Stress-Corrosion Cracking in Aluminum Beverage Can Ends-Issues, Observations, and Potential Solutions
By DeWeese, S K Ningileri, S T; Das, S K; Green, J A S
This paper examines the current status of stress-corrosion cracking (SCC) in the manufacturing and use of 5182 aluminum beverage can ends. Stress-corrosion cracking failures of the aluminum easy-open ends is an infrequent and undesirable failure encountered with filled beverage containers. Although SCC is responsible for a relatively small percentage of aluminum beverage can failures each year, it is still a concern to can manufacturers, fillers, and distributors. A change in SCC failure mode from transgranular to intergranular SCC has been observed in 5182 aluminum beverage can ends. This has been observed since the introduction of new end designs with higher stress promoting SCC fractures. The transition of cracking mode is attributed to both anodic dissolution and hydrogen embritttement. In this paper, these mechanisms are discussed including fractography evidence, identifying the probable causes and solutions. Some suggestions are made to mitigate the tendency for SCC at the fillers. INTRODUCTION
Stress-corrosion cracking (SCC) is an insidious and serious form of metallurgical attack that has compromised the quality and reliability of metallurgical products and has frustrated the best efforts of numerous scientists and engineers for decades. The first manifestation of SCC was reported by E.S. Sperry1 at the beginning of the 20th century. This was the failure of brass cartridge cases of the British army during the monsoon season in India. This seasonal cracking, as it was then known, was later shown to be aggravated by the fact that the ammunition was stored adjacent to the horse stables. Since then, an ammoniacal environment has become one of the standard test environments to evaluate the stress- corrosion behavior of alpha brasses.
The field of stress corrosion became especially active during the 1960s and 1970s, driven by the space program, the need for greater system reliability especially with manned space systems, and the fact that there were some spectacular and unexpected failures. The conference proceedings Fundamental Aspects of Stress Corrosion Cracking2 from a meeting held at Ohio State University in September 1967 provides an excellent basis for understanding the breadth and perverse nature of the phenomena. The unexpected failures of several titanium propellant tanks from the Apollo program, while being pressure tested in supposedly inert environments of methanol and nitrogen tetroxide, added impetus to the research efforts. At this time, it was not appreciated that trace quantities of certain species (e.g., chloride ions) could cause rapid crack propagation in titanium alloys.
Prior to the 1967 conference, many investigators had argued that there must be some overall general mechanism of stress corrosion. However, the large number of metallurgical systems that have been shown to be susceptible such as copper and silver alloys, Fe-Ni-Cr alloys, low-carbon and high-strength steels, aluminum alloys, and titanium alloys, together with the huge variation of environments that cause susceptibility, ranging from ammonia to traces of chlorides and even moist air, belies this mechanistic assumption. Now it is generally agreed that SCC can occur by a variety of mechanisms when a susceptible metallurgical structure is subjected to a sufficient stress level in the presence of a specific corroding environment. And as metal systems are engineered to higher performance levels and thinner metal sections, as has been occurring with the aluminum beverage can, the likelihood of SCC failure becomes more prevalent. see the sidebar for details on can lid design.
SIGNIFICANT SCC MECHANISMS
Brittle Film Mechanism
According to the brittle film mechanism model, SCC is propagated by the repeated formation and rupture of some sort of brittle film. Probably the most successful application of this model is to explain the observations in the cracking of brass in ammoniacal environments where a heavy oxide film, or tarnish, is observed.5 In this model, the applied stress causes a crack to run through the brittle oxide film and then the crack blunts out upon reaching the underlying ductile metal. This generates slip bands and exposes fresh metal to the environment, which causes additional oxidation and the cycle repeats. This mode of propagation is predicted to cause a series of steps, or striations, on the fracture surfaces and indeed such striations are observed. Interestingly, the path of the cracking in the brass ammonia system can be varied from intergranular to transgranular cracking by changes in the chemistry of the ammonia environment and by changes in the tarnish film.
Stress-Assisted Anodic Dissolution
Stress-assisted anodic dissolution has been most closely associated with SCC failures in aluminum alloys, which are generally intergranular in nature. Here, the grain boundary regions are considered to dissolve preferentially because they are more anodic in nature, or contain anodic magnesiumcontaining phases such as Mg^sub 2^Al^sub 3^ in the case of the 5XXX alloys, or MgZn^sub 2^ for the 7XXX alloys. The applied stress continually breaks down the oxide film in the grain boundary region, enabling and enhancing anodic dissolution, and crack propagation occurs along the boundaries. This mechanism has been the basis for the development of stresscorrosion-resistant alloys and generally has stood the test of time. Undoubtedly anodic dissolution does occur in aluminum alloys prone to SCC.
More recently, it has been suggested that high-strength aluminum alloys are stress cracked by some sort of hydrogen embrittlement mechanism.6,7 In this model, hydrogen, either in the alloy or introduced by the corrosion process, forms a brittle layer or region which, when stressed, causes a crack to advance rapidly, generally in a transgranular fashion. In some variations of this model, hydrogen is proposed to form an actual brittle hydride with aluminum or magnesium, but there is limited evidence for this in the literature. In detail, this model relies on the fact that crack tip environmental conditions in aluminum are acidic (pH ~3.5) in nature. This implies that the cathodic corrosion reaction is the recombination of hydrogen ions to hydrogen gas, which could be absorbed by the metal. From the fracture-mechanics viewpoint, the applied stress plays a role of enhancing the diffusion of hydrogen to a region of stress triaxiality ahead of the crack tip, which then becomes embrittled. This model can also explain the influence of a change of loading mode on the SCC susceptibility (mode 1 tensile loading causes more severe attack than mode 3 torsional loading-in the latter case, there is no, or much less, stress triaxiality to cause concentration of the hydrogen near the crack tip). It also explains the greater susceptibility in the presence of cathode inhibitors, such as contaminating oils, etc. Assuming that hydrogen is generated as part of an overall corrosion process, one implication of this model is that anodic dissolution and hydrogen embrittlement may occur simultaneously in a susceptible aluminum alloy.
Another model that has been advocated is that a critical species (e.g., chloride ion) is adsorbed at a crack tip and there it reduces the bond strength of the metallic bonds to the extent that applied stress causes a crack to propagate. This model, which attempted to account for the pervasive role of chlorides in causing so many SCC failures, has fallen out of favor in the literature, however, given the enormous number of different metal-environment systems that can suffer SCC.
CAN END DESIGN FACTORS
When filled with a typical soft drink, an aluminum beverage can has an internal pressure of ~370 kPa at room temperature. The internal pressure is developed from the gases in the head space being in equilibrium with the dissolved carbon dioxide (carbonation) in the liquid. The internal pressure causes a doming strain in the end which induces biaxial tensile stresses in the end panel of the beverage can.
The scoring groove in the easy-open end (as described in the sidebar) is formed by incising a trapezoidalshaped die into the pre- coated aluminum from the external side of the end. This operation displaces the aluminum into compressive strains that may be either symmetric or non-symmetric to the score groove, depending on the design of the score aperture (Figure 1).
Stress-corrosion cracking initiates at the score groove flank- land junction and propagates from the exterior side of the can end toward the pressurized or interior side of the end. The material remaining after scoring, known as the score residual, is typically 35 to 40 percent of the non-deformed end stock thickness. The score groove is formed by compressively incising the groove into the metal (strain, epsilon = ~ 0.60-0.70).
The stress induced by the scoring process is partly responsible for one of the three components (see Figure 2a) necessary for SCC to occur, namely: a susceptible alloy, environmental conditions specific to cause SCC, and tensile forces (from residual and applied stresses) exceeding the threshold stress for initiation. The strain field around the incised score is very complex and consists of a compressive strain gradient decreasing from the score land-flank juncture (score groove edge) and tapering to non-deformed metal adjacent to the score. The strain field results in residual stresses in the deformation zone surrounding the score. This is shown in Figure 2b.
Figure 3 shows the effects of score width and flank angle on the lateral strain at the score residual. At a given score residual, the narrower the score width and score flank angle, the lower the lateral strain. Historically, this was used to minimize the degree of internal coating breaks (metal exposure) by having beverage- style and beer-style score knives. The beverage-style score knife had a lower included angle than the beer-style score knife, at the expense of tooling service life.
For a given score geometry (score width and flank angle), the lateral strain within the score residual is inversely proportional to residual thickness. As the score residual becomes thinner, both the lateral strain and the residual stresses increase.
Can Internal Pressure Issues
The aluminum beverage can requires an internal pressure for strength and handling resistance during transportation. The amount of internal pressure in a beverage can depends on several factors: the carbonation or pressurization levels used at the filler, the nature of the product (sugared or diet products), head space volume, and temperature.
The pressure inside a beverage can is directly proportional to temperature. The increase in pressure with increasing temperature results from two components: the increase in the gas pressure with temperature for a constant head space volume, and the decrease in the amount of carbon dioxide dissolved in the liquid with increasing temperature.
The solubility of carbon dioxide is greater in sugared soft drinks than in diet soft drinks. Therefore, cans of sugared soft drinks have lower internal pressures than cans of diet soft drinks, especially at temperatures above 210C. This is shown in Figure 4. Most cans and ends in North America are produced with a guaranteed minimum strength to resist an internal pressure of 620 kPa.
Pressures in cans of beer are generally less than soft drink products due to lower carbonation levels. Beer usually has a carbonation level of about 2.8 volume of CO2 per volume of liquid, as opposed to soft drinks with typical carbonation levels of 3.6 to 4.0 volumes of CO2 per volume of liquid. Most beer products in bottles and cans are pasteurized after filling at temperatures of around 60[degrees]C. Internal pressure in beer cans during pasteurization usually exceeds 620 kPa. Stress-corrosion cracking failures are very rarely seen in beer products, probably because of the selfdrying effect of the warm cans following pasteurization.
SCC FAILURES AND FRACTOGRAPHIC EXAMINATION
Stress-corrosion cracking failures are predominately intergranular in aluminum alloys. However, transgranular SCC has been observed for a few alloys under highly specific environmental conditions.11,12
Stress-corrosion cracking failures in the older end designs (stay- on-tab [SOT] or ecology end) were always observed to be transgranular. With the introduction of the newer end design (large opening end [LOE] design), failures were observed to be either intergranular or transgranular, depending on specific end and environmental factors.
The fracture surface for transgranular SCC failure in 5182 ends has a unique appearance, and is shown in the SEM fractograph (Figure 5). Fracture surfaces display a delta or fan type morphology that propagates across the grain structure from the exterior to the interior surfaces of the end. The fracture propagated from a series of simultaneous initiation sites. A typical feature seen in these transgranular fractures is that the stable crack growth phase can propagate up to 75 to 80 percent of the score residual before final fracture occurs. This indicates a relatively low stress state for transgranular SCC crack propagation for this system.
Location of SCC Initiation Sites
The stresses along the perimeter of the score panel also vary. The older SOT score design typically showed the initiation of SCC failures at the 4 o’clock and/or 8 o’clock positions (with the 12 o’clock orientation being the tab rivet). Stress-corrosion cracking failures in this SOT end were always transgranular.
With the introduction of the newer large-opening end (LOE) score design, the typical site for SCC initiation changed to the 6 o’clock position. The LOE design is approximately 38 percent larger than the SOT design. This end design experiences either transgranular or intergranular SCC failure, depending on the level of stress.
These SCC initiation sites are the areas of highest tensile stresses (residual plus applied stress). These sites appear to be associated with the positions on the score panel that are adjacent to the end countersink with the lower radii of curvature (the straightest sections of the perimeter). The SCC initiation sites are shown in Figure 6.
The high strain region for the LOE is the 6 o’clock location, and the high strain regions for the SOT are the 4 o’clock position, followed by the 8 o’clock position. Also, the degree of maximum lateral strain in the LOE design is significantly greater than shown in the SOT design. Note that these locations for maximum lateral strain coincide with the initiation sites for SCC fractures in each of the two end designs.
The SCC Test Results
The intergranular failure in 5182H39 alloy was first observed with the introduction of the LOE end design in 2001. The stress concentration factor for initiating SCC failures depends on the design of the end itself, the score residual, and score knife wear (now less of a factor with current manufacturing practices). In addition, the “openability” of the end is enhanced by stresses (residual and applied) along the score. Stress-corrosion cracking occurs at specific high-stress points along the score panel.
Figure 7 shows the effect of stresses (applied stress in combination with residual stresses) on the SCC failure rate for the LOE end design. The test used a series of ends made on the same conversion press using the same tooling. Also, all ends were produced from a single coil of coated end stock (5182-H39). Five different residual ratios (the ratio of the score residual, t, to the starting end stock gauge, T) were produced and varied from 0.250 to 0.477. For each residual ratio, two different metallurgical conditions were tested: as-formed (5182-H39) and stress-relieved (5182-H28,163[degrees]C for 30 min.)
The testing consisted of an exposuretype test using ends double- seamed onto filled beverage cans, carbonated to an internal pressure of 276 kPa. The ends were exposed to a weak solution of NaOCl (140 ppm) at room temperature (22[degrees]C). The duration of the test was 40 days. The aggregate or mean time to failure of the test samples were determined for the 40 day test. Mean time to failure for the as-formed ends varied from 11 days to 24 days. The mean time to failure for the stress-relieved samples varied from 35 to 40 days,
The most significant outcome of this test was the change of failure mode to intergranular SCC for the as-formed samples with the lower residual ratios (Figure 7). Prior to the introduction of the LOE end design, only transgranular SCC failures had been observed in aluminum beverage can ends. This is one of the rare occurrences of transgranular SCC in aluminum alloy systems.9 The transition from transgranular to intergranular SCC corresponds to an increase in effective stress at the point of initiation (thinner cross- sectional area for a constant internal pressure).
Figure 8 shows the rate of failure for the as-formed ends experiencing intergranular SCC, as shown in Figure 7. Note the dramatic shift in the failure rate and the total percentage failed by the slight change in the residual ratio from 0.307 to 0.250. For the residual ratio of 0.250, all the ends failed within 3 days. But for the residual ratio of 0.307, only about 70 percent of the samples failed over a period of 13 days. This correlates to an increase in critical stresses at the point of initiation.
The fractograph in Figure 9 shows the intergranular SCC failure mode. The failure initiates at the score residual and score flank junction and then propagates toward the internal surface by intergranular dissolution. The elongated grain structure is clearly apparent. By contrast, the fractograph in Figure 10 shows the unique combination of mixed-mode SCC in the 5182 lid. The fracture surface shows adjacent areas of transgranular and intergranular crack propagation. The observation of two different paths of crack propagation implies a change of crack mechanism in the SCC process.
The internal pressure within the headspace of the can produces a biaxial tensile stress within the can end. This pressure produces a nominal tensile stress within the end (sigma^sub non^). At the score, the stress intensity factor, K^sub t^, yields a stress within the residual (sigma^sub max^) that is significantly higher than the nominal stress. This is schematically illustrated in Figure 11. Based on typical cross-sectional dimensions, K^sub t^ has a value of about 3. This does not include a shape factor contribution. Note that a lower score residual (i.e., a thinner cross section) results in higher stresses and higher susceptibility to SCC.
The contribution of metallurgical structure is only now being appreciated in the sensitivity of SCC failures in aluminum beverage cans. The high magnesium 5000 alloy series (> -3% magnesium) is sensitive to SCC whereas the 3000 alloy series (manganese-based body stock) is not considered sensitive to SCC. With reference to alloy composition, material with a higher Si:Fe ratio appears to be more sensitive to SCC failure.
In summary, a change in SCC failure mode from transgranular SCC to intergranular SCC has been observed in 5182 aluminum can beverage ends. This occurred with the introduction of a new end design. Intergranular SCC in this end design is associated with higher stress conditions (from both applied and residual stresses) for the initiation and propagation phases of the SCC fractures. Stress- corrosion cracking failure in aluminum beverage cans is an ongoing industrial issue. Much effort continues to be made to heip determine the safe operating procedures for the aluminum producers, aluminum can manufacturers, beverage fillers, and distributors.
RELATING OBSERVATIONS TO SUGGESTED THEORY
Observations of Fracture Surfaces
Figure 7 indicates that the stress conditions can modify the apparent failure mode between intergranular SCC to a mixed-mode fracture and then to transgranular SCC as one moves from left to right on the abscissa, while all other test and environmental conditions are maintained constant. This change of fracture path suggests that two distinct mechanisms of SCC may be operating in the lid material.
The fractograph shown in Figure 9 is typical of the appearance of the intergranular mode of SCC failure. In this cross-sectional SEM fractograph, it is easy to distinguish the scored flank region and the score residual section, and finally the internal coating of the can. The residual metal section is characterized by very elongated and deep intergranular fissures. The elongated fissures reflect the “pancake” grain structure that has resulted from extensive rolling deformation that the lid material has undergone. The walls of the fissures appear to be fairly smooth and featureless.
The fractograph shown in Figure 10 illustrates the appearance of the mixed mode of SCC. This fracture surface contains some smaller intergranular cracks, but has more pronounced features that appear as “fans” or delta shapes that originate from one initiation site and propagate in a transgranular mode. These are typical cleavage facets and suggest rapid crack propagation from an initiation site. The crack advances rapidly through an embrittled region and then blunts out when the crack outruns the embrittling species. The appearance of the cleavage facet is generally smooth near the initiation point, and then becomes more fan-like as the propagating crack jumps to different crystallographic planes as it moves forward, hence the fan appearance. In certain cases, there are some intergranular cracks observed at the periphery of the large cleavage facet in lower left corner. This latter observation suggests that intergranular cracking took place fairly rapidly after the cleavage event occurred, suggesting that under these conditions the lid material is prone to intergranular dissolution.
The intergranular corrosion of aluminum alloys containing alloying additions of zinc, magnesium, silicon, and copper is well known. These additions form intennetallic precipitates at the grain boundaries, such as MgZn^sub 2^ in the case of the 7XXX alloys, CuAl^sub 2^ in the 2XXX alloys, and Mg^sub 2^Al^sub 3^ in the 5XXX alloys, especially when the magnesium content is > ~3.0%. These precipitates tend to react galvanically with the bulk of the alloy and especially with the precipitate-free zone adjacent to the grain boundary. For example, Mg^sub 2^Al^sub 3^ is anodic to the aluminum matrix and will dissolve preferentially. This galvanic action led to the development of the electrochemical theory of stress corrosion of aluminum alloys, and explains the preferred intergranular attack of these alloys. Many of the successful SCC-resistant alloys and tempers have been developed based on this theory.
Additional Critical Factors
Several other factors are considered to play a role in exacerbating the SCC attack on the can lids.
The applied stresses are complex and the stress levels are becoming more severe due to the continued reduction of the gauge of aluminum, the new design of the lids, and the higher internal pressures of the beverage brought about by transport and packaging considerations. In addition, the scoring process in effect introduces a sharp notch to further concentrate stress, develops extensive dislocation fields, introduces some local residual stress, and can be expected to enhance dissolution or aid hydrogen transport. Stresses can cause a region of triaxiality ahead of a crack tip. Absorbed hydrogen can diffuse to this region or be transported there by slip systems, and can be concentrated there.
The local crack chemistry in propagating cracks in aluminum alloys has been found to be acidic with a pH value of ~3.5. This is because the local chemistry is controlled by the solubility product of the least soluble corrosion product and in the case of aluminum alloys this is aluminum hydroxide, Al(OH)^sub 3^. This acidic region means that an additional cathodic reaction can come into play during dissolution. In addition to the reduction of dissolved oxygen that commonly occurs in neutral pH conditions, the recombination of hydrogen ions can occur, 2H^sup +^ + 2e [arrow right] H^sub 2^. In fact, hydrogen recombination is the dominant cathodic corrosion reaction in acidic environments. Some of the nascent hydrogen formed immediately adjacent to the crack tip can be expected to be absorbed by the metal.
Studies of grain boundary regions in 7XXX alloys by Auger electron spectroscopy have demonstrated that there is extreme magnesium segregation at the grain boundaries. This is due not only to the MgZn^sub 2^ particles themselves but also to the presence of magnesium in solid solution in the atomic layers immediately adjacent to the boundaries. The magnesium content at the boundary is several times that of the bulk alloy.13 While this observation has been found on the 7XXX alloys, there is little reason to doubt that the same segregation would also occur in the 5XXX alloys and specifically in 5182 used for the can lids, since the magnesium content is relatively high. This segregated magnesium can be expected to be highly soluble in the acidic crack tip environment since magnesium hydroxide is not ihermodynamically stable in acid media.14
The segregated magnesium at the grain boundaries has been shown to diffuse rapidly along the boundaries and to adversely impact the protective qualities of overlaying aluminum oxide films. Specifically, alumina films containing magnesium impurities have been shown to be less protective, permit faster electrochemical dissolution, and to permit more rapid hydrogen recombination on their surfaces. There has been speculation on Mg-H interactions or the formation of a brittle magnesium hydride during SCC crack propagation, though there is little evidence to suggest actual hydride formation.
Proposed SCC Mechanisms
Given all of the fractographic observations and considering the describedfactors, two fracture mechanisms are considered to be operative in the lid material. Specifically, the deep intergranular fissures are considered to be indications of a dissolution mechanism. This dissolution is anticipated to be rapid in view of the intense magnesium segregation that is assumed. Further, the observation that intergranular attack overlaps the cleavage facet (Figure 10) indicates that the attack is rapid.
The transgranular cleavage facets are considered to indicate that a hydrogen embrittlement mechanism is also operative. Hydrogen absorbed during the corrosion process has been concentrated by me stress triaxiality at some initiation site and initiated a rapid cleavage crack, as indicated by the fanshaped feature. The crack then propagated rapidly until it outran the embrittling specie and then blunted out. This embrittlement may be due to a locally high concentration of hydrogen which weakens the metal-metal bonds (normally called decohesion) or to the formation of some brittle Mg- H complex or an actual metal hydride.
The change of crack path in the lid material is considered to occur when the local conditions of grain orientation, stress concentration, defect or other active site, and hydrogen content and other factors, all combine to change the cracking mechanism to the transgranular hydrogen mechanism over the ongoing dissolution process.
SUGGESTED SCC MINIMIZATION STRATEGIES
Guidelines to Fillers
One primary control of SCC losses in aluminum beverage cans is to educate the fillers in proper procedures. These procedures strive to keep the exterior of filled cans dry and well ventilated to prevent the formation of corrosion cells.
The process of filling beverage cans with soft drinks consists of, first, mixing the flavor syrup and other ingredients (Sugar or non-sugar sweetener, acidifiers, stabilizers, etc.) with treated and filtered water and carbonating the mixture with dissolved carbon dioxide. The refrigerated mixture, at a temperature of 4[degrees]C, is piped into the cans at the filler and the can is then closed by attaching the end at the double seam.
The filled and closed cans must be warmed to room temperature and preferably higher to aid drying and prevent sweating after packaging. Such warming is accomplished by the rinser/warmer unit. This unit also serves to wash away any beverage that may have spilled on the can during filling.
Excess water is removed from the can ends and the exteriors of the cans are dried by using air blowers or compressed-air wiping knives. The cans are then automatically inspected for fill volume and date coded, usually by ink jet marking on the can dome. Packaging for the filled cans has a wide variety of designs, from Hi- Cones (into six packs) to fiberboard cases. The packaged cans are stacked into pallets for warehousing and distribution to consumers.
Several professional trade groups, including the International Society of Beverage Technologists and the Beverage Can Makers of Europe, have issued guidelines on preventing SCC failures in the filling, warehousing, and distribution systems for beverage cans. These guidelines include: * Can rinsing: use rinse water low in sulfate and chloride concentration
* Can warming: warm cans above the dew point to prevent “can sweating”
* Can and end drying: the objective here is to ensure zero residual moisture on the can prior to secondary packaging. This applies especially to ends and the score area. Ensure the can and end is dry -this is the critical element of the beverage-filling operation.
* Secondary packaging: perforated shrink wrap helps moisture evaporate; vent holes in the ends of each tray pack help moisture evaporate; and fiberboard packaging will help absorb retained moisture.
* Storage: pallets of filled product should be stored to allow good air circulation and ease of inspection. Inspect on a regular basis. Warehouse temperature variations should be avoided.
Given the critical role of the scoring operation in SCC initiation, it is logical to explore whether any process modification, or scoring redesign, might be adopted to reduce SCC susceptibility. For instance, concerning the transgranular SCC, hydrogen is concentrated by a region of stress triaxiality ahead of a propagating crack tip. Accordingly, a score design or tool might be conceived to modify the region of triaxiality to focus the hydrogen more into a direction parallel to the elongated grain structure. This may reduce the susceptibility to transgranular SCC.
FUTURE METALLURGICAL RESEARCH
Two additional sets of experiments are suggested to further validate the above mechanisms.
First, it would be worthwhile to establish, by Auger electron spectroscopy studies, that the intense magnesium segregation that has been demonstrated in the case of 7xxx alloys is also valid for the 5xxx alloys. A parallel study to that reported in Reference 10 would be recommended. Observations of strong magnesium segregation would add considerable support to the above mechanisms.
The second set of experiments would be to repeat the SCC sensitivity-compressive strain evaluation shown in Figure 7, keeping all the conditions the same but with the addition of a hydrogen recombination inhibitor. The presence of such a recombination inhibitor would reduce the kinetics of the dissolution process, since the corrosion process is cathodically controlled and the recombination of hydrogen ions is the preferred cathodic reaction. However, the inhibitor would facilitate the absorption of hydrogen by the metal, increase its concentration in the metal and facilitate embrittlement. Thus, referring to Figure 7, the intergranular SCC would be slowed and time to failure increased, whereas the transgranular SCC would be enhanced and the time to failure decreased. Accordingly, the data curve would be skewed across the region of mixed cracking. A convenient hydrogen recombination inhibitor would be arsenic added in the form of sodium arsenite, NaAsO^sub 3^, at the 10 ppm level.5 An observation of a skewed curve from the strain versus time to failure curve would add considerable credibility to the above mechanistic discussions.
In this regard, it is interesting to note that T.D. Burleigh and coworkers8 found that the presence of oil contamination in the rinse water caused a pickup of hydrogen (by ~30%) in their exposure tests on 5182 alloy. oil can be considered to be a hydrogen recombination inhibitor and to increase the hydrogen content of the metal, which, in turn, would increase susceptibility to the transgranular cleavage failure. This observation is entirely consistent with the SCC mechanism proposed here.
The authors acknowledge the support of the Ball technical staff in conducting this study and the Ball Corporation for permission to publish it.
How would you…
…describe the overall significance of this paper?
The transition in stress-corrosion-cracking (SCC) failures from a transgranular to intergranular mode at higher stresses has been observed in recent design changes in the easy-open ends on aluminum beverage cans. This is proposed to be a shift from hydrogen embrittlement to anodic dissolution activated by stresses.
…describe this work to a materials science and engineering professional with no experience in your technical specialty?
Hydrogen embrittlement and anodic dissolution are two competing mechanisms used to explain SCC failures. The shift from transgranular to intergranular mode is a somewhat unique occurrence. This article will give additional insight to reconciling these opposing mechanisms.
…describe this work to a layperson?
The all-aluminum beverage can is a consumer product encountered daily. The easy-open ends are highly engineered, balancing ease of opening with various strength requirements. This article discusses aspects of a significant corrosion failure found in this everyday object.
EVOLUTION OF CAN LID DESIGN
Stress-corrosion cracking failures of the aluminum easy-open ends is an infrequent and undesirable failure experienced with filled beverage containers. Although SCC is I thought to be responsible for a relatively small percentage of aluminum beverage can failures each year, it is a concern to can manufacturers, fillers, and distributors.
In the United States, on average some 375 aluminum cans are used annually per person.3 The two-piece can is fabricated from two alloys, 3004 (Al-Mn alloy) for the walls and 5182-H38 (Al-Mg alloy) for the end lid. Recent trends have continued to downgauge the cylinder wall thickness, reduce the diameter, and modify the design of the lid. Also, the lid end design has been mechanically strengthened and the yield strength of the end alloy has been increased to 380 MPa. Finally, the level of carbonation and the internal pressure in the can has been increased, This increase of can pressure facilitates stacking and transportation of filled cans. All these changes have resulted in the use of less aluminum and a lower-cost package, but have also led to the characterization that the beverage can is an engineered “pressure vessel.”4
For the benefit of the consumer, the lid design has been modified to retain the tab following opening of the can (the ecology end), the opening aperture itself has been enlarged to enhance pourability, and the Hd has an incised score panel to facilitate can opening. This scoring cuts significantly into the thickness of the lid to ensure that the can opens readily. Unfortunately, under adverse circumstances, it may become an initiation j site for a stress-corrosion crack.
The evolution of can end designs over the past 50 years is shown in Figure A. The two general trends have been a reduction of approximately 25% in end diameter and 30% in metal gauge used for the ends. These two trends decrease the cost of aluminum in the ends, while simultaneously maintaining buckle resistance to the internal pressures within the can.
1. E.S. Sperry, Brass Wbrfd, 2 (1906), p. 39.
2. R.W. Staehte, editor, Fundamental Aspects of Stress Corrosion Cracking (Houston, TX: Nat. Assoc. of Corrosion Engineers, 1969).
3. Stress Corrosion-Technical Bulletin (Berlin, Germany. Beverage Can Maters of Europe, 2004).
4. Russell H. Jones, editor, Stress Corrosion Cracking (Materials Park, OH: ASM International, 1992).
5. E.N. Pugh, J.V. Craig, and AJ. Sedriks, The Stress Corrosion Cracking of Copper, Silver and Gold Alloys,” Fundamental Aspects of Stress Corrosion Cracking, ed. R.W. Staehle (Houston.TX; Nat. Assoc. of Corrosion Engineers, 1969), pp. 118-149.
6. M.O. Spiedel, “Hydrogen Embrittlement of Aluminum Alloys,’ Hydrogen in Metals, ed. L.M. Bernstein and A.W, Thompson (Metals Park, OH: ASM, 1974), p. 249.
7. J.R. Pickens, J.R. Gordon, and JAS. Green, The Effect of Loading Mode on the Stress Corrosion Cracking of Aluminum Alloy 5083,” Met Trans. A, 14A (May 1983), pp. 925-930.
8. Metal Improvement Company (Roseland, NJ: Curtis Wright Corp., 2004), www.melaMprovemenl.com.
9. Stress Corrosion Follow-Up, Stolle E-News (Centennial, CO: Stolle Machinery Co., 2004), www. stollemachinery.com.
10. L. Kudamatsu et al., “Development of Steel Sheet for Easy Open Ends” (Paper presented at the Fourth International Tinplate Conference, London, Oct 10-14 1988).
11. ID. Burieigh, E.H. Gillespie, and S.C. Biondich, “Blowout of Aluminum Alloy 5182 Can Ends Caused by Transgranular Stress Corrosion Cracking,” Aluminum Alloys tor Packaging, ed. J.G. Morris et al. (Warrendate, PA: TMS, 1993), pp. 323-332.
12. J.R. Davis, editor, Corrosion of Aluminum and Aluminum Alloys (Materials Park, OH: ASM International, 1999).
13. J.M. Chen et al., “Grain Boundary Segregation of an AI-Zn-Mg Ternary Alloy,” Met. Trans. A, 8A (December 1977), p. 1935.
14. M. Pourbaix et al. Atlas of Electrochemical Equilibria in Aqueous Solutions (Oxford, U.K.: Pergamon Press and Brussels, Belgium: CEBELCOR, 1966).
S.K. DeWeese Is with Ball Corporation, Broomfield, CO; S.T. Nlngiteri is project manager end S.K. Das is president and CEO with Secat Inc., Lexington, KY; JAS. Green is a consultant with JASG Consulting in Ellicott City, MD. Mr. DeWeese can be reached at firstname.lastname@example.org.
Copyright Minerals, Metals & Materials Society May 2008
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