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By P. de Vitry, Oliver Bowles

The conventional wire saw, introduced in the slate district of Pennsylvania by the Bureau of Mines in 1927, and used thereafter with remarkable success, consists of a three-strand steel cable having a diameter of 93/16 or 54 in. The equipment, methods used, and accomplishments have been described in previous reports.'-I Similar wire saws have been used extensively in Europe, in quarries and in shaping blocks and slabs. During the period 1932-1937, competition in European marble quarries was keen, prices were declining, and orders were not abundant. When orders were received, there was generally an urgent call for speedy delivery. Because of the diversification of demands, it was impossible to maintain in stock marbles of the color, size, and thickness called for, therefore the material generally had to be fabricated after the order was received. The wire saw enabled stone mills to fill such orders promptly. With skillful handling, 7/8-in. slabs could be cut accurately and rapidly with this equipment. It was found, however, that, as the three-strand wire became worn, the cutting speed dropped rapidly and the cut became less accurate. To overcome these difficulties the junior author of this paper, after much experimentation, developed a new type of wire consisting of a single ribbon-shaped strand, twisted as shown in Fig. I. To ensure straight sawing, the twist of the ribbon is reversed from right to left and vice versa every 25 ft. The reversal of the twist is shown in Fig. 2. This is not a new feature; it has been used to advantage with the three-strand wire both in Europe and America for several years. However, some difficulties arise in splicing the three-strand reverse-twist wire, and the claim is made that cutting is less accurate as the wire becomes worn. To make an endless belt, the ends of the single-strand wire are brazed together with Pyrophyllite is a hydrous aluminum silicate (A1²Si4O10(OH)²)1 that, occurs in both the foliated and the massive forms. The foliated variety resembles talc in that it has a greasy feel, a pearly luster, perfect basal cleavage, and usually is white, although in some deposits apple green, gray, brown, russet, and nearly black specimens may be found. The massive deposits of pyrophyllite yield pearl gray to light tan aggregates, some of which are so friable that they may be crushed with the hands into a fine gritty powder. Frequently specimens of the massive variety are studded with crystals of radial structure. The mineral has a specific gravity of about 2.7 and a hardness of 1 on Mohs' scale.' The chemical composition of commercial pyrophyllite reflects the mineralogy of the deposits, which, according to Stuckey12 contain pyrophyllite and quartz, together with chloritoid, sericite, and a few other minerals of negligible importance. The latter, he says, ('are noticed in but small quantities to the extent they might occur as accessory constituents of an igneous rock or as products of regional metamorphism or weathering. . . . Quartz is abundant everywhere except in the very best grades of pyrophyllite." Analyses of specimens from several deposits are listed in Table 1, showing that the silica content varies from 57 to 73 per cent and the alumina from 22 to 33 per cent. Oxides of iron and the alkaline metals are collectively present, with one exception, to an extent of about 1 to 1.5 per cent. Sources At present, North Carolina is the only commercial source of pyrophyllite, although the mineral has been found in small quantities in many parts of the world. The North Carolina deposits are in the "Carolina Slate Belt," which traverses the center of the state in a southwesterly direction. This belt varies from 8 to 50 miles in width. Two major deposits of pyrophyllite lie along its eastern boundary, in what is known as the Deep

Jan 1, 1942

By J. Lindbo, B. Vigeholm

IN transmission electron microscopy, specimen thicknesses quoted are frequently based upon either traditionally familiar values or a few evaluations representing the whole material. Taking aluminum in 100 kv TEM as an example, average and maximum values are widely accepted as being 2000 to 2500 and 6000 to 7OOOÅ, respectively. It is the purpose of this note to revise these figures and to point out the importance of evaluating individual specimens carefully. All evaluations have been based upon slip trace widths and lattice orientations as determined from diffraction patterns. In many cases where two or more slip systems have been operating the geometrical method described by Crocker and Bevis1 has been applied as a check. The foils were usually oriented to within 2 deg from the exact Bragg angle. Kikuchi patterns were used as indications of correct orientations but not applied in the evaluations. The tilt of the foil surface relative to the beam normal may with certain geometrical relations between the tilt axis and the operating slip plane lead to substantial errors.2'3 Where these geometries were inevitable the samples were omitted in the investigations. Various measures were taken in order to overcome ambiguities, e.g., comparison with neighboring grains, tilts, evaluations based upon two or three slip systems,' and comparison with extinction contour evaluations where s = 0 conditions were obtained. Errors caused by misinterpretation of the slip traces on a micrograph are easily avoided by considering the bright-dark sequence on opposite surfaces and by taking into account the tapering of the foil usually leading to a distinguishable divergence of the pair of traces. In the investigations described here the interpretation was further facilitated as the slip trace formation was observed on the microscope screen and sketches taken down. In a few observations slipping was so violent that the distinction of pairs was considered dubious and consequently given up. To assess average values, thickness evaluations carried out in two independent investigations were analyzed: I) 99.998 pct A1 extruded, cold-drawn, and recrys-tallized to grain sizes between 35 and 500 um were deformed in tension at room temperature. In order to determine the dislocation densities as a function of strain,4 specimens prepared by spark machining and electropolishing in a PTFE holder were analyzed in the electron microscope at 100 kv. Thicknesses were evaluated for 569 selected areas. Histogram I in Fig. 1 shows the results, the average thickness being 4900Å. II) Electropolished specimens prepared from 99.9999 pct A1 with magnesium contents from 0.005 to 0.5 pct rolled to foil and annealed were subjected to fission fragment irradiationa2 For volumetric dose and defect concentration determinations thickness evaluations were performed on 123 areas. Histogram II in Fig. 1 shows the thicknesses of these areas averaging 4250Å. Whereas in investigation I no selection related to thickness has been made, some thin as well as thick areas have been omitted from I1 for experimental reasons. This and the application of two different microscopes although of the same type may account for the difference in average thickness. Assuming a mean error of 10 pct for a single evaluation, the error of the average should be negligible. In order to investigate the maximum transmissible foil thickness in aluminum, specimens were prepared from single crystals with (111) and (110) orientations parallel to the beam, and from polycrystalline material. Maximum values of thickness that could be measured were given by the stopping of the dislocations. The visibility often extended far beyond the end of the slip trace, but with no means of knowing the

Jan 1, 1970

By E. J. Ripling, S. I. Liu

IT has been shown in a number of recent publications that much information on the mechanical behavior of metals can be gleaned by first deforming test specimens under one set of conditions, and then evaluating the effect of this deformation by means of a subsequent tensile test. A straining procedure of this type has been applied to study the flow behaviors of a number of metals.1, 2 In these studies the prestraining and testing were both tensile strains, while the straining temperature was varied between the prestraining and testing steps. These data indicate that there is a more or less general flow behavior exhibited by all metals which are pre-stretched at one temperature and then tested in tension at a second temperature. Prestraining a metal at a high temperature produces a lower yield strength during the subsequent low temperature strain than would the same total strain at the lower temperature. On reversing the order of the strains, low temperature prestraining produces a higher yield strength during the subsequent high temperature straining than would a single high temperature strain of the same magnitude. Unlike the flow behaviors, the fracturing characteristics of metals when prestrained under one set of conditions and tested under a second set of conditions is quite material dependent. The fracture behavior of a number of steels2" as well as of a commercially pure zinc,4 have been investigated by these two-step tests and was found to depend upon the position of the transition temperature with respect to the prestraining and testing conditions. Since these fracturing phenomena are complicated by the appearance of the transition temperature, it has been difficult to reliably define the basic effect of a two-step straining procedure in fracturing. Metals crystallizing in the face-centered cubic system do not exhibit transition temperature effects so that the mechanical characteristics obtained on these materials could serve as a base for two-step straining behaviors. Consequently groups of specimens of the aluminum alloy 24s-T4 were prestrained various amounts at room temperature in either tension or compression after which these strained specimens were further strained to failure, or tested in tension, at either room or some lower temperature. Other groups of 24S-T4 specimens were prestretched at various low temperatures after which these were tested in tension at room temperature. These straining techniques then made it possible to compare the behaviors of aluminum with the behaviors previously presented for steels. Material and Procedure The commercial aluminum alloy 24S-T4 in the form of 3/4 in. rods was used in this investigation. The material was subjected to the following treatments before machining: 1—resolution treating at 920" & 10°F for 45 min, 2—quenching in water at room temperature, and 3—aging at room temperature for four days. The re-heat treated material was then machined to either of the specimen shapes shown in Fig. 1. Fig. la shows the threaded-end specimen used for prestraining in compression. Fig. 1b shows the specimen used for both prestrain in tension and final testing. The material subjected to precompression in the form of the specimen shown in Fig. la was re-machined to the specimen shape in Fig. 1b for final testing. Prestraining in tension and final testing were done without any intermediate machining. All straining was performed on a 10,000 1b Riehle tensile testing machine. Tensile strains were conducted in a specially designed concentric fixture which produced an eccentricity of less than 0.001 in." The change in diameter at the minimum section of

Jan 1, 1954

By J. B. Cohen

EARLY investigators1-4 reported that additions of beryllium to silver caused embrittlement. The in-termetallic phases which are now known to exist5,' were not known at the time. Solidification appeared to occur by a simple eutectic reation.4,7 (The primary beryllium-rich phase appeared globular, when the liquidus temperature was near the melting point of pure beryllium, and angular when the liquidus was much lower; in discussion of Sloman's work,4 it was suggested that this difference was due to an intermetallic compound which was not revealed by the polishing technique.) In the investigation reported here, it was found that if the compounds now known to exist were avoided, certain beryllium-silver alloys containing large amounts of the primary beryllium phase possessed sufficient ductility to be cold formed. Alloys containing 14 and 20 at. pct Ag were prepared from vacuum-cast beryllium containing 1.0 to 1.2 pct impurities and silver, 99.99+ pct pure. The elements were melted together by induction in beryllia or alumina crucibles under argon at less than 1 atm pressure. The cooling rate was adjusted by regulating the temperature of the melt and by reducing the power to zero without stopping the water flow in the induction coils, which were in contact with the crucible. Ingots were about 0.75 in. in diam with a grain size of 0.12 to 0.25 in. Microscopic investigation of transverse or longitudinal sections of an ingot cast from a temperature well above its melting point showed that, in general, the compounds did not form, presumably because of rapid cooling; the alloys consisted of beryllium-rich particles in a silver-rich matrix as shown in Fig. 1. A eutectic structure was not observed, probably .due to the low beryllium concentration of the eutectic composition and eutectic divorcement. Within a given grain, the small globules of the primary phase were all oriented in the same direction, as indicated by examination under polarized light; the orientation varied from grain to grain. This suggested that the globules were branches of the same dendrites. If cooling was too slow, a third constituent was observed in the microstructure; because it formed around the primary beryllium-rich globules it was Fig. 1—Ag-Be alloy, 18.6 at. pct Ag, as-cast. Bright field illumination. Etched with an aqueous solution of NH,OH and H2O2. X500. Reduced approximately 21 pct for reproduction. probably the 6 compound. A thin ring around an ingot often contained this constituent. When this compound was present in large quantities, or the silver segregated so that large regions occurred with the beryllium-rich globules in contact, sections from the ingot were brittle. When the compound was initially absent, it was possible to carry out reductions in thickness of 50 pct by cold rolling with only limited edge cracking; failure did not occur until reductions of 80 pct. A typical cross section after cold rolling is shown in Fig. 2. After simple bend tests, microcracks were found; these apparently were difficult to propagate in the two-phase structure, as the specimens continued to behave in a ductile manner after the cracks appeared. Similar cracks would probably be closed during rolling. It is apparent in Fig. 2 however, that the primary beryllium particles undergo extensive plastic deformation. Hardness values of the alloys indicated that some work-hardening occurred during cold working. Fig. 2—Ag-Be alloy, 18.6 at. pct Ag, reduced 83 pct in thickness by cold rolling. Bright field illumination. Etched with an aqueous solution of NH,OH and H2O2. X500. Reduced approximately 21 pct for reproduction.

Jan 1, 1961

By W. C. Knoblaugh

ROTARY kilns are adaptable to many fuels, but this paper deals principally with the use of powdered coal. The observations and conclusions presented are based on rotary kilns used in the manufacture of basic refractories for steel furnaces. Solid fuels are pulverized for two principal reasons: (I) ease of control, and (2) rapidity of combustion. By the use of pulverized coal, the operator can increase or decrease the fuel supply by small increments and with very little time lag between the change of control and the resultant change in the flame. This is extremely important in this type of furnace, as the temperature range for maturing the clinker is very narrow. By fine grinding, it is possible to increase the rate of burning of the coal, thereby liberating great quantities of heat within a small space in a short time. Under given conditions, temperature is a measure of the heat liberated per unit of time. To attain the temperatures of 2950° to 3150°F., which are encountered in the manufacture of various steel-furnace refractories, it is estimated that it is necessary to liberate 125,000 to I75,000 B.t.u. per cubic foot of the burning zone per hour. The manufacture of dolomitic refractories is a ceramic process and, as is true in most such processes, the material must be subjected to a certain critical temperature for a certain length of time in order to complete the reaction. If the time of reaction is reduced, the temperature must be increased; in other words, if the rate at which the material flows through the hot zone of the kiln is increased, the temperature of the hot zone must be increased to compensate for this shorter exposure to the flame. FIRING WITH PULVERIZED COAL Fig. I is a line drawing of a rotary kiln with the attendant equipment indicated in their relative positions. The kiln proper consists of a sloping, refractory-lined steel tube rotating slowly about its longitudinal axis (0.75 r.p.m.), and is countercurrent in operation in that the raw feed is introduced at the stack end and gradually approaches the zone of highest temperature at the lower end, where the fuel is introduced. Before 1930, the usual method of firing a rotary kiln was to store the coal in the pulverized form and then feed from this storage to the kiln as needed. This proved unsatisfactory and dangerous because of dust explosions and the inability to obtain an even feed to the kiln caused by arching of the coal in storage and flooding of the coal feeders. In later years the practice has been to store the raw coal and grind it as it is needed. Such a layout is indicated in Fig. I. The coal is stored in bins, charged from the top and drawn from the bottom into recording scales and then into some type of unit pulverizer. There are a number of good pulverizers on the market, each incorporating features that the manufacturer considers outstanding.

Jan 1, 1946

By W. M. Weigel

Salt production in Louisiana and Texas at the present time is entirely from salt domes of the interior and coastal groups. Before and during the Civil War salt was recovered from numerous salines and salt springs by evaporation in open kettles. The location of these in nearly every instance was near what is now known to be a salt dome, and the source of the salt in the salines and spring waters was that in the salt core. The tops of many of the shallower domes have been subjected to erosion1. The discovery that the salt occurred as large masses of rock salt in the form of domes was made in May, 1802, at Avery Island, La., by John Marsh Avery, grandson of the original owner, John C. Marsh, who built the old salt works in 18122. The first shaft for mining was sunk on the Avery dome in 1867 by Chouteau and Price. It was 90 ft. deep, about 58 ft. being in solid salt. It was abandoned in 1870. There are three general groups or fields of salt domes; interior Louisiana in northern Louisiana, interior Texas in northeastern Texas, and the coastal domes nearer the Gulf Coast and extending from the Mississippi River to the Rio Grande. This last group contains by far the greatest number of domes. As many of the coastal domes especially are prolific producers of petroleum and a few of sulfur, their geology has been intensively studied and fully covered in the literature. The depth to the top of the salt varies from near the surface, where shallow ground waters have dissolved the salt to just below recent deposits, to several thousand feet, where the depth has been cstimated only by geophysical methods. The diameter of the salt core varies from less than 1/2 mile to over 2 miles. Only domes where the salt is comparatively near the surface have been developed for commercial salt. Three deeper domes, which will be described later, have recently been developed or are in process of development, as a source of brine for chemical use. Fig. 1 shows the location of the domes so far developed for commercial salt and brine. Nos. 1, 2, 4, 7, 8, 9 and 10 are in the Gulf Coastal field,

Jan 1, 1938

By W. M. Weigel

Salt production in Louisiana and Texas at the present time is entirely from salt domes of the interior and coastal groups. Before and during the Civil War salt was recovered from numerous salines and salt springs by evaporation in open kettles. The location of these in nearly every instance was near what is now known to be a salt dome, and the source of the salt in the salines and spring waters was that in the salt core. The tops of many of the shallower domes have been subjected to erosion1. The discovery that the salt occurred as large masses of rock salt in the form of domes was made in May, 1802, at Avery Island, La., by John Marsh Avery, grandson of the original owner, John C. Marsh, who built the old salt works in 18122. The first shaft for mining was sunk on the Avery dome in 1867 by Chouteau and Price. It was 90 ft. deep, about 58 ft. being in solid salt. It was abandoned in 1870. There are three general groups or fields of salt domes; interior Louisiana in northern Louisiana, interior Texas in northeastern Texas, and the coastal domes nearer the Gulf Coast and extending from the Mississippi River to the Rio Grande. This last group contains by far the greatest number of domes. As many of the coastal domes especially are prolific producers of petroleum and a few of sulfur, their geology has been intensively studied and fully covered in the literature. The depth to the top of the salt varies from near the surface, where shallow ground waters have dissolved the salt to just below recent deposits, to several thousand feet, where the depth has been cstimated only by geophysical methods. The diameter of the salt core varies from less than 1/2 mile to over 2 miles. Only domes where the salt is comparatively near the surface have been developed for commercial salt. Three deeper domes, which will be described later, have recently been developed or are in process of development, as a source of brine for chemical use. Fig. 1 shows the location of the domes so far developed for commercial salt and brine. Nos. 1, 2, 4, 7, 8, 9 and 10 are in the Gulf Coastal field,

Jan 1, 1938

INTRODUCTION The regional characteristics of porphyry copper deposits in South America southward from Pantanos and Pegadorcito, Columbia, will be summarized. The age of formation of deposits spans the period from the Permian to the present. Where possible, the general characteristics and the individual structural features of the deposits are integrated into the plate tectonic model. Individual deposits cited in this chapter, located in Fig. 1, are the best known and most fully described, and form the Andean copper belt (Peterson, 1958). Smaller porphyry copper deposits fall both east and west of the line of large, high grade deposits (e.g., Chuquicamata and Toquepala) that make up the spine of this belt. Most of those smaller deposits (e.g., Mi Vida [Koukharsky and Mirre, 19761) are omitted from Fig. 1 and Tables 1 and 2, however. Neither the Andean orogen nor the individual deposits within it are as well de- scribed in the literature as are their counter- parts in the Cordilleran orogen of North America. Therefore, generalizations relying on regional geologic mapping, geophysics, and isotopic geochemistry lack details pertinent to this type of deposit in the Cordilleran orogen. The South American porphyry copper province has been described by Hollister ( 1974) to include most of the Andean orogen excepting the tin province of Bolivia. Porphyry copper deposits have not been commonly found within the tin belt, although copper-tin (e.g., San Rafael, Peru) and copper-tungsten (e.g., Llamuco, Chile) de- posits are known. James (1971) shows the depth to the Moho to be appreciably greater where the tin belt exists, but disclaims any genetic significance from the coincident appearance of the tin-tungsten province and the inhibition of porphyry copper development where the sialic crust thickens. Dates on tin mineralization in this belt (Clark and Farrar, 1973) are similar to some dates for porphyry copper mineralization known else- where in the Andean orogen. It seems that some mechanism is needed to explain the geographic separation of porphyry copper- molybdenum and simultaneous tin-tungsten- molybdenum metallization within the same tectonic system and appearing at similar distances from the trench. The thickened crust of the tin belt is one of the few known differences. Tables 1 and 2 list some of the most important porphyry copper deposits explored to date, but do not include Antaminas, Peru. or a number of other large skarn deposits

Jan 1, 1978

By Frederick Vickery

THE discovery of oil on the Goleta anticline has extended the petroliferous areas of California 18 miles west of the previous limit at Summerland and 35 miles south of the fields at Santa Maria. In each of these localities the oil is found in the Miocene-Pliocene strata, but as the oil at Goleta is found in the Sespe formation, which is generally considered as Oligocene or lowermost Miocene, the Goleta field extends the production of that horizon 52 miles west of South Mountain, the previous western outpost. Two horizons of high-gravity, paraffin-base oil have been found at shallow depths, and the exploratory well, Goleta No. 1, is now being drilled to test the productivity of an expected horizon in the Meganos formation. The Goleta structure is 12 miles west of the city of Santa Barbara and 2 miles north of the coast. Although it is within part of the old Spanish Rancho, Los Pueblos, it corresponds to T. 4 N., R. 29 W., S. B. B. and M., Santa Barbara county, California. The discovery well, Goleta No. 1 was drilled by the Miley Petroleum Ex-ploration Co. on the Tecolote ranch in the part of the Spanish grant corresponding to the sw. 1/4 of the nw. 1/4 of Section 3. Despite the apparent isolated position of the Goleta field, it is within the Ventura basin. The idea that the Ventura basin terminates at the ocean beach near the city of Ventura is neither sup-ported by the geologic structures nor the sequence of strata west of that city. This basin at the present time is known on the land as the Santa Clara valley and beneath the sea as the Santa Barbara channel. It extends westward to Point Concepcion. The Ventura basin, before the Miocene and Pliocene epochs of deformation, was much broader than at the present time. It is bounded by two lines of uplift, the Santa Ynez range on the north and the Santa Monica mountains with their extension, the Channel islands, on the south. Both of these ranges trend east and west. The Ventura basin and the mountains that bound it are part of a remark-able east-west orogenic unit, which interrupts the pre-vailingly northwest ranges along the Pacific Coast.

Jan 10, 1927

By Roman Šejnoha, Karel Mazanec

A pronounced tendency for delayed fracture zoas observed in the martensite structure of low-alloy steels in the as-quenched condition. Cracks of predominantly intercrystalline nature nucleated and propagated at stresses substantially lower than the short-titne tensile strength of- the steels, and were observed even in hydrogen-free speciunens. The occurrence of delayed fractures in the as-quenched condition is interpreted as a supevposition of the effect of free dislocations generated during the martensite transforlnation, and of a weakening of the prior austenite grain boundaries caused by dynamic shocks exerted on the boundaries by growing martensite plates. AN important metallurgical problem, connected with the production and use of high-strength steels, is delayed fracture in martensitic structures. Such fractures occur in quenched steels as a result of prolonged loading with external stresses far lower than the short-time tensile strength of the steel. The same effect can be produced by sufficiently high internal stresses: it is well-known that quenched structures will in some cases crack intergranularly at room temperature if not tempered quickly enough even in the absence of any externally applied stress. In Troiano's papers1,2 on delayed fractures in high-strength steels, the fundamental role of hydrogen migrating under the influence of multiaxial stress towards the nuclei of cracks was found. Fig. 1 shows the typical relationship between stress and the time to fracture found by Troiano in hydrogen-charged high-strength steels. On the other hand, Shurakov 3,4 explains the realization of delayed fractures in martensite as due to a "quasi-viscous" behavior of the austenite grain boundaries immediately after the martensitic transformation. 1) MATERIALS AND EXPERIMENTAL METHODS Two low-alloy steels of commercial purity were used for the experiments reported in this paper. Their chemical composition is shown in Table I. The heats were melted in a laboratory furnace, cast to ingots, forged to bars of 3/4 in. diameter, and annealed. Machined test pieces were heat-treated as follows: Ni-Cr-Mo steel held at 900°C for 30 min and quenched in oil; Mn-Si-Cr steel held at 900°C for 30 min and quenched in water. For the delayed-fracture tests, notched cylindrical test pieces, Fig. 2, stressed in torsion, and unnotched cylindrical tensile-test pieces, 1/4 in. diameter, were used. The specimens were normally loaded with a constant static load 45 min after quenching for the torsional tests and 10 min

Jan 1, 1965

By L. G. Austin, R. R. Klimpel

This article demonstrates that the Gilvarry and Klimpel-Austin equations for the random fracture of solids are incorrect by deriving intuitively correct expressions for simple cases and showing that the foregoing equations do not reduce to right answers. It is the purpose of this article to point out serious errors in some existing treatments of the random fracture of solids. Gilvarry ' gives the equation where B(y,x) is the cumulative fraction by weight of particles which fall below size y when particles of size x are broken. The constants i, j, k contain shape factors and densities of flaws leading to propagating fracture surfaces. The authors believe this equation to be incorrect. Klimpel and usttin' give where r~ is the number of fracture planes originating from edge flaws (per L where L is the edge length of the original particle), rs is the number of fracture planes originating from flaws on the particle surface (per S where S is the surface of the original particle) and rv is the number of fracture planes originating from flaws within the particle volume. This equation is also incorrect. The failures of these equations will be demonstrated by deriving intuitively correct expressions for simple cases and showing that the Eqs. 1 and 2 do not reduce to the correct answer. In this process, the flaws in statistical reasoning inherent in the derivations of Eqs. 1 and 2 become apparent. THE KLIMPEL-AUSTIN EQUATION Eq. 2 was derived by considering a set number of fracture planes occurring spatially at random within a given particle. This case will be treated here in its simplest form: a line with 1, 2, 3, etc., fractures occurring at random. This is illustrated in Fig. 1, with the original length of the line as Vo. The breakage of a long, thin, brittle fiber is a corresponding physical case, with the volume of a fragment proportional to its length. Assuming independence of the occurrence of the fractures, the probability of any one fracture occurring in length dV is dV/V0. The probability of the first fracture occurring at a length V to V + dV is dV/Vo. For the case of only one fracture per particle, formation of a number-size distribution of the left-

Jan 1, 1969

The following specifications, proposed by a sub-committee of the American Society for Testing Materials, are still to be adopted and subsequently submitted to the International Association for Testing Materials. Standard Specifications for Pig-Iron and Castings. I. Analysis.—It is recommended that all purchases be made by analysis. II. Sampling.—In contracts where pig-iron is sold by chemical analysis, each carload, or its equivalent, shall be considered as a unit. At least one pig shall be selected at random from each two tons of every carload, and so as to fairly represent it. Drillings shall be taken so as to fairly represent the fracture surface of each pig, and the sample analyzed shall consist of an equal quantity of drillings from each pig, well mixed and ground before analysis. III. Allowances and Penalties.—III all contracts, in the absence of a definite understanding to the contrary, a variation of 10 per cent. of silicon, either way, and of .01 in sulphur above the standard is allowed. A deficiency of over 10 per cent. in the silicon, up to 20 per cent., and a further increase in sulphur up to .01 over the above allowance subjects the shipment to a penalty of 1 per cent. in the price for each element so affected. IV. Base Analysis OF Grades. —In the absence of specifications the following numbers, known to the trade, shall represent the appended analyses for standard grades of foundry pig-iron, irrespective of the fracture, and subject to allowances and penalties as above. Silicon Sulphur Grade. Per Cent. Per Cent. No. 1,...2.75 0.035 No. 2,.2.25 0.045 No. 3,........1.75 0.055 No. 4,.1.25 0.066 Cast-Iron Pipe and Special Castings. I. Description of Pipes.—The pipes shall be made with hub- and spigot-joints, and shall accurately conform to the dimensions given in Tables Nos. I. and 11. They shall be straight and shall be true circles in section, with their inner and outer surfaces concentric, and shall be of the specified dimeneions in outside diameter. They shall be at least 12 feet in length, exclusive of socket. For pipes of each size from 4-inch to 24-inch, inclusive, there shall be two standards of outside diameter, and for pipes from 30-inch to 60-inch, inclusive, there shall be four standards of outside diameter, as shown by Table No. 11.

Jan 1, 1905

By R. Kossowsky

BRIDGMAN has listed five possible transitions in bismuth at 25°C which were observed as volume changes. He was unable however to detect the last three transitions by electrical-resistivity changes. Kennedy and LaMori have verified the pressures at the Bi 1-11 and Bi 11-111 transitions. They suggested however that the 125-kbar transition previously reported by Bundy3 actually corresponds to the 90,000 kg per sq cm transition observed by Bridgman.1 Jura4 reported a well-defined drop in resistivity at 92 kbar obtained by a four-lead method. The Bi 111-IV and Bi IV-V transitions have not been reported in any of the above-mentioned references.'-' Bundy3 observed a small dent in the low-temperature boundary of the Bi Vll region which is consistent with the presence of a 111-IV boundary. Zeitlin and Bray-man5 were the only ones to report small but well-defined drops in resistance of bismuth at 44.0 and 63.8 kbar, corresponding to Bridgman's 111-IV and IV-V transitions, respectively. High-purity (99.99 pct) 0.020-in.-diam 0.5-in.-long bismuth wire was used to calibrate the Uniram tetra-hedral high-pressure apparatus used in the present work. The bismuth phase changes at 25°C were correlated with changes in electrical resistance as a function of pressure. The output of a sensitive bridge constructed of carbon resistors and kept at a constant temperature of ±°C during the measurements was fed to the Y axis of a MoseleyX-Y recorder and the X axis was used for time scans. Specimens were pressurized using a hydraulic press with maximum capacity of 200,000 lb. Load was applied in increments of 2000 lb and held for about 3 min. The resistance was recorded at the end of this time period. Good contact was established below 20,000 lb. Typical recorder charts are reproduced in Figs. 1 and 2 which show that resistance fluctuations were practically nil during each time scan. Well-defined discontinuities in resistance were observed at 58,000 and 82,000 lb, Fig. 1. Two small but definite "blips" are noticed at 92,000 and 122,000 lb, Fig. 2. At 156,000 lb another drop is observed which

Jan 1, 1967

By Pieter W. Greeff, Graeme W. Mitchell

INTRODUCTION This paper describes the underground environmental planning parameters developed for Pancontinental Mining Limited's Jabiluka uranium deposit. The deposit is located 230 km east of Darwin in the Alligator Rivers region of the Northern Territory, Australia. The deposit is situated within land subject to an Aboriginal land claim, adjacent to a National Park and within a proposed extension to that National Park, as well as being close to a long- established Aboriginal reserve. In addition, being the first large underground uranium mine planned in Australia, the environmental considerations, both on surface and underground are considerable. Pancontinental is committed to the highest practicable standard of environmental management in the operation of the mine. In addition, Government regulations and requirements are particularly stringent by world standards. The underground environmental problems are discussed later in the paper with emphasis placed upon radiation aspects, both gamma and radon daughters, and dust. Consideration has also been given to heat and humidity due to the location of the deposit in a sub-tropical region 12° south of the equator. The paper explains how these environmental factors influence the design and selection of stoping methods. GEOLOGY AND ORE RESERVES The Jabiluka Two orebody is completely overlain by flat-bedded Kombolgie Sandstone with a thin layer of surficial colluvial sands. The orebody occurs in an east-west striking folded succession of Lower Proterozoic metasedimentary rocks of the Cahill Formation. Uraninite and pitchblende constitute essentially all the uranium mineralisation and are contained in the four lithological series, as shown in Fig. 1. - ,Upper Graphite Series - Main Mine Series - Lower Mine Series 1 - Lower Mine Series 2 The succession at Jabiluka has been folded and forms an east-south-east striking, asymmetric syncline-anticline which dips south and plunges to the east. A major fault bisects Jabiluka Tw8 having a north-south trend and dipping about 80 to the west. A near vertical tourmaline granitic pegmatite dyke up to 30 m in thickness has intruded the Jabiluka Two orebody at the western end. Current drilled ore reserves for Jabiluka Two are 52.0 million tonnes at 0.39% U308. Included in these reserves are 1.1 million tonnes of gold ore at 10.7 g/tonne. A large proportion of the orebody can be mined in thick sections (over 80% of tonnage) either as single sections or by bulking a number of thin sections which are close together.

Jan 1, 1981

By J. E. Cannaday, R. J. Austin, R. K. Saxer

PREVIOUS investigators have determined activation energies and have postulated various controlling mechanisms for creep.'-' Recently Barrett et 01.I5 have suggested, as the result of their work with cadmium, that the apparent activation energies for creep (AHc) above 0.5 T, could be correlated by means of activation energies for self-diffusion (AHSD) and modulus of elasticity functions: AHC = AHSD - (5RT/E) dE/dT. The wide disparity of apparent activation energies previously obtained with nickel compared to those reported for other fcc metals and the desire to test the postulate of Barrett el al. prompted this study. The as-received 17.5-mil nickel wire used during this work assayed 99.87 pct pure, Table I. X-ray analysis indicated that the wire was in a stress-re- lieved state and metallographic examination revealed an equiaxed grain size of approximately 0.01 mm. Specimens degreased in acetone and cut to 1-1/2 in. lengths were tested in a cam-mechanized constant-stress creep machine. A vacuum of at least lo-5 mm Hg was maintained during all tests. Stresses were applied when thermal equilibrium was established. When a measurable (at least 0.5 pct) secondary creep was observed, the temperature was pulsed to a higher temperature. Pulsing time varied from 30 sec at 480°C to 2 min at 1300'C and above. Slopes of creep curves were determined after the initial acceleration in the strain rate had subsided and steady-state creep was achieved. To confirm the results obtained, additional tests were performed incorporating a temperature pulsing method suggested by sherby. In this method the load was removed during temperature pulsing and reapplied after thermal expansion had occurred. No appreciable differences in results were obtained. Tests were performed over a wide range of temperature (470" to 1310°C) and stresses (250 to 17,000 psi). The expression was used to calculate activation energies. In this expression T1 and T2 are the initial and final temperatures, and el and i2 are the creep rates at T, and T2. The variation of the apparent activation energies with the mean test temperature calculated by means of Eq. [I] are shown in Fig. 1. The mean test temperature (Tmean) was calculated from the relationship

Jan 1, 1967

By Howard B. Bradley, Charles F. Weinaug

The phase behavior of a naturally occurring hydrocarbon system whose critical temperature is near the reservoir temperature has been described. The same volume per cent liquid was observed for the firkt time at three different pressures for isotherms immediately below the critical temperature. The shapes of the isothermal equilibrium constant curves necessary to predict this phenomena are discussed and illustrated. INTRODUCTION The phase behavior of a number of hydrocarbon systems has been reported; however, very little data are available on natural occurring hydrocarbon systems whose critical temperature is near the reservoir temperature. This appeared to be the case for the fluid studied here; therefore the pressure-volume-temperature data for this system were determined for preqentation. METHOD The PVT data were obtained with equipment similar to hat described by Weinaug and Katz,1 consisting principally of a two-section, double-windowed gauge mounted in a conktant temperature air bath. The procedure followed was essentially that used by Katz and Kurata5 except that a cathetom-eter, reading direct to 0.01 cm, was used to determine the height of each extreme phase limit. The system studied was formed by recombining -.vapor and liquid samples from the first stage separator of the well. Portions of these samples were displaced with mercury into the cell which had previously been filled with mercury. A predetermined amount of the liquid sample was introduced, while an excess of vapor was admitted. The resulting mixture was brought to equilibrium at the condition of the separator during sampling. By adding mercury to maintain the pressure, enough vapor was displaced to give a gas/oil ratio within the cell equal to that produced by the separator at the time of sampling. The remaining material was then considered to be equivalent to the reservoir fluid. An analysis of material in the cell, computed from analysis of vapor and liquid samples and the separator gas/oil ratio. is given in Table I. Table I — Composition of Hydrocarbon System Component Mole Per Cent Carbon Dioxide 0.13 Nitrogen 0.76 Methane 53.91 Ethane 14.20 Propane 9.64 Iso Butane 1.25 Normal Butane 4.29 Iso Pentane 1.12 Normal Pentane 1.87 Hexane 2.72 Heptanes + 10.11 hloleculal. Weight of Heptanes + 178.50

Jan 1, 1951

By Howard B. Bradley, Charles F. Weinaug

The phase behavior of a naturally occurring hydrocarbon system whose critical temperature is near the reservoir temperature has been described. The same volume per cent liquid was observed for the firkt time at three different pressures for isotherms immediately below the critical temperature. The shapes of the isothermal equilibrium constant curves necessary to predict this phenomena are discussed and illustrated. INTRODUCTION The phase behavior of a number of hydrocarbon systems has been reported; however, very little data are available on natural occurring hydrocarbon systems whose critical temperature is near the reservoir temperature. This appeared to be the case for the fluid studied here; therefore the pressure-volume-temperature data for this system were determined for preqentation. METHOD The PVT data were obtained with equipment similar to hat described by Weinaug and Katz,1 consisting principally of a two-section, double-windowed gauge mounted in a conktant temperature air bath. The procedure followed was essentially that used by Katz and Kurata5 except that a cathetom-eter, reading direct to 0.01 cm, was used to determine the height of each extreme phase limit. The system studied was formed by recombining -.vapor and liquid samples from the first stage separator of the well. Portions of these samples were displaced with mercury into the cell which had previously been filled with mercury. A predetermined amount of the liquid sample was introduced, while an excess of vapor was admitted. The resulting mixture was brought to equilibrium at the condition of the separator during sampling. By adding mercury to maintain the pressure, enough vapor was displaced to give a gas/oil ratio within the cell equal to that produced by the separator at the time of sampling. The remaining material was then considered to be equivalent to the reservoir fluid. An analysis of material in the cell, computed from analysis of vapor and liquid samples and the separator gas/oil ratio. is given in Table I. Table I — Composition of Hydrocarbon System Component Mole Per Cent Carbon Dioxide 0.13 Nitrogen 0.76 Methane 53.91 Ethane 14.20 Propane 9.64 Iso Butane 1.25 Normal Butane 4.29 Iso Pentane 1.12 Normal Pentane 1.87 Hexane 2.72 Heptanes + 10.11 hloleculal. Weight of Heptanes + 178.50

Jan 1, 1951

By C. A. Siebert, G. Sandoz

Oil quenching of steel, in good commercial practice, is carried out using a great deal of agitation in the quenching bath. Many of the tests reported in the literature represent the results obtained on steels quenched into still oil. Scott1 describes three stages of cooling in liquid quenching baths, namely: "State 'A'— Vapor Blanket Cooling—In this, the first stage, a thin stable vapor film surrounds the hot metal. Cooling is by conduction and radiation through the gaseous film and is therefore relatively slow. "Stage 'B'—Vapor Transport Cooling— Termination of the 'A' stage is marked by wetting of the metal surface. Vapor forms copiously in bubbles and is carried away by gravity and convection currents to condense in cooler surrounding liquid. This is the fastest stage of cooling. "Stage 'C'—Liquid Cooling—The 'B' stage ends as the surface temperature of the metal approaches the boiling point of the quenching liquid. Vapor no longer forms, so cooling is by conduction and convection and the temperature difference is greatly decreased. This mode of cooling is slower than that of the 'B' stage." Mechanical agitation of the quenching medium will accelerate all three stages of cooling. The greatest effect, however, will be noted on stage ' B' because the removal of the bubbles from the surface of the steel will not be entirely dependent upon gravity and natural convection currents. The method employed in this investigation for evaluating the oils was the end quench test. The various oils were substituted for water as the quenching liquid and hardness determinations were made at various distances from the quenched end. Most of the tests were run using a 3/8-in. orifice with a 21/2-in. stream height. However, to check the effect of other velocities two additional conditions were used, namely: (I) A 3/8-in. orifice and I-in. stream height, and (2) a 1/2-in. orifice and a 31/2-in. stream height. In all cases the samples were suspended so that the quenched end was 1/2-in. above the orifice. Some tests were run on an NE 8740, .3,312, and a 4340 steel. The results obtained on these steels confirmed those obtained on the NE 9450 and have therefore been Omitted from this paper. The test bars were machined from 1/8-in. stock to I-in. diam by 4 in. long.

Jan 1, 1949

By G. Sandoz, C. A. Siebert

Oil quenching of steel, in good commercial practice, is carried out using a great deal of agitation in the quenching bath. Many of the tests reported in the literature represent the results obtained on steels quenched into still oil. Scott1 describes three stages of cooling in liquid quenching baths, namely: "State 'A'— Vapor Blanket Cooling—In this, the first stage, a thin stable vapor film surrounds the hot metal. Cooling is by conduction and radiation through the gaseous film and is therefore relatively slow. "Stage 'B'—Vapor Transport Cooling— Termination of the 'A' stage is marked by wetting of the metal surface. Vapor forms copiously in bubbles and is carried away by gravity and convection currents to condense in cooler surrounding liquid. This is the fastest stage of cooling. "Stage 'C'—Liquid Cooling—The 'B' stage ends as the surface temperature of the metal approaches the boiling point of the quenching liquid. Vapor no longer forms, so cooling is by conduction and convection and the temperature difference is greatly decreased. This mode of cooling is slower than that of the 'B' stage." Mechanical agitation of the quenching medium will accelerate all three stages of cooling. The greatest effect, however, will be noted on stage ' B' because the removal of the bubbles from the surface of the steel will not be entirely dependent upon gravity and natural convection currents. The method employed in this investigation for evaluating the oils was the end quench test. The various oils were substituted for water as the quenching liquid and hardness determinations were made at various distances from the quenched end. Most of the tests were run using a 3/8-in. orifice with a 21/2-in. stream height. However, to check the effect of other velocities two additional conditions were used, namely: (I) A 3/8-in. orifice and I-in. stream height, and (2) a 1/2-in. orifice and a 31/2-in. stream height. In all cases the samples were suspended so that the quenched end was 1/2-in. above the orifice. Some tests were run on an NE 8740, .3,312, and a 4340 steel. The results obtained on these steels confirmed those obtained on the NE 9450 and have therefore been Omitted from this paper. The test bars were machined from 1/8-in. stock to I-in. diam by 4 in. long.

Jan 1, 1949

By William Hutton Blauvelt

The type of gas producer in which the ashes are fluxed and run off as slag was among the very earliest made. Ebelmen built the first one in 1840 at Audincourt, France, only a year after the installation of the first gas producer of which we have record. Charcoal was used as the fuel, with blast-furnace slag and clay as flux. My interest in this type of producer began when as a boy I saw the fluxing producer at Chester, N. J., which was invented by W. J. Taylor, and described by him in 1881.' This producer was invented independently by Mr. Taylor to meet the difficulties he experienced in making producer gas for roasting sulphurous iron ores. It will be of interest to review briefly his description of his producer. It was built like a small blast furnace, having a hearth 24 in. in diameter and 24 in. high. The bosh angle was 25' from the vertical; the diameter at the bosh 4 ft. and at the top 3 ft.; the total height 12 ft. The producer contained one water-cooled tuyere 12 in. above the bottom, with a 1.5-in. nozzle. Depth of the coal above the tuyere, 6 ft. Mr. Taylor mixed the coal with from 30 to 40 per cent. of basic blast-furnace slag to flux the ash. Occasionally some limestone was also used, but never limestone alone, as the use of the furnace cinder gave a larger volume of slag and made it easier to maintain a proper fluidity. The producer was blown with a small Weimer blowing engine, delivering 300 ft. of air per minute, at from 1 to 1.5 lb. pressure. The slag was tapped every 2 hr. and was black and glassy in appearance. Broken or egg size anthracite was used. The fine sizes of anthracite could not be made to work. Mr. Taylor expressed the belief that bituminous coal could be used if not too fine, but no experiments were made with this coal. Runs were made of four weeks' duration, with no stops or changes. Mr. Taylor reported the following advantages: 1. Excellent gas, very uniform in quality. The high fuel bed permitted no air to pass unconsumed, and the gas was almost entirely free from carbonic acid.

Jan 1, 1914