Volume 13 - Corrosion Part 3 pot

Fig. 72 Variation in ductility of polycrystalline cadmium as a function of indium content of mercury- indium surface coatings. Specimens were tested at 25 °C (75 °F) in air and in mercury-indium solution. The concept of inert carrier LME is useful in failure analysis where LME has occured in a liquid solution for example, solder rather than a pure liquid metal. The susceptibility of various species in the solder in causing LME can be investigated by incorporating each species in an inert liquid. This may suggest a means of preventing embrittlement by replacing the most potent embrittler with a nonembrittling species. Fatigue in Liquid-Metal Environments Most studies of LME have been concerned with the effects of tensile loading on fracture. Investigations of fatigue behavior are important because this test condition is more severe than other test conditions, including the tensile testing of metals. Thus, a solid tested in fatigue may become embrittled, but the same solid tested in tension may not exhibit embrittlement. This may be because at the high stress level corresponding to yield stress, which is a prerequisite for the occurrence of embrittlement in tough metals, the solid may be sufficiently ductile to prevent initiation of a crack in the liquid. For example, smooth specimens of high-purity chromium-molybdenum low-alloy steel (yield stress: 690 MPa, or 100 ksi) are not embrittled by liquid lead, but the same steel specimens containing a fatigue precrack are severely embrittled by liquid lead. The fatigue life in liquid lead is reduced to 25% of that in an inert argon environment. Crack propagation in a liquid-metal environment under fatigue and tensile loading can be significantly different. The stress intensity at failure of 7.7 MPa (7 ksi ) for a 4340 steel specimen containing a fatigue precrack tested in cyclic fatigue in liquid lead was five times lower than that for the same specimen tested in tension in static fatigue and was twenty times lower than that in an inert argon environment. Furthermore, the 7.7-MPa (7-ksi ) stress intensity was the same whether the specimen had a machined notch (0.13-mm, or 0.005-in., root radius) or had a fatigue precrack at the root of the notch; that is, embrittlement was independent of root radius. These results indicate that embrittlement is very severe and even blunt cracks can propagate to failure. Similar results have also been reported for the same steel tested in liquid mercury. These results clearly indicate that fatigue testing of a notched or a fatigue precracked specimen provides the most severe test condition and therefore causes maximum susceptibility to embrittlement in a given liquid-metal environment. To determine whether a solid is susceptible to LME in a particular liquid, it is advisable to test the solid metal with a stress raiser or a notch in a tension-tension fatigue test. Decrease or Elimination of LME Susceptibility The reduction in the cohesion mechanism of embrittlement indicates that the electronic interactions, resulting in possible covalent bonding due to electron redistribution between the solid- and the liquid-metal atoms, reduces cohesion and thus induces embrittlement. Such interactions at the electronic level are the inherent properties of the interacting atoms and are therefore difficult or impossible to change in order to reduce the susceptibility to LME. One possibility is to introduce impurity atoms in the grain boundary that have more affinity for sharing electrons with the liquid-metal atoms than for sharing electrons with the solid-metal atoms. For example, additions of phosphorus to Monel segregate to the grain boundaries and reduce the embrittlement of Monel by liquid mercury. Additions of lanthanides to internally leaded steels reduce lead embrittlement of steel. However, in general, the best alternative is to electroplate or clad the solid surface with a metal as a barrier between the embrittling solid-liquid metal couple, making sure that the barrier metal is not embrittled by the liquid metal. Another possibility is that a ceramic or a covalent material coating on the solid surface will inhibit embrittlement. Apparently, only materials with metallic bonding are susceptible to LME. The severity of embrittlement could be reduced by decreasing the yield stress of the solid below the stress required to initiate a crack or by cladding with a high-purity metal of the alloy that is embrittled but that has a very low yield stress. Thus, Zircaloy, which is clad with a high-purity zirconium, becomes immune to embrittlement by cadmium. The obvious possibility is to replace embrittling liquid metal or solutions by nonembrittling metals or solutions. Embrittlement of Nonferrous Metals and Alloys * Zinc is embrittled by mercury, indium, gallium, and Pb-20Sn solder. Mercury decreases the fracture stress of pure zinc by 50% and dilute zinc alloys (0.2 at.% Cu or Ag) by five times that in an inert environment. The fracture propagation energy for a crack in zinc single crystals in mercury is 60% and in gallium is 40% of that in air. Zinc is embrittled more severely by gallium than by indium or mercury. Aluminum. Mercury embrittles both pure and alloyed aluminum. The tensile stress is decreased by some 20%. Fatigue life of 7075 aluminum alloy is reduced in mercury, and brittle-to-ductile transition occurs at 200 °C (390 °F). Additions of gallium and cadmium to mercury increase the embrittlement of aluminum. Delayed failure by LME occurs in mercury. Dewetting of aluminum by mercury has been found to inhibit embrittlement. The possible cause of dewetting is the dissolution of aluminum by mercury and oxidation of fine aluminum particles by air and formation of aluminum oxide white flowers at the aluminum/mercury interface. Aluminum alloys are embrittled by tin-zinc and lead-tin alloys. The embrittlement susceptibility is related to heat treatment and the strength level of the alloy. Gallium in contact with aluminum severely disintegrates unstressed aluminum alloys into individual grains. Therefore, grain-boundary penetration of gallium is sometimes used to separate grains and to study topographical features and orientations of grains in aluminum. There is some uncertainty about whether zinc embrittles aluminum. However, indium severely embrittles aluminum. Alkali metals, sodium, and lithium are known to embrittle aluminum. Aluminum alloys containing either lead, cadmium, or bismuth inclusions embrittle aluminum when impact tested near the melting point of these inclusions. The severity of embrittlement increases from lead to cadmium to bismuth. Copper. Mercury embrittles copper, and the severity of embrittlement increases when copper is alloyed with aluminum and zinc. Antimony, cadmium, lead, and thallium are also reported to embrittle copper. The apparent absence of embrittlement of copper, if observed, should be attributed to the test conditions and metallurgical factors discussed in this article. Mercury embrittlement of brass is a classic example of LME. Embrittlement occurs in both tension and fatigue and varies with grain size and strain rate. Mercury embrittlement of brass and copper alloys has been extensively investigated. Lead and tin inclusions in brass cause severe embrittlement when tested near the melting point of these inclusions. Lithium reduces the rupture stress and elongation at failure of solid copper. Sodium is reported to embrittle copper; however, cesium does not. The embrittling effects of bismuth for copper and their alloys are well documented. Gallium embrittles copper at temperatures ranging from 25 to 240 °C (75 to 465 °F). Gallium also embrittles single crystals of copper. Indium embrittles copper at 156 to 250 °C (313 to 480 °F). Other Nonferrous Materials. Tantalum and titanium alloys are embrittled by mercury and by Hg-3Zn solution. Refractory metals and alloys, specifically W-25Re, molybdenum, and Ta-10W, are susceptible to LME when in contact with molten Pu-1Ga. Cadmium and lead are not embrittled by mercury. However, indium dissolved in mercury embrittles both polycrystalline and single crystal cadium (Fig. 60). The embrittlement of titanium and its alloys by both liquid and solid cadmium is well recognized by the aircraft industry. Cadmium-plated fasteners of both titanium and steel are known to fail prematurely below the melting point of cadmium. Cadmium-plated steel bolts have good stress-corrosion resistance, but cadmium is known to crack steel. Therefore, such bolts are not recommended for use. Zinc is reported to embrittle magnesium and titanium alloys. Silver, gold, and their alloys are embrittled by both mercury and gallium. Nickel is severely embrittled by cadmium dissolved in cesium. The fracture mode is brittle intergranular with bright grain-boundary fracture. Failure of Zircaloy tubes used as cladding material for nuclear fuel rods has been suspected to result from nuclear- interaction reaction products, such as iodine and cadmium carried by liquid cesium, which is used as a coolant in the reactor. Systematic investigation in the laboratory has shown that cadmium, both in the solid and the liquid state or as a carrier species dissolved in liquid cesium, causes severe liquid and solid metal induced embrittlement of zirconium and Zircaloy-2. Embrittlement of Zircaloy by calcium, strontium, zinc, cadmium, and iodine has also been reported. Note cited in this section * *Adapted from M.G. Nicholas, A Survey of Literature on Liquid Metal Embrittlement of Metals and Alloys, in Embrittlement by Liquid and Solid Metals, M.H. Kamdar, Ed., The Metallurgical Society, 1984, p 27-50 Embrittlement of Ferrous Metals and Alloys Embrittlement by Aluminum. Tensile and stress rupture tests have been conducted on steels in molten aluminum at 690 °C (1275 °F). For short-term tensile tests, a reduced breaking stress and reduction of area were found as compared to the values in air. In the stress rupture tests, the time to failure was dependent upon the applied stress. Embrittlement by Antimony. AISI 4340 steel tested in fatigue in liquid Pb-35Sb at 540 °C (1000 °F) and in antimony at 675 °C (1250 °F) was very severely embrittled. The embrittlement in lead-antimony occurred 165 to 220 °C (300 to 400 °F) higher than that observed for high-purity lead. Small additions of antimony ( 5 wt%) to lead had no effect on embrittlement when tested in fatigue, although 0.002 to 0.2% Sb additions have caused a significant increase in the embrittlement of smooth AISI 4140 steel specimens tested in tension. Embrittlement by antimony increases with temperature and is thought to occur by the grain-boundary diffusion of antimony in steel. Embrittlement by Bismuth. Upon testing in liquid bismuth at 300 °C (570 °F), no embrittlement was noted in bend tests on a quenched-and-tempered steel. The stress rupture data obtained on the low-carbon steel showed that the time to failure and reduction of area increased with decreasing load, but no intercrystalline attack was noted. Embrittlement by Cadmium. In several studies of the embrittlement of low-alloy AISI 4340 steel, embrittlement occurred at 260 to 322 °C (500 to 612 °F) but not at 204 °C (399 °F) for high-strength steel. Cracks were observed in samples loaded to 90% of their yield stress at 204 °C (399 °F). The threshold stress required for cracking decreases with an increase in temperature. Cracking at 204 °C (399 °F) was strongly dependent on the strength level of the steel, and embrittlement was not observed for strength levels less than 1241 MPa (180 ksi). Delayed failure occurs in cadmium- plated high-strength steels (AISI 4340, 4140, 4130, and an 18% Ni maraging steel). Failures were observed at 232 °C (450 °F), which is about 90 °C (160 °F) below the melting point of cadmium. Static failure limits of 10% and 60% of the room-temperature notch strength have been reported for electroplated and vacuum-deposited cadmium, respectively, at 300 °C (570 °F). A discontinuous crack propagation mode was observed, consisting of a series of crack propagation steps separated by periods of no apparent growth. This slow crack growth region was characterized by cracks along the prior- austenite grain boundaries. Once the cracks reached a critical size, a catastrophic failure occurred that was characterized by a transgranular ductile fracture. Embrittlement of high-strength steels occurs when they are stress rupture tested in liquid and solid cadmium. Cadmium produces a progressive decrease in the reduction of area at fracture of AISI 4140 steel over the temperature range of 170 to 321 °C (338 to 610 °F). Cadmium was identified as a more potent solid-metal embrittler than lead, tin, zinc, or indium. Embrittlement by Copper. The embrittlement of low-carbon steel by copper plate occurred at 900 °C (1650 °F) during a slow-bend test. The embrittling effects of the copper plate exceeded those encountered with brazing alloys. Similar observations have been made for plain carbon steels, silicon steels, and chromium steels at 1000 to 1200 °C (1800 to 2190 °F). The surface cracking produced during the hot working of some steels at 1100 to 1300 °C (2010 to 2370 °F) also has several characteristics of LME. It is promoted by surface enrichment of copper and other elements during oxidation and subsequent penetration along the prior-austenite grain boundaries. Elements such as nickel, molybdenum, tin, and arsenic that affect the melting point of copper or its solubility in austenite also influence embrittlement. A ductility trough has also been noted, with no cracking produced at temperatures above 1200 °C (2190 °F). Steel plated with copper and pulse heated to the melting point of copper in milliseconds was embrittled by both liquid and solid copper. Hot tensile testing in a Gleeble testing machine at high strain rates produced severe cracking in AISI 4340 steel by both the liquid and solid copper. Embrittlement by Gallium. The alloy Fe-3Si is severely embrittled by gallium, as are the solid solutions of iron and AISI 4340 steel. Embrittlement by Indium. Indium embrittles pure iron and carbon steels. Embrittlement depends on both the strength level and the microstructure. Pure iron was embrittled only at temperatures above 310 °C (590 °F), appreciably above the melting point of indium. However, other steels, such as AISI 4140 (ultimate tensile strength of 1379 MPa, or 200 ksi), were embrittled by both solid and liquid indium. Surface cracks were detected at temperatures below the melting temperature of indium. This was interpreted as a local manifestation of the underlying embrittlement mechanism, and it was assumed that the cracks must reach a critical size before the gross mechanical properties are affected. Embrittlement by Lead. The influence of lead on the embrittlement of steel has been extensively investigated and has been found to be sensitive to both composition and metallurgical effects. The studies fall into two major classifications: • LME due to contact with an external source of liquid lead • Internal LME in which the lead is present internally as inclusions or a minor second phase, as in leaded steels Both external and internal lead-induced embrittlement exhibit similar characteristics, but for the purpose of simplicity, each will be treated individually. LME Due to External Lead. AISI 4145 and 4140 steels exposed to pure lead exhibit classical LME, as shown by substantial decreases in both the reduction of area and the elongation at fracture. The fracture stress and the reduction of area decreased at temperatures considerably below the melting point of lead and varied continuously through the melting point, suggesting that the same embrittlement mechanism is operative for both solid- and liquid-metal environments. Additions of zinc, antimony, tin, bismuth, and copper increase the embrittling potency of lead. Additions of up to 9% Sn, 2% Sb, or 0.5% Zn to lead increased the embrittlement of AISI 4145 steel. In some cases, the embrittlement and failure occurred before the UTS was reached. The extent of embrittlement increases with increasing impurity content. No correlation was observed between the degree of embrittlement and wettability. The lead-tin alloys readily wetted steels, but the more embrittling lead-antimony alloys did not. LME Due to Internal Lead. Leaded steels are economically attractive because lead increases the machining speed and the lifetime of the cutting tools. The first systematic investigation of embrittlement of leaded steels was reported in 1968 (Mostovoy and Breyer). Embrittlement characteristics similar to those promoted by external lead were observed. The degradation in the ductility began at approximately 120 °C (215 °F) below the melting point of lead, with the embrittlement trough present from 230 to 454 °C (446 to 849 °F). This was followed by a reversion to ductile behavior at about 480 °C (895 °F). The severity of the embrittlement and the brittle-to-ductile transition temperature T R have been shown to be dependent upon the strength level of the steel, with the degree of embrittlement and T R increasing with strength level. In these studies, intergranular fracture was produced, which was propagation controlled at low temperatures and nucleation controlled at high temperatures. The degree of embrittlement was critically dependent on the lead composition, and the influence of trace impurities completely masked any variations due to different carbon and alloy compositions of the steel. Lead embrittlement of a steel compressor disk was induced by bulk lead contents of 0.14 and 6.22 wt%. The lead is associated with the nonmetallic inclusions, and upon yielding, microcracks form at the weak inclusion/matrix interface, releasing a source of embrittling agent to the crack tip that aids subsequent propagation. An electron microprobe analysis of the nonmetallic inclusions identified the presence of zinc, antimony, tin, bismuth, and arsenic. With the exception of arsenic, all these trace impurities have been shown to have a significant effect on the external LME of steel. The two most promising methods of suppressing LME are the control of sulfide composition and morphology and the cold working of the steel. The addition of rare-earth elements to the steel melt modifies the sulfide morphology and composition and can eliminate LME. Embrittlement by Lithium. Exposure of AISI 4340 steel to lithium at 200 °C (390 °F) resulted in static fatigue, with the time to failure depending on the applied stress. A decreasing fracture stress and elongation to fracture were noted with increasing UTS of variously treated steels, and catastrophic failure occurred for those steels with tensile strengths exceeding 1034 MPa (150 ksi). The tensile ductility of low-carbon steel at 200 °C (390 °F) was drastically reduced in lithium, with intergranular failure after 2 to 3% elongation, but there was no effect on the yield stress or the initial work- hardening behavior. The fracture stress was shown to be a linear function of d , where d is the average grain diameter, in accordance with the Petch relationship. Embrittlement by Mercury. It has been shown that mercury embrittlement is crack nucleation controlled and can be induced in low-carbon steel samples by the introduction of local stress raisers. The fracture toughness of a notched 1Cr- 0.2 Mo steel was significantly decreased upon testing in mercury. The effective surface energy required to propagate the crack was 12 to 16 times greater in air than in mercury. The fatigue life of 4340 steel in mercury is reduced by three orders of magnitude as compared to that in air. The addition of solutes (cobalt, silicon, aluminum, and nickel) to iron, which reduced the propensity for cross-slip by decreasing the number of active slip systems and changed the slip mechanism from wavy to planar glide, increased the susceptibility to embrittlement. Iron alloys containing 2% Si, 4% Al, or 8% Ni and iron containing 20% V or iron containing 49% Co and 2% V have been shown to be embrittled by mercury in unnotched tensile tests. The degree of embrittlement behavior was extended to lower alloy contents by using notched samples. No difference in the embrittlement potency of mercury or a saturated solution of indium in mercury was noted. Effect of Selenium. Selenium had no embrittling effect on the mechanical properties of a quenched-and-tempered steel (UTS: 1460 MPa, or 212 ksi) bend tested at 250 °C (480 °F). Embrittlement by Silver. Silver had no significant effect on the mechanical properties of a range of plain carbon steels, silicon steels, and chromium steels tested by bending at 1000 to 1200 °C (1830 to 2190 °F). However, a silver-base filler metal containing 45% Ag, 25% Cd, and 15% Sn has been reported to embrittle A-286 heat-resistant steel in static- load tests above and below 580 °C (1076 °F), the melting point of the alloy. Effect of Sodium. Unnotched tensile properties of low-carbon steel remained the same when tested in air and in sodium at 150 and 250 °C (300 and 480 °F). Similarly, Armco iron, low-carbon steel, and type 316 steel were not embrittled by sodium at 150 to 1600 °C (300 to 2910 °F). Embrittlement by Solders and Bearing Metals. A wide range of steels are susceptible to embrittlement and intercrystalline penetration by molten solders and bearing metals at temperatures under 450 °C (840 °F). Tensile tests have revealed embrittlement as a reduction in ductility. The embrittlement increased with grain size and strength level of the steel, except for temper-embrittled steels. The tensile strength and ductility of carbon steel containing 0.13% C were decreased upon exposure to the molten solders and bearing alloys. The embrittlement was concomitant with a change to a brittle intergranular fracture mode and penetration along prior-austenite grain boundaries. Ductile failure was observed with samples tested in air or in the liquid metal at temperatures exceeding 450 °C (840 °F). No intercrystalline penetration of solder was noted in carbon steels containing 0.77% and 0.14% C at 950 °C (1740 °F). It has been reported that solder embrittles steel more than Woods metal (a bismuth-base fusible alloy containing lead and tin), particularly if it contains 4% Zn. The bearing metals produced embrittlement similar to the solder containing 4% Zn alloy. Embrittlement by Tellurium. Tellurium-associated embrittlement has been reported for carbon and alloy steels. Hot shortness occurs in AISI 12L14 + Te steel, with the most pronounced loss in ductility between 810 and 1150 °C (1490 and 2100 °F), embrittlement being most severe at 980 °C (1795 °F). The embrittlement has been shown to occur by the formation of a lead-telluride film at the grain boundary, which melts at 923 °C (1693 °F). The mechanical test data and the examination of fracture surfaces by Auger electron spectroscopy (AES) and scanning electron microscopy (SEM) indicated LME of steel by the lead-telluride compound. Effect of Thallium. Thallium had no embrittling effects on the mechanical properties of a quenched-and-tempered steel (UTS: 1460 MPa, or 212 ksi) tested in bending at 325 °C (615 °F). Embrittlement by Tin. The embrittling effect of tin has been observed in a range austenitic and nickel-chromium steels, the degree of embrittlement increasing with their strength level. Embrittlement depends on the presence of a tensile stress and is associated with intercrystalline penetration. The embrittlement by solid tin occurs at approximately 120 °C (215 °F) below its melting point. The fracture surfaces exhibited an initial brittle zone perpendicular to the tensile axis that followed the prior-austenite grain boundaries. Layers of intermetallic compound present at the steel/tin interface did not impede the embrittlement process. Embrittlement has been observed in delayed-failure tests down to 218 °C (424 °F) (14 °C, or 25 °F, below the melting point of tin); however, in tensile tests, embrittlement by solid tin was effective at temperatures as low as 132 °C (270 °F), which is 100 °C (180 °F) below melting point. AISI 3340 steel doped with 500 ppm of phosphorus, arsenic, and tin has been tested in the presence of tin while in the segregated (temper embrittled) and the unsegregated states. Temper-embrittled steels were found to be more susceptible to embrittlement than steel heat treated to a nontempered state. A lower fatigue limit and lifetime at stresses below the fatigue limit for low-carbon steel and for 18-8 stainless steel have been noted when tested in tin at 300 °C (570 °F). The exposure time to the tin before testing had no influence on the fatigue life of the steel. Embrittlement of Austenitic Steels by Zinc. Two main types of interaction of zinc and austenitic stainless steel have been observed. Type I relates to the effects on unstressed material in which liquid-metal penetration/erosion is the major controlling factor, and Type II relates to stressed materials in which classic LME is observed. Type I Embrittlement. Zinc slowly erodes unstressed 18-8 austenitic stainless steel at 419 to 570 °C (786 to 1058 °F) and penetrates the steel, with the formation of an intermetallic nickel-zinc compound at 570 to 750 °C (1060 to 1380 °F). At higher temperatures, penetration along the grain boundaries occurs, with a subsequent diffusion of nickel into the zinc-rich zone. This results in a nickel-exposed zone adjacent to the grain boundaries, reducing the stability of the phase and causing it to transform to an -ferrite; the associated volume change of the transformation produces an internal stress that facilitates fracture along the grain boundaries. Similar behavior has been observed in an unstressed 316C stainless steel held 30 min at 750 °C (1380 °F), in which penetration occurred to a depth of 1 mm (0.4 in.), and in an unstressed type 321 steel held 2 h at 515 °C (960 °F), in which a penetration of 0.127 mm (0.005 in.) was observed. Type II embrittlement occurs in stainless steel above 750 °C (1380 °F) and is characterized by an extremely fast rate of crack propagation that is several orders of magnitude greater than that of Type I, with cracks propagating perpendicular to the applied stress. In laboratory tests, an incubation period was observed before the propagation of Type II cracks, suggesting that they may be nucleated by Type I cracks formed during the initial contact with zinc. At 800 °C (1470 °F), a stressed type 316C stainless steel failed catastrophically when coated with zinc. Cracking was produced at a stress of 57 MPa (8 ksi) at 830 °C (1525 °F) and 127 MPa (18 ksi) at 720 °C (1330 °F), but failure was not observed at a stress of 16 MPa (2 ksi) at 1050 °C (1920 °F). Liquid-metal embrittlement may be produced by the welding of austenitic steels in the presence of zinc or zinc-base paints. Intercrystalline cracking has been observed in the heat-affected zone in areas heated from 800 to 1150 °C (1470 to 2100 °F), and electron microprove analysis has been used to identify the grain-boundary enrichment of nickel and zinc, together with the formation of a low-melting nickel-zinc compound. The embrittlement of sheet samples of austenitic steel coated with zinc dust dye and zinchromate primer occurs at stresses of the order of 20 MPa (3 ksi). Embrittlement of Ferritic Steels by Zinc. Embrittlement of ferritic steels and Armco iron by molten zinc has been reported in the temperature range of 400 to 620 °C (750 to 1150 °F). Long exposures and intercrystalline attack were needed to cause a reduction in the elongation to fracture; an iron-zinc intermetallic layer was formed that inhibited embrittlement until the layer was ruptured. High-alloy ferritic steels exhibit embrittlement by zinc at temperatures above 750 °C (1380 °F). Delayed failure occurs in steel in contact with solid zinc at 400 °C (750 °F), which is 19 °C (34 °F) below the melting point of zinc. The slow crack growth region is characterized by intergranular mode of cracking. Exposing AISI 4140 steel to solid zinc results in a decrease in the reduction of area and in fracture stress at 265 °C (510 °F), with no significant changes in the other mechanical properties. The tensile fracture initially propagates intergranularly, with the final failure occurring by shear. Liquid zinc embrittles 4140 steel at 431 °C (808 °F), and it has been shown that zinc is present at the crack tip. Figure 73 shows a compilation of liquid-metal embrittling and nonembrittling couples based on the theoretical calculations of the solubility parameter and the reduction in the fracture- surface energy. The embrittlement curve is separated by brittle and ductile fracture. A concise summary of embrittlement couples is provided in Table 5. Both pure and alloyed solids are listed. Table 5 Summary of embrittlement couples P, element (nominally pure): A, alloy: C, commercial: L, laboratory Fig. 73 Calculated reduction in the fracture surface energy relating to solubility parameter for many solid- liquid embrittlement couples. Note that the curve separates embrittlement couples from nonembrittled solid- liquid metal couples. Solid Metal Induced Embrittlement M.H. Kamdar, Benet Weapons Laboratory, U.S. Army Armament Research, Development, and Engineering Center Embrittlement occurs below the melting temperature of the solid in certain LME couples. The severity of embrittlement increases with temperature, with a sharp and significant increase in severity at the melting point, T m , of the embrittler (Fig. 74). Above T m , embrittlement has all the characteristics of LME. The occurrence of embrittlement below the T m of the embrittling species is known as solid metal induced embrittlement of metals. Liquid Hg Cs GA Na In Li Sn Bi Tl Cd Pb Zn Te Sb Cu Solid P A P P A P P A P P A P A P P P A P P P P Sn P X . . . . . . X . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Bi P X . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Cd P . . . . . . X X X . . . . . . . . . . . . X . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . P X X . . . X X . . . X X . . . X . . . . . . . . . . . . . . . X . . . . . . . . . . . . . . . Zn LA . . . X . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Mg CA . . . . . . . . . . . . . . . X . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . X . . . . . . . . . P X X . . . X . . . . . . X . . . . . . . . . X X . . . . . . X X . . . . . . . . . . . . . . . Al CA X X . . . X . . . X X . . . . . . X . . . X X . . . . . . X . . . X . . . . . . . . . Ge P . . . . . . . . . X . . . . . . X . . . . . . X . . . X . . . X X X . . . . . . . . . X . . . P X X . . . X X . . . . . . . . . X . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Ag LA . . . . . . . . . . . . . . . . . . . . . . . . X . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . CP X X . . . . . . . . . X . . . . . . X . . . . . . X X . . . . . . X . . . . . . . . . . . . . . . Cu LA X . . . . . . X . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Liquid Hg Cs GA Na In Li Sn Bi Tl Cd Pb Zn Te Sb Cu Solid P A P P A P P A P P A P A P P P A P P P P CA X X . . . X . . . . . . X X X X (?) X . . . . . . . . . X . . . . . . . . . . . . . . . P X . . . . . . . . . . . . . . . . . . . . . X . . . . . . . . . . . . . . . . . . X . . . . . . . . . . . . . . . LA . . . . . . . . . . . . . . . . . . . . . . . . X . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Ni CA . . . . . . . . . . . . . . . . . . . . . . . . . . . X . . . . . . . . . . . . . . . X . . . . . . . . . . . . . . . P . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . X LA X X . . . X . . . . . . X X X . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Fe CA . . . . . . . . . . . . . . . . . . X . . . X X . . . . . . . . . . . . X X X X X X X P . . . . . . . . . . . . . . . . . . . . . . . . X . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . PD LA . . . . . . . . . . . . . . . . . . . . . . . . X . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Ti CA . . . X . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . X . . . . . . . . . . . . . . . . . . [...]... 1762 °F) 20 4-2 32 39 9-4 50 BE Au (10 53 °C, or 1927 °F) 20 4-2 32 39 9-4 50 BE Cu-Bi(b) Hg ( -3 9 °C, or -3 8 °F) -8 4 -1 19 ST Cu-Bi(c) Hg -8 7 -1 25 ST Cu-3Sn(c) Hg -4 8 -5 4 ST Cu-1Zn(c) Hg -4 6 -5 1 ST Tin bronze Pb (32 7 °C or 621 °F) 200 39 2 IM Zinc Hg -5 1 -6 0 ST Inconel ln (156 °C, or 31 3 °F) Room temperature RE Zircaloy-2 Cd 30 0 ST Ti-6Al-4V Ti-8Al-1Mo-1V 572 (a) DF, delayed-failure tensile test; BE, bend test;... calculated vapor-transport times at the embrittler-melting temperatures Embrittler Vapor pressure Time, s Pa torr Zn 20 1.5 × 1 0-1 5 × 1 0 -3 Cd 10 7.5 × 1 0-2 1 × 1 0-2 Hg(a) 0 .3 2.25 × 1 0 -3 3 × 1 0-1 Sb 4 × 1 0-4 3. 0 × 1 0-6 3 × 102 K 1 × 1 0-4 7.5 × 1 0-7 1 × 1 03 Na 2 × 1 0-5 1.5 × 1 0-7 5 × 1 03 Ti 4 × 1 0-6 3. 0 × 1 0-8 3 × 104 Pb 5 × 1 0-7 3. 75 × 1 0-9 2 × 105 Bi 2 × 1 0-8 1.5 × 1 0-1 0 5 × 106 Li 2 × 1 0-8 1.5 × 1 0-1 0 5 ×... delayed-failure tensile test (c) S, smooth specimen; N, notched specimen Table 7 Occurrence of SMIE in nonferrous alloys All test specimens were smooth type Base metal Embrittler (melting point) Onset of embrittlement Test type(a) °C °F Cd (32 1 °C, or 610 °F) 38 100 DF Cd 149 30 0 BE Cd 38 100 DF Cd 149 30 0 BE Ti-3Al-14V-11Cr Cd 149 30 0 BE Ti-6Al-6V-2Sn Cd 149 30 0 BE Ag (961 °C, or 1762 °F) 20 4-2 32 39 9-4 50... 93 199 DF S Sn 204 39 9 ST S ln 121 250 ST S Pb-4Sn (NA) 204 39 9 ST S Pb-Sn (NA) 204 39 9 ST S Pb-Sb (NA) 204 39 9 ST S Pb 288 550 ST S 4145 leaded Pb 204 39 9 ST S 434 0 Cd 260 500 DF N Cd 30 0 572 DF N Cd 38 100 DF S Zn 400 752 DF N 434 0M Cd 38 100 DF S 8620 Pb 288 550 ST S 8620 leaded Pb 204 39 9 ST S A-4 Pb 288 550 ST S A-4 leaded Pb 204 39 9 ST S D6ac Cd 149 30 0 DF N Courtesy of Dr A Druschitz (a) NA, data... solid-metal embrittlers are also known to cause LME Table 6 Occurrence of SMIE in steels Base metal Embrittler (melting point) Onset of embrittlement °C Test type(b) Specimen type(c) °F 1041 Pb (32 7 °C, or 621 °F) 288 550 ST S 1041 leaded Pb 204 39 9 ST S 1095 ln (156 °C, or 31 3 °F) 100 212 ST S 33 40 Sn ( 232 °C, or 450 °F) 204 39 9 ST N Pb 31 6 601 ST N 4 130 Cd (32 1 °C, or 610 °F) 30 0 572 DF N 4140 Cd 30 0... Pb 204 39 9 ST S Pb-Bi (NA)(a) ST S Pb-Zn (NA) Below solidus ST S Zn (419 °C, or 786 °F) 254 489 DF N Sn 218 424 DF N Cd 188 37 0 DF N Pb 160 32 0 DF N ln Room temperature DF N Pb-Sn-Bi (NA) Below solidus ST S ln 80 176 DF S Sn 204 39 9 ST S Sn-Bi (NA) Below solidus ST S Sn-Sb (NA) Below solidus ST S ln 110 230 DF S ln 93 199 DF S ln-Sn (118 °C, or 244 °F) 4145 Below solidus 93 199 DF S Sn 204 39 9 ST S... Materials, 19 73, p 65 K Sieradski and R.C Newman, Philos Mag A, Vol 5 (No 1), 1985, p 95 T Cassaigne, E.N Pugh, and J Kruger, National Bureau of Standards and Johns Hopkins University, unpublished research, 1987 C.M Chen, M.H Froning, and E.D Verink, in Stress Corrosion New Approaches, STP 610, American 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 Society for... 2 × 1 0-8 1.5 × 1 0-1 0 5 × 106 ln 3 × 1 0-1 9 2.25 × 1 0-2 1 3 × 1017 Sn 8 × 1 0-2 1 6.0 × 1 0-2 3 1 × 1019 Ga 6 × 1 0 -3 6 4.5 × 1 0 -3 8 2 × 1 035 (a) At room temperature Delayed Failure and Mechanism of SMIE If a metal in contact with the embrittling species is loaded to a stress that is lower than that for fracture and is tested at various temperatures, then either the environment-induced fracture initiates and... failure in SMIE and LME systems Base metal Liquid Solid Type A behavior: delayed failure observed 4140 steel Li Cd 434 0 steel Cd ln 4140 steel ln Cd 4140 steel Pb 4140 steel Sn 4140 steel ln 4140 steel Zn 2024 Al Hg 2424 Al Hg-3Zn 7075 Al Hg-3Zn 50 83 Al Hg-3Zn Al-4Cu Hg-3Zn Cu-2Be Hg Cu-2Be Hg Type B behavior: delayed failure not observed Zn Hg Cd Hg Cd Hg + ln Ag Hg + ln Al Hg Courtesy of... Vol 14A, 19 83, p 2 23 R.W Staehle et al., Corrosion, Vol 26 (No 11), 1970, p 451 H.W Pickering and P.R Swann, Corrosion, Vol 19 (No 3) , 19 63, p 37 3 A.W Thompson and I.M Bernstein, Advances in Corrosion Sciences and Technology, Vol 7, M.G Fontana and R.W Staehle, Ed., Plenum Press, 1980, p 53 R.M Riecke, A Athens, and I.O Smith, Mater Sci Technol., Vol 2, 1986, p 1066 E.H Dix, Trans AIME, Vol 137 (No 11), . 39 9-4 50 BE Ti-6Al-6V-2Sn Au (10 53 °C, or 1927 °F) 20 4-2 32 39 9-4 50 BE Cu-Bi (b) Hg ( -3 9 °C, or -3 8 °F) -8 4 -1 19 ST Cu-Bi (c) Hg -8 7 -1 25 ST Cu-3Sn (c) Hg -4 8 -5 4 ST Cu-1Zn (c) . (32 1 °C, or 610 °F) 38 100 DF Ti-6Al-4V Cd 149 30 0 BE Cd 38 100 DF Ti-8Al-1Mo-1V Cd 149 30 0 BE Ti-3Al-14V-11Cr Cd 149 30 0 BE Cd 149 30 0 BE Ag (961 °C, or 1762 °F) 20 4-2 32 39 9-4 50. 1041 Pb (32 7 °C, or 621 °F) 288 550 ST S 1041 leaded Pb 204 39 9 ST S 1095 ln (156 °C, or 31 3 °F) 100 212 ST S Sn ( 232 °C, or 450 °F) 204 39 9 ST N 33 40 Pb 31 6 601 ST N 4 130 Cd (32 1 °C,