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By J. A. Thompson, M. A. Mayers

IN recent years, the problem of damage to coke-oven walls by expanding coal charges undergoing carbonization has engaged great attention on the part of research workers in this field, and has led to the development of many devices for estimating these pressures.1-14 Only in Russell's method11 and in the Koppers ovens7 are conditions quite similar to those in commercial ovens. Tests even in these ovens are empirical because no information has been published to permit an evaluation of "scale effects." Owing to the relatively large wall area in these test ovens, the average wall pressure determined may be lower than that existing in localized areas. This factor may be compensated for in part by the testing of coal charged into the test oven at a bulk density that is as high as the highest that is believed to exist in commercial ovens. Both of these lacks might be filled by the development of an adequate gauge for the direct measurement of wall pressure exerted by the charge in commercial coke ovens in normal operation. Previous attempts to elaborate such instruments7,15 have resulted in units which, because they had to be water-cooled, probably produced changes in the conditions of carbonization, and which had to be removed from the oven before carbonization was complete, to avoid their destruction. Late in 1941 the Coal Research Laboratory undertook to develop such a gauge, in which gas pressure was balanced against the charge pressure and which thus would not require water cooling, on a suggestion made in 1937 by one of the present authors and William B. Warren, a former member of the Laboratory staff. The gauge heads themselves were to be simple and cheap enough so that they might be discarded after each run if seriously damaged. This would permit them to be left in place throughout the carbonizing period, when they could be pushed out with the coke. The basis of the design, of which the latest exemplification is shown in Fig. I, was that as the internal gas pressure within a flexible metal envelope was increased, the flexible wall would move away from a stop when the gas pressure just overbalanced the charge pressure. If the stop were an insulated electrical contact, the point of balance could be indicated by the breaking of an electric circuit. It was anticipated that, while the instrument must operate at high temperature, the presence of a strongly reducing atmosphere throughout the carbonizing period would so reduce-the danger of scaling that mild steel could be used for the gauge. This expectation has been justified. Danger of attack on silica oven walls by the steel gauge body can be eliminated by cementing asbestos paper or an alundum or silica plate to the back of the gauge body. The gauge head itself has a relatively heavy mild-steel body, A, into which standard ¼ in. pipe B is screwed to conduct gas and electrical leads. The body is closed by a single bellows C. consisting of two annuli of 0.011-in. (0.28-mm.) cold-rolled shim-steel spot-welded together at their outside edges. The lower one is

Jan 1, 1944

By F. W. C. Boswell, R. L. Cunningham, H. M. Weld

Chemical analysis shows the partition of the magnesium addition between the metal and the graphite is independent of whether the graphite is in the form of nodules or flakes. The bright central spot observed in nodular graphite polished sections is shown to be due to the structure of the nodules. DURING the past few years numerous papers have been published on the production1-7 and heat treatment" of nodular cast irons. However, very little information is available on the structure of the nodules and the reason for their formation. It would appear that an explanation of the phenomenon of nodulation must await a detailed knowledge of the nature of the nodule. Disagreement exists among various workers on such a fundamental point as to whether or not a nodule actually contains a microscopically visible, nongraphitic nucleus.9-11 Work was initiated with a view to obtaining definite information on this and other aspects of the problem. Chemical Separation and Analysis Microscopic appearance had indicated9-11 that a nongraphitic core might exist in nodules (Fig. 1). It appeared that a chemical comparison of separated nodules and flake graphite would be profitable, for if the core is nongraphitic a relatively large amount of certain elements would be found in the nodules as compared to the flake. This comparison could best be made by obtaining nodules and flake from the same melt of cast iron, part of which had been in-noculated with a suitable addition reagent for the formation of nodules. A 300 lb melt was produced. About half of this was innoculated in the ladle with an Fe-Si-Mg alloy to produce nodules and cast into keel blocks; the remaining half of the melt had silicon added in an effort to compensate for the silicon in the innoculant used in the other half of the melt and this was cast into similar keel blocks. The analyses of the resultant irons are given in Table I. Nodules and flake were chemically extracted from 2 kg samples of metal as described in the appendix. While this method dissolved the iron and chromium carbides, it did not dissolve the titanium, vanadium, and molybdenum carbides. These latter carbides could have been removed by nitric acid, but in other experiments it was found that it attacked pools of metallic iron entrapped in the nodules which had, until the nitric acid treatment, been protected from the attack of the sulphuric, hydrochloric, and hydrofluoric acids. It will be shown in the next section that entrapped iron exists in nodules but not in flake. On microscopic examination many tiny irregular particles were observed to accompany both the nodules and the flake. These were identified by tabling a sample of nodules from which a sufficient fraction was obtained for X-ray analysis. A good pattern of titanium carbide was obtained indicating that this was the predominant compound present. Since it was not possible to determine how completely the carbides could be removed mechanically and since their removal from flake is much more difficult, it was decided not to resort to any separation as this would render further comparisons more difficult. The samples were ashed in a combustion

Jan 1, 1953

By Lauren A. Wright

SINCE the early nineteen-thirties the production of talc* in California has increased five-fold to a yield in 1947 of about 76,000 tons (fig. 1); conse- .' Unless otherwise qualified, the term "talc" in this paper will imply a mixture of the pure talc with such related minerals as commonly occur in the commercial material. quently the state is now one of the nation's three principal talc producers. The 1947 output, which represented about one-sixth of the national total for that year, was approximately equal to production from Vermont, and was second in value to that from New York. New York has an annual output of about 130,000 tons. Three factors have been mainly responsible for the expansion of talc production in California. (1) Technical advances and innovations, particularly in the ceramic industry, have increased the demand for certain California talcs. The nation's high-qual- LAUREN A. WRIGHT, Junior Member AIME, is Associate Geologist, California Division of Mines, San Francisco, Calif. Los Angeles Meeting, October 1948. San Francisco Meeting, February 1949. TP 2622 H. Discussion of this paper (2 copies) may be sent to Transactions AIME before Feb. 28, 1950. Manuscripts received Dec. 7, 1948; revision received Aug. 29, 1949. This paper is Contribution No. 508 of the Division of Geological Sciences, California Institute of Technology, Pasadena, Calif. the industry, mainly through demand for the urgently needed steatite. Fibrous talc, used as a paint extender by the armed forces, also was produced in quantity. (3) Expanding western industry has continued to draw heavily on California talc deposits, and has extended the upward production trend into the postwar period. Considerably more than half of California's current talc output is used as ceramic raw material. The paint, rubber, paper, and textile industries are the more important consumers of the remainder. Smaller amounts are used as insecticide carriers, and in cosmetics, and various minor industrial applications. During World War I1 the principal steatite deposits were studied in detail by members of the U. S. Geological Survey; other sources of commercial talc in the state have in the past received but scant geologic attention. The following discussion is essentially a preliminary report based on research currently in progress by the California Division of Mines, and includes a brief outline of previous steatite studies by the U. S. Geological Survey. Distribution of California Talc Deposits: The principal talc resources of both California and Nevada are confined to an elongate belt that extends from the vicinity of Baker, in north-central San Bernardino County, northwestward to include the southern Death Valley region, the Inyo Range, and the Palmetto-Oasis district of western Nevada (fig. 2). This belt, of which all but the northernmost tip lies in California, is approximately 200 miles long and 30 miles in average width. Included in the belt are four contrasting areas. The deposits within each area are geologically similar and yield talcs that have characteristic properties and applications. Three of these local "talc provinces" lie within California's portion of the talc-

Jan 1, 1951

By Lauren A. Wright

SINCE the early nineteen-thirties the production of talc* in California has increased five-fold to a yield in 1947 of about 76,000 tons (fig. 1); conse- .' Unless otherwise qualified, the term "talc" in this paper will imply a mixture of the pure talc with such related minerals as commonly occur in the commercial material. quently the state is now one of the nation's three principal talc producers. The 1947 output, which represented about one-sixth of the national total for that year, was approximately equal to production from Vermont, and was second in value to that from New York. New York has an annual output of about 130,000 tons. Three factors have been mainly responsible for the expansion of talc production in California. (1) Technical advances and innovations, particularly in the ceramic industry, have increased the demand for certain California talcs. The nation's high-qual- LAUREN A. WRIGHT, Junior Member AIME, is Associate Geologist, California Division of Mines, San Francisco, Calif. Los Angeles Meeting, October 1948. San Francisco Meeting, February 1949. TP 2622 H. Discussion of this paper (2 copies) may be sent to Transactions AIME before Feb. 28, 1950. Manuscripts received Dec. 7, 1948; revision received Aug. 29, 1949. This paper is Contribution No. 508 of the Division of Geological Sciences, California Institute of Technology, Pasadena, Calif. the industry, mainly through demand for the urgently needed steatite. Fibrous talc, used as a paint extender by the armed forces, also was produced in quantity. (3) Expanding western industry has continued to draw heavily on California talc deposits, and has extended the upward production trend into the postwar period. Considerably more than half of California's current talc output is used as ceramic raw material. The paint, rubber, paper, and textile industries are the more important consumers of the remainder. Smaller amounts are used as insecticide carriers, and in cosmetics, and various minor industrial applications. During World War I1 the principal steatite deposits were studied in detail by members of the U. S. Geological Survey; other sources of commercial talc in the state have in the past received but scant geologic attention. The following discussion is essentially a preliminary report based on research currently in progress by the California Division of Mines, and includes a brief outline of previous steatite studies by the U. S. Geological Survey. Distribution of California Talc Deposits: The principal talc resources of both California and Nevada are confined to an elongate belt that extends from the vicinity of Baker, in north-central San Bernardino County, northwestward to include the southern Death Valley region, the Inyo Range, and the Palmetto-Oasis district of western Nevada (fig. 2). This belt, of which all but the northernmost tip lies in California, is approximately 200 miles long and 30 miles in average width. Included in the belt are four contrasting areas. The deposits within each area are geologically similar and yield talcs that have characteristic properties and applications. Three of these local "talc provinces" lie within California's portion of the talc-

Jan 1, 1951

By T. D. Tiemann

The extraction of iron oxide from several types of siliceous iron ore by digestion in sulfuric acid solutions was investigated in the temperature range from 80 to 258°F at concentrations from 10-60% H2SO4. Degree and rate of extraction were found to be critical with respect to both temperature and acid concentration, requiring temperatures in excess of 250° F and acid concentrations of 50% to effect 80-90% extraction of iron oxide in 60 min. Rate of extraction is directly related to surface area and apparently controlled by the rate of chemical reaction. The activation energy was found to be approximately 25,000 cal per g mol. Extraction characteristics of the several ores were found to be similar. The Department of Minerals and Metals Engineering at the University of Wisconsin conducts a continuing research program to develop processes for the utilization of low-grade iron ore deposits such as the nonmagnetic taconites of the Wisconsin Gogebic range. Methods under study include high-intensity magnetic and high-tension electrical concentration, flotation, direct reduction and chemical extraction. Concentration of a number of low-grade siliceous iron ores by the extraction of silica in sodium hydroxide solutions has been reported by the writer and others in the literature.1,2,3,4 Residual concentrates containing as much as 65% iron have been obtained from taconites containing less than 30%. Currently an alternative approach is under study, namely the dissolution of iron oxides from siliceous gangue by extraction in acid, with subsequent recovery of the iron from solution. Such a process for extracting iron oxide from the low-grade oclitic deposits of Alberta with hydrochloric acid was investigated and described by Gravenor, Govett and Rigg.5 The current investigation, which is entirely devoted to extraction with sulfuric acid, is the subject of this paper. In part, the objective of this research was to determine the possibility of leaching iron ores in situ with sulfuric acid. The investigation of a recovery method is left to later research. The ores investigated contained hematite (Fe2O3) as the principal iron-bearing mineral and minor amounts of goethite (Fe2O3 .H2O), magnetite (Fe3O4), and iron-bearing silicates. The gangue was essentially quartz occurring as single crystals and microcrystalline aggregates. Detailed description of the ores may be found in the literature.1,2,4,6 A comparison of the specific ores used in this investigation is shown in the partial analyses of Table I. EXPERIMENTAL METHODS All extractions were made at atmospheric pressure using beakers with a mechanical stirrer for those tests run below the boiling point, and Erlenmeyer flasks for those at the boiling point. A thermostatically controlled hot plate was used for heating, and this maintained the mean temperature within the solution to within ±3°F during extraction. The sample of ore for extraction was weighed to + 0.02 g. For tests below the boiling point, the sample was added to the beaker containing the sulfuric acid solution at the desired temperature. For the boiling-point tests, the sample was added to the Erlenmeyer flask followed by distilled water and the

Jan 1, 1968

By C. A. Bruch, J. F. Duffey

On the basis of metallographic and cooling-rate studies of 15 U-Zr alloys, a new version of the 6 phase region of the U-Zr phase diagram is proposed. IN several published diagrams of U-Zr alloy system:' there is good agreement that at elevated temperatures complete solid solubility exists between y uranium and ß zirconium. Similarly, there is reasonably good agreement concerning the uranium-rich and zirconium-rich ends of the diagram. However, there is no agreement concerning the nature of the diagram at intermediate compositions and at temperatures below about 650°C. The early work of Saller et al.8 hows an intermediate phase, 6. In a later diagram by Saller and Rough' shown in Fig. 1, the d phase field is drawn with dotted lines, indicating doubt that this phase exists. The diagram of Summers-Smith2 does not show the d phase. More recently Holden and Seymour' have shown that in the composition range 40 to 60 wt pet Zr, the high-temperature ? phase, which is body-cen-tered-cubic, transforms to an ordered structure, 8. They found that the order-disorder transformation occurs between 595° and 610°C, and presented a diagram based on the study of four different alloy compositions. Most recently Saller et a1.5 presented a new diagram which differs with respect to the composition limits of the 8 field from the others proposed by these authors. The purpose of the present investigation was to determine the boundaries of the d phase field. Experimental Procedures and Results Alloys—Buttons, weighing 80 g of 15 different alloy compositions were prepared by arc melting Mal-linckrodt uranium and iodide-process zirconium in a water-cooled crucible. A controlled atmosphere of argon and helium was maintained in the furnace. Prior to making each button, a sacrificial button of zirconium was melted to increase the purity of the furnace atmosphere. Each alloy was then melted six times to improve its homogeneity. Subsequently the alloy buttons were jacketed in copper and hot rolled to slabs at 775°C. From these slabs, specimens were machined for chemical, thermal, and metallographic analyses, Table I gives both the estimated chemical composition of the alloys, and the values which were obtained by analysis. The oxygen contents varied from 0.013 to 0.032 pet, nitrogen from 0.0014 to 0.014 pet, and carbon from 0.0013 to 0.011 pet. The purity of the alloys ranged from 99.947 to 99.975 pet. Thermal Analysis—The specimens for thermal analysis were blocks 11/4 x 1/4 x % in. The bead of a calibrated thermocouple was placed between two of these blocks and molybdenum wire was wrapped around the assembly. Each assembly was sealed in a quartz capsule having an argon atmosphere. The capsule was placed inside a refractory crucible which was heated in a furnace. Using a method described by Smith,' the emf of a differential thermocouple, with junctions inside and outside of the refractory crucible, was used to activate the

Jan 1, 1959

By Graham John Mitchell

In May, 1919, a crosscut on the 1400-ft. level of the Briggs mine, a Calumet, & Arizona property at Bisbee, Ariz., penetrated a deposit of pyrite and chaleopyrite that had replaced quartzite and limestone and was cut, especially in the quartzite, by numerous fractures. As the drift advanced through these fractured portions, they were found to be free from copper-sulfate minerals although '(sooty" chalcocite was present in the fractures and also as coatings and replacements of pyritc. Howcver, sulfate soon made its appearance in certain fractures, filled them, and began to form stalactites. Several centers of incipient growth were noted for subsequent study, observations in the drift at frequent intervals during the past 17 months affording the opportunity of determining the rate of formation of the stalactites. The one to attain the greatest growth during the period of observation measured 27 in. long and averaged 1 in. in diameter. A close study showed it to be composed of small interlocked crystals of chalcanthite radiating from a small central tube-like opening. The trielinic form of the crystals was evident with prism and pinacoid faces recognixable. Assuming that the supply of copper-sulfate solution was constant (luring the 17 months—-and such an assumption is supported by the fact that the number of drops of excess solution falling, per minute, from the tip of the stalactite at the times of observation checked closely—thc stalactite, it has been estimated, grew at the rate of 1.58+ in. per month. DISCUSSION James I?. Kemp," New Yolk City.—J. S. Curtiss, in the monogr:tpll on Eureka, Ncv., say:: that a stalactite formed in limestonc grew 5/8 in. in three weeks and continued to grow as long as it had a drop of water or solution at the lower end. Another grew at the rate of 3/8 in. in the same time. In the case cited by Mr. Mitchell, the growth was about 136 in. in three weeks. A year ago we had an instance of the formation of cubes of galena on a railroad spike. A mine in northeastern Oklahoma had been idle about two years but on account of pumping of neighboring properties, the circulation of water was active throughout the drift. A cube of galena,

Jan 1, 1922

By A. W. Hesse

Ventilation of all types of coal mines is fundamentally the same, in that sufficient air must be provided to properly dilute and remove dangerous and obnoxious gases and leave the oxygen content of the air not under 1 9 per cent in all places where men and animals must work. Perhaps animals should not be mentioned, as the speed of mechanical loading leaves livestock behind and outmoded. Since this paper has to do with the influence of mechanical loading equipment on ventilation, the problems will be confined to the splits and the coal-producing places of the mine. The removal of coal dust from the air, or its rapid transmission to the return air courses; the possible increase in the liberation of gases by the multiplication of cuts in the coal made during one shift or multiple shifts; and the additional electrical equipment at the faces, are the factors that cause the added importance to good ventilation in the permissible equipment area of gassy bituminous coal mines. Ventilation tor Mechanical Loading in Gassy Mines In gassy mines, the air currents must be maintained across the faces in such a way as to prevent short cuts to the returns and leaving pockets in which accumulation of gas might occur. To do this, two or more doors are required on almost every entry containing a track to the face, so located that in no event will a trip of cars block open the doors and permit the air to bypass some place or places on the air-approach side. This method frequently requires the use of unusual lengths of duplex cables to reach the mechanical equipment at the faces, particularly where the distances between parallel working places are 100 ft., as at the Nemacolin mine. Fig. I shows a pillar section being worked with mechanical loaders. This layout requires carrying the air in on the right and releasing it on the left, because the track clearance is on the right side and the line brattice is built on this side from the last crosscut to the face. Thus, the air travels to the face with greater velocity and gives more assurance of sweeping out the gas that might be exuding from the crevices of the coal. All working places, in driving through solid coal, are thus treated alike. They are all potentially gassy and must be swept clear of whatever may contaminate the air. Once the air starts into the working section, it is not released until it does a good job. Of course the volume of air must be large enough not only to dilute the gases but to carry away explosive fumes and fine coal dust quickly. Mechanical loading entails cutting and drilling of the coal during the same working period as the loading when two-shift and three-shift operations are employed, and all three processes create finely pulverized coal, which in itself is hazardous if not controlled by some means. If the ventilation is inadequate and gas accumulates, and a spark from defective electrical

Jan 1, 1942

By A. W. Hesse

Ventilation of all types of coal mines is fundamentally the same, in that sufficient air must be provided to properly dilute and remove dangerous and obnoxious gases and leave the oxygen content of the air not under 1 9 per cent in all places where men and animals must work. Perhaps animals should not be mentioned, as the speed of mechanical loading leaves livestock behind and outmoded. Since this paper has to do with the influence of mechanical loading equipment on ventilation, the problems will be confined to the splits and the coal-producing places of the mine. The removal of coal dust from the air, or its rapid transmission to the return air courses; the possible increase in the liberation of gases by the multiplication of cuts in the coal made during one shift or multiple shifts; and the additional electrical equipment at the faces, are the factors that cause the added importance to good ventilation in the permissible equipment area of gassy bituminous coal mines. Ventilation tor Mechanical Loading in Gassy Mines In gassy mines, the air currents must be maintained across the faces in such a way as to prevent short cuts to the returns and leaving pockets in which accumulation of gas might occur. To do this, two or more doors are required on almost every entry containing a track to the face, so located that in no event will a trip of cars block open the doors and permit the air to bypass some place or places on the air-approach side. This method frequently requires the use of unusual lengths of duplex cables to reach the mechanical equipment at the faces, particularly where the distances between parallel working places are 100 ft., as at the Nemacolin mine. Fig. I shows a pillar section being worked with mechanical loaders. This layout requires carrying the air in on the right and releasing it on the left, because the track clearance is on the right side and the line brattice is built on this side from the last crosscut to the face. Thus, the air travels to the face with greater velocity and gives more assurance of sweeping out the gas that might be exuding from the crevices of the coal. All working places, in driving through solid coal, are thus treated alike. They are all potentially gassy and must be swept clear of whatever may contaminate the air. Once the air starts into the working section, it is not released until it does a good job. Of course the volume of air must be large enough not only to dilute the gases but to carry away explosive fumes and fine coal dust quickly. Mechanical loading entails cutting and drilling of the coal during the same working period as the loading when two-shift and three-shift operations are employed, and all three processes create finely pulverized coal, which in itself is hazardous if not controlled by some means. If the ventilation is inadequate and gas accumulates, and a spark from defective electrical

Jan 1, 1942

By A. W. Hesse

VENTILATION of all types of coal mines is fundamentally the same, in that sufficient air must be provided to properly dilute and remove dangerous and obnoxious gases and leave the oxygen content of the air not under 19 per cent in all places where men and animals must work. Perhaps animals should not be mentioned, as the speed of mechanical loading leaves livestock behind and outmoded. Since this paper has to do with the influence of mechanical loading equipment on ventilation, the problems will be confined to the splits and the coal-producing places of the mine. The removal of coal dust from the air, or its rapid transmission to the return air courses; the possible increase in the liberation of gases by the multiplication of cuts in the coal made during one shift or multiple shifts; and the additional electrical equipment at the faces, are the factors that cause the added importance to good ventilation in the permissible equipment area of gassy bituminous coal mines. VENTILATION FOR MECHANICAL LOADING IN GASSY MINES In gassy mines, the air currents must be maintained across the faces in such a way as to prevent short cuts to the returns and leaving pockets in which accumulation of gas might occur. To do this, two or more doors are required on almost every entry containing a track to the face, so located that in no event will a trip of cars block open the doors and permit the air to by-pass some place or places on the air-approach side. This method frequently requires the use of unusual lengths of duplex cables to reach the mechanical equipment at the faces, particularly where the distances between parallel working places are 100 ft., as at the Nemacolin mine. Fig. I shows a pillar section being worked with mechanical loaders. This layout requires carrying the air in on the right and releasing it on the left, because the track clearance is on the right side and the line brattice is built on this side from the last crosscut to the face. Thus, the air travels to the face with greater velocity, and gives more assurance of sweeping out the gas that might be exuding from the crevices of the coal. All working places, in driving through solid coal, are thus treated alike. They are all potentially gassy and must be swept clear of whatever may contaminate the air. Once the air starts into the working section, it is not released until it does a good job. Of course the volume of air must be large enough not only to dilute the gases but to carry away explosive fumes and fine coal dust quickly. Mechanical loading entails cutting and drilling of the coal during the same working period as the loading when two-shift and three-shift operations are employed, and all three processes create finely pulverized coal, which in itself is hazardous if not controlled by some means. If the ventilation is inadequate and gas accumulates, and a spark from defective electrical

Jan 1, 1941

By A. C. Losekamp, R. M. Fincel, J. B. Conway

TANTALUM like several other metals exhibits a great affinity for or reactivity with certain gases. Tested in atmospheres which are not completely pure this metal becomes contaminated by certain impurities (notably nitrogen, oxygen, and hydrogen) with a resultant change in various properties. Well-known are the drastic changes in hardness, ductility, and electrical resistivity. Also previously recorded are the distortions of the lattice1,2 accompanying the increased content of oxygen and nitrogen in tantalum. This pronounced affinity for certain gases along with the significant property variations thereby produced makes it necessary to observe additional experimental precautions in making various property determinations with tantalum. Some recent measurements of the linear thermal-expansion characteristics of tantalum have made this point quite emphatically. Failure to observe the greatest precautions in reducing the contaminating influence has led to some extremely distorted results. Not only does such operation lead to different heating and cooling curves during the thermal-expansion measurements, but upon cooling to room temperature the sample is found to have suffered a permanent (actually semipermanent since the specimen can be returned to its original length by a 3-hr degassing treatment in vacuum at 2300°C) elongation. In addition, subsequent heatings will not follow the initial heating curve unless the expansion measurements are based on the actual specimen length at the beginning of the heating cycle. These thermal-expansion measurements were made by positioning the specimen horizontally at the longitudinal midpoint of a 20-in.-long, 4-in.-diam resistively heated tungsten tube furnace. The specimens were in the form of 20-mil sheet formed into a sharp U-shape about 0.3 in. in width and 0.3 in. in depth. Fiducial marks in the form of 0.010-in. holes drilled at either end of the 2.5-in.-long sheet specimens were viewed by a paired filar micrometer telescope system mounted on an Invar bar on the top of the furnace. Tests were made in helium with temperatures measured by a Pt/Pt-lORh thermocouple to 1000°C; above 1000°C temperature was measured by means of an optical pyrometer sighted into a black-body hole drilled into the specimen-holding fixture in the vicinity of the specimen. Previous tests with black-body holes drilled in a 0.25-in.-diam rod specimen had confirmed the temperature uniformity between specimens and specimen-holding fixture in this furnace. A typical thermal-expansion curve for arc-cast tantalum in helium is shown in Fig. 1 for the case where no attempt is made to remove all the impurities from the cover-gas atmosphere. The initial heating curve appears smooth and has a shape characteristic of all metals. However, it is known that the sample begins absorbing impurities at temperatures above 1000°C which are sufficient to distort the lattice and lead to specimen lengths which are somewhat larger than would be obtained in the absence of contamination. During the cooling cycle the expansion curve is seen to be significantly different from the heating curve. This is due to the impurity content causing a semipermanent (i.e., permanent until the impurities are removed by vacuum degassing) deformation which will not allow the speci-

Jan 1, 1965

By James A. McGurty, Walter J. Koshuba, Samuel G. Gordon, Gilbert E. Klein

ONE of the problems encountered in working with metals at elevated temperatures is the instability resulting from solid-solid diffusion at a -common interface. A determination of the nature and magnitude of this phenomenon is important as related to structural properties of materials potentially useful in high temperature applications. This paper is concerned with the solid state reaction between metallic molybdenum and beryllium at elevated temperatures. In the older literature,' there were no known inter-metallic compounds in the system Be-Mo, and the list of beryllium intermetallics was rather small. New compilations have been made by Smithells' and by Taylor," listing all the then known beryllium binary intermetallic compounds. In addition to these, some new phases have just been reported' in the system Zr-Be and some further compounds of the type MBe,3 have been currently described. The new compound MoBe,, which is reported upon in this paper differs considerably from the above mentioned MBe13 group, and will be discussed in detail. Diffusion Study The nature of the solid diffusion at a Mo-Be interface was determined by sealing a beryllium rod inside a molybdenum capsule and heating at a temperature of 2000°F' (1100°C) for 100 hr. The molybdenum capsule was fabricated from % in. swaged molybdenum rod obtained from the Fansteel Metallurgical Corp. The beryllium metal was obtained from Brush Beryllium Corp. as % in. cold-drawn rod. This material assayed 99.05 pct Be and contained as impurities 0.84 pct Be0 and 0.19 pct Be2C. The molybdenum capsule was % in. in OD x 2 in. long. To prepare it, a concentric hole 0.250 in. + 0.001 in. — 0.000 in. in diam was drilled and reamed in one end of a rod to a depth of 1 '/2 in. A 1 in. long beryllium rod 0.250 in. + 0.000 in. — 0.001 in. in diam was pressed into the molybdenum capsule. A molybdenum plug 0.250 in. & 0.001 in. in diam x 9/16 in. long was then used to seal the beryllium in place. The capsules were placed in a molybdenum resistance furnace and heated at 2000°F (1100°C) for 100 hr. On cooling, the specimens were cut and examined metallographically. Fig. 1 is a micrograph at X250 of a portion of the interface from a transverse section. Four phases are clearly visible, with sharp boundaries between them. Phase A is a section of the molybdenum capsule; phase B is the molybdenum-rich intermetallic composition MoBe,; phase C is the beryllium-rich compound MoBe13; phase D is a section of the beryllium rod. All phases were readily polished to a high metallic luster. The molybdenum-rich phase has a pronounced purple tinge. The beryllium-rich phase (C) is only slightly darker than the pure beryllium. In polarized reflected light, however, it shows striking birefringence between crossed Nicols. The black areas represent flaws or pores in the material. A small representative amount of powdered material from each of the phases was carefully removed under the microscope with a diamond tool. With this material, powder X-ray diffraction patterns were made for the identification of each phase. Table I lists the X-ray diffraction spacings and corresponding indexes characteristic of these intermetallic compounds. Synthesis To establish the formula of the new intermetallic compound, X-ray diffraction patterns were made from powders obtained from solidified melts of known composition. The syntheses were made by

Jan 1, 1952

By P. Niessen, S. J. Bercovici

RECENTLY Aust et al.&apos; proposed a model of non-equilibrium grain boundary segregation based on a vacancy gradient induced uphill diffusion process of solute to grain boundaries. According to their proposal, solute atoms which are attracted to vacancies would, during the cooling of a sample, migrate with vacancies toward grain boundaries (vacancy sinks) and thereby enhance the solute concentration in boundary regions. For solute atoms which are not attracted to vacancies no such boundary solute enrichment would be expected. These concepts were given experimental support by boundary and bulk mi-crohardness measurements of various thermally treated dilute zinc and lead alloys and also doped aluminum oxide. The underlying assumption in their work was that solute enrichment in boundary regions is indicated by higher microhardness values. From thermodynamic considerations1&apos;2 it appears that the solid-liquid solute distribution coefficient (K), defined as the concentration of solute in the solid to the concentration of solute in the liquid at equilibrium, indicates whether attractive or repulsive forces exist between vacancies and solute atoms. K < 1 indicates a solute-vacancy attractions, K > 1 indicates no, or a repulsive, vacancy-solute interaction. This note is to report on observations of different grain boundary etching phenomena in high purity zinc alloys. These phenomena appear to depend on whether a K < 1 or K > 1 solute is present. In experiments designed to measure directly the solute enrichment in grain boundary regions, samples of zone refined zinc with 0.1 at. pct solute additions were cooled from 380°C at a rate of 3°C per min. A surface layer of approximately 50 was then removed from the samples by chemical polishing. The samples&apos; surfaces were then protected by a chromating treatment such that in a subsequent etch only grain boundary regions were dissolved. The dissolved regions were approximately semicircular troughs, 10 µ in width. The grain boundaries were usually in the middle, i.e., at the bottom, of this trough. For zinc containing the K < 1 solutes cadmium and aluminum, the grain boundary was always indicated by a ridge as shown in Fig. 1. For pure zinc, and zinc containing the K > 1 solutes copper and gold, a groove at the bottom of the trough indicated the boundary position, Fig. 2., To test if these different grain boundary etching behaviors were possibly connected with the segregation process discussed above, the Zn-Cd alloy was again heated to 380°C and quenched into brine. After surface protection and etching it was found that this specimen no longer exhibited the grain boundary ridges but rather grain boundary grooves as shown in Fig. 2. It would appear, therefore, that the boundary ridge is a consequence of a process occurring during the slow cooling of the samples. From absorption spectrophotometric analysis it was found that for slowly cooled samples containing K < 1 solutes, the solute/zinc ratios of the material from the boundary trough were consistently considerably higher than the solute/zinc ratios of the bulk alloys. This boundary region solute enrichment was not observed in quenched samples. It is concluded therefore, that the boundary ridge shown in Fig. 1 is directly or indirectly connected with nonequilibrium solute segregation.

Jan 1, 1970

By B. S. Gottfried

A generalized mathematical model is presented, which describes the thermal recovery of oil in linear systems with convective external heat loss. Three-phase fluid flow, conduction-convection heat transfer, chemical reaction between oxygen and oil and aqueous phase change are included, making the model applicable to a variety of thermal recovery processes. The complexity of the model requires that numerical solutions be obtained with a digital computer. The numerical approximations of the partial differential equations are presented and the algorithm is discussed in detail. The analysis is applied to the detailed simulation of a forward-combustion laboratory tube experiment. The results of the simulation exhibit with reasonable accuracy all of the principal thermal and bydrodynamic characteristics observed in the laboratory (i.e., the propagation of a combustion zone, a steam plateau, and oil and water banks). Numerical variations in the parameters appearing in the external heat-loss term, the aqueous phase-change term, and the Arrhenius combustion term are examined, and their effects on the predicted temperature profiles, oil and water production histories, and the air-oil ratio are presented. In spite of an oversimplification in the chemical reaction mechanism and the considerable computer time required to obtain solutions, the model represents a major advance in the ability to simulate the phenomena observed in thermal recovery experiments. INTRODUCTION Thermal oil recovery refers to a class of recovery processes where heat is supplied to a reservoir to provide the necessary expulsive energy. This thermal energy can be supplied externally as steam or hot water, or it can be generated in situ by forward or reverse combustion. In either case, however, thermal recovery processes are charac- terized by the simultaneous flow of two or three fluid phases in a variable - temperature field, accompanied by possible chemical reaction or phase-change effects.2-5, 7- 10, 14,16-18,20-22 Although a physical understanding of the thermal recovery processes is far from complete, it is possible to construct mathematical models which describe approximately all of the principal physical and chemical phenomena. However, attempts to solve such models, even with high-speed computers, involve formidable mathematical difficulties. Consequently, theoretical solutions have been obtained only for idealized cases in which important physical phenomena are neglected. For example, consider the process of forward in situ combustion. All such theories which have been developed consider only certain aspects of the process, such as heat transfer, 2-4,8,9,l6,l8 heat transfer with phase change,10 heat transfer with chemical reaction,7 or the hydrodynamics of three-phase flow.21 A general theory including all of the above phenomena has not been developed to date. This paper presents a unified theory of thermal recovery processes in linear systems. A mathematical model is developed which explicitly includes conduction-convection heat transfer with convective external heat loss, chemical reaction between air and oil, aqueous phase change, and the hydrodynamics of three-phase flow. A system of equations is developed which can be solved numerically on a high-speed digital computer, resulting in predicted temperature, pressure, and saturation histories in space and time. The model allows a more detailed simulation of thermal recovery tube experiments than had previously been possible. THEORETICAL DEVELOPMENT Consider the linear flow of gas, water and oil in a homogeneous porous medium. Assume that the oil will react with gaseous oxygen, and that mass is transferred between the water and gas phase by evaporation or condensation. In the absence of hydrocarbon phase change, intraphase diffusion,

Jan 1, 1966

By H. R. Ogden

The notch sensitivity of titanium alloys is affected by impurity content, microstructure, and heat treatment. Using notch tensile properties to evaluate notch sensitivity, three commercial titanium-base alloys were investigated over the temperature range of -320o to 390°F. The alloys studied were Ti-5Al-2.5Sn, Ti-6Al-4V, and Ti-2Fe-2Cr-2Mo. Six interstitial levels were studied for each alloy in four mi-crostmctural conditions. The ff- 0 alloys were studied in both the annealed and heat-treated condition. All tests were conducted using 1/4-in. diam round tensile samples. The notched samples had a 50 pct notch with a notch sharpness (a/r) of 53, corresponding to Kt 6.3. Notch sensitivity was increased by increased interstitial content, low temperature, and acicular-type microstructures. The heat-treated, high-interstitial content - 13 alloys were more notch sensitive than the same alloys in the annealed condition. In addition, minima in the notch tensile-temperature curves were observed at temperatures near 32°F. These minima became more pronounced with increasing interstitial content. 1 HE increaseduse of titanium alloys in high strength applications at subzero temperatures requires that a better under standing of the low-temperature behavior of the alloys be obtained. At low temperatures, metals generally exhibit increased strength with lowered ductility and tohness. Notch sensitivity and the factors which influence it become more important as the strength level is increased and the temperature is decreased. Some of the factor s which affect notch sensitivity in ti- tanium alloys are composition, interstitial content, microstructure, and heat-treatment condition. In this paper, the results of a study of these variables on the notch sensitivity of three titanium alloys at temperatures from —320o to 390o are reported. The notch geometry selected for this program was chosen to provide a severe test of notch sensitivity. The ability of the metal to flow without initiating fracture under these severe conditions is reflected in the notch-strength values. Since notch strength is, in part, determined by the constraint exerted by the combined tensile stresses, the behavior of notch specimens reflects the properties of the material under combined stresses. For example, a material with a high notch-unnotch strength ratio may be expected to withstand the biaxial stresses encountered in a pressure vessel better than a material with a low ratio of notch-unnotch strengths. This implies that maximum strength alone is not a sufficient criterion for applications of this type. It further may be concluded that the various conditions that promote poor notch behavior also will affect adversely the performance of materials under combined stresses. In the alloys studied, these included low temperatures, high interstitial levels, and hasformed microstructures. EXPERIMENTAL PROCEDURES Three alloys, Ti-5Al-2.5Sn, Ti-6Al-4V, and Ti-2Fe-2Cr-2Mo, were purchased from commercial producers as 2-in.-dam rod stock. Oxygen and nitrogen were added to these alloys by remelting to obtain a variety of interstitial contents as shown in Table I. In order to Compensate for the much more potent strengthening effect of nitrogen compared with oxygen, an oxygen equivalent content is used to describe the total oxygen plus nitrogen contents studied. It has been suggested previously1 thatnitro-gen is approximately twice as effective in strengthening as oxygen, and carbon, to the extent it is in so-

Jan 1, 1962

G. D. Kneip, Jr., and J. 0. Betterton, Jr. (Union Carbide & Carbon Corp., Oak Ridge, Tenn.)—The authors have contributed to the theory of zone melting by considering the effects of the solidification of the final zone on the distribution curves for the finite length bar. However, they do not consider in sufficient detail the restrictions imposed by the particular metals involved. Eq. 4 has the analytical solution which is equivalent to Pfann&apos;s&apos; equation for normal solidification of the final zone. For distribution coefficients less than one, it is apparent from this form of the equation that, as the solidification approaches completion, the concentration attains extremely large values which are inconsistent with the density of physical materials. The necessary restrictions then are that the solute concentration cannot exceed the density of the alloy, or more frequently, the solubility limit in the solid phase. Secondly, the distribution coefficient cannot be constant from 0 to 100 pet solute unless the liquidus and solidus coincide over this whole region. It would thus be more correct to impose limits on the solubility such as would be indicated by a typical eutectic diagram. Furthermore, in this case the assumption that the distribution coefficient remains constant is more likely to be realized. Eq. 4 should then be replaced by the following C,,(x) =k/L-x [ ?11 3 C11 (x)Ddx - ?11 4 Cn(x)dx= k[Cn (L-l)/k] [ L-x/l]x-1 L-l = x = L-l [kCenterlie/Cn (L-l)]1/k-1 The typical shape for the distribution curve in the final zone is indicated schematically in Fig. 12 for the simple eutectic case. The solidification proceeds according to Pfann&apos;s&apos; expression until the concentration reaches the maximum solubility in the solid phasc. At this point, the concentration changes abruptly to the eutectic composition and remains at this value for the duration of the solidification. The width of the flat region, and the back reflection of this effect into the concentration curve on subsequent passes, depends upon the ratio of the eutectic concentration to the original concentration, upon the distribution coefficient, and upon the number of passes. Similar effects on the concentration curves in the first zone length would be expected in many systems for which the distribution coefficient is greater than one, and where the maximum solubility is not too much larger than the original concentration. The discussers agree with the authors that the assumption that the liquid is uniform in concentration should be considered with caution. Kneip and Better-ton&apos; have shown, however, that in floating zone refining of zirconium, using induction heating, the distribution of iron agrees with theory which assumes a uniform liquid composition. Hence, in this case, the experimentally realized distribution coefficient agrees with the present phase diagram within the experimental uncertainty. L. Burris, Jr., C. H. Stockman, and I. G. Dillon (authors&apos; reply)—The authors are happy to have received the comments of G. D. Kneip, Jr., and J. 0. Betterton, Jr., which provide additional insight into the zone refining process. It is true that the equations developed are inapplicable if the solute solubility in the solid phase is exceeded. However, this restriction on the use of the equations was clearly stated in the paper. A closer approach to

Jan 1, 1957

By C. D. Cochran, F. H. Deily, M. M. White, Victor Horner

A laboratory study of the effects of physical and easily measured rheological properties upon dynamic filtration was carried out. The purpose was to determine the factors affecting dynamic filtration and thus learn how to control this important property. The recommended procedure for determining the amount of fluid loss prescribed by API Code RP 29 "Field Procedure for Testing Drilling Fluids,"&apos; is very useful for mud conditioning purposes, but inadequate for comparing filtration properties of different muds under dynamic conditions. Where dynamic filtration control becomes important for purposes such as logging and coring, other means of evaluating filtration properties are needed. A special device and testing technique were developed to provide reliable data on filtration characteristics of different fluids under dynamic conditions. Liquid saturated rock samples were drilled with a microbit, and the amount of connate liquid displaced by the fluid being investigated was measured. The empirical testing procedure was considered reliable. A large number of laboratory-prepared and field-mixed dritling fluids with a wide range of physical properties were studied. These properties included: (I) shape of consistency curves, (2) apparent viscosity, (3) plastic viscosity, (4) yield value, (5) density, and (6) API filter loss. Test results indicate that very little, if any, trend could be established between changes in these fluid properties and dynamic filtration. INTRODUCTION During the process of drilling, excessive loss of liquid from the mud to the formation is undesirable for many reasons. Filtration can reduce production-zone permeability, cause sloughing of formations, affect electric log curves, and render the fluid data obtained from cores unreliable. Thus, it is very important that we gain a knowledge of the factors that control filtration. It is generally agreed that in any type of filtration (static, dynamic, or combinations thereof), the mechanics of filtering depend in great part upon the composition of the fluid. Moreover, it is also true that the physical and rheological properties are influenced by the composition of the fluid. Indirectly, then, it was hypothesized that a relationship existed between dynamic filtration and physical-rheological properties, two effects of a common cause—fluid composition. This paper discusses an attempt to relate filtration-while-drilling to physical and easily measured rheological properties. It is believed that the results of this study may prove useful to others in the industry who are working on the problem of understanding filtration. The industry accepted practice for determining fluid loss from drilling mud is described in API Code RP 29. The practice is to measure the amount of filtrate that passes during a given interval of time from a quiescent drilling fluid under applied pressure through a piece of filter paper. This is primarily a test of static filtration properties. Other investigators have studied filtration under dynamic conditions. Williams5 and Pro-kop5 studied radial filtration and filter-cake buildup by circulating fluid through a porous cylinder under conditions similar to actual mud flow. Nowak,4 Krueger,3,4 and Vogel3 studied the effect of filtration upon permeability by flowing mud past the face of rock Samples. In their tests, the bit action was simulated by mechanical scrapers that kept the filter cake from building up. Ferguson and Klotz2 studied three types of filtration: static, dynamic, and beneath-the-bit. They investigated each type separately, using a model well in which bottom-hole hydraulics and mechanics were closely approximated.

Jan 1, 1958

By C. F. Floe, M. Cohen, M. B. Bever, V. G. Paranjpe

NITROGEN is becoming recognized as one of the important elements in ferrous physical metallurgy. Several investigations indicate that nitrogen plays a significant part in such phenomena as strain aging and temper brittleness. Nitrogen is known to be a powerful austenite stabilizer and may be a useful alloying element in stainless steels. The hard-enability of steels has been found to be increased by nitrogen. The nitriding and carbonitriding pro- cesses are examples of successful industrial application of nitrogen in ferrous materials. For a complete understanding of the various roles played by nitrogen in ferrous metallurgy a knowledge of the phase relations in the binary iron-nitrogen system is essential. The literature contains a variety of iron-nitrogen phase diagrams. Those suggested by Fry, Sawyer,² Epstein and his coworkers,³,4 Lehrer,5 Eisenhut and Kaupp, and Hägg7,8 differ significantly from one another. Moreover, the most recent of these investigations dates back about twenty years. A re-determination of the entire diagram, therefore, appeared desirable. In addition to the published phase diagrams, investigations have been carried out on portions of the iron-nitrogen system. Bramley and Haywood determined the composition and temperature of the iron-iron nitride eutectoid. The maximum solubility of nitrogen in a iron was investigated by Portevin and Seferian10 and by Diikstra. 11 Several investigators have determined the crystal structure of the various iron-nitrogen phases. In fact, Eisenhut and Kaupp6 and Hägg7 based their entire phase diagrams on the results of such studies. Epstein&apos; and Emmett12 reported lattice-parameter measurements on the E (hexagonal close-packed) phase. Osawa and Iwaisumi13 also studied the change in lattice parameters of the a, ? and E phases with nitrogen content. Brill" and Hägg7 showed that the y&apos; phase has an ordered structure, and Hendricks and Kostingsl5 suggested a structural relation between the E and phases. In a recent investigation Jackl6 arefully studied the super-lattice reflections from the y&apos; and phases. Experimental Procedure A large variety of experimental methods has been used by previous investigators in studying the iron-nitrogen system. These methods include thermal analysis, microscopic examination, X ray diffraction, magnetic measurements, dilatometric analysis and anelastic experiments. Thermal analysis is of low accuracy in determining changes in the solid state. This accuracy is further decreased for iron-nitrogen alloys since composition changes will occur during heating or cooling unless extreme care is taken to use a proper atmosphere. Metallographic methods are of small value . because homogeneous iron-nitrogen alloys of sufficient size can be produced only with great difficulty. The dilatometric method is open to the same objections as thermal analysis in that changes in composition may occur during heating and cooling. Lattice parameter measurements by X ray diffraction can be used satisfactorily for determining phase boundaries if cooling is fast enough to prevent structural changes. In view of the limitations of the various other techniques two methods only were adopted in this investigation, (1) X ray diffraction measurements, and (2) a controlled nitrogenizing method. The second method is an innovation for investigating the iron-nitrogen system, and will be described at length. Materials: All iron-nitrogen alloys were prepared from carbonyl iron powder of the analysis shown in table I. The average particle size was about 20

Jan 1, 1951

By Kun Li, Alex Simkovich

The rate of dissolution of alumina in carbon-saturated liquid iron has been studied experimentally in a system where alumina was in the form of a cylindrical rod immersed in an iron bath contained in a graphite crucible. Data obtained consisted of the concentrations of aluminum in the melt as a function of time. In the case of static experiments, the data are shown to agree with theoretical prdictions based on the diffusion of aluminum.. The rate of dissolution was greatly increased by the rotation of the alumina rod. It is concluded that the diffusion of aluminum from the alumina/metal interface is the rate-controlling step. In the past, thermodynamic investigations of systems encountered in ferrous process metallurgy have received widespread attention. More recently, considerable work has been devoted to the study of kinetics associated with these systems in an effort to determine their rate controlling mechanisms. The alumina-iron system is of great importance in ferrous metallurgy. Yet information concerning kinetics of reaction in this system is seriously limited. The present study was made in order to establish the rate-controlling step for dissolution of solid alumina in liquid iron. LITERATURE REVIEW A number of papers concerning dissolution of solid metals in liquid metals have been reported in the literat~re. Generally, for these simple systems, dissolution is controlled by mass transfer of the dissolving species. Complex systems involving dissolution of solid metal carbides and oxides in liquid metals and slags have been studied to a much lesser extent. Skolnick5,6 reported on the reaction between liquid cobalt and poly-crystalline cylinders of tungsten carbide, in which the cylinders were dissolved while being rotated about their longitudinal axes at various speeds and temperatures. As a result of unexpected preferential grain boundary attack by the liquid cobalt, large errors in the measured dissolution rates occurred because of loss of tungsten carbide grains to the liquid cobalt. Nevertheless, it was possible to establish that the liquid Co-W carbide reaction was not controlled by mass transfer. In a similar approach, cooper7 was able to show that artificial sapphire rods, (alumina single crystals) dissolving in lime-alumina-silica slags obeyed a mechanism of mass transfer control. Here, again, the rods were rotated at various speeds and temperatures, and the process was followed as a function of these variables. Forster and Knacke8 took a practical approach to reaction between slags and refractories. By blowing argon through refractory cylinders of silica, silli-manite, or dolomite and directing the gas to rise along the slag-refractory interface, it was possible to increase the rate of mass transfer. Although the method was admittedly crude, it nevertheless permitted an evaluation of the relative stabilities of refractories with respect to slag attack. Data were interpreted on the basis of mass transfer control. EXPERIMENTAL TECHNIQUE Apparatus. An illustration of the apparatus used in this study is shown in Fig. 1. The furnace consisted of a Morganite recrystallized alumina tube wound with a molybdenum coil. A secondary molybdenum heater was mounted around the upper half of the primary coil to aid in controlling the thermal gradient within the furnace. The primary heater tube was 3 in. in ID and 30 in. long. A reducing mixture of 95 pct N and 5 pct H was maintained around the heating elements. Thermal insulation was provided by alumina powder. The chamber within the primary combustion tube contained a boron nitride block near the top to assist in controlling the thermal gradient to the furnace and also to provide a bearing surface for the rotating graphite shaft. The outside diameter of the graphite shaft was $ in. A separate threaded graphite specimen holder was screwed into the end of the shaft. The holder contained a tapered hole drilled into the end to guide the oxide specimens as they were pressed into it for mounting. Additional guidance for the rotating graphite shaft was furnished by a water-cooled bronze bushing attached to the top of the furnace. A steel clamp was fastened to the upper end of the graphite shaft and rested on a thrust bearing; the shaft and clamp were driven by a dc motor through a set of gears. Two O-rings located immediately above the bronze bushing maintained a gas-tight seal about the graphite shaft. The lower half of the alumina tube housed the crucible and charge, which were placed on a 3/4-in. diam movable alumina support tube. With this arrangement, charges could be inserted into or removed from the furnace while the hot zone was maintained at or above 1000°C. To control the temperature of the furnace, the thermocouple was mounted inside the support tube and in contact with the crucible bottom. Stray electric fields in the furnace were of sufficient intensity to cause erratic indications by the thermocouple. By enclosing the thermocouple protection tube in a molybdenum sheath and grounding this shield, the problem was eliminated. Output of the thermocouple went to an automatic continuous balance controller. Procedure. A typical run was as follows. First, electrolytic iron was premelted in graphite crucibles and cast into graphite molds with the same configura-

Jan 1, 1970

By B. C. Wonsiewicz, G. Y. Chin

THE presence of deformation bands in cold-worked metals is known to influence the development of deformation textures1 but has not been considered in any theoretical treatments. At the present time the most rigorous theory of deformation textures appears to be based on Taylor&apos;s concept of homogeneous deformation of a crystal within a polycrystalline aggregate.&apos; For each crystal the Taylor theory allows the operation of several equivalent combinations of five or more slip systems. Hence deformation bands can arise from the operation of a separate combination of slips in each of several portions within a crystal. On the other hand, a second type of band may occur as a consequence of slip on systems other than those selected on the basis on homogeneous deformation. This type appears to be associated with crystals of symmetrical orientations, such as the [loo] and [lll.] axial orientations in drawn wires of fcc single crystals studied by Ahlborn.3 In this note we propose a simple mechanism of deformation banding of the second type, and compare the theoretical results for wire drawing of [loo] and [Ill] with Ahlborn&apos;s observations. In outlini the model requires that the work to accomplish a given overall shape change by banding be less than or equal to the work to accomplish the same change by homogeneous deformation. While the work by homogeneous deformation, WH = ?T ;dri, can be calculated exactly by the equivalent methods of Taylor2 or of Bishop and Hill,4 the work by banding, WB, is less certain. The latter consists of 1) Wi, work by slip within a band, 2) Wb, work associated with formation of a boundary between bands, and 3) Wc, work to correct the difference in shape change arising from banding and homogeneous deformation. The Wi term can be evaluated explicitly for a band with an assumed set of slip systems, but a satisfactory analysis of Wb and Wc has not been obtained. Since WB = Wi + Wb + Wc WH, for banding to occur, however, a necessary condition is Wi = WH . Hence it seems reasonable to choose a band with the lowest value of Wi among competing bands consistent with Wi = WH, and then minimize Wb and Wc by arranging the physical configuration of the selected band. Wc is minimized by combining neighboring bands such that their net strain matches the macroscopic strain as closely as possible, and Wb is minimized by enforcing strain compatibility across the boundary between two neighboring bands, similar to the case of a grain boundary.5 It is realized that this procedure is not completely satisfactory, but it does seem to explain Ahlborn&apos;s observations reasonably well. The procedure is now outlined for the two corner orientations. a) [l00]. Fig. 1 shows a (100) stereographic projection. The results of the Taylor analysis for axi-symmetric flow by homogeneous deformation6 show that, while [loo] itself is unstable, the rotation of the tensile axis will nevertheless be confined to within about 15 deg from [loo]. Alternatively, less work may be expended by banding with slip on a pair of systems with primary-conjugate relationship, e.g., (ll1)[10l.I + (li1.)[110]. The latter represents minimum Wi for all possible bands.* Since there are four symmetrically

Jan 1, 1970