br − vs. Cl − solutions ( Kaneko & Isaacs, 2000 ; Pahlavan et al., 2019a , b ), confirming that the higher aggressiveness of Br − vs. Cl − ions can be explained on the basis of pit growth kinetics.
Ferritic ( Kaneko & Isaacs, 2002 ) and austenitic ( Guo & Ives, 1990 ; Kaneko & Isaacs, 2002 ) stainless steels alloyed with Mo can be more susceptible to pitting in Br − vs. Cl − solutions. A higher pitting potential was measured in Br − vs. Cl − solutions when Mo content was low ( Guo & Ives, 1990 ; Kaneko & Isaacs, 2002 ; Pahlavan et al., 2019a ); however, Mo additions had a stronger inhibiting effect on the pitting potential in Cl − than in Br − solutions ( Horvath & Uhlig, 1968 ). Hence, above a certain Mo content that depended on stainless steel microstructure ( Kaneko & Isaacs, 2002 ), Br − became more aggressive than Cl − , causing stable pitting at lower potentials. Those results were in good correlation with anodic polarization curves measured with the artificial pit electrode ( Kaneko & Isaacs, 2002 ). In Cl − solutions, studies with the artificial pit electrode reveal that Mo increases pitting resistance by shifting the bare metal anodic dissolution curve to higher potentials ( Laycock & Newman, 1997 ). This displacement of the anodic curve for a given increment of Mo is lower in Br − solutions vs. Cl − solutions ( Kaneko & Isaacs, 2002 ). This lower effect of Mo in preventing Br − pitting was attributed to the formation of soluble molybdenum complexes ( Newman, 2001 ), a lower production of inhibiting polymolybdate species in bromide solutions ( Domínguez-Aguilar & Newman, 2006 ), or to the fact that when Br − displaces the anodic dissolution curve to a higher potential, as schematized in Figure 1 , Mo dissolves more easily, and it is less effective as an inhibitor ( Ernst & Newman, 2008 ).
For commercial steels, E pit in Cl − is lower than E pit in Br − for type 304 stainless steel ( Ernst & Newman, 2008 ), type 301 stainless steel (0.19% Mo as impurity) ( Guo & Ives, 1990 ), type 316 stainless steel (2.5% Mo) ( Guo & Ives, 1990 ; Ernst & Newman, 2008 ; Pahlavan et al., 2019a ), and type UNS S31260 (3% Mo) duplex stainless steel ( Yamamoto & Hosoya, 1995 ). The reverse behavior is observed for type 904L (UNS N08904) with 4.5 wt.% Mo ( Guo & Ives, 1990 ; Abd El Meguid, 1997 ) and for 6% Mo superaustenitic stainless steel (UNS S31254) ( Guo & Ives, 1990 ). Some of those results are summarized in Figure 2 for measurements of E pit near room temperature.
Effect of molybdenum content on the pitting potential, E pit , in Cl − (filled symbols) vs. Br − (hollow symbols) solutions near room temperature for various commercial alloys as indicated in the top scale ( Guo & Ives, 1990 ; Abd El Meguid, 1997 ; Ernst & Newman, 2008 ; Pahlavan et al., 2019a ).
With the exception of Mo, how effective the different alloying elements are for preventing localized corrosion in Br − and I − solutions was not as deeply and systematically studied as in Cl − solutions. For Cl − solutions, the beneficial effect of Cr, Mo, W, and N on pitting and crevice corrosion resistance of stainless steels is often summarized with the PRE N (pitting resistance equivalent with nitrogen) number ( ISO 2010 ):
PRE N = wt . % Cr + 3.3 ( wt . % Mo + 0.5 wt . % W ) + 16 wt . % N
In the literature, there are similar versions of this equation that do not consider the beneficial effect of W or with different factors for N (Malik et al., 1994 , 1995 ; Sedriks, 1996 ; Szklarska-Smialowska, 2005 ). E pit in seawater ( Malik et al., 1994 , 1995 ) and CPT in ferric chloride solutions ( Sedriks, 1996 ) both increase with an increase in PRE N . Similar correlations exist for crevice corrosion critical parameters ( Sedriks, 1996 ). CPT increases with PRE N in 35,500 ppm Br − solutions, but the slope in the CPT vs. PRE N graph was lower in Br − than in Cl − solutions ( Ozturk & Grubb, 2012 ). Considering this finding, thumb or engineering rules valid in chloride solutions, like PRE N >40 for resistance to localized corrosion in seawater at room temperature ( Norsok, 2014 ; Francis & Hebdon, 2015 ), will not be valid for bromide solutions of similar concentrations. The critical pitting temperature of austenitic UNS N08904 (PRE N =34.9) and UNS S31254 (PRE N =43.5) stainless steel is near room temperature in Br − solutions ( Guo & Ives, 1990 ; Abd El Meguid, 1997 ), i.e. more than 30°C lower than values measured in Cl − solutions ( Guo & Ives, 1990 ). Given those results, the use of high PRE N (and in particular, high Mo) stainless steels for mitigation of pitting corrosion is less effective in Br − than in Cl − solutions.
Austenitic stainless steels suffer SCC in the presence of Cl − ions ( Scully, 1968 ; Hänninen, 1979 ; Cragnolino & Macdonald, 1982 ; Sedriks, 1996 ; Streicher, 2011 ), both in the fully solubilized and sensitized microstructures. The crack path in Cl − solutions can be either transgranular (TG) or intergranular (IG), as reviewed by Cragnolino and Macdonald (1982 ). Similar to the situation for pitting corrosion, the SCC in halides other than Cl − solutions received comparatively less attention. Some cases of SCC were reported for austenitic ( Griess et al., 1985 ; Itzhak & Elias, 1994 ; Itzhak et al., 1996 ) and martensitic ( Downs et al., 2007 ) stainless steels in the presence of bromides or other bromine species ( Lee et al., 1983 ; Nordin, 1983 ). Similar to Cl − solutions, sensitization is not required for SCC occurrence in Br − solutions. Finally, a number of researchers conclude that I − act as an inhibitor of Cl − SCC of austenitic stainless steels ( Overman, 1966 ; Uhlig & Cook, 1969 ; O’Dell & Brown, 1978 ; O’Dell et al., 1980 ; Pinkus et al., 1981 ; Itzhak & Eliezer, 1983 ).
While pitting corrosion was not observed in F − solutions, sensitized stainless steels suffer IGSCC in this environment ( Ward et al., 1969 ; Theus & Cels, 1974 ; Takemoto et al., 1985 ; Trabanelli et al., 1988 ; Zucchi et al., 1988 ; Shibata et al., 1993a , b ; Whorlow et al., 1997 ), a problem that was researched mainly because the presence of F − is common in thermal insulation materials ( Whorlow & Hutto, 1997 ) and in fluxes used in welding processes ( Ward et al., 1969 ; Takemoto et al., 1985 ). In contrast to Cl − solutions, no instances of F − -induced SCC were reported up to date in fully solubilized microstructures, as will be discussed in depth below.
A short discussion of different tests used to quantify the aggressiveness of the different halide anions toward the SCC of stainless steel is presented. It has first to be noticed that sensitized type 304 suffers IGSCC in slow strain rate tests (SSRT) even in oxygen-containing “pure water” at a temperature above 50°C (Ford & Povich, 1979; Ford & Silverman, 1980 ; Cragnolino & Macdonald, 1982 ; Congleton & Sui, 1992 ). By “pure water,” it is meant a solution with less than 5 ppb Cl − ( Congleton & Sui, 1992 ) or conductivity <0.3 μS/cm ( Ford & Silverman, 1980 ). It was proposed that strain can expose bare metal to the environment, and if the potential is in a range where repassivation of grain boundaries is slower than the matrix, crack propagation can proceed even in the absence of aggressive ions ( Ford & Povich, 1979 ; Ford & Silverman, 1980 ; Cragnolino & Macdonald, 1982 ). The presence of halide ions increases the SCC susceptibility, for example, the minimum oxygen concentration and the minimum potential for SCC occurrence in SSRT decrease with an increase in Cl − concentration ( Congleton & Sui, 1992 ).
Halide ions are typically required for cracking in constant deflection tests, like those depicted in Figure 3 . Dana and Delong (1956 ) and Dana (1957 ) proposed in the 1950s the SCC test shown in Figure 3 , left. The test mimics the leaching of chlorides in the insulation followed by concentration by evaporation at the stainless steel surface, simulating the environment in contact with a stainless steel pipe under thermal insulation. This test was standardized in 1971 in the ASTM C692 standard, and it is one of the alternatives listed in the current version of this standard ( ASTM C692-13, 2018 ). This test, however, is limited to “wicking-type” thermal insulation. Those materials wet completely when partially immersed in water. In this test, chlorides and other ions leach in the deionized (DI) water and concentrate on the surface of a U-bend stainless steel specimen. Temperature is controlled by the Joule effect with a transformer, connected either to a resistance heater taped to the specimen, as originally proposed by Dana (1957 ) or directly to the specimen, as later standardized by ASTM C692 standard ( Figure 3 , left). The 304 stainless steel is sensitized by heating for 3 h at 649°C. This temperature is close to the nose of the time-temperature-sensitization curve of type 304 stainless steel ( Sedriks, 1996 ). The specimen is then bent into a U shape and stressed to 30 ksi (207 MPa) with a bolt and nut, and placed in the testing apparatus depicted in Figure 3 , left. An insulation specimen passes the test if there are no cracks in the stainless steel surface after a 28-day period. In an alternative method ( Hutto et al., 1985 ; Whitaker et al., 1990 ), temperature is controlled with a steam-heated pipe, and the DI water is dripped over the insulating material with a peristaltic pump ( Figure 3 , right). This alternative method was incorporated to ASTM C692 in the 1990 edition, and it has the advantage that it is applicable to both wicking and non-wicking thermal insulators ( Hutto et al., 1985 ), while also reproducing more closely the type of wetting of an insulator most likely to be encountered in “real life” service. Prior to using either of the tests depicted in Figure 3 for qualification of insulating materials, the standard ( ASTM C692-13, 2018 ) requires that the test method and sensitized stainless steel must be tested with pure water (less than 0.1 ppm Cl − ) and with a 1500-ppm Cl − solution. Four coupons must crack in the Cl − solution, and none of the four coupons should crack in pure water. Notice that similar solution conditions than those of the “blank” test might produce cracking of a sensitized stainless steel tested in a SSRT. In other words, the observed or apparent threshold conditions for cracking are dependent on the test method used.
Alternative tests for laboratory studies of stress corrosion cracking under thermal insulation, after ASTM C692 ( ASTM C692-13, 2018 ).
“Thresholds” for SCC, i.e. a threshold aggressive anion concentration, potential, temperature, and stress intensity factor, are supposedly critical parameters above which SCC is possible and below which immunity is granted. Threshold values are usually defined considering material performance in short-term laboratory tests, performed under aggressive conditions ( Andresen, 2019 ). Immunity is associated with successful performance in an accelerated laboratory test, which often implies the absence of crack initiation after a certain time. For constant load and constant deflection tests, exposure time is often set to 30 days ( Sridhar et al., 2017 ). After finishing the test, specimens are analyzed; however, there could be ambiguity in the minimum dimensions that a flaw or defect must have to be considered an initiated crack. In addition, it could be argued that cracks could have initiated if the exposure time was longer than the selected test duration.
It was proposed that rather than attempting to define thresholds for SCC, dependencies of crack growth rate with SCC variables should be characterized ( Andresen, 2013 , 2019 ). The SCC crack growth rate has a complex dependence on many variables, i.e. temperature, species in solution and their concentration, pH, stress, electrochemical potential, material microstructure, and strength ( Staehle & Gorman, 2003 ). Careful experiments of in situ crack-growth rate measurements of stainless steels and nickel-based alloys in regions of supposed “immunity” to SCC in high-temperature water revealed in many instances crack propagation at a low growth rate ( Andresen, 2019 ). It was proposed that a similar situation might be valid in other material-environment systems. If crack growth rate dependency with SCC variables and the actual defect size (or resolution of non-destructive testing methods) are known ( Andresen & Ford, 1988 ), the residual life of the component or the optimum inspection interval can be assessed.
Much of the research conducted on the effect of F − , I − , and Br − on SCC was based on accelerated tests, where the effect of one of those variables is explored while keeping the others constant. The identified region of “immunity” or crack growth at a low rate will not guarantee crack propagation at a low rate if one or more of the rest of the SCC variables or loading conditions are changed. Ideally, the dependency of crack growth rate with each of the variables that control SCC should be studied, but the equipment required is more sophisticated, and tests are more expensive.
Most cases of SCC occur when the bulk surface of the stainless steel is in the passive state. However, it has been shown ( Acello & Greene, 1962 ); Asawa, 1971 , 1987 ) that dilute additions of Cl − , Br − , and I − to a H 2 SO 4 solution cause the SCC of the solution annealed type 304 stainless steel. This occurs in a specific range of halide ion concentration and at a potential close to the corrosion potential, where the alloy is actively dissolving. Outside this potential range, the alloy failed by uniform attack or water-line corrosion ( Asawa, 1987 ). Cl − , Br − , and I − acted as dissolution inhibitors when their concentrations were below 10 −1 , 2×10 −3 , and 3×10 −4 m , respectively; beyond these concentrations, they promoted uniform corrosion( Asawa, 1987 ). SCC was observed in a potential and halide ion concentration where the uniform dissolution rate was lower than 1 nm/s, regardless of the nature of the halide ion ( Asawa, 1987 ).
A few studies addressed the effect of halides other than Cl − on high temperature (around 300°C) water, an environment relevant for cooling water of nuclear power plants. The SCC in high-temperature water of types 304 and 316 stainless steel in the presence of Br − ions ( Kumada, 1996 ) is characterized by a transgranular crack path, and the reported threshold amounts of Br − and dissolved oxygen required for cracking in accelerated tests at 250°C were 50 ppm and 0.1 ppm, respectively. Fluorides at a concentration level of 1–10 ppm favored intergranular cracking in constant load tests of thermally sensitized type 304 and type 316 stainless steel in high-temperature water with 10 ppm O 2 ( Berry et al., 1973 ). Finally, it was reported ( Chung et al., 1996 ) that fluorides favor IGSCC of irradiated type 304 stainless steel, where Cr depletion at the grain boundary was caused by irradiation.
The rest of this review will mainly address SCC of stainless steels in halide solutions at lower temperatures and at bulk pH values where a passive film is stable, but might rupture locally giving localized corrosion phenomena like pitting or crevice corrosion. Under such conditions, Cl − SCC initiates from localized attack in pits and crevices ( Newman, 2001 ). It was proposed that a lower bound for the critical potential for Cl − SCC of stainless steels and other corrosion-resistant alloys is the repassivation potential for localized corrosion, E rp ( Tsujikawa et al., 1994 ; Cragnolino et al., 1996 ; Sridhar et al., 2017 ). E rp might correspond to the pitting or crevice repassivation potential, depending on the type of localized corrosion process that is occurring. Electrochemical techniques for measuring E rp are described in the literature ( Sridhar et al., 2017 ). Below this potential, any initiated crack or flaw in the passive film would be arrested by repassivation of the film ( Cragnolino et al., 1994 ). This condition of E corr > E rp for SCC crack propagation is necessary, but not a sufficient condition. Notice that an excessively high potential would cause crack blunting by pitting or crevice corrosion ( Tsujikawa et al., 1994 ; Newman, 2001 ). Crack velocity has to be larger than the rate of localized corrosion, and because an increase in temperature causes a higher increase in crack velocity due to a larger activation energy ( Che-Sheng Chen et al., 1997 ; Newman, 2001 ), this results in a critical temperature for SCC ( Che-Sheng Chen et al., 1997 ). Finally, plastic strain rate is required at the crack tip, for a local disruption of the passive film ( Andresen & Ford, 1988 ; Sridhar et al., 2017 ; Andresen, 2019 ).
Extensive testing by Tsujikawa et al. (1994 ), using spot-welded specimens of various austenitic stainless steels exposed to Cl − solutions confirmed that initiation of cracks required a potential greater than E rp . In those tests, a small sheet of stainless steel was spot-welded on top of a larger coupon, generating both a crevice and weld nuggets with residual stresses around them. The SSRT of Fe-Ni-Cr-Mo alloys 316L and 825 in Cl − solutions confirmed that at potentials below E rp , specimens failed in a ductile fashion ( Cragnolino et al., 1996 ; Pan et al., 2000 ). For the same system, using pre-cracked wedge-loaded fracture mechanics specimens, crack growth was only detected when the potential was above E rp ( Pan et al., 2000 ), and no cracks were observed below E rp in constant deflection tests ( Cragnolino et al., 1996 ). Recently, in situ measurement of crack growth rate on pre-cracked specimens confirmed ( Gui et al., 2014 ; Sridhar et al., 2017 ) a decrease in crack growth rate of almost two orders of magnitude as the potential decreased below E rp , for a 13-Cr supermartensitic stainless steel in a 0.3 m NaCl solution at 85°C.
Summarizing, depending on the test method used, it is inferred that for stainless steels in Cl − solutions at a potential below E rp , SCC initiation or crack growth rate is extremely difficult or decreases abruptly, respectively. This dependence of SCC with potential, assuming a similar cracking mechanism in solutions of the other halides, is useful for interpreting SCC results reported for the rest of the halides. For stainless steels with no or little content of Mo, the E rp ( Tzaneva et al., 2006 ) in halide solutions increases in the same order as E pit , i.e. Cl − <Br − <I − . In this regard, chlorides are often reported ( Scully, 1968 ; Davis, 1994 ; Whorlow et al., 1997 ) as the most effective of the halide ions to promote SCC of stainless steels. Likewise, considering the absence of stable localized corrosion of stainless steels in F − solutions ( Tzaneva et al., 2006 ; Macdonald & Lei, 2016 ; Pahlavan et al., 2016 ), SCC should not be expected in F − solutions. As previously mentioned, fluoride SCC is exclusively observed in sensitized microstructures ( Zucchi et al., 1988 ), probably because the chromium-depleted region near the grain boundary repassivates slower than the matrix in the presence of F − ions.
Sensitized stainless steels suffer IGSCC in the presence of F − ions ( Ward et al., 1969 ; Theus & Cels, 1974 ; Takemoto et al., 1985 ; Trabanelli et al., 1988 ; Zucchi et al., 1988 ; Shibata et al., 1993a , b ), even in dilute 5×10 −5 m (1 ppm) F − solutions at room temperature, as determined by the SSRT ( Trabanelli et al., 1988 ; Zucchi et al., 1988 ). IGSCC was also reported in constant load tests in 10 −4 m (2 ppm) F − solutions at room temperature ( Trabanelli et al., 1988 ). By comparison of results available in the literature ( Takemoto et al., 1985 ; Trabanelli et al., 1988 ; Zucchi et al., 1988 ), it is observed that IGSCC occurrence requires more aggressive conditions (higher T or higher F − concentration) in constant load or constant deflection tests than in SSRT.
The breakdown of the passive layer was not observed in conventional polarization curves of sensitized stainless steel in F − solutions ( Zucchi et al., 1988 ), as previously discussed. In contrast, a breakdown potential ( E b ) associated to intergranular cracking was observed when stressed specimens were anodically polarized in F − solutions ( Theus & Cels, 1974 ) at 65°C. According to the authors ( Theus & Cels, 1974 ), IGSCC occurred when E corr > E b . E b decreased with pH and with F − concentration in the range of 1–1000 ppm. E corr decreased with pH, but it was fairly independent of F − concentration. Therefore, a decrease in pH or an increase in F − concentration increased cracking susceptibility. Notice the difference of this cracking condition with E corr > E rp for the SCC of stainless steel ( Tsujikawa et al., 1994 ; Cragnolino et al., 1996 ; Sridhar et al., 2017 ). The E b determined by Theus and Cels (1974 ) is a potentiodynamically determined potential for initiation , and it might vary with experimental parameters like crack incubation time and potential scanning rate. While specimens exhibited cracks when the potential was above E b , some specimens exhibited cracking when polarized below E b . A more conservative potential to prevent cracking might be the repassivation potential, E rp , measured during a backward scan in the active direction after controlled localized corrosion propagation ( Szklarska-Smialowska, 2005 ; Sridhar et al., 2017 ). Theus and Cels (1974 ) did not conduct cyclic potentiodynamic tests for E rp determination, but the existence of an arrest potential for intergranular attack in F − solutions is hypothesized. The existence of a crack arrest or critical potential in F − solutions can be inferred from tests conducted by Zucchi et al. (1988 ). The authors performed an SSRT of sensitized stainless steel wires and then after crack initiation; the deformation was kept constant. Crack propagation could be detected by monitoring the load as a function of time: crack propagation caused a decrease in load, and the load remained constant when the crack was arrested ( Zucchi et al., 1988 ). It was verified that crack propagated and arrested as the potential was switched within or outside the crack propagation potential range, respectively. For example, for a 200-ppm F − solution at 25°C, crack propagation occurred within −0.4 and +0.4 V SCE . The upper potential limit might correspond to crack blunting, but this was not discussed in depth by the authors. Finally, the minimum crack propagation rate that can be resolved with this technique was not reported in the paper.
In contrast to Cl − SCC, a solution annealing heat treatment completely prevents F − IGSCC ( Ward et al., 1969 ; Theus & Cels, 1974 ). Transgranular branching and pitting attack commonly observed in tests of Cl − SCC of sensitized type 304 SS are not observed in F − solutions ( Ward et al., 1969 ). Those results suggest that F − attack is restricted to the intergranular region. Theus and Cels (1974 ) argue that the dependence of IGSCC with parameters of the sensitizing thermal cycle is difficult to assess. This issue could be solved with a systematic study of the dependence of F − IGSCC with the degree of sensitization (DOS). The thermal cycle affects DOS, which can be defined as the extent of Cr depletion near the grain boundary ( Parvathavarthini & Mudali, 2014 ). DOS can be quantified by coverage (fraction of total grain boundary depleted in Cr), depth (minimum concentration of Cr), and width (distance from grain boundary with a depleted Cr content) ( Parvathavarthini & Mudali, 2014 ). DOS can be estimated with electrochemical techniques, as recently reviewed by Parvathavarthini and Mudali (2014 ).
Similar to the case for Cl − SCC ( Speidel, 1981 ), an increase in temperature from 25°C to 80°C increases the kinetics of the attack, according to Zucchi et al. (1988 ) and Ward et al. (1969 ). Measurements of time to failure at constant load and the ductility loss in the SSRT suggest that the aggressiveness of F − increases with its concentration in the solutions, from 10 −5 to 10 −2 m (0.2–200 ppm F − ) ( Trabanelli et al., 1988 ; Zucchi et al. 1988 ). This is in accord with the decrease in breakdown potential in anodic polarization curves of stressed specimens reported by Theus and Cels (1974 ), from 5×10 −5 to 5×10 −2 m (1–1000 ppm F − ). Some authors claim that a maximum susceptibility to IGSCC occurs at an intermediate F − concentration. Using SSRT, Shibata et al. (1993 b ) showed that the region of maximum susceptibility is around 450 ppm F − (0.02 m ). Likewise, Ward et al. (1969 ) reported that solutions with a concentration around 1 m F − (20,000 ppm F − ) do not cause IGSCC. Adding Cl − ions to a the F − solution did not change the kinetics or mode of attack; hence, a synergism between those ions was discarded ( Ward et al., 1969 ). Whorlow et al. (1997 ) drew similar conclusions, after adding F − ions to a Cl − solution. This solution had a Cl − concentration just below the threshold amount required for SCC in constant deflection tests, as will be discussed in detail in the next section.
SCC in F − and F − +Cl − solutions were mainly studied ( Takemoto et al., 1985 ; Whorlow et al., 1997 ; Whorlow and Hutto, 1997 ) because those ions are common impurities found in materials for thermal insulation. This failure mechanism is also known as ESCC, where the “ E ” means that the water and halides involved in the cracking mechanism are external to the pipe, tank, or vessel. Pipes conveying high-temperature fluids are commonly wrapped with thermal insulation to minimize heat losses and to protect nearby personnel. Water from rain, nearby processes, or arising from steam condensation can wet the thermal insulation ( Dana, 1957 ; Ahluwalia, 2006 ). Alternatively, vapor from the atmosphere can condensate on a surface that is temporarily below the dew point ( Ahluwalia, 2006 ). Water lixiviates impurities in the thermal insulation, and upon contact with the hot stainless steel surface, those impurities are concentrated by evaporation of the solvent ( Dana, 1957 ; Ahluwalia, 2006 ). If aggressive species, namely, F − +Cl − , are originally present in the thermal insulation material, this leaching and concentration process can create optimum environmental conditions for the SCC of the stainless steel. Most failures occur in equipment with the stainless steel surface at a temperature between 50°C and 175°C ( Ahluwalia, 2006 ; NACE, 2017 ). Below this temperature, the concentration by evaporation is not significant, and the kinetics of SCC is low, and above 175°C, liquid water presence on the stainless steel surface is less frequent. More conservatively ( McIntyre, 1985 ), the upper temperature is listed as 260°C, the temperature below which some chloride salts retain their hydration water. However, even in equipment operating above the maximum temperature, it is necessary to consider possible SCC occurrence during start-up and shutdown ( McIntyre, 1985 ).
The SCC of stainless steel under insulation, like any other SCC problem, requires tensile stresses on susceptible materials exposed to a given environment. Most stainless steel products contain sufficient residual stresses to cause SCC in an aggressive environment even without applied stresses ( NACE, 2017 ). This situation worsens if welding or cold work is involved in the fabrication process. Decreasing stress can be impractical, and changing the material is economically infeasible, especially in operating plants. Therefore, most strategies for prevention of ESCC focus on the environment. Standard specifications, like ASTM C795 ( ASTM C795-08, 2018 ), provide limits for the maximum concentration of Cl − and F − in thermal insulation in contact with stainless steel. Thermal insulation has to be qualified in a preproduction standard test (ASTM C692-13, 2018), Figure 3, where it is tested to determine if the amount of leachable Cl − , F − , and inhibitors in the thermal insulator can cause SCC in a sensitized type 304 stainless steel stressed specimen at high temperature. Once the chemical composition and production process of the insulator is thus qualified, standard test method ASTM C871 ( ASTM C871-182018 ) is used to measure the chemical composition, specifically the concentration of chlorides, fluorides, sodium, and silicates of subsequent production lots. The maximum acceptable concentrations of Cl − and F − in the insulator are indicated in standard ASTM C795 ( ASTM C795-08, 2018 ), as a function of the concentration of inhibiting species (sodium and silicates). Regulatory guide US NRC 1.36 ( US Nuclear Regulatory Commission 2015 ), applicable to thermal insulation to be used in nuclear power plants, contains similar guidelines as those of ASTM C795. The historical evolution of those standards is briefly discussed below, and it is of interest because it shows changes in perception on the aggressiveness of F − .
Karnes in the 1960s determined that SCC can be inhibited if a certain amount of sodium and silicates are present in the thermal insulation material ( Whitaker et al., 1990 ), and the results were reported in a Cl − vs. Na + +silicates graph, known as the acceptability curve ( Figure 4 ). In this graph, an arbitrary line separates the compositions of the insulator that cause the SCC in the stainless steel from those that do not. The line was traced considering that if no more than one out of four specimens failed, the test was considered a “pass” ( Whitaker et al., 1990 ). While Karnes originally published the acceptability curve with Cl − in the ordinate, the United States Atomic Energy Commission (USAEC) published the Regulatory Guide (RG) 1.36 in 1973 ( US Atomic Energy Commission, 1973 ), where the ordinate was changed to Cl − +F − . Accordingly, the maximum Cl − +F − content in the insulator was limited to 600 ppm. According to Whitaker et al. (1990 ), F − was added to this regulatory guide because of the “chemical similarity and normally greater chemical aggressiveness of the fluoride,” rather than based on experimentally measured effects of F − on SCC. The United States military standard MIL-I-24244 ( MIL-I-24244A 1974 ) issued in 1974 did not contain the F − requirement, but this was added in a subsequent revision, to be consistent with RG 1.36 ( Whitaker et al., 1990 ). Furthermore, the minimum sodium+silicate content was fixed at 50 ppm. ASTM C795, first published in 1977, did not contain the fluoride requirement in the acceptability curve ( Figure 4 ), until the 2008 version, where it was added to “be consistent with other standards” (ASTM C795-08, 2018).
Acceptability curve of thermal insulation material, based on analysis of leachable halides (Cl − or Cl − +F − , depending on the standard or specification as indicated) and leachable inhibitors (sodium+silicate). The diagonal line was first proposed by Karnes ( Whitaker et al., 1990 ), and below this line, one out of four or zero out of four stainless steel specimens failed in an accelerated SCC test. Karnes line was adopted by subsequent standards and specifications, with major modifications as indicated.
The F − +Cl − vs. Cl − controversy in the ordinate of the Karnes acceptability graph was studied in depth by Whorlow and Hutto (1997 ), in a report prepared for the US Nuclear Regulatory Commission and later published by ASTM ( Whorlow et al., 1997 ). The experimental setup was similar to the one depicted in Figure 3 (right), but instead of dripping DI water to leach ions from the insulator block, different solutions were dripped directly over the type 304 sensitized stainless steel surface, at a fixed flow rate for 28 days. The authors ( Whorlow & Hutto, 1997 ) provided a chart to convert the concentration of ions in the solution (in mg/l) to the equivalent concentration in the insulator (in mg/kg), considering that in a real qualification test, the same flow rate would drip over the insulator but with DI water. In the absence of inhibiting species like sodium and silicates, the authors concluded that F − ions cause SCC when they are in a concentration above 20 mg/kg (20 ppm) in the insulator. This is equivalent ( Whorlow & Hutto, 1997 ) to 0.8 ppm of F − in solution, which is further concentrated at the stainless steel surface by water evaporation. Therefore, this susceptibility to SCC of type 304 sensitized stainless steel in F − solutions is in accord to the results reported elsewhere ( Ward et al., 1969 ; Theus & Cels, 1974 ; Zucchi et al., 1988 ) with different experimental setups.
In comparison to Cl − ions, SCC in F − can be inhibited with considerably lower amounts of sodium and silicates, as summarized in Figure 5 ( Whorlow & Hutto, 1997 ). Furthermore, when F − ions were added to a solution of Cl − , sodium, and silicate originally below the Karnes acceptability curve, SCC was not observed, despite the total Cl − +F − content of the solution was in the cracking region of the acceptability graph ( Figure 6 ). A single exception was the test run at 1000 ppm sodium+silicate and 100 ppm Cl − +35 ppm F − , marked with an arrow in Figure 6 . Based on those results, the authors ( Whorlow & Hutto, 1997 ) discarded a synergistic effect between Cl − and F − , in accord with research previously published by Ward et al. (1969 ). In other words, the amount of silicate and sodium required to inhibit Cl − SCC seems to be sufficient to inhibit the effect of further additions of F − ( Whorlow & Hutto, 1997 ). Despite those results, when the US Nuclear Regulatory Commission (NRC) revised RG 1.36 in 2015 ( US Nuclear Regulatory Commission, 2015 ), both Cl − and F − were considered as equally aggressive species and placed in the ordinate of Karnes acceptability graph. By this time, ASTM had already included F − in the ordinate of Karnes graph, in the 2008 version of ASTM C795 ( ASTM C795-08, 2018 ). The International Atomic Energy Agency (IAEA) published in 2011 the report “Stress corrosion cracking in light water reactors: good practices and lessons learned” ( International Atomic Energy Agency, 2011 ), where the Karnes figure with Cl − +F − in the ordinates was reproduced. In summary, current standards and regulatory guides on ESCC of austenitic stainless steels consider F − as aggressive as Cl − , despite contrary conclusions drawn in laboratory tests.
SCC of type 304 sensitized stainless steel in fluoride and sodium+silicate solutions; results are presented superimposed to ASTM C795 criteria for acceptability of thermal insulators ( Whorlow & Hutto, 1997 ).
Effect of fluoride addition to a chloride, sodium, and silicate solution that passed the 28-day SCC tests. Karnes line given as reference. Except for the solution concentration marked with an arrow, no SCC was observed in Cl − +F − solutions after 28 days. Adapted from Whorlow KM, Woolridge E0, and Hutto FB, Effect of halogens and inhibitors on the external stress corrosion cracking of type 304 austenitic stainless steel, Insulation materials: testing and applications: third volume, ASTM STP 1320, R.S. Graves and R.R. Zarr, Eds., copyright ASTM International, 100 Barr Harbor Drive, West Conshohocken PA19428, www.astm.org .
Whorlow and coworkers ( Whorlow & Hutto, 1997 ; Whorlow et al., 1997 ) reported that the inhibitor effectiveness was dependent on the type of silicate; the order of increasing effectiveness was sodium disilicate (Na 2 Si 2 O 5 ), sodium metasilicate (Na 2 SiO 3 ) and sodium orthosilicate (Na 4 SiO 4 ). In other words, summarizing criticism on the acceptability curve of thermal insulation material ( Figure 4 ), neither the aggressive ions plotted in ordinate or the inhibiting ions in the abscissa seem to be equally effective in promoting or preventing ESCC, respectively. Whorlow and coworkers reported the SCC of sensitized stainless steels in environments with chloride and sodium+silicate concentrations below the Karnes line, in other words, in the zone of acceptable analysis ( Whorlow & Hutto 1997 ; Whorlow et al., 1997 ). The occurrence of SCC failures below the Karnes curve is not surprising because this curve was traced following the criterion that if none or one out of four SCC specimens failed for a given thermal insulator composition, the point was considered a pass ( McIntyre, 1985 ; Whitaker et al., 1990 ). Current recommendations of ASTM C 795 require that the preproduction SCC test (ASTM C692-13, 2018) is passed if none of the specimens exhibit cracks. Lots produced with this validated production method and using similar ingredients are acceptable if their chemical composition falls inside the acceptable region ( Figure 4 ) or if their chemical composition was validated with a pass in a preproduction SCC test. US REG 1.36 ( US Nuclear Regulatory Commission, 2015 ) is more conservative because it not only requires that the chemical composition of the insulator falls in the acceptable region ( Figure 4 ) but also that the chemical analysis of the Cl − +F − content of the lot does not exceed 150% of the average values measured during preproduction qualification tests and that the sodium and silicate content is not below 50% of the average amount of those inhibitors measured during successful preproduction qualification tests.
Takemoto et al. (1985 ) questioned the 600 ppm high chloride limit in the acceptability graph ( Figure 4 ). Using constant deflection tests with an initial stress above that required by ASTM C692 ( ASTM C692-13, 2018 ), they proposed a threshold of 2000 ppm for a sensitized type 304 stainless steel and 3000 ppm for a solution annealed stainless steel, exposed to 95°C solutions. The difference in apparent Cl − threshold concentration values could be explained considering its strong dependence with the rest of the testing variables.
In contrast to other studies ( Ward et al., 1969 ; Whorlow & Hutto, 1997 ), Takemoto et al. (1985 ) suggest a synergistic effect of F − and Cl − , indicating that the 600 ppm high Cl − limit can be much lower in the presence of F − , decreasing to 200 ppm when the F − concentration is 100 ppm. The authors state that F − aggressiveness is maximum at 50°C, which is in conflict with results later published by Zucchi et al. (1988 ), where aggressiveness of F − was higher at 80°C. Figure 7 summarizes the effect of F − and Cl − on the SCC of sensitized stainless steels at 50°C, according to Takemoto et al. (1985 ). The line that separates the SCC vs. No SCC zone was constructed based on laboratory tests in NaF and NaCl solutions without added silicates.
Summary of the effect of fluorides and chlorides on SCC at 50°C, for a solution without silicates and a type 304 sensitized stainless steel under constant deflection. Adapted from Takemoto et al. (1985 ). © NACE International, 1985.
A major limitation of ASTM standards for ESCC management is that they only measure the capacity of inhibitors in the insulating material to guard against SCC failures caused by aggressive species in the insulating material. As reviewed by McIntyre (1985), Takemoto et al. (1985 ), and Hutto et al. (1985 ), and NACE International Standard SP0198-2017 ( NACE, 2017 ), other potential sources of Cl − of industrial relevance are the atmosphere, especially near coastal or industrial areas, or rainwater, processed, or potable water that might accidentally contact the thermal insulation. Insulation materials might have an excess of inhibitors to mitigate the effect of aggressive ions introduced from external sources. The ability of the insulation material to inhibit the effect of aggressive ions from outside sources can be tested using any of the experimental setups in Figure 3 , if a Cl − solution is used instead of DI water. In the “accelerated Dana test” a 1500-ppm (mg/l) Cl − solution is used ( Hutto et al., 1985 ; Whitaker et al., 1990 ), and an additional benefit is that the testing time is reduced from 28 to 6 days. However, it has not yet been included as an alternative procedure in ASTM C692. While inhibitors in the thermal insulator can offset the effect of externally introduced aggressive ions, it has to be noticed that if the equipment is severely wetted, the inhibitor might be leached and transported away from the surfaces needing inhibition ( NACE, 2017 ).
If water reaches the stainless steel surface through cracks in the thermal insulator without soaking through the insulation ( McIntyre, 1985 ), the amount of leachable aggressive and inhibiting species in the insulator is not so relevant, and cracking can be controlled by the Cl − or F − concentration in the water. Guidelines for preventing water ingress to the thermal insulation are listed in the NACE International Standard SP0198-2017 ( NACE, 2017 ). This standard addresses corrosion under insulation (CUI) under both stainless and carbon steels. Despite differences in their corrosion mechanisms, preventing water contact to the external surface of equipment is a common corrosion control strategy. An additional ESCC control strategy is the use of coatings or aluminum foil wrapping ( Richardson & Fitzsimmons, 1985 ; NACE, 2017 ). The aluminum foil provides an additional barrier that prevents water and ions to be in contact with the stainless steel surface ( Richardson & Fitzsimmons, 1985 ). More important, aluminum provides cathodic protection and decreases the corrosion potential of the stainless steel surface ( Richardson & Fitzsimmons, 1985 ), which is effective for preventing Cl − and F − SCC ( Smialowski & Rychcik, 1967 ; Rhodes, 1969 ; Uhlig & Cook, 1969 ; Theus & Cels, 1974 ).
Finally, it is noticed that the methodology of ESCC management with ASTM standards was developed based on results obtained with accelerated tests with smooth stainless steel coupons. However, if flaws or cracks are detected on a stainless steel component, it cannot be assured that they will not propagate during further use, even if new thermal insulation is installed, and its composition lays within the acceptability zone of Figure 4 . Under such a scenario, knowledge of dependence of SCC growth rate with stress and environmental and material variables would be of much greater use for assessing the remaining life of the component ( Andresen & Ford, 1988 ).
Pitting and crevice corrosion of stainless steels is observed in Br − solutions, even in the fully solubilized condition. Therefore, considering a similar mechanism for SCC in Br − vs. Cl − , one of the necessary ( Tsujikawa et al., 1994 ; Che-Sheng Chen et al., 1997 ; Newman, 2001 ; Sridhar et al., 2017 ) conditions for SCC would be fulfilled if E Corr > E rp . Br − SCC was reported in environments of industrial interest, like completion fluids for the oil and gas industry ( Downs et al., 2007 ; Sridhar et al., 2017 ) and brines for absorption refrigeration systems ( Griess et al., 1985 ; Itzhak & Elias, 1994 ; Itzhak et al., 1996 ). With regard to laboratory tests, Rhodes (1969 ) reported SCC in 304 stainless steel exposed to 59% MgBr 2 solutions at 150°C, but cracking was not observed when the solution was more dilute than 36%. As a comparison, the SCC of 304 stainless steels occurs readily in accelerated laboratory tests in 20% MgCl 2 at 105°C ( Brauns & Ternes, 1968 ), and more generally, at a temperature above 100°C, just a few ppm of Cl − can cause SCC ( Haruyama 1982 ) in industrial equipment. Therefore, albeit probably not as aggressive as Cl − , the evidence reported suggests that concentrated, hot Br − brines can cause SCC of stainless steels even in the fully solubilized condition.
Mo increases resistance to both localized corrosion and SCC of stainless steels in Cl − solutions ( Speidel, 1981 ; Sedriks, 1996 ; Prosek et al., 2009 ). This higher resistance is reflected in material selection guidelines; for example, to prevent Cl − SCC in marine atmospheric environments, Norsok standard M-001 ( Norsok 2014 ) sets the maximum operating temperature of type 6 Mo stainless steel in 120°C, higher than the 60°C set for a type 316 stainless steel (2.5 wt.% Mo). As illustrated in Figure 2 , for alloys with a high concentration of Mo like stainless steel type 904L, Br − can be more aggressive toward pitting than Cl − . This higher aggressiveness of Br − is also evidenced by a large decrease in CPT and CCT ( Guo & Ives, 1990 ; Abd El Meguid, 1997 ; Ozturk & Grubb, 2012 ). An unsolved question is whether the higher aggressiveness of Br − vs. Cl − toward pitting and crevice corrosion in high Mo stainless steels implies also a higher SCC susceptibility in Br − vs. Cl − solutions.
Completion fluids used in high pressure, high-temperature oil, and gas wells might contain large concentrations of Br − . When contaminated with O 2 , or acid gases like CO 2 and H 2 S, those brines can cause SCC in corrosion-resistant alloys like martensitic stainless steels (13 Cr-1Mo, 13 Cr-2Mo) ( Downs et al., 2007 ). SCC was reported in laboratory studies of 13 Cr-2Mo stainless steel exposed to CO 2 and H 2 S solutions at both 160°C and 40°C in concentrated Br − solutions. Formate brines are often used in replacement of bromide brines to mitigate this problem ( Downs et al., 2007 ; Sridhar et al., 2017 ).
LiBr commercial brines for use in absorption refrigeration systems can contain lithium chromates and lithium hydroxide ( Griess et al., 1985 ; Igual Muñoz et al., 2003 ), added as corrosion inhibitors. These systems operate under deaerated conditions, and besides the problem of integrity of materials, the release of hydrogen by the cathodic reaction at the corrosion potential produces non-condensable gases that decrease efficiency ( Chandler, 1999 ). Desired corrosion rates in LiBr brines are below 1 mpy (25 μm/year). Various stainless steels and carbon steels in presence of chromates or other inhibitors fulfill this constraint ( Chandler, 1999 ). The presence of chromates and deaerated conditions contribute to lower the corrosion potential ( Igual Muñoz et al., 2003 ) and mitigate pitting corrosion. However, the high operating temperature (~150°C) might favor SCC.
With regard to SCC of stainless steels in LiBr brines, in a preliminary report, Griess et al. (1985 ) studied the SCC of austenitic type 304 stainless steel in 68% LiBr solutions at 160°C, conditions similar to those encountered in absorption refrigeration systems. The presence of oxygen was required for cracking of U-bend specimens, and none out of 20 specimens cracked under deaerated conditions ( Griess et al., 1985 ). On the other hand, the SCC of types 304 and 316 stainless steels was reported by Itzhak et al. in 55% LiBr brines at 120°C and 140°C in deaerated conditions, both with the constant load ( Itzhak & Elias, 1994 ) and slow strain rate technique ( Itzhak et al., 1996 ). Specimens were studied in the “as-received” state ( Itzhak & Elias, 1994 ; Itzhak et al., 1996 ), which included a 15% cold work estimated by comparison to fully annealed specimens of the same heat. The SCC crack path was generally transgranular, but depending on loading mode and temperature, mixed IG and TG cracks were observed for type 316 stainless steel. Itzhak et al. studied SCC in brines with pH varying from 6 to 11.6, and time to failure generally increased with increasing pH. The content of chlorides present as potential impurities in LiBr brines was not measured, and two different grades of LiBr salts were used by the authors, i.e. “analytical” ( Itzhak & Elias, 1994 ) and “commercial” ( Itzhak et al., 1996 ) grade.
Considering the concentration of LiBr solutions used in studies of evaluation of stainless steel for absorption refrigeration systems, it could be argued that trace amounts of chlorides in the salt could explain the SCC process. For example, Griess et al. (1985 ) reported that 37 ppm of chlorides were present in the brine as impurities from the LiBr salt, enough to cause chloride SCC at 160°C even in absence of Br − ( Sedriks, 1996 ). Despite this, the authors argue that Cl − was not responsible for cracking because the Cl − SCC inhibitors like chromate did not inhibit cracking in the concentrated LiBr salt ( Griess et al., 1985 ). However, the role of chromate as a Cl − SCC inhibitor is disputed ( O’Dell & Brown, 1978 ) because despite stabilizing the passive film, it can raise the electrochemical potential. Stronger evidence to support the observation that cracking was actually due to Br − is that according to U-bend constant deflection SCC tests conducted by Rhodes (1969 ), intentional additions of Cl − to 36% MgBr 2 solutions did not induce cracking in type 304 stainless steel specimen, but SCC was observed in 59% MgBr 2 solutions at 150°C.
The effect of Br − on SCC of stainless steel under thermal insulation was not as deeply studied as the effect of F − and Cl − , probably because they are not as widely found in insulation as the lighter halides ( Whorlow et al., 1997 ). The SCC of sensitized type 304 stainless steels in Br − solutions was reported with an experimental setup similar to that shown in Figure 3 (right) ( Whorlow et al., 1997 ). The presence of SCC cracks and pits was reported in sensitized type 304 stainless steel U-bend coupons heated to 100°C and exposed to 1500 ppm Br − solutions (added as KBr), which was directly dripped over the stainless steel coupon. Under this experimental setup, the solution was concentrated by water evaporation at the hot stainless steel surface, so it is difficult to compare those experiments with those conducted in LiBr brines. The Br − SCC of sensitized stainless steel could be inhibited with sodium orthosilicate, with a concentration of inhibitor that was much lower than required to inhibit Cl − SCC ( Whorlow & Hutto, 1997 ). Br − ions were never included in ordinates of the acceptability curve of thermal insulation material ( Figure 4 ). However, some analytical techniques ( ASTM C871-18, 2018 ) to determine the Cl − concentration in thermal insulation cannot discriminate between Cl − , Br − , and I − , thus providing a false higher concentration of Cl − . The concentration of inhibitor required in the thermal insulator would be increased accordingly, and considering that Cl − is the most effective of those halides to cause ESCC in sensitized stainless steel, the possible effect of Br − and I − would be inhibited as well ( Whorlow & Hutto, 1997 ).
The presence of Br − ions in seawater can lead to corrosion issues in stainless steels used for multi-stage flash (MSF) desalination plants ( Oldfield & Todd, 1981 ). Despite Cl − being more concentrated in seawater than Br − , Table 1, two different mechanisms ( Oldfield & Todd, 1981 ; Lee et al., 1983 ) were proposed for the enrichment of bromine species in the vapor phase. Chlorine (Cl 2 ) is usually added to feed-water as a biocide, and it can oxidize Br − to Br 2 , a reaction that is favored when pH is lower than 6 ( Oldfield & Todd, 1981 ). Alternatively, bromamines can evolve from chlorinated seawater, if ammonia is present in seawater ( Lee et al., 1983 ). Bromine and bromamines are then removed from the MSF plant together with non-condensable gases. Bromamines can decompose to hydrobromic acid and free bromine ( Lee et al., 1983 ). Bromine is more oxidizing than oxygen ( Lee et al., 1983 ), thus it raises the corrosion potential and favors pitting, crevice corrosion, and SCC. The SCC of type 316L stainless steel due to the presence of bromine species was reported in air ejector condenser systems ( Black & Morris, 1981 ; Nordin, 1983 ) and venting pipes of MSF plants ( Lee et al., 1983 ). Pitting and uniform corrosion of 316L were also reported in Br 2 solutions ( Hodgkiess et al., 1985 ), and it was related to bromine reduction reaction, Eq. (1) that yields bromides. To prevent this type of failure, it was proposed to use higher alloyed stainless steel like 904L ( Black & Morris, 1981 ; Nordin, 1983 ) or strict control of pH, chlorination level, and ammonia content in feedwater ( Oldfield & Todd, 1981 ; Lee et al., 1983 ).
The SCC of zirconium alloys used for nuclear power reactor fuel cladding in iodine, a byproduct of the uranium fission reaction, is a well-documented problem ( Cox, 1972 ; Wood, 1972 ; Sidky, 1998 ). While zirconium alloys are widely used as cladding in water-cooled reactors operating at temperatures close to 300°C, studies ( Lobb & Jones, 1976 ; Lobb & Nicholson 1976 ; Lobb, 1978 ; Kiselevskii et al., 1993 ) conducted at 650°C and 750°C report that iodine vapor decreases creep resistance of austenitic stainless steels, a problem relevant for fuel cladding integrity of fast and advanced cooled reactors. Likewise, I 2 and I − can be generated in aqueous homogeneous reactors, where uranium salts are dissolved in an aqueous solution that serves as moderator and coolant ( Lillard, 2010 ). As reviewed by Lillard (2010 ), pitting and SCC of stainless steel in iodine species can affect the integrity of the off-gas extraction system of those reactors, required to remove the gases produced by radiolysis.
E pit and E rp of stainless steels in I − solutions are larger than in Br − or Cl − solutions ( Rhodes, 1969 ; Tzaneva et al., 2006 ; Macdonald & Lei, 2016 ). Therefore, the E Corr > E rp ( Tsujikawa et al., 1994 ; Che-Sheng Chen et al., 1997 ; Sridhar et al., 2017 ) criterion predicts SCC resistance in a wider range of potential in I − vs. the rest of the halide solutions. Furthermore, under open-circuit conditions, the requirement that E Corr > E rp ( Tsujikawa et al., 1994 ; Che-Sheng Chen et al., 1997 ; Sridhar et al., 2017 ) is usually fulfilled by the presence of oxygen in the solution. However, iodide solutions are unstable in the presence of oxygen, reacting with it to give iodates or iodine (Eqs. (2) and (11) ( Pinkus et al., 1981 ; Itzhak & Eliezer, 1983 ):
Given those considerations, in the case of I − solutions, the reactants required for halide stress cracking might not be present simultaneously under common open-circuit conditions.
Nevertheless, SCC was observed in sensitized type 304 stainless steel. Whorlow and Hutto (1997 ), using an experimental setup similar to the one in Figure 3 (right), reported SCC in I − solutions, added as KI. However, cracks were much shorter than those observed in chlorides, and cracking could be inhibited with silicates. In a previous attempt, SCC was not observed in iodide solutions ( Whitaker et al., 1990 ), but this was explained based on a possible low carbon content of the stainless steel used. Current recommendations of ASTM C692 ( ASTM C692-13, 2018 ) for evaluation of thermal insulation require a solution annealed type 304 stainless steel with carbon content in the range of 0.05–0.06%, which is then sensitized for 3 h at 649°C.
Some researchers propose that iodide and iodine can act as inhibitors of SCC in Cl − and Br − solutions. Slow strain rate test (SSRT) studies ( Huang et al., 1993 ) of solution annealed Ti-stabilized 321 stainless steel (UNS S32100) concluded that I − and I 2 inhibit SCC of stainless steels in 0.5 m NaCl+0.5 m HCl solutions at 55°C, with I − being more effective than I 2 . The fracture surface in the presence of 5 mmol/l KI was ductile with the presence of dimples. Similarly ( Uhlig & Cook, 1969 ), additions of NaI to the boiling MgCl 2 SCC test solution increased the time to failure of solution annealed and cold-worked specimens of 304 stainless steel under constant load conditions, and no cracks were observed after 200 h of testing when more than 3.7% of NaI were added to the MgCl 2 solution. According to Itzhak et al. ( Pinkus et al., 1981 ; Itzhak & Eliezer 1983 ), I − acts as a cathodic inhibitor of chloride SCC, reacting with oxygen and protons according to Eq. (11). Evidence for the occurrence of this reaction was that the solution turned to yellow-brown color. The corrosion potential in boiling 38% MgCl 2 +15% NaI was 200 mV lower than the corrosion potential in boiling 38% MgCl 2 ( Pinkus et al., 1981 ). A decrease in corrosion potential should contribute to inhibition of chloride SCC in boiling MgCl 2 solutions ( Smialowski & Rychcik, 1967 ; Rhodes, 1969 ; Uhlig & Cook, 1969 ). Additions of 1% KI to a 55% LiBr brine of pH=4 also inhibited Br − SCC of type 316 stainless steel ( Itzhak et al., 1996 ). Finally, some constant deflection tests performed on type 304 stainless steel with a setup similar to that shown in Figure 3 (right) ( Whorlow et al., 1997 ) suggest that I − is an ineffective inhibitor of Br − and F − SCC when the stainless steel is sensitized.
Pitting and SCC of stainless steels is possible in halides other than chlorides. Several examples where this problem manifested in the industry were presented. Chloride is not only the halide most commonly encountered in environments of industrial relevance but also in most conditions the most aggressive. There is general agreement that the localized corrosion initiation ( E pit ) and repassivation potential ( E rp ) of stainless steels increase in the following order, Cl − <Br − <I − . The intensity of SCC in halide solutions decreases in the same order, which can be understood considering that a necessary condition for SCC is that E Corr > E rp . A notable exception is that pitting in Br − solutions can be more aggressive than in Cl − solutions, for alloys with a high content of molybdenum. However, it is yet unsolved if high molybdenum stainless steels are more susceptible to SCC in Br − than in Cl − solutions. Detailed knowledge of the effect of alloying elements on Br − localized corrosion and SCC would allow more rationality in the selection of stainless steels for bromide service.
For solubilized stainless steels, SCC was not observed in F − and I − solutions. For F − solutions, this could be related to the pitting and localized corrosion immunity in those solutions. However, a sensitizing heat treatment decreases chromium content at the grain boundary, favoring localized rupture, and intergranular SCC is possible in the presence of stress. Passivity breakdown of stainless steels in iodide solutions would require excessively high potentials, and furthermore, iodides are unstable in the presence of dissolved oxygen, reacting with it to yield iodates and iodine. Hence, several researchers propose that iodides act as inhibitors for SCC of stainless steels in bromides and chlorides.
The author thanks Dr. Martín Rodríguez for helpful comments on the manuscript.
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Review of prediction of stress corrosion cracking in gas pipelines using machine learning.
2. methodology and analysis, 2.1. research framework, 2.2. research definition, 2.3. search strategy.
(“Machine Learning” <OR> “Machine Learning Application”) |
[AND] |
(“Oil and Gas” <OR> “Energy Industry” <OR> “Energy”) |
[AND] |
(“Stress Corrosion Cracking” <OR> “SCC”) |
[AND] |
(“Prediction Techniques” <OR> “SCC Prediction” <OR> “Pipeline Condition Assessment”) |
[AND] |
(“Non-Destructive Testing” <OR> “Pipeline Integrity Management” <OR> “Integrity Management”) |
3. critical research and analysis, 3.1. stress corrosion cracking (scc) failure events, 3.2. statistics of corrosion incidents and factors leading to scc in energy pipelines.
3.4. stress corrosion cracking (scc) management program, 3.5. stress corrosion cracking (scc) detection techniques.
Detection Method | Brief Details |
---|---|
Linear polarization resistance (LPR) | This method quantifies the electrochemical resistance of a corroding metal working electrode in close proximity to its open circuit potential. The process entails the polarization of a voltage range of ±10 mV relative to the corrosion potential [ , ]. |
In-line inspection (ILI) | ILI tools, commonly known as “smart pigs”, are devices that are inserted into the pipeline and travel with the product flow. They use various technologies such as magnetic flux leakage (MFL) or ultrasonic sensors to detect anomalies, including SCC, within the pipeline. These tools provide a comprehensive assessment of the pipeline’s condition [ ]. Currently, there are two primary methods used for crack detection: ultrasonic testing and test using electromagnetic acoustic transducers (EMATs) [ ]. They are readily accessible for the on-line inspection of SCC in commercial settings. Further ILI technologies can be seen in , and their merits and demerits are in [ ]. |
Electrochemical noise (EN) | This technique continually monitors corrosion potential and variations in current. It is utilized to acquire corrosion current by measuring noise resistance [ ]. |
Acoustic emission (AE) | AE monitoring involves detecting and analyzing the high-frequency acoustic signals emitted by crack growth or propagation. It can provide real-time information on the occurrence and progression of SCC [ ]. AE sensors are placed on the pipeline, and any acoustic emissions resulting from crack activity are captured and analyzed [ ]. |
Electromagnetic testing | Electromagnetic techniques, such as eddy current testing (ECT) and magnetic particle inspection (MPI), can be employed to detect SCC [ ]. ECT utilizes electromagnetic induction to detect surface and near-surface cracks, while MPI uses magnetic fields and iron particles to locate cracks or defects that are magnetically visible. |
Ultrasonic testing (UT) | UT uses high-frequency sound waves to detect internal defects or cracks in the pipeline. It involves transmitting ultrasonic waves into the material and analyzing the reflected waves to identify any indications of SCC [ ]. UT can be performed on both the external and internal surfaces of the pipeline. |
Cyclic potentiodynamic polarization | This process entails applying an over-potential greater than the corrosion potential toward the noble side until a current of 5 mA is reached. Then, the potential is reversed until the corrosion potential is achieved [ ]. |
Radiographic testing (RT) | RT uses X-rays or gamma rays to detect internal defects in the pipeline. It involves passing the radiation through the material and capturing the transmitted radiation on a film or detector. Any cracks or indications of SCC can be identified by examining the resulting radiographic image [ ]. |
Electrochemical impedance spectroscopy (EIS) | This process entails the use of an alternating current (AC) potential with a magnitude of ±10 mV around the corrosion potential. This is achieved throughout a broad range of frequencies, generally spanning 0.1 to 106 Hz. The purpose of this is to obtain the corrosion current [ ]. |
Electromagnetic acoustic transducer (EMAT) | The electromagnetic acoustic transducer (EMAT) is a modern non-destructive testing (NDT) device employed in in-line inspection (ILI) equipment to detect SCC in gas pipelines [ ]. EMATs operate by utilizing a magnetic field to create an ultrasonic compression wave on the inner surface of the pipe wall [ ]. |
Hydrostatic testing | Hydrostatic testing is a method employed to detect SCC in pipelines. When conducted correctly, this approach ensures that any significant flaws present during the test are discovered. Hydrostatic testing is a frequently employed technique to ensure the preservation of pipeline integrity in the presence of developing flaws, such as pitting corrosion, fatigue, corrosion fatigue, or SCC [ , ]. |
Magnetic flux leakage | Magnetic flux leakage (MFL) is a non-destructive testing (NDT) method employed for the identification of SCC. A high-strength magnet is employed to magnetize the steel in areas prone to corrosion or potential metal degradation. This method has been employed to identify corrosion flaws, fractures, and mechanical impairments [ ]. |
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Technology | Pros | Cons | Run in Operating | |
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Oil | Gas | |||
Shear Wave (Liquid-coupled) Ultrasound | No | Yes | ||
EMAT | Yes | Yes | ||
FMFL | Yes | Yes | ||
LWUT | Yes | Yes | ||
New Technologies | Yes | Yes | ||
SEEC (Self-Excited Eddy Current) | Yes | Yes |
3.7. scc prediction through machine learning, 3.8. research analysis, 3.9. critical review analysis, 3.10. gaps and challenges in implementing machine learning.
Gap/Challenge | Brief Details |
---|---|
Data Availability | ML models require large amounts of high-quality data to effectively learn and make accurate predictions. Obtaining sufficient labeled data related to SCC in pipelines to validate ML models is a major challenge. |
Data Quality | A lack of good quality data is one of the major problems that machine learning experts are facing in obtaining the required outcomes. Data quality is an issue in developing a good model to predict SCC in energy pipelines. As a result, we must make sure that data pre-processing is carried out to achieve the highest degree of accuracy possible, which involves eliminating outliers, the imputation of missing values, and eliminating undesired characteristics. |
Data Variability | The data collected for SCC detection can vary in terms of pipeline materials, environmental conditions, stress levels, and other factors. This variability makes it challenging to develop a machine learning model that can effectively handle different data types. Ensuring a diverse and representative dataset is crucial for training models that can handle the various conditions encountered in pipeline systems. |
Data Privacy | Data protection, data security, and privacy are some of the issues connected with the application of machine learning. For instance, the General Data Protection Regulation (GDPR) was developed in 2016 to provide people with more control over their data while also protecting the personal information of those living in the European Union and the European Economic Area. The California Consumer Privacy Act (CCPA) launched in 2018 mandates businesses to tell customers about the acquisition of their data. It is one example of a state policy being developed in the United States [ ]. Process data, operation data, and other inspection and maintenance data are some of the most important information that pipeline companies need to secure to avoid any interruption in their businesses and operations. |
Required Skillset | To obtain the best results from the data collected over the years, the oil and gas sector is facing difficulty in obtaining the right skills. Machine learning techniques and approaches are relatively new to people working in the energy pipeline industry. There are not enough ML specialists in this field, which hinders the potential to develop successful models that will bring benefits to business or predict issues to control unwanted events. |
Affordability | To develop a significantly advanced data analytics system in order to use machine learning techniques, pipeline owners will require data engineers/scientists with sound technical knowledge of data analytics, modeling, and mathematics. Without these skills, companies are not able to start with a good digital transformation system. |
Understanding the Algorithms | Given the complexity of machine learning, data scientists are required to have expertise in this particular field and an in-depth understanding of science, technology, and mathematics to develop ML models to achieve the best results. Many businesses lack the internal expertise necessary to comprehend algorithms and how they operate, which can cause them to lose out on crucial insights. |
Class Imbalance | SCC occurrences in pipelines are typically rare events in the overall dataset. This class imbalance can lead to biased models that struggle to accurately detect SCC instances. Techniques such as oversampling, undersampling, or synthetic data generation can be employed to address the class imbalance issue and ensure that the model is trained on a balanced dataset. |
ML Model Generalization | Developing a generalized ML model that can be applied to the detection of unseen pipeline conditions is difficult. A model should be capable of detecting SCC across different pipeline sections, varying stress levels, and diverse corrosion environments. Adequate model evaluation and validation of unseen data are necessary to assess the generalization capability of a model. |
Obtaining the Right Data/Information | Obtaining the appropriate data to train ML models is one of the biggest challenges we are facing. ML models may not perform as well as they should since data are frequently siloed, erroneous, or incomplete. Therefore, this requires careful data gathering, processing, and curation for the purpose of model training. |
Lack of Training Data | The most crucial step in ML model development is training the model using enough data in order to let the model obtain reliable outputs. Less training data will result in model outputs that are biased or erroneous. |
Infrastructure Requirements | In some oil and gas companies, the data infrastructure is inadequate, which makes it difficult to find the required data in the data retrieval process. Therefore, it is an essential requirement to maintain an appropriate data management infrastructure in a company for the easy use of available data to dig out embedded values. This will make testing various tools easier and also make data transfers easier. |
Feature Selection | Identifying the most informative features or input parameters for predicting SCC in gas pipelines can be a challenge. Different factors, such as pipe material, temperature, pressure, pH, environmental variables, and pipe geometry, can influence SCC. |
Incorporating Time-Dependent Factors | SCC in gas pipelines is a complex phenomenon that can evolve over time because of various factors, including aging, environmental changes, and operational conditions. Capturing and incorporating the temporal aspect of SCC into ML models can be a research gap. |
Lack of Labeled Data | ML models typically require labeled data for training and validation. However, obtaining labeled data for SCC in oil and gas pipelines can be difficult given the complex and expensive nature of conducting inspections and assessments. |
5. conclusions, author contributions, data availability statement, conflicts of interest.
Author Names | Reference No. | Year | Brief Details | Comments |
---|---|---|---|---|
Al-Sabaeei et al. | [ ] | 2023 | A systematic review highlights the complexity and effectiveness of ML methods in predicting pipeline failures, emphasizing factors such as dataset variations, data sources, and model complexity. It underscores the success of ANNs, SVMs, and HML in detecting defects, focusing on corrosion while also identifying a need for more diverse research on other failure types. | |
Ma et al. | [ ] | 2023 | A novel hybrid approach is presented to effectively estimate the burst pressure of corroded pipelines. It incorporates a feature space with physical importance and a fusion mechanism that combines empirical formula and collective learning. The suggested model, which uses the light gradient-boosting machine, exhibits better interpretability through feature importance analysis. | |
Alamri A. H. | [ ] | 2022 | This review summarizes the current state of ML applications in SCC for risk assessment. It identifies existing knowledge gaps, discusses challenges, and outlines future perspectives on utilizing ML and AI in corrosion risk assessment. | |
Liu and Bao | [ ] | 2022 | Explores the application of ML in automated pipeline condition assessment, leveraging advanced sensing technologies to analyze routine operations, NDT, and computer vision data. | |
Soomro et al. | [ ] | 2022 | Emphasizes the limitations of existing probabilistic models; this research advocates for Bayesian network approaches, offering insights, methodologies, and dataset considerations for risk analysis in evaluating corroded hydrocarbon pipelines. | |
Soomro et al. | [ ] | 2022 | Emphasizes the emerging role of machine learning in predicting pipeline corrosion, mainly through hybrid models like ANNs and SVMs, while also addressing current research gaps and proposing future directions for enhancing accuracy and validation in this evolving field. | |
Coelho et al. | [ ] | 2022 | This study emphasizes that localized corrosion and inhibition efficiency prediction is recommended, requiring large, high-quality training data and collaboration for systematic ML integration into the corrosion community. | |
Wasim and Djukic | [ ] | 2022 | This review includes an analysis of monitoring tools; models for corrosion prevention, prediction, failure occurrence, and remaining life; and insights into external corrosion management, reliability-based and risk-based models, and integrity assessment using machine learning and fuzzy logic approaches. | |
Khakzad et al. | [ ] | 2022 | Using a Bayesian network and an empirical corrosion simulation model, this research estimates corrosion rates based on factors like pipe diameter and flow conditions, subsequently converting these predictions into a distribution of failure probabilities. | |
Seghier et al. | [ ] | 2022 | Presents a robust ensemble learning approach for accurate internal corrosion rate prediction in oil and gas pipelines, utilizing four models: random forest, adaptive boosting, gradient boosting regression tree, and extreme gradient boosting. | |
Soomro et al. | [ ] | 2021 | This study proposes ML-based algorithms to estimate the probability of failure, leveraging extensive simulations to generate a rich dataset for comprehensive validation and providing insights into improved reliability assessment in the industry. | |
Sheikh et al. | [ ] | 2021 | Employing a hybrid approach integrating machine learning techniques, this research successfully predicts corrosion severity levels with high accuracy based on distinct features extracted from acquired acoustic emission data. | |
Rachman et al. | [ ] | 2021 | Explores the integration of ML in pipeline integrity management (PIM). This review covers ML applications across PIM elements such as inspection, monitoring, maintenance, and analysis techniques and addresses current challenges while also highlighting future research opportunities. | |
Reddy et al. | [ ] | 2021 | Emphasizing the importance of early detection and prevention, this review explores sensor technologies employing physical and electrochemical techniques, discussing their recent developments, sensitivity, selectivity, and standard inspection methods for corrosion monitoring. | |
Ossai, C. I. | [ ] | 2020 | This study uses a data-driven methodology to estimate increased corrosion defect depth (CDD) in oil and gas pipelines using a subspace clustering neural network (SSCN) and particle swarm optimization (PSO). | |
Jiang, P. | [ ] | 2018 | This thesis addresses the growing global demand for risk analysis in corrosion- and SCC-related failure events. It introduces an innovative method utilizing machine learning, including ensemble methods and support vector machines (SVMs) for automatic risk analysis. | |
Ratnayake and Antosz | [ ] | 2017 | Presents a novel approach to risk-based maintenance (RBM) analysis using fuzzy logic. The proposed approach extends the traditional RBM framework by incorporating fuzzy sets to represent the uncertainty associated with risk factors. | |
Aljaroudi et al. | [ ] | 2016 | Addresses the critical issue of offshore pipeline leak-detection system failures, emphasizing their potential operational and environmental consequences. Introduces a risk-based assessment methodology to evaluate system integrity, quantify associated risks, and guide decision-makers in determining appropriate preventive measures based on an acceptable risk threshold. | |
Hasan, A. | [ ] | 2016 | Introduces a risk-based security management method utilizing an analytic hierarchy process (AHP) model to assess the likelihood of pilferage in different pipeline sections, aiding in prioritizing security measures for effective prevention. | |
Guo et al. | [ ] | 2016 | Introduces a robust risk evaluation method utilizing a fuzzy Petri net (FPN) model to assess potential hazards in long-distance oil and gas transportation pipelines. | |
Parvizsedghy and Zayed | [ ] | 2016 | This work employs a neuro-fuzzy technique; the study develops a model utilizing historical data to predict and assess the financial consequences of potential failures, offering an 80% accurate tool for practitioners and academics involved in the risk assessment of gas pipelines. | |
Zhou et al. | [ ] | 2016 | Provides an analytical model based on fuzzy logic to determine the probability of corrosion-related issues in energy pipelines, considering corrosion cracking and thinning to be important variables. This model offers important insights into corrosion failure likelihood by considering variables like inspection efficacy and timing. | |
Lu et al. | [ ] | 2015 | This study offers a new method of assessing the possible risks related to natural gas pipeline leaks. The approach makes use of a risk matrix in addition to a bowtie model. | |
Zhou et al. | [ ] | 2015 | Provides a novel strategy for estimating the service time of subterranean gas pipelines before corroding under the cyclically loading condition. The methodology employs cumulative damage rates, models corrosion defect depths as an exponential function of elapsed time, and computes remaining life by using an iterative approach. | |
De Masi et al. | [ ] | 2015 | Addresses the growing challenge of maintaining the integrity of hydrocarbon pipelines over long distances because of aging plants and components in the oil and gas industry. By leveraging an ensemble of artificial neural networks (ANNs), the proposed ML approach demonstrates promising results in predicting the complex evolution of corrosion, outperforming traditional deterministic models and single-ANN models. | |
El-Abbasy et al. | [ ] | 2015 | Proposes a condition assessment model and uses both an analytic network process and a Monte Carlo simulation to consider the uncertainty of factors affecting pipeline conditions and the interdependency relationships between them. | |
De Masi et al. | [ ] | 2014 | Highlights the role of reliable corrosion predictions in pipeline integrity management, reducing economic impact, and preventing environmental damage. | |
Ismail et al. | [ ] | 2011 | Explores SCC in austenitic stainless steel in high-temperature aquatic surroundings, employing fact-based techniques such as classical statistics, machine learning, and fuzzy logic. The decision tree approach was found to be highly effective, demonstrating superior performance and intelligibility in addressing the investigated problem. |
Searching Index | Specific Content |
---|---|
Article Type | Publications in books, journals, and conferences |
Database | Web of Science, IEEE Xplore, Elsevier, Springer |
Classification | By the type of publication (i.e., concept, case study, and review), nationalities, application segments, enabling technologies, and affiliations (i.e., universities and industries) |
Focus | Determine opportunities and challenges related to SCC detection and prediction in the context of oil and gas |
Types of Learning | Data | Goal |
---|---|---|
Supervised | Labeled | Learn a mapping function |
Unsupervised | Unlabeled | Find patterns |
Semi-supervised | Labeled and unlabeled | Define a mapping function |
Reinforcement | Trial and error | Maximize rewards |
Corrosion Sensor Detection Technique | Type of Corrosion | Corrosion Phenomena/Parameter Assessed | Sensitivity | Field Monitoring Use |
---|---|---|---|---|
Acoustic emission (AE) | Stress corrosion cracks, pitting corrosion | Acoustic energy (impingement, leaks, and cracks) | Medium | Yes |
Image processing techniques (IPT) | General, localized corrosion, SCC, erosion–corrosion | Morphology of the corroded surface (image color, texture, and shape characteristics) | High | Yes |
Electrochemical noise (EN) | Uniform corrosion, localized corrosion (pitting, crevice), SCC | Electrical noise on the corrosion potential or current | High | Yes |
Hydrogen monitoring (HM) | Erosion–corrosion, stress corrosion cracking | Hydrogen diffusion through metal | High | Yes |
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Hussain, M.; Zhang, T.; Chaudhry, M.; Jamil, I.; Kausar, S.; Hussain, I. Review of Prediction of Stress Corrosion Cracking in Gas Pipelines Using Machine Learning. Machines 2024 , 12 , 42. https://doi.org/10.3390/machines12010042
Hussain M, Zhang T, Chaudhry M, Jamil I, Kausar S, Hussain I. Review of Prediction of Stress Corrosion Cracking in Gas Pipelines Using Machine Learning. Machines . 2024; 12(1):42. https://doi.org/10.3390/machines12010042
Hussain, Muhammad, Tieling Zhang, Muzaffar Chaudhry, Ishrat Jamil, Shazia Kausar, and Intizar Hussain. 2024. "Review of Prediction of Stress Corrosion Cracking in Gas Pipelines Using Machine Learning" Machines 12, no. 1: 42. https://doi.org/10.3390/machines12010042
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Citation: Roman I. Bogdanov, Yaakov B. Unigovski, Emmanuel M. Gutman, Iliya V. Ryakhovskikh, Roni Z. Shneck. Stress corrosion cracking of pipeline steels in near-neutral pH solutions: the role of mechanochemical and chemomechanical effects[J]. AIMS Materials Science, 2019, 6(6): 1065-1085. doi: 10.3934/matersci.2019.6.1065
沈阳化工大学材料科学与工程学院 沈阳 110142
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Bamford, WH. "Fatigue Crack Growth, Fatigue, and Stress Corrosion Crack Growth: Section XI Evaluation." Online Companion Guide to the ASME Boiler & Pressure Vessel Codes. ASME Press, 2020.
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Fatigue has often been described as the most common cause of failure in engineering structures, and designers of pressure vessels and piping have incorporated fatigue considerations into design requirements since the first edition of Section III in 1963. The development of this technology and its application in Section III are discussed in Chapter 39 of the third edition of this publication. Its application in Section XI is discussed in Section 32.3 of this chapter.
With the advancement of the state of the art, the ability to recognize how the growth and size of a crack can lead to failure has been enhanced. This technology has been a key aspect of the Section XI flaw evaluation procedures since the 1974 edition was published and will be discussed thoroughly herein.
Further advancements in the state-of-the-art in fracture evaluation resulted in the introduction of models for stress corrosion cracking into Section XI in 2001, and since that time additional models have been added and more are planned. This addition gives Section XI flaw evaluations a complete treatment of all the potential modes of cracking that can occur in components, and details will be provided in Section 32.2.
Since additional crack growth data, for both fatigue and stress corrosion cracking, is periodically made available on a regular basis from additional research testing, the models contained in Section XI are regularly reviewed and updated to ensure their applicability, accuracy, and reflection of the latest known material behaviors. This improved knowledge may lead to more conservative models for crack growth over time. To accommodate such situations, Section XI has adopted an approach to conservatively account for potential increases in crack growth predictions with augmented examinations. The rules of IWB-3600 require that augmented (additional) examinations must be performed at an accelerated pace for situations where a flaw indication was accepted by analytical evaluation. This rule also protects against NDE uncertainties at the location of interest.
The nomenclature used here will be identical to that used in Section XI; any nomenclature that is not present in Section XI will be explained as it appears.
This chapter has been updated using the 2019 Code edition.
Warren H. Bamford is the original author of this chapter and has since revised it for all of the subsequent editions including the current online edition.
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Austenitic alloys have been extensively used in the nuclear industry as structural components due to their combination of excellent mechanical properties and high corrosion resistance. Although these alloys have a good service record, many can become susceptible to stress corrosion cracking (SCC) under pressurized water reactor (PWR) primary water conditions. After several decades of study, a considerable number of factors have been revealed that affect SCC susceptibility, such as material...
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Student thesis : Phd
Date of Award | 31 Dec 2011 |
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Original language | English |
Awarding Institution |
File : application/pdf, -1 bytes
Type : Thesis
Experiments to explore the mechanisms of stress corrosion cracking., url to cite or link to: http://hdl.handle.net/1802/14819.
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To capture the stress corrosion cracking (SCC) in metallic materials, a phase-field framework considering localized plastic deformation and stress states is established. A new function of critical energy release rate is proposed to describe the degradation of cracking resistance of materials in the SCC process. This proposed framework can ...
Stress corrosion reactions only affect the cement; they do not affect the grains. Therefore, each parallel bond is a potential reaction site. (b) Stress corrosion reactions occur at the bond surface and remove bond material at a uniform rate that is proportional to the crack velocity in Eq. (3), which is, in turn, proportional to the reaction ...
Factors Affecting Stress Assisted Corrosion Cracking of Carbon Steel under Industrial Boiler Conditions A Dissertation Presented to The Academic Faculty by Dong Yang In Partial Fulfillment Of the Requirements for the Degree Doctor of Philosophy in the School of Mechanical Engineering Georgia Institute of Technology Atlanta, GA August, 2008
Stress corrosion cracking (SCC) is a form of material damage which relies on coupling between a microstructure, a chemical (and physical) environment and a stress state; it results in cracking of the material but a certain number of characteristics differentiates it from pure mechanical rupture. ... PhD thesis. Université de Toulouse (2015 ...
Stress Corrosion Cracking and Corrosion of Carbon Steel in Ethanolic Environments Doctor of Philosophy, 2022 Ali Ashrafriahi Department of Chemical Engineering and Applied Chemistry University of Toronto Several studies have recently claimed that the Stress Corrosion Cracking (SCC) of carbon steel in
This work addresses the risk of Stress Corrosion Cracking (SCC) in CO₂transport pipelines. The susceptibility of X80 pipeline steels in aqueous CO₂environments in the presence of nitrates and sulphites is investigated using electrochemical potentiodynamic tests and Slow Strain Rate Tests (SSRT) at 23 and 75°C. ... PhD Thesis: URI: http ...
Machine learning methods for corrosion and stress corrosion cracking risk analysis of engineered systems Author: Jiang, Peng ... A thesis in fulfillment of the requirements for the degree of ... First name: Peng . Other name/s: Abbreviation for degree as given in the University calendar: PhD . School: Materials Science and Engineering ...
His PhD thesis was distinguished with the Morris Cohen Award, awarded annually by The Electrochemical Society to outstanding graduate research in the field of Corrosion. ... Stress corrosion cracking of sensitized stainless steel in oxygenated high temperature water. Corrosion 1973; 29: 451-469. 10.5006/0010-9312-29.12.451 Search in Google ...
Stress corrosion cracking of P235 and P265 steels is investigated through slow strain rate and constant deformation tests on samples exposed in a cell reproducing the underground conditions of the French deep geological nuclear disposal, at 25 and 90°C. ... [PhD thesis]. France: Université Technologique de Compiègne; 2008 December 7. ...
Abstract: In this chapter, we present a few examples of stress corrosion cracking (SCC) studies with an emphasis on the various approaches implemented to reproduce, understand and propose solutions. Based on a few material/environment couples, we attach particular importance to the various study methodologies aiming to reproduce, at laboratory ...
Pipeline integrity and safety depend on the detection and prediction of stress corrosion cracking (SCC) and other defects. In oil and gas pipeline systems, a variety of corrosion-monitoring techniques are used. The observed data exhibit characteristics of nonlinearity, multidimensionality, and noise. Hence, data-driven modeling techniques have been widely utilized. To accomplish intelligent ...
Thesis (PhD)--University of Pretoria, 2013. tm2015. Materials Science and Metallurgical Engineering. PhD. ... The critical stress intensity for stress-corrosion cracking - K1scc - was determined by measuring the current flow from an external cathode, with comparisons with the crack growth rate. Crack propagation was measured with the potential ...
The review presents brief theoretical foundations of the mechanochemical and chemomechanical surface effects (MCE and CME) associated with corrosion-mechanical destruction of metals, considers the results of various scientific publications that studied the mechanochemical behavior of steels in near-neutral pH media, and analyzes the application of the results of such studies with the purpose ...
PhD School: Mining Engineering Faculty: Engineering Title: Laboratory-Based Investigation of Stress Corrosion Cracking Of Cable Bolts Abstract 350 words maximum Premature failure of cable bolts due to stress corrosion cracking (SCC) in underground excavations is a worldwide problem with limited cost-effective solutions at present.
Materials and Corrosion is a leading journal at the interface of materials science, metallurgy, and metallurgical engineering, with a focus on corrosion science. Modelling of intergranular stress corrosion cracking (SCC) in pipelines is an important field of study as predictive techniques are integral to pipeline integrity and management.
Further advancements in the state-of-the-art in fracture evaluation resulted in the introduction of models for stress corrosion cracking into Section XI in 2001, and since that time additional models have been added and more are planned. ... PhD Dissertation, Ohio State University, 1978. 49. Jones, R. H., Stress-Corrosion Cracking, ASM ...
Although these alloys have a good service record, many can become susceptible to stress corrosion cracking (SCC) Logos Skip to main NEW SEARCH Collections About Deposit HELP 0. Back to Search. CONTACT. Name. Email-Comment. Send message Actions ... [PhD thesis]. University of Oxford. Copy APA Style MLA Style. Shen, Z. Understanding the ...
Student thesis: Phd. Abstract The new generation of highly alloyed super duplex stainless steels such as Zeron 100 are preferable materials for industrial applications demanding high strength, toughness and superior corrosion resistance, especially against stress corrosion cracking (SCC). SCC is an environmentally assisted failure mechanism ...
Type II hot corrosion combined with static stress in Ni-based superalloys has not been extensively studied. However, stress corrosion cracking (SCC) is a well-documented failure mechanism especially in aqueous systems [ 12,13 ]. Studies have been conducted on the effects of stress on corrosion pitting growth in aluminium alloys [ 14 ].
It assumes the existence of a threshold of stress intensity factor. In the local model for intergranular stress corrosion cracking, Couvant et al. [CAB 16, COU 17] propose that the propagation step of a long crack can be modeled using a sigmoid law: [14.2] da / dt = A × K n 1 + exp − λ × K − K 0 × exp − Q RT.
Description : Thesis (Ph. D.)--University of Rochester. Materials Science Program, 2011. Abstract : Stress corrosion cracking (SCC) is a type of subcritical cracking of materials that occurs when a SCC susceptible material is simultaneously stressed in tension (applied or residual) and exposed to a specific corrosive environment.
Stress Corrosion Cracking of High-Strength Pipeline Steels. Y. Frank Cheng, Y. Frank Cheng. Search for more papers by this author. Book Author(s): Y. Frank Cheng, ... Mechanoelectrochemical Effect of Corrosion of Pipelines Under Strain. REFERENCES, () , , , - ...
This paper presents an analysis of Stress Corrosion Cracking (SCC) based on the Theory of Critical Distances (TCD). The research is based on an experimental program composed of fracture specimens with notch radius varying from 0 mm (crack-like defect) up to 1 mm, and tensile specimens.
Corrosion fatigue and stress corrosion cracking of magnesium alloys in a simulated physiological environment. Magnesium (Mg) alloys have attracted great attention as potential materials for temporary implants in uses such as pins, screws, plates and stents. ... Accordingly, this PhD thesis has attempted to evaluate CF and SCC resistance of one ...