Intergranular corrosion testing of austenitic stainless steels in nitric acid solutions G O H Whillock and B F Dunnett BNFL, B170, Sellafield, Seascale, Cumbria CA20 1PG U.K. Tel. +44 19467 79297 Fax. +44 19467 79003 e-mail [email protected]
Abstract In hot strong nitric acid solutions, stainless steels exhibit intergranular corrosion. Corrosion rates are often measured from immersion testing of specimens manufactured from the relevant material (e.g. plate or pipe). The corrosion rates, measured from weight loss, are found to increase with time prior to reaching steady state, which can take thousands of hours to achieve. The apparent increase in corrosion rate as a function of time was found to be an artefact due to the surface area of the specimen’s being used in the corrosion rate calculations, rather than that of the true area undergoing active corrosion i.e. the grain boundaries. The steady state corrosion rate coincided with the onset of stable grain dropping, where the use of the surface area of the specimen to convert the weight loss measurements to corrosion rates was found to be appropriate. This was confirmed by sectioning of the specimens and measuring the penetration depths. The rate of penetration was found to be independent of time and no induction period was observed. A method was developed to shorten considerably the testing time to reach the steady state corrosion rate by use of a pre-treatment that induces grain dropping. The long-term corrosion rates from specimens which were pre-treated was similar to that achieved after prolonged testing of untreated (i.e. initially ground) specimens. The presence of cut surfaces is generally unavoidable in the simple immersion testing of specimens in test solutions. However, inaccuracy in the results may occur as the measured corrosion rate is often influenced by the orientation of the microstructure, the highest rates typically being observed on the cut surfaces. Two methods are presented which allow deconvolution of the corrosion rates from immersion testing of specimens containing cut surfaces, thus allowing reliable prediction of the long-term corrosion rate of plate surfaces.
Keywords : intergranular corrosion, stainless steel, weight loss, nitric acid Introduction
Stainless steels generally exhibit good passivity in many nitric acid (HNO3) solutions and are accordingly routinely specified for this duty. In sufficiently oxidising conditions, however, generated either in strong aqueous solutions of nitric acid at or close to boiling or in weaker 4+ 2- 2- solutions incorporating powerful oxidants such as Ce or chromates (CrO4 or Cr2O7 ), stainless steels suffer intergranular attack and the corrosion rate can become appreciable. Intergranular attack occurs owing to the fact that grain boundaries dissolve much faster than the metal matrix due to metallurgical effects associated with impurities present in the steel [1-16]. The intensity of intergranular attack depends on the oxidising power of the solution and the exposure time; it may range from light etching of grain boundaries to intense penetration accompanied by grain dropping. In the latter case, the result is uniform metal loss even though the underlying mechanism is highly localised.
There are a number of difficulties associated with measuring the corrosion rate of stainless steels undergoing intergranular corrosion. The first and most obvious of these is that the corrosion rate appears to increase with immersion time until a steady state corrosion rate is reached. This is illustrated in Figure 1. The & R
immersion time required to reach a U U R constant corrosion rate depends on the V L R corrosivity of the solution and can be Q U D
very long (several thousand hours) in W H 7LPH K
some cases. Thus, unless testing is P continued for sufficiently long to P Figure 1 Gravimetric corrosion rate of 18/13/Nb ss \
reveal the steady state rate, it is in 8M HNO3 + 25 g/L Fe at 100 °C. possible that the rate of plant corrosion could be significantly underestimated. It is shown below that the apparent initial increase in corrosion rate is actually an artefact resulting from initial overestimation of the area of metal undergoing attack, which is a direct consequence of the progression of intergranular corrosion. The second difficulty is that the progression of intergranular corrosion of stainless steels in nitric acid solutions can be strongly anisotropic, depending on the orientation of the exposed grains. This originates from microstructural effects arising from the steel-making process. As an example of this, Figure 2 shows the corrosion penetration into each surface of a typical stainless steel plate, thus demonstrating the strong anisotropy which can be exhibited. This anisotropy creates a significant problem in estimating the lifetime of industrial plant, since cut surfaces are not usually exposed to the process liquor on plant, whereas in most corrosion testing, specimens are used which contain cut surfaces. In this connection, it is noted that techniques which involve the testing of a wire loop [17] are not generally applicable owing to the fact that wire made from the particular cast of metal under investigation is not normally available. In addition, the sewing-tape method [18] is not readily adapted to hot nitric acid. Overlaying weld metal on cut surfaces (which is known as buttering) is an effective method [19,20], but this requires large specimens (otherwise the heat input is excessive) which is not always appropriate owing to the fact that, depending on the corrosivity of the test liquor, an unacceptably high rate of release of corrosion products to the test liquor may occur leading to unrealistically high corrosion rates [1,21,22]. Methods which involve sealing a specimen into a holder [23] or Rolling direction clamping a plate onto a chamber are SODWH 50 µm not generally suitable in hot nitric acid Plate owing to the difficulty in maintaining an effective seal coupled with the VLGH inevitable introduction of a crevice 50 µm E which can seriously affect the Side nd measured corrosion rates [24]. Immersion testing of coupons HQG containing cut surfaces and hence 50 µm exposed end and side grain is thus often the most practical option, but Figure 2 Dependence of intergranular penetration on accordingly runs the risk of generating surface orientation (18/13/Nb ss). unrealistically high corrosion rates when compared with plant conditions. Finally, end grain pitting, which constitutes preferential attack parallel to the hot-working direction of plate, tube and forgings [25], is the most pronounced effect of microstructural orientation and can result in rapid perforation. Figure 3 shows an example of this. End-grain pitting in stainless steels has been linked with 500 µm the dissolution of aligned manganese sulphide inclusion stringers and Figure 3 End grain pit in NAG 18/10L ss (1500 h in niobium carbide precipitates in 8M HNO + 3 g/L Fe + 0.7 g/L Ru at 105 ºC). niobium-stabilised steels [26] and 3 sulphide inclusions in 304L steels [27]. However, for modern low-carbon steels where second phase precipitates are largely absent, end-grain pitting is primarily a consequence of the exposure of segregated material arising in the original production ingot, which has subsequently been distributed by hot-working as bands containing grain boundaries with enhanced susceptibility to attack as compared with the boundaries present in adjacent material [28,29]. During the course of studies directed at shortening the time required to determine the long-term corrosion rate, methods became apparent for determining the corrosion rate of the individual faces of a coupon. Accordingly, this paper presents a method for shortening the test time required to determine the long-term corrosion rate of stainless steels in hot nitric acid, whilst also allowing deconvolution of the corrosion rate obtained from gravimetric testing of a coupon with respect to each of its faces.
Experimental Table 1 Composition (wt%) of stainless steels used.
Three austenitic stainless Material C S P Si Mn Ni Cr Nb steels, used in the 18/13/Nb 0.090 0.006 0.070 0.51 0.99 12.60 16.60 0.92 construction of plant on the Sellafield site, were NAG 18/10L 0.015 0.008 0.015 0.32 1.54 9.64 18.77 <0.01 used in this study; Uranus 65 0.006 0.0006 0.138 0.12 1.01 20.04 24.58 0.13 18/13/Nb, NAG 18/10L and Uranus 65. The composition of these is given in Table 1. All were used in the solution annealed condition. The test solutions used were made from Analar grade reagents and deionized water and were selected to be representative (in terms of corrosivity) of various process streams encountered in nuclear reprocessing. As described later, test coupons were prepared either by wet grinding with successively finer SiC paper to 800 grit finish, or were pre-corroded in boiling 8M nitric acid containing 1 g/L Cr added as CrO3. Full experimental details are given elsewhere [30,31].
Penetration measurements
A number of 18/13/Nb stainless steel coupons (initially ground to 800 grit on all SODWH VLGH faces) were tested in 8M HNO3 containing P 25 g/L Fe at 100 °C. At various intervals, Ã coupons were removed and measured to WK HS G determine any thickness loss; macroscopic [Ã D dimension changes were found following 0 the onset of grain dropping as evinced by the appearance of fine soot-like particles in the bottom of the test vessel. The removed coupons were then metallographically 7LPHÃ K mounted for microscopic examination to determine the extent of intergranular Figure 4 Intergranular penetration vs. time for penetration into each face (including the 18/13/Nb ss in 8M HNO3 + 25 g/L Fe at 100 °C. thickness loss if detected). At the same time, the weight loss of four nominally identical coupons was measured EURDGÃIURQW periodically (these data are presented as SLWOLNH Figure 1). Figures 4 and 5 shows the progression of intergranular corrosion revealed by the
above measurements. For penetration into 0 D the plate and side faces, the data can be [ G H fitted reasonably well to linear S W K relationships, indicating a constant rate of
penetration into either surface for the entire P
7LPH K period of testing. This is in marked contrast to the corrosion rate calculated Figure 5 Intergranular penetration of the end from weight loss (see Figure 1), which face of 18/13/Nb ss in 8M HNO3 + 25 g/L Fe at increased linearly until the onset of stable 100 °C. grain dropping, thereafter becoming constant. Linear regression of the data presented in Figure 4 gives average penetration rates of 0.78 and 1.27 mm/y for plate and side faces respectively. For penetration into the end grain, the data are more difficult to interpret owing to the presence of deeper pit-like penetrations extending beyond the broad corrosion front. Figure 5 shows that the maximum penetrations associated with the broad front can be fitted reasonably well to a linear relationship, giving an average rate of 1.49 mm/y. However, the pit-like penetrations, which became more pronounced with increased test time, show considerably more variation. Notwithstanding this, fitting a line to the data using linear regression gives an average rate for pit-like penetration of 2.82 mm/y. Since the plate, side and end surfaces respectively contributed 60%, 24% and 16% of the whole nominal surface area of the test coupons, the average penetration rate is accordingly calculated as 1.01 mm/y or 1.22 mm/y, depending on whether the rate for broad-front penetration or pit-like penetration is used for the end grain. Figure 1 shows that the steady state corrosion rate of this stainless steel, obtained by weight loss normalised with respect to nominal surface area, was 1.18 mm/y. It is thus evident that measurement of the intergranular corrosion rate by weight loss is correct, but only if the metal is undergoing stable grain dropping. The above results also suggest that end grain pitting, for this steel at least (which has a high susceptibility), contributes significantly to weight loss. Prior to steady-state grain dropping, the apparent initial increase in corrosion rate obtained gravimetrically (Figure 1) is clearly an artefact; this is most probably due to normalisation of the weight loss with respect to the nominal surface area of the test specimen, which is clearly incorrect owing to the localised nature of the attack. This, in our view, accounts for the continuously increasing corrosion rate reported by others in relatively short-term tests carried out in nitric acid solutions [1,3,24,32-35]. Since the penetration rate is constant, it is clearly the case that the actual area of the metal undergoing corrosion must increase with time, but reaches a constant value after the onset of stable grain dropping. This explains the constant corrosion rate observed gravimetrically after the onset of stable grain dropping. In addition, there is clearly a weight loss contribution due to the loss of grains which will increase with time, becoming constant once steady state conditions are achieved. However, in our experience, the grain dropping contribution amounts to at most 30% of the corrosion rate, and hence is insufficient to account for the totality of the observed effect. Accelerated testing – pre-corrosion to induce grain dropping The above findings suggested that a pre-corrosion treatment to induce 3UHFRUURGHG \ *URXQG grain dropping could provide the P P means to shorten the test time à WH D required to obtain reliable ÃU Q R corrosion rates by weight loss. The VL R U selected treatment was carried out R ÃF by placing the test coupons in LF G R boiling 8 M nitric acid containing 1 UL H 3 g/L Cr added as (CrO3) for 24 hours. This pre-corrosion treatment is sufficiently aggressive to initiate 7LPHà K wholesale grain dropping on 18/13/Nb stainless steel, although repeated treatments were found to Figure 6 Effect of initial condition on corrosion rate of be required for NAG 18/10L and NAG 18/10L ss in 9M HNO3 + 0.9 g/L Fe + 0.19 g/L V at Uranus 65 stainless steels. Figure 6 95 °C. demonstrates that the pre-corroded coupons displayed a constant corrosion rate (ie. were at steady state from the outset of exposure to the test liquor). More importantly, it is clear that the corrosion rate obtained from them was the same as that eventually reached by initially ground coupons. Extensive testing carried out by ourselves has confirmed this finding for both 18/13/Nb and Uranus 65 stainless steels tested in different liquors having previously been pre-corroded by the Cr(VI) treatment. However, on a cautionary note, we have found that the Cr(VI) treatment is too aggressive for some other stainless steels and leads to initial over-estimation of the true long-term corrosion rate; in such circumstance, a less aggressive treatment employing V as the accelerator has been found to be more appropriate. Deconvolution of coupon corrosion rates The first method devised by us exploits the large short-term difference in weight loss between pre-corroded and ground surfaces, this difference arising not from different actual corrosion rates, but from different effective areas undergoing active corrosion. From a series of coupons prepared to include variously pre-corroded and ground faces, it is possible to calculate the steady state corrosion rate of each face if it is assumed that the corrosion loss contributed by ground faces is not anisotropic (which is not actually correct, but appears to be insignificant, at least for relatively short-term testing). Figure 7 illustrates the simplest series of coupons required. Additional coupons are needed if it is required to compensate for the effect of a hole drilled through the coupon to allow it to be QRQH DOO SODWH VLGH HQG suspended from a hook [30] or to separate the contributions 5' made by the inner and outer surfaces of a tube. The coupons containing both pre- VKDGHGÃIDFHVÃDUHÃSUHFRUURGHG corroded and ground faces were made slightly oversize so Figure 7 Simplest series of coupons required to determine the that, following pre-corrosion, corrosion contribution made by plate, side and end faces. the faces which were to be ground could be ground back to virgin metal whilst leaving these coupons with the same nominal dimensions as the entirely ground and entirely pre-corroded coupons. The deconvolution method is simple: the weight loss from the entirely ground coupon is normalised with respect to its entire surface area; this value is then used to calculate the weight loss contributed by the ground faces present on coupons containing pre-corroded faces; the weight loss contributed by the pre-corroded faces is then found by subtraction and normalised with respect to the nominal surface area of those faces; the percentage contribution made by each face is then calculated from the weight loss given by the entirely pre-corroded coupon. Table 2 presents the results obtained for 18/13/Nb tested in 8M HNO3 + 12.5 g/L Ru Table 2 Deconvoluted corrosion rates and face - + 8.48 g/L Pd + 15 g/L Fe + 172.5 g/L NO3 contributions for 18/13/Nb ss. at two temperatures (and hence different liquor corrosivity). The fact that the Rate (mm/y) Contribution (%) percentages obtained at the two different 80 °C 100 °C 80 °C 100 °C temperatures were similar (for plate, and Plate 0.095 0.638 74 70 side, but not end grain) is noteworthy, since Side 0.134 0.916 104 100 the corrosion rates in the two tests were significantly different. This suggests that the End 0.258 1.277 200 140 percentages obtained should be expected to Coupon 0.129 0.914 n/a n/a be relatively insensitive to the test conditions producing intergranular attack and hence Table 3 Comparison between deconvoluted and should be readily applicable (for the same actual corrosion rates (values in italics are directly material and for coupons of the same measured). nominal dimensions, although the required factors can readily be calculated for different Deconvolution Penetration coupon geometries [30]). To illustrate this, Plate 0.83 0.78 Table 3 compares the calculated face Side 1.18 1.27 corrosion rates (obtained from the measured End 1.65 1.49/2.82 whole coupon rate and the contribution Coupon 1.18 1.01/1.22 factors given in Table 2 for 100 °C) with those measured directly by the penetration method, for 18/13/Nb tested in an different liquor (8M HNO3 + 25 g/L Fe at 100 °C). The agreement is very close. Where two values are given in Table 3, this relates to the difference between broad-front and pit-like attack of the end grain. For exposed end grain, Table 2 shows that proportionally more attack occurred at 80 °C than at 100 °C (ie. at lower overall corrosivity). It is considered that this may be due to relatively facile dissolution of exposed inclusion stringers or bands of material containing more highly-segregated grain boundaries, not greatly affected by increase in solution corrosivity. Thus, at lower temperature and hence lower corrosivity, a proportionally larger end-grain effect might be expected owing to the lower corrosion rate of the remaining end grain surface. The deconvolution method described above is labour intensive involving the preparation of a complex series of test coupons. In order to overcome that, a second deconvolution method was devised which uses entirely pre-corroded coupons. The corrosion rate (W), of a coupon for which the area fractions of plate, side and end faces are denoted by p, s and e respectively is given by: W = Pp + Ss + Ee (1) where P, S and E are the corrosion rates of the plate, side and end faces respectively. It follows then that P, S and E can be determined unequivocally by measuring W for three coupons selected to have different values of p, s and e and then solving the three simultaneous equations generated. It is of course required that the test conditions are such that it can reasonably be assumed that P, S and E are constant; in practice, this requires all three specimens to be taken from the same plate (and from adjacent areas) and to be tested simultaneously in the same test vessel in order to ensure exposure to identical conditions. In order to ensure that measurement of weight loss accurately reflects the intergranular penetration rate, stable grain dropping is required, as explained above. Although not strictly required, we have found that this method works best if the three coupons are sized such that one contains a large proportion of plate surface, the next is predominantly side surface and the last is mainly end surface. Table 4 illustrates the results obtained for NAG 18/10L tested in 8M HNO3 + 0.7g/L Ru + 3g/L Fe at 103 °C. Within the scatter, which is considered to be largely a consequence of grain dropping, but may also be influenced by the variable period times (ie. by variable Table 4 Deconvoluted faces rates obtained from testing corrosion product build-up), the 3 coupons, each consisting primarily of end, side and plate periodic rates are time independent. faces respectively. This is consistent with expectation for material undergoing stable grain- W (mm/y) Face rates (mm/y) dropping, demonstrating the efficacy t (h) ‘End’ ‘Side’ ‘Plate’ ESP of the pre-corrosion treatment. Using an identical method, the corrosion 71 1.74 1.30 0.96 1.89 1.31 0.86 rate of Uranus 65 stainless steel in 144 2.22 1.76 1.48 2.36 1.76 1.38 the same liquor was deconvoluted. 236 1.74 1.30 1.22 1.84 1.28 1.16 Figure 8 compares the results 313 2.16 1.58 1.45 2.30 1.57 1.37 obtained for the two steels. The 408 1.72 1.25 1.15 1.83 1.24 1.08 individual face rates for Uranus 65 476 1.47 1.01 0.87 1.59 1.01 0.80 are all lower than for the 567 1.78 1.51 1.08 1.91 1.54 0.98 corresponding NAG 18/10L faces, Mean 1.83 1.39 1.17 1.96 1.39 1.09 indicating improved corrosion resistance. This is most probably due Std dev 0.27 0.25 0.23 0.27 0.25 0.23 ) 2 to the increased chromium content and the y / NAG 18/10L relatively low oxidising power of the m Uranus 65 m ( liquor (otherwise the reverse effect would 1.5 e t be expected, as has previously been a r
pointed out [28]). For NAG 18/10L, the n o 1 i corrosion rate increases in the order plate < s o r side < end, which is consistent with r o
c 0.5 general expectation. For Uranus 65, the n a
corresponding order is plate < side = end. e These data show that cut surfaces of M 0 Uranus 65 display significantly inferior Plate Side End-grain corrosion resistance to the plate surface, but the orientation of the hot-working Figure 8 Deconvoluted corrosion rates for (rolling) direction is apparently not NAG 18/10L and Uranus 65 tested in 8M HNO3 important. For NAG 18/10L, however, + 0.7g/L Ru + 3g/L Fe at 103 °C. exposed end grain is clearly the least corrosion resistant face, significantly worse than side-grain. This deconvolution method is now standard practice in our laboratory where the main interest is focussed on the determination of plate and tube corrosion rates from coupon testing. The coupon corrosion rate is in all cases affected by the cut surfaces, such surfaces being unavoidably present and capable of distorting the corrosion rate even when their area fraction is minimised as far as is reasonably practicable. This is not a new finding, as clearly implied by the coupon-design requirement of the Huey test [36], even though 25 this was not part of Huey’s original specification 1$*Ã/ 8UDQXVÃ [22]. However, by application of the 20 deconvolution method, it is possible to arrive at a r
e 15 convenient coupon design for which a correction b m factor can be calculated to covert the coupon rate u 10 directly into the contribution provided by the N plate or tube surfaces. This of course provides 5 more realistic data with which to assess plant 0 lifetimes. Once a particular batch of material is 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 1.1 1.2 1.3 deconvoluted, the results from testing in a wide Depth (mm) variety of solutions (capable of supporting (a) 8 intergranular corrosion) can be readily 1$*Ã/ interpreted. 8UDQXVÃ 6 r e
End grain pits b 4 m u
Figure 9 shows the distribution of end grain pit N depths found in random cross-sections which 2 were 24 mm in total length for both NAG 18/10L 0 and Uranus 65 stainless steels tested under the 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 1.1 1.2 1.3 same conditions detailed above. Figure 9(a) Depth (mm) shows the distribution present immediately (b) following the pre-corrosion treatment (2 x 24 h treatments for NAG 18/10L and 3 x 24 h Figure 9 Distribution of end grain pits treatments for Uranus 65), whereas Figure 9(b) after: (a) pre-corrosion treatment; (b) shows the distribution found following prolonged prolonged testing in 8M HNO3 + 3 g/L Fe exposure to the test solution. The distribution of + 0.7 g/L Ru at 105 ºC pit depths appears broadly similar for either stainless steel, although clearly many more pits occurred in the NAG 18/10L. Comparison between Figures 9(a) and 9(b) indicates that considerable end grain pit growth occurred in the test solution for both metals. The high incidence of pits < 0.1 mm in depth in NAG 18/10L following the pre-corrosion treatment may possibly be over-estimated (ie. it was difficult to distinguish which features were very small pits and which were just general intergranular penetration). Including the broad-front corrosion loss given by the deconvoluted end grain rates found previously (Figure 8), the maximum pit penetrations found were 1.6 mm for NAG 18/10L and 1.4 mm for Uranus 65. The maximum end grain pitting rates observed in the test solution were thus 7.0 mm/y and 4.5 mm/y respectively. These results are not considered significantly different and it is noted that it is unlikely that the maximum extent of penetration was actually found (systematic, repetitive, sectioning would be required, which was not done). Hence, it is considered that no significant difference has been found in the end grain pitting rate, but that NAG 18/10L gives a much larger number of end grain pits compared with Uranus 65. The underlying cause of this behaviour is not known unequivocally to us, although it is likely to be due to more extensive segregation (probably positive phosphorus) in NAG 18/10L than is the case in Uranus 65 (probably negative chromium segregation) [28,29]. In comparison with the broad-front intergranular corrosion rate (Figure 8), the maximum observed end grain pitting rates are ~3.6 times larger for NAG 18/10L and ~4.3 times larger for Uranus 65. It is clear that the weight loss deconvolution method does not predict the end grain pitting rate. No reliable method, other than careful sectioning or radiography, is presently known to us for measuring the end grain pitting rate. Conclusions 1. The intergranular corrosion rate of stainless steels in nitric acid solutions is time-independent; the apparent initial increase in rate, as measured by weight loss, is an artefact due to incorrect use of a constant nominal area in calculating the corrosion rate. 2. Sectioning to determine the penetration rate directly works well, but is laborious. 3. Weight loss gives the corrosion rate correctly if the metal being tested is at steady state (ie. undergoing stable grain dropping). The time required to achieve this can be long for initially ground coupons. 4. The test time can be considerably shortened by the use of a more aggressive pre-corrosion treatment, appropriately selected for the stainless steel being tested, which induces stable grain dropping. 5. The deconvolution method described can be used to determine the corrosion rates of individual faces of test coupons, thus eliminating variability due to the use of different coupon geometries by different workers. 6. End grain pitting can be measured at present only by careful sectioning or radiography.
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