Weldability and Solidification Phenomena of Cast Stainless Steel
Weldability and Solidification Phenomena of Cast Stainless Steel
The weldability of cast alloy CF-8M is found to be largely dependent on its initial solidification mode, with primary ferrite heats exhibiting greater weldability
BY M. J. CIESLAK AND W. F. SAVAGE
Introduction tion established during solidification vations has been generated and of weld metal is far better preserved. reported by other investiga 4 5 15 1 29 30 31 34 35 41 Stainless steel casting alloys have Therefore, a more accurate interpreta tors - - - "- - - - . - been employed successfully for many tion of the mechanics of solidification Several theories have been sug years in the severe environments can be obtained from examination of gested to explain weld metal hot encountered in the chemical, petro the microstructure and the pattern of cracking. Early work by Bochvar,6 leum, and nuclear industries. An entire microsegregation in weld metal at Pumphrey and Jennings,7 Pumphrey class of alloys has been developed for room temperature. and Lyons,8 Medovar,'1 and Toropov10 use in corrosive environments at work The redistribution of solute during resulted in the formulation of the ing temperatures below 1200°F solidification has a strong influence on Shrinkage-Brittleness Theory. This the (649°C). Unfortunately, in many cases both the weldability and the castabili ory postulates a semi-rigid, "coher where acceptable mechanical proper ty of many stainless steels. "Hot tears" ent," dendritic network which is sus ties and corrosion resistance have in castings and "hot cracks" in welds ceptible to hot cracking if shrinkage been achieved, both castability and appear to occur by similar mechanisms strains exceed some critical level. The weldability have been poor. Since as a direct result of the microsegrega theory also states that, since diffusion there is often a correlation between tion of certain solute elements. is limited in welding, microsegregation castability and weldability, a better is augmented and the resulting low melting constituents aggravate the hot understanding of the solidification Hot Cracking Studies mechanics of these materials should cracking problem. This theory also make the design of improved alloys Studies of the solidification of aus recognizes the possibility of "healing" or "backfilling" of cracks with low possible. tenitic stainless steels have been con melting point constituents, such as Solidification research based on ducted for many years. It has been 1 eutectic liquids, if they are present in commercial alloy systems is difficult. observed since the early 1940's that large enough quantities. This is both because of the complexity hot cracking in austenitic stainless of the alloys and because the high steel weldments is reduced or pre A second theory—the Strain Theory temperatures involved make direct vented when a small amount of delta of Hot Cracking—offered by Pellini11 observation of the solidification pro ferrite is present. Early studies by Bor and Apblett1- proposes that a contin 2 3 cess difficult. As a result, most infor land and Hull supported the fact that uous, thin liquid film present at some mation on the solidification mechan 5-10 vol-% of delta ferrite in weld stage of the solidification process can ics of metal systems has been based ments and castings afforded the best not support the strains induced in the upon room temperature metallograph resistance to hot cracking and hot weldment or casting without hot ic examination of the solidification tearing. Since that time, a voluminous cracking. substructures of castings and weld amount of data which is in general The Generalized Theory of Super- ments and upon mathematical theo agreement with these initial obser- Solidus Cracking as proposed by Bor ries. land13 is probably the most widely In castings, the cooling times are accepted theory of hot cracking and normally long enough for diffusion- hot tearing today. Borland postulates Paper presented at the AWS 61st Annual that a semi-continuous liquid film controlled processes to modify both Meeting held in Los Angeles, California, exists in the weldment or casting sepa the microstructure and pattern of during April 14-18, 1980. microsegregation during cooling from rated by bridges of solidified material. M. j. CIESLAK is the Steel Founders' Society the solidus to room temperature. Con The solid bridges, when subjected to a of America Graduate Fellow, and W. F. critical level of induced strain, will versely, the cooling times for most SAVAGE is Professor of Metallurgical Engi weldments are so short as to suppress fracture, forming hot cracks. Borland neering and Director of Welding Research, uses a composition-dependent surface most diffusion-controlled reactions in School of Engineering, Materials Division, energy argument in conjunction with the solid state. Thus, the microstruc Rensselaer Polytechnic Institute, Troy, New ture and the pattern of microsegrega York. rapid solidification rates to describe
136-sl MAY 1980 Figure 1 consists of a projection of the solidus and liquidus curves on a plane of the Fe-Cr-Ni ternary dia gram.22 According to Fig. 1, composi tions on the nickel-rich side of the dashed liquidus curve should solidify as primary austenite while composi tions on the chromium-rich side should solidify as primary delta ferrite. Masumoto et al-3 reported that welds solidifying on the nickel-rich side of the liquidus were extremely suscepti ble to hot cracking whereas welds solidifying on the chromium side did not hot crack. The DeLong24 and Schaeffler' dia grams attempt to predict room tem perature ferrite contents from the compositions of the alloys. Predicting weldability based on these diagrams has met with mixed success. Masumo to21 found no correlation between weldability and ferrite content based on these diagrams. Recent work by Lippold' has shown that the form and volume of room temperature ferrite is a function of composition, the shape Fig. 1— Solidus and liquidus curves oi the Fe-Cr-Ni system of the Fe-Ci-Ni phase diagram, and the cooling rate experienced by the mate rial. the form and composition of the last- jected to the liquid during primary to-solidify material. delta ferrite solidification. Elements Materials and Procedures In addition to theories postulating a such as silicon (Si), molybdenum liquid phase as a prerequisite for hot (Mo), sulfur (S), and phosphorus (P), Materials for this investigation were cracking, a solid-state theory of hot which are more soluble in delta ferrite, supplied by member foundries of the cracking has been proposed by Mov- are rejected to the liquid during pri Steel Founder's Society of America. chan.14 mary austenite solidification. Elements Seven heats of CF-8M containing vary Several tests have been developed such as carbon (C) and manganese ing amounts of delta ferrite were to rate the relative hot cracking ten (Mn) are more soluble in austenite and received. All heats were annealed for dencies of commercial austenitic are thus rejected to the liquid during one hour at 2050°F (1121°C) and water stainless steels. These include the Fis primary delta ferrite solidification. quenched. The chemical compositions sure Bend Test,18 the Cast Pin Tear Sulfur and phosphorus form low of the heats are given in Table 1. Th Test,3-19 the Varestraint Test,5-*1*-0-21 and melting eutectics between sulfides compositions shown were determined many others. These tests generally and phosphides, and silicon tends to by wet analysis and are the average of conclude that austenitic stainless form glassy silicate films. In alloys sol three analyses. It should be noted that steels containing a few volume-per idifying as primary austenite, sulfur, the carbon content of heat 6 and the cent delta ferrite offer a much higher phosphorus, and silicon are rejected to nickel content of heat 3 were greater resistance to hot cracking than do the liquid, form low melting constit than allowed by CF-8M specifica wholly austenitic stainless steels. uents in the interdendritic regions dur tions. All fusion welding processes involve ing the terminal transient period, and Ferrite contents as measured by microsegregation of the component thus cause hot cracking. Alloys solid point counts taken of the as-cast alloying elements. In stainless steel ifying as primary delta ferrite reject microstructure in accordance with weldments, chromium (Cr) is rejected fewer of these elements to the liquid ASTM E-562 are listed in Table 1. to the liquid during primary austenite and consequently are more resistant to Magne-gage readings for both the as- solidification and nickel (Ni) is re hot cracking. cast and weld microstructures are also
Table 1-Chemical Compositions and Ferrite Contents of Seven CF-8M Heats
Phase anal /SIS Magne- gage ferrite no. Heat . Vol-% ferri e by no. C Mn Si Cr Ni Mo S P N point count in castir g Casting Weld 1 .08 .60 1.05 18.32 13.20 2.26 .016 .035 .041 0.2 0.2 0.2 2 .05 .26 .83 17.65 12.03 2.05 .017 .021 .06 1.5 1.2 2.0 3 .04 .21 .38 19.55 15.38 2.88 .025 .036 .04 2.0 2.6 2.6 4 .06 1.17 .60 18.18 12.08 2.48 .016 .027 .05 4.2 3.0 5.8 5 .07 .57 .99 18.21 9.57 2.39 .019 .024 .06 7.6 5.0 8.7 6 .10 .30 .69 20.31 10.62 2.34 .032 .046 .05 14.2 14 17 7 .06 .44 1.08 21.10 9.55 2.52 .008 .021 .048 22.8 > 25 >25
WELDING RESEARCH SUPPLEME NT I 137-s I-IV-
\— l"— I Fig. 3—Perspective view of position of Var estraint specimen within the leg of the casting keel block
mm) wide, voltages and currents were adjusted as shown in Table 3 for each heat. Following Varestraint testing, the as- welded surface was examined under a stereo-microscope equipped with a filar eyepiece. At each level of aug mented strain, the total crack length was determined for each specimen by summing the lengths of the individual cracks. The maximum crack length and number of cracks at each strain level were also determined. Metallographic specimens were re moved from the test bars as shown in Fig. 4. A section parallel to the weld surface, as well as longitudinal and transverse cross-sections of the weld
Fig. 2—Front view of mold design showing the location of Varestraint speci mens Table 2—Constant Parameters
Electrode composition: tungsten + 2% listed. Test specimens were machined specimens were cleaned with acetone thorium from selected positions in a 12 in. (305 and demagnetized with an a-c de Electrode: conical, 90 deg included angle mm) long keel block casting as shown gaussing coil. Electrode extension from collet ~ '-''As in. schematically in Figs. 2 and 3. The welding parameters for this (16.7 mm) The Varestraint Test26*27 was used to investigation are given in Tables 2 and Electrode diameter: Va in. (3.18 mm) determine the weldability of the 7 3. In order to maintain an elliptical Arc length ~ %2 in. (2.4 mm) (measured heats received. Prior to testing, all weld pool approximately % in. (12.7 cold) Travel speed: 4 ipm (1.7 mm/s) Gas flow rate: 40 cfh argon (18.9 L/min) Current: DCSP LOCATION OF WELD PUDDLE AT INSTANT OF APPLICATION OF AUGMENTED STRAIN Table 3--Variable Parameters
Energy Heat Current, Voltage, input, no. A V kj/in,
1 330 14.0 69.0 2 330 14.0 69.0 3 300 15.0 67.5 SOLID-LIQUID INTERFACE ^-ARC LOCATION 4 325 14.5 71.0 5 305 14.5 66.0 SECTION REMOVED FOR OBSERVATION HAZ CRACKS 6 310 14.5 67.5 Fig. 4—Top surface of weld showing location of arc, weld puddle and solid-liquid 7 310 14.5 67.5 interface at instant of straining
138-sl MAY 1980 J HEAT 1
/ i / i 500 500
Q. o —I >UJ /% 250 250 o ce 0 0 05 10 2 0 05 10 0 05 10 20 o. PERCENT APPLIED STRAIN PERCENT APPLIED STRAIN PERCENT APPLIED STRAIN o Fig. 5—Total fusion zone crack length vs. percent applied strain for heats 1-7 > UJ Q\ X o graph indicates which etchant was 500 throughout. The point count method 500 o of analysis was employed. Ten-second counts were repeated at least twice for each point to reduce the possibility of electrical error. Raw data were re X duced to weight-percentages using o the MAGIC IV computer algorithm DC 28 250 < developed by Colby. LU Crt Results and Discussion Varestraint Results 0. Figure 5 shows the total crack length o 0 0-5 1-0 vs. percent augmented strain relation 0 05 10 20 PERCENT APPLIED STRAIN ships of the seven heats. Where addi PERCENT APPLIED STRAIN > tional specimens were available, du UJ Q plicate tests were run to substantiate X the data. o tx were prepared. Specimens were Several heats of CF-8M were found temperature ferrite present in the as- < mounted in epoxy and polished to be extremely resistant to hot crack cast microstructure. LU WELDING RESEARCH SUPPLEME NT I 139-s \ 2% AUGMENTED STRAIN 500 \ \ •PRIMARY S E x \ \- \ •z. \ LLI \ _l R \ < 6 \ c\j oir \ 250 h- _i 2 ID iS O p V The data indicate that the sensitivity to increased, alloys in the iron-rich cor hot cracking tends to decrease with ner of the ternary system experience a increasing volume-percent ferrite. transition from peritectic to eutectic However, the lack of consistency in behavior. Reference to Fig. 1 reveals the trend suggests that other factors that this transition occurs at the point are involved. where the dashed liquidus crosses the solidus curve at the 90% iron line. Thus, alloys which contain 90% or Analysis of Microstructure more iron and lie between the dashed Pseudo-binary diagrams, consisting liquidus and the righthand solidus of sections of the Fe-Cr-Ni ternary at exhibit peritectics. Alloys containing constant iron content,23*32*36 are helpful less than 90% iron which fall between in understanding the microstructure of the solidus curves exhibit ternary these alloys. However, it should be eutectics. emphasized that pseudo-binary dia Comparison of Figs. 7 and 8 illustrate grams can only provide qualitative this fact. information relative to the solidifica In Fig. 7, the pseudo-binary section tion of commercial alloys, because the at 93% Fe, the liquidus is located to the synergistic effects of carbon, molybde Ni-rich side of both solidus lines to // I / t num, manganese and silicon on solid form a typical skewed peritectic trian ification phenomena are poorly under gle. On the other hand, in Fig. 8, with I / / stood. Therefore, the pseudo-binary the pseudo-binary section at 70% Fe, 4""-*-^ I / \ approach will be used only as an aid in the liquidus lies between the two soli ""-^X^J ' visualizing primary effects rather than dus lines as is typical of a ternary T\ \ \ as a basis for making quantitative pre eutectic triangle. Since commercial dictions. stainless steels contain significantly t The Fe-Cr-Ni system is complicated less than 90% Fe, it seems likely that %Cr- 10 20 30 by the fact that while the binary Fe-Cr the invariant reaction associated with 30 20 10 %Ni system exhibits a minimum at 1516°C the terminal transient period of solidi Fig. 8—Fe-Cr-Ni pseudo-binary diagram at (2761°F) and 21% Cr, the binary Fe-Ni fication will be an eutectic and the 70% Fe exhibiting a ternary eutectic typical system exhibits a peritectic at 1517°C pseudo-binary will resemble that ot most commercial austenitic stainless (2763°F) and 4.3% Ni. The Fe-Cr-Ni shown in Fig. 8. steels system exhibits a ternary eutectic at Figure 9 shows such a pseudo-binary < 1300°C (2372°F) and 9 Fe-48 Cr-43 diagram in schematic form with the 1 through 7 indicated by the num Ni. Thus, as the chromium level is approximate relative locations of heats bered arrows. 140-s I MAY 1980 ferrite, the divorced phase at the grain and subgrain boundaries is austenite. The ferrite which forms during the initial transient period for alloys that solidify as primary ferrite is sufficiently enriched in chromium and depleted in nickel to be stable as ferrite at room temperature. Thus, the cores of the subgrains remain as delta ferrite while the remainder of the subgrains experi ence a transformation to austenite. By contrast, in alloys which solidify as primary austenite, the divorced eutec tic ferrite that is retained is located at the grain and subgrain boundaries. UJ CE According to their estimated loca tion on Fig. 9, heats 4-7 would be Si expected to solidify as primary 8- ferrite, as has been confirmed by met cr LU allographic observation. CL Heats 4-6 exhibit the vermicular S- UJ ferrite structure characteristic of weld ments with ferrite numbers between 4 and 20. The views of Fig. 10 are repre sentative of this vermicular 5-ferrite morphology. Inspection of Varestraint results for these materials shows that heats 4 and 5 are highly resistant to hot cracking, which is in agreement with the usual theories. However, heat 6 shows anomalous behavior. Inspec / 23 4 5 tion of Table 1 indicates that the levels I II l l I t of sulfur and phosphorus in this heat were the highest of the seven heats %Cr - %Ni tested in this study. Microprobe analy Fig. 9—Generalized pseudo-binary diagram showing approximate relative posi sis across a backfilled crack in this tions of heats 1-7 material, shown at X250 in Fig. 11A, revealed a significant increase in phos Within the triangular eutectic re sient period, the last remaining liquid phorus over the matrix levels. Appar gion, a three-phase equilibrium exists is of eutectic composition. However, ently, the nominal phosphorus con between austenite, delta ferrite, and since the amount of eutectic liquid is tent of this heat is too large to be liquid. Alloys having compositions on extremely small, the interdendritic partitioned effectively to the primary the nickel-rich side of the apex of the regions consist of a divorced eutec 5-ferrite, and the resulting segregation three-phase eutectic triangle solidify tic.5 pattern lowers the melting tempera as primary austenite. Alloys having In alloys which solidify as primary ture in intergranular regions. compositions on the chromium-rich austenite, the divorced phase at the The cracking which occurred in the side solidify as primary delta ferrite. In grain and subgrain boundaries is fer heat 6 material was located exclusively either case, during the terminal tran rite. In alloys which solidify as primary at grain boundaries as no intercellular * • S. ._ J "*f * * -'••>-< ... j. Fig. 10-Morphology of retained primary delta ferrite dendrite spines: A—heat 4, high magnification, electroetch; B-heat 5, electroetch; C—heat 6, electroetch WELDING RESEARCH SUPPLEMENT I 141-s ' POSITION, 5,± STEPS Fig. 11—Intergranular hot crack in heat 6: A—electroetch; B—Microprobe trace across the Fig. 12-Heat 7, Widmanstatten austenite intergranular material showing phosphorus segregation. The trace initiated in the vicinity ot (dark etching phase) nucleating on grain a S-dendrite and ended in a region of interdendritic austenite; electroetch boundary austenite; Nomarski technique, electroetch cracking was found. Metallographic shown at X250 in Fig. 12, indicates that examination revealed that the vermi the composition of this alloy lies in a interdendritic and intergranular solid cular ferrite structure was absent in the chromium-rich region far to the right ification sites. Lippold5 found this grain boundaries of heat 6. Micro of the three-phase triangle. Primary eutectic ferrite in backfilled cracks in probe analysis detected an increase in 5-ferrite solidification consumed al Type 310 stainless steel Varestraint nickel and a decrease in chromium at most the entire liquid, with austenite specimens and reported a Cr/Ni ratio the grain boundaries; this is in agree forming only at the final intergranular of approximately 2.3 in this region. ment with the postulate that the last sites. During cooling to room tempera David etal.31 report finding no inter material to solidify in this heat would ture, much of the primary ferrite trans dendritic ferrite using a magnetic etch be austenite. formed to austenite by a Widmanstat- ing technique in interrupted solidifica Figure 11B summarizes the results of ten-type nucleation and growth mech tion samples of Type 310 stainless the microprobe analysis made of the anism. Inspection of Figure 12 reveals steel. They also found no indication of backfilled crack shown in Fig. 11A. No that the austenite formed at the grain the formation of an eutectic ferrite difference in molybdenum contents boundary provides nucleation sites phase with the differential thermal could be detected between the back from which the dark etching Widman analysis techniques used in their filled region and the matrix. The statten austenite plates grow into the studies. However, they used a cooling ferrite grains. Inspection of Varestraint behavior of this heat indicates that the rate of 6°C/min (11°F/min), which is results for this heat shows that heat 7, presence of primary 6-ferrite will not at least three orders of magnitude which also solidified as primary ferrite, necessarily ensure good weldability. slower than those found in normal is highly resistant to hot cracking. Control of the residual elements GMA, GTA, and SMA weldments. It is known to have a deleterious effect on Heats 1-3 solidified as primary aus also questionable if differential ther the weldability of stainless steels is tenite. Inspection of Varestraint data mal analysis can resolve the crystalliza also important. indicates that the three heats are high tion of a second phase when less than Heat 7 also solidified as primary ly susceptible to hot cracking. In theo 1 vol-% of this phase is formed during delta ferrite. The microstructure, ry, ferrite should form in the final the solidification process. V % i y«N b K ! ; ' y ' * *mmdQ$: f I® I Fig. 13-Morphology of retained eutectic ferrite in primary austenite heats: A-heat 1, intercellular eutectic ferrite, electroetch; B~heat 2, intercellular eutectic ferrite, electroetch; C—heat 3, semi-continuous eutectic ferrite exhibiting a fusion zone ferrite number of 2.6 142-si MAY 1980 -% POSITION, 1^ STEPS Fig. 15— Backfilled crack area in Heat 1: A—electroetch (Contamination spots and beam focusing areas can be seen.); B—Phosphorus Ka X-ray map of area in 15A; C—Sulfur Ka Fig. 14—Heat 1 exhibiting intercellular (Mo-La) X-ray map of area in 15A eutectic ferrite: A—electroetch; B—Micro• probe trace across eutectic ferrite in heat 1 Close inspection of Fig. 13C, taken of ture of the backfilled crack shown in heat 3 at X250, reveals an almost con Fig. 15A made with the microprobe Finally, Lippold5 has shown that the tinuous interdendritic eutectic ferrite using characteristic X-rays. Figure 15B nominal composition within the pri network to be present in this weld was made with phosphorus Ka radia mary austenite region greatly in ment. Microprobe analysis was per tion. Figure 15C represents a summa fluences the amount of eutectic ferrite formed across the eutectic ferrite tion of the sulfur Ka and molybdenum formed. The closer the composition is shown in Fig. 14A. The results, summa La which are almost identical in wave to the ternary liquidus, the greater is rized in Fig. 14B, indicate the ferrite to length. Phosphorus segregation to this the amount of eutectic ferrite formed. be enriched in chromium and molyb region is clearly evident in Fig. 15B. This result agrees qualitatively with denum and depleted in nickel. Note Although the segregation shown in Fleming's analysis38 of a similar situa that the molybdenum content of the Fig. 15C is obviously greater, the sulfur tion in binary Al-Cu alloys. The frac ferrite is more than three times that of Ka peak is so close to the Mo La peak tion of eutectic is an exponential func the adjacent austenite. The effect of that it is not possible to obtain a tion of the ratio between the nominal molybdenum as a ferritizer has been quantitative estimate of the amount of alloy composition and the eutectic well documented,29'39*40-30 but usually sulfur present in the backfilled crack. composition. with respect to its influence on the The compositions of commercially ferrite content of materials which so Cast Structures available Type 310 stainless steels are lidify with ferrite as the primary phase. sufficiently removed from the appro Undoubtedly, the partitioning of mo Inspection of the microstructure priate ternary liquidus points to ensure lybdenum to interdendritic volumes found in the as-cast CF-8M alloy that only a very small amount of terna greatly aids in ferrite formation. revealed that more ferrite was present ry eutectic may form. Hot cracking in primary austenite than would be expected in compara Figure 13 clearly establishes the exis alloys has been related to many fac ble wrought materials, such as Type tence of eutectic ferrite at final solidi tors, including the partitioning of 316 stainless steel. The segregation in fication sites between primary auste harmful ferritizing elements to final castings is on a far coarser scale, and nite cellular dendrites and in inter solidification sites. Views B and C of the solidification substructure size are granular volumes in heats 1, 2, and 3. Fig. 15 are images of the microstruc several orders of magnitude larger WELDING RESEARCH SUPPLEMENT 1143-s Fig. 16—Cast and annealed microstructures of Heats 1-7, (A-G, respectively) all electro- etched ferrite in the primary austenite heats is in the form of a divorced eutectic, rather than primary, ferrite. Comparison of Magne-gage read ings for heats 3 and 4 illustrates the danger of using this type of ferrite determination to predict weldability. The difference in ferrite numbers (3.0 vs. 2.6) does not correlate with the weldability as revealed by the Vare straint test results shown in Fig. 6. Thus, comparison of weld solidifica tion substructures must be made before the weldability can be pre dicted accurately. Massive Transformation Lippold5 has recently postulated that the conversion of primary delta ferrite to austenite could occur by a diffu- sionless massive transformation as a result of the rapid cooling through the two-phase delta-plus-gamma region. This would apply to alloy composi tions which lie on the chromium-rich side of the three-phase triangle on the pseudo-binary diagram where the pri mary solid phase is delta ferrite. Figure 17 compares ferrite morpho logies at the fusion boundary of auto genous welds in alloys which solidi fied as primary delta ferrite and as primary austenite, respectively. The than in weldments. The slower solidifi ferrite-austenite interfaces in the pri have a composition approaching CF. In cation and cooling rates experienced mary austenite heats exhibited none of fact, it is possible for regions within by castings result in the retention of the angular facets which would be the ferrite to have any composition considerable ferrite even after a one- expected with a massive transforma between C„ and CF. hour anneal at 2050°F (1121°C). tion. The liquidus temperature for this Figure 16, taken at X250, shows the Solidification as primary delta ferrite cast structure will have a range deter base metal microstructure of the vari results in chromium segregation at the mined by the slopes of the solidus and ous heats studied. Notice that signifi spines of the ferrite dendrites. Thus, liquidus and the nominal composition cant ferrite is present in both the pri even with slow cooling, retained fer of the alloy. The result is that the mary austenite and the primary ferrite rite has a higher chromium content ferrite islands in the casting will have a heats. The morphology of the ferrite than the austenite. Figure 18 shows higher melting temperature than the appears to be unaffected by the pri that for a nominal alloy composition surrounding matrix. When the alloy is mary solidification mode, although the C0, primary dendrite spines should welded, it is probable that ferrite 144-sl MAY 1980 /. •x /J # >• • *• v. .->:*/ !•• * /-.•/ t • Fig. 18—Pseudo-binary diagram showing possible range of composition oi primary S-ferrite dendrites Recent STEM microanalysis by Ly tion range appears to straddle the ter man33 has shown a compositional vari nary liquidus curve, either ferrite or ation of less than 1 wt-% of chromium austenite may serve as the primary and nickel across the massive transfor phase, depending on the actual com mation interface. STEM micrographs position. indicate a wall of dislocations located 5. Divorced eutectic ferrite is ob at the interface, providing a mecha served at final interdendritic and inter Fig. 17—Fusion line ferrite morphology: nism for the diffusionless shear trans granular solidification sites when aus A—primary S-ferrite heat 5, showing angular formation. tenite is the primary phase. facets, electroetch; B—primary austenite 6. Microanalysis reveals this ferrite heat 2, electroetch to be enriched in both molybdenum and chromium. Conclusions 7. Ferrite is present in the as-cast 1. The results of Varestraint testing and annealed base metal structure in islands in the base metal at the solid- support the observations that stainless amounts greater than would normally liquid interface do not melt, while steels which solidify with delta ferrite be found in wrought materials of simi adjacent austenite of roughly nominal as the primary phase are less suscepti lar composition, regardless of the pri composition does melt. ble to hot cracking than those in mary solidification mode. The ferrite has an immense tempera which austenite is the primary phase. 8. Because of the coarseness of the ture gradient impressed upon it as a 2. Segregation of phosphorus to as-cast and the annealed microstruc result of its intimate contact with the final solidification sites has been tures, it is not possible to identify liquid on one side of the island and the shown to occur in heats in which whether ferrite or austenite was the heat sink provided by the solid base austenite is the primary phase; thus primary phase from the ferrite mor- plate on the other side. This rapid supporting the hypothesis that this pholoy. cooling of ferrite through the two- element is partitioned to the liquid 9. Magne-gage analyses alone do phase austenite-plus-ferrite field per during primary austenite solidifica not appear to be sufficient to deter mits the occurrence of the massive tion. mine whether or not a specific heat of delta to gamma-phase transformation 3. Phosphorus may also be detri material is readily weldable. previously mentioned. Inspection of mental to heats which solidify as pri 10. The angular, faceted ferrite mor the ferrite islands in Fig. 17A shows the mary delta ferrite if its concentration is phology observed at the fusion angular crystallographic facets charac in excess of that which may be effec boundary in heats in which ferrite is teristic of some massively transformed tively absorbed by the ferrite. the primary phase resembles that asso materials. 4. Since the CF-8M alloy composi ciated with massive diffusionless WELDING RESEARCH SUPPLEMENT I 145-s transformations found in other alloy 4806. 24. DeLong, W. T., and Ostrom, C. A., systems. 7. Pumphrey, W. F., and Jennings, P. H.,). WRC Bulletin 1204. 11. Analysis of available ternary Fe- inst. Met., 75, 1948. 25. Schaeffler, A. L, Metal Progress, Vol. 8. Pumphrey, W. F., and Lyons, |. V., /. 56, 1949, pp. 680, 680B. Cr-Ni phase diagrams leads to a con Inst. Met, 74, 1948. 26. Savage, W. F., and Lundin, C. D., clusion that an eutectic reaction is the 9. Medovar, B. I., Avtom. Svarka., 7(4), Weld. Jour., Oct. 1965, pp. 433-s to 442-s. invariant reaction characterizing the 1954, Brutcher Trans. No. 3400. 27. Savage, W. F., and Lundin, C. D., end of solidification in commercial 10. Toropov, V. A., Metallovendenie I Weld. Jour., Nov. 1966, pp. 497-s to 502-s. stainless steels. Obrabatka Metallov, June, 1957, Brutcher 28. Colby, |. W., MAGIC IV, Bell Tele Trans. No. 3982. phone Laboratories, Allentown, PA. Ac/enow/edgments 11. Pellini, W. S., Foundry, 80(11), 1952. 29. Thomas, R. D., Metal Progress, 1946, 12. Apblett, W. R., and Pellini, W. S Vol. 50, pp. 414-479. The authors gratefully acknowledge Weld. Jour., 33(2), 1954. 30. Lyubauskii, K. V., and Toropov, V. A., the financial support of the Steel 13. Borland, J. C, Brit. Weld, lour., 7(8) Paton Commemoration Volume Mashgiz, Founder's Society of America and the 1960. 1956, pp. 117-134. Welding Research Council in this 14. Movchan, B. A., and Poznyak, L. A. 31. Heuschkel, )., Weld. Jour., 1955, Vol. investigation. Special thanks are also Avtom, Svarka., 8(6), 1954, Brutcher Trans 34, pp. 484-s to 504-s. offered to Dr. Peter Wieser and the No. 3572. 32. Pugh, J. W., and Nisbet, J. D., Trans. 15. Matsuda, F., Nakagawa, H., Vehara AIME, Vol. 188, Feb. 1950. other members of the High Alloy T., Katayama, S., and Arata, Y., Trans. JWRI 33. Lyman, C. E., unpublished research at Technical Research Committee of the Vol. 8, No. 1, 1979, pp. 105-112. Rensselaer Polytechnic Institute. SFSA for their guidance and insight 16. Cooch, T. C, and Honeycombe, J. 34. Parks,). M., Weld. Jour., 1955, Vol. 34, throughout this project. Metal. Const, March 1975, pp. 146-148. pp. 568-570. 17. Arata, Y., Matsuda, F., and Katayama 35. Lashko, N. F., and Lashko, A., Autom. References S., Trans. JWRI, Vol. 5, No. 2, 1976, pp. Svarka., 1958, Vol. 11, pp. 99-110. 1. Scherer, R., Riedrich, C, and Hock, G., 35-51. 36. Metals Handbook, Eighth Edition. Arch. Eisenhut, Vol. 13, 1939-1940, pp. 53- 18. Lundin, C. D., and Spond, D. F., Weld. Vol. 8, pp. 291 and 304. 57. Jour., Nov. 1976, pp. 345-s to 367-s. 37. David, S., Goodwin, C, and Braski, 2. Borland, I. C, and Younger, R. N., Brit. 19. Hull, F. C, Weld. Jour., April 1959, pp N., Weld. Jour.. Nov. 1979, pp. 330-s to Weld. Jour., Jan. 1960, pp. 22-59. 176-181. 336-s. 3. Hull, F. C, Weld. Jour., Sept. 1967, pp. 20. Gooch, T. G, and Honeycombe, ). 38. Flemings, M., Solidification Process 399-s to 409-s. Metal. Const., Sept. 1970, pp. 375-380. ing, 1974, McGraw-Hill, Inc., p. 81. 4. Arata, Y., Matsuda, F., and Katayama, 21. Arata, Y., Matsuda, F., and Saruwatari 39. Tremlest, H. F., Trans. Inst. Weld., F., Trans. JWRI, Vol. 6(1), pp. 105-116. S., Trans. JWRI, Vol. 3(1), 1974, pp. 79-88. 1953, Vol. 16, pp. 143-150. 5. Lippold, )., Ph.D. Thesis, Rensselaer 22. Metals Handbook, Eighth Edition, 40. Feild, A. L, Bloom, F. K., Linnert, C. Polytechnic Institute, August 1978. Vol. 8, p. 424. E., NDRG Report OSRD 3034, 1943. 6. Bochvar, A. A., and Sidevskaya, Z. A., 23. Masumoto, I., Tamaki, K., and Kutsu- 41. Lundin, C, DeLong, W., and Spond, Izvest. Akad., NAVK SSSR, Otcl. Teckhn. na, M., Trans. JWS, Vol. 41(11), 1972, pp. D., Weld. Jour., August 1975, Vol. 54, pp. NAVK, No. 3, 1947, Brutcher Trans. No. 1306-1341. 241 -s to 246-s. WRC Bulletin 254 November 1979 (1) A Critical Evaluation of Plastic Behavior Data and a Unified Definition of Plastic Loads for Pressure Components by J. C. Gerdeen (2) Interpretive Report on Limit Analysis and Plastic Behavior of Piping Products by E. C. Rodabaugh (3) Interpretive Report on Limit Analysis of Flat Circular Plates by W. J. O'Donnell These three reports summarize a four-year effort by the PVRC Task Group on "Characterization of the Plastic Behavior of Structures" to meet the need for unified and standardized methods for limit analysis on plastic collapse determinations. Publication of this report was sponsored by the Pressure Vessel Research Committee of the Welding Research Council. The price of WRC Bulletin 254 is $13.50 per copy. Please include $3.00 for postage and handling. Orders should be sent with payment to the Welding Research Council, 345 East 47th St., Room 801, New York,-NY 10017. 146-sl MAY 1980