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By L. D. Jaffe, D. C. Buffum, F. L. Carr, M. Goldman

IN an earlier investigation by Jaffe and Buffum,' a time-temperature transformation diagram for isothermal temper embrittlement of an SAE 3140 steel showed some suggestion of two overlapping "noses." These authors mentioned, as a possible explanation, that their specimens embrittled at 525°C were slightly softer than expected, perhaps because of some slight unknown variation in tempering treatment. Because a double nose would be of considerable importance in elucidating the mechanism of the temper brittleness reaction,' it appeared worthwhile to redetermine the portion of the diagram near 525°C. The same heat of SAE 3140 steel was used as in the previous work, and experimental procedures were the same.' Blanks were austenitized 1 hr at 900°C, water-quenched, tempered 1 hr at 675°C, and water-quenched. Groups were then embrittled isothermally at the temperatures and for the times shown in Table I, and water-quenched. For comparison, one series was tempered various times at 675°C, water-quenched, and given no further treatment. V-notched Charpy impact specimens machined from the blanks were tested over a range of temperatures. Energy absorption and percentage of fibrous fracture were plotted vs test temperature for each group, utilizing both the new data and those for the same treatment obtained in the earlier work,' except that the earlier data at 525°C were discarded.* are given in Table I. After a small correction (-0.28°C/BHN) for the variation of hardness" from Rockwell C24, the transition temperatures were plotted against embrittling time for each embrittling temperature, using the data of Table I for the range of treatments covered there, and using previous dataL for other treatments. By cross-plotting, Fig. 1 was finally obtained. It appears from this figure that if there are two noses to the time-temperature-transformation diagram for isothermal temper embrittlement of SAE 3140 steel, the upper nose must lie above 575°C. As interest had been primarily in the region near 525°C, a thorough study of embrittlement above 575°C was not made. Fig. 1 shows that such a study is needed, especially since preliminary measurements at the tempering temperature of 675°C reveal much more rapid embrittlement than would be expected from the data at lower temperatures. The transition temperature increased, approximately, as the logarithm of the tempering time (about 81/2°C, after correction for hardness, for each doubling of the time).& Since three tensile specimens tempered 240 hr at 675°C and water-quenched had an average yield-tensile ratio of 0.844, with 70.7 pct reduction of area, it is assumed1 that the tempering temperature is not above Ae1. The omission of a quench-and-reheat operation below Ae1, for the groups tempered at 675°C, as compared to groups embrittled at lower temperatures, should not affect results." More detailed investigation of embrittlement above 575°C is now under way.

Jan 1, 1954

By F. Weinberg

INTEREST is currently being directed toward grain boundary morphologies in zinc in terms of grain boundary facetingl and grain boundary energies.' Some years ago the present author attempted to measure grain boundary energies in 99.999 pct purity Zn as a function of boundary angle, using the standard procedure of measuring equilibrium angles of the boundary traces at triple points. This followed an investigation he carried out on preferential melting of |H (112) Fig. 1—Electron transmission micrograph of a stacking fault, grain boundary intersection in copper. GRAIN BOUNDARY STACKING FAULT Fig. 2—Schematic representation of the intersection of Fig. 1, indicating inclination angles 0 and trace angles $. specimens were prepared by the Bollman technique in an electrolyte of 70 pct methanol, 30 pct HNO3 main-tained at -30°C. The interfacial free energy of the high angle grain boundary and of the stacking fault is assumed to be thermodynamically equivalent to the corresponding in-terfacial tensions. In terms of the true dihedral angles & between the intersecting interfaces, the equilibrium of interfacial tensions in the direction parallel to the stacking fault yields: ysF = -ygb(cos F1+ cos ±cosF2) Here, ysF is the stacking fault free energy and ygb is the grain boundary free energy, assumed equal on both sides of the intersection. Forces acting normal to the grain boundary, of the form ??gb/?F,l are assumed negligible as the grain boundary remains straight for a considerable distance on either side of the intersection. The micrograph of Fig. 1 may actually consist of several closely spaced stacking faults on the same slip plane. The actual point of intersection between the grain boundary and the first stacking fault, and the equilibrium configuration of the intersection, should not be affected by the presence of the other faults, as

Jan 1, 1970

By A. H. Redfield

BITUMENS have been defined by Abraham1 as substances of variable color, hardness, and volatility; composed principally of 'saturated hydrocarbons, substantially free from oxygenated bodies, sometimes associated with mineral matter; the nonmineral constituents being fusible and largely soluble in carbon disulphide. The principal bituminous substances other than petroleums known to commerce are: (1) native asphaltites, such as gilsonite, grahamite, and glance pitch, distinguished by their hardness and relatively high fusing point; (2) native asphaltic pyrobitumens, such as wurtzilite, distinguished by their infusibility and comparative insolubility in carbon disulphide; (3) mineral waxes, such as ozokerite, distinguished by their high content of crystallizable paraffins; (4) native asphalts, containing varying amouqts of mineral matter, distinguished by the comparative absence of crystallizable paraffins; and (5) bituminous residues obtained from the distillation of petroleum. GILSONITE Gilsonite, a black native asphaltite 98 to 99.9 pct pure, characterized by a hardness of 2 (Mohs), a specific gravity of 1.01 to 1.10, and melting point of 230° to 400°F (K and S), is mined commercially only in northeastern Utah. It occurs in parallel, almost vertical veins, from a few inches to 18 ft wide and 8 to 30 miles long, that cut across the Tertiary limestone and shale of the broad Uinta Basin. The bitumen is believed to have been distilled by heat from oily material in the underlying Green River shale. At the Rainbow mine of the American Gilsonite CO.,† the gilsonite, formerly extracted in an opencut by a system of slopes and benches, is now mined by shafts and horizontal drifts. Little timbering is required. The mining methods of the Raven Mining Co., Chicago, and of the American Asphalt Association and the Utah Gilsonite Co., both of

Jan 1, 1949

By J. J. Jakosky

THE fundamental hydrodynamic principles governing the production of oil from wells have been carefully studied and evaluated by many investigators. These prior studies are quite complete and cover virtually the entire field of oil production and recovery.1 One of the most impor-tant measurements in production studies is that of the bottom-hole pressure at various rates of production.2 It is the purpose of this paper to outline the technique and economics of fluid-level measurements as applied to pumping wells and to review briefly other methods for obtain-ing bottom-hole data. The applications of bottom-hole pressure or fluid-level measure-ments are based on simple hydrodynamic principles. Under equilib-rium conditions, pbh = hd + pg + pch [ l] when pbh = bottom-hole pressure, lb. per sq. in. absolute, h = height of fluid in well, ft., d = average fluid density, lb. per sq. in. per foot, pg = weight of section of gas column 1 in. square from fluid level to casing head, pch = casing-head pressure, lb. per sq. in. absolute. In words, equation 1 states that the bottom-hole pressure is equal to the weight hd of the column of fluid plus the weight of the column of gas between the fluid and the casing head plus the casing-head pressure. At moderate rates of flow, the rate of production is approximately proportional to the difference between the reservoir, or static bottom-hole, pressure and the flowing or producing pressure; i.e.: Q = c(pr - pbh) [2] pr = the reservoir pressure, c = the constant of proportionality. However, at production rates approaching zero and also at high rates the proportionality may break down. By combining equations 1 and 2 the

Jan 1, 1939

By Charles Luner

ALTHOUGH the kinetics of the atmospheric oxidation of tin have been studied,1-3 the kinetics in pure oxygen have not been reported. This note presents some results of the kinetics of the oxidation of tin foil in pure oxygen. The tin foil used in this investigation was 99.78 pct pure, the two major impurities being Pb (0.17 pct) and Sb (0.024 pct). The foil was degreased and air-dried prior to oxidation. The air-formed film was estimated by coulometric reduction4 of the surface oxides to be about 25A. The samples, which were in strip form and had a total geometric surface area of about 300 sq cm, were first evacuated at room temperature in an oxidation bulb to 1 X 10 -5 mm Hg and subsequently degassed for 1 hr at the temperature of the experiment. Spectroscopically pure oxygen was then introduced into the bulb. The rates of oxidation were measured by following the rate of pressure change due to oxygen consumption in a closed system, the volume of which was sufficiently large so that the pressure change during an experiment was 10 to 15 per cent of the total pressure. An oil-mercury manometer,5 with a sensitivity of 3 X x mm Hg (which corresponds approximately to 5 X 10 -7 gm of oxygen) was used to measure the pressure change. The oxidation bulb was heated by an oil-bath and controlled to within -10.3 C. The reproducibility of the rates was good, as determined by the coincidence of rates at 190oC for duplicate experiments. Fig. 1 shows the oxidation curves in the temperature range 168 to 211.5oC at 8 mm Hg pressure. The effect of pressure on the rate of oxidation at 190°C is also shown. The rate of oxidation was not markedly affected by changes in pressure in the range 4 to 9 mm Hg, except perhaps at the lowest pressure, 4 mm Hg, where after about 70 min there appeared to be a decline in the rate, as compared with the rate at the higher pressures. At the lower tem- perature, 168oC, and for short reaction times, it is difficult to distinguish between a linear relationship and a slight curvature in the oxidation rate. However, for longer times, there was a somewhat larger indication of curvature. The samples after oxidation were gold-colored in appearance. The oxide that was formed at 170°C was identified as a-SnO by reflection electron diffraction. This is in agreement with results obtained by other workers.6, 7 The rate data best fitted the logarithmic-oxidation rate law, w = K1n ln(1 + -t/to) [1] where w is the weight of oxygen consumed, t is the time, and K In, and to are constants. First estimates to were obtained according to the method of Taylor Thon8 and were then employed to obtain the best simuItaneous estimates of to and K In, that would fit the logarithmic expression. These estimates were found numerically, so that the residual variance could be minimized. The calculations were performed on an

Jan 1, 1961

By Nicholas J. Grant, Bill C. Giessen

The phase diagram Ta-Ni has been treated repeatedly; investigations up to 1958 are summed up in Ref. 1. Since then, an equilibrium diagram has been presented by Kornilov and Pylaeva.2 They found the following intermediate phases: Ta2N; (CuA12 type),3 "TaNi" (p phase, Mo8Co7 type), TaNiz, and TaNi3 (Ticus type).5 Of these, the crystal structure of TaNin remained unknown, and no further literature reference was found. According to Kornilov and pylaeva,&apos; TaNil melts peri-tectically at 1420°C. A 5-g alloy of Ta + 66.7 at. pct Ni was prepared by inert-gas arc melting using 99.95 pct Ta rod (NRC-MG) and 99.9 pct Ni shot (Fisher). The weight loss after three melts was <2 mg; a microscopically homogeneous ingot resulted. As shown in Ref. 6, no oxygen or impurity pickup needed to be considered under these conditions. Specimens were annealed for 15 hr at 1350°C, 15 hr at 1205oC, and 15 hr at 875°C with subsequent powdering and stress relief for 10 min at 1350°C, 15 min at 1205oC, and 1 hr at 875oC, respectively. Micrographs and X-ray analysis (114.6 mm Norelco camera, filtered CuK, radiation) showed homogeneous one-phase alloys in the as-cast state and at all three temperatures, all having the same powder pattern except for broad lines in the as-cast state. From this it is concluded that either TaNi, melts congruently or its melting behavior is a transition between the eutectic and peritectic cases. Since it was found that very slight surface contamination on stress relieving can considerably alter the structure of similar phases (e.g.,TaPt3),7 patterns before and after stress relief were also compared; they were identical except for line broadening in the non-stressrelieved condition, and two faint impurity lines in the stress-relieved samples. The powder pattern is presented in Table I. All lines were accounted for by indexing with a tetragonal lattice, 0Yh- I4/mmm , with a, = 3.154 ± 0.001A, co = 7.905 * 0.002i and c/a = 2.506. The atomic positions are: 1000; 1/2, 1/2, 1/2] + 2Ta(a): 000; + 4Ni(e): *(002); z = 0.333 * 0.002 This proves TaNi2 to be a representative of the MoSiz (Cllb) type. The variable parameter was determined by grouping the visually estimated intensi- ties according to their 2 indices, plotting their F values, and varying z for best fit, which was obtained at z = 0.333= 1/3. Since absorption was not considered, only the relative intensities with similar 28 can be expected to agree. The interatomic distances are presented in Table 11; the standard error is *o.o2A, resulting chiefly from the uncertainty in z. These distances indicate a coordination number 10 for both kinds of atoms and show that the pseudohexagonal layers along (110) are rather un-distorted. The mean atomic volume of 13.1A3 is close to the Vegard&apos;s law value of 13.3A3, showing

Jan 1, 1964

By Herbert Everest

THE idea of economical production is usually associated with large operations, tonnages, and mines, with even larger capital behind them. Nevertheless many small mines operate in the shadow of large competitors and make a good showing on the capital invested despite larger overhead expenses. For the purpose of this discussion I will divide the cost of production into classes, and specify opposite each the approximate percentage expended thereon: 1. Labor, including miners and company men; 60 to 75 per cent. 2: Development, including all necessary yardage, room turning, crosscuts, etc.; 7 to 12 per cent. 3. Deadwork, covering payment to miners for handling falls, drawslate, or faults, rock work, water, and in general, all nonproductive labor; 3 to 7 per cent. 4. Supplies: mine timbers, oil, brattice material, lumber, cement, and repairs to equipment; 2 to 6 per cent. 5. Expense: management, selling, office, taxes, etc.; 1 to 5 per cent. 6. Depreciation of coal reserves and royalties; 6 to 10 per cent. 7. Depreciation of equipment and interest on capital invested; 1 to 3 per cent. 8. Fuel; ½ to 1 ½ per cent. The labor cost for a small mine is relatively much lower than for a large mine, particularly the sum paid to company men. Idle-pay expense is cut to a minimum; the labor item being larger than all the others together, a saving in labor makes a decided showing on the total cost. The development charges in small and large mines are about the same. There should be a small showing on the Beadwork item in favor of the small mine. The unit cost for supplies is usually small in the small mine. The operator of the small mine finds, owing to a limited tonnage, that the expenses of management, selling, etc., are abnormally high.

Jan 1, 1916

By W. C. Wright

A PRACTICAL program of education for workers of the mining industries is being formulated by the Federal Board for Vocational. Education in cooperation with the States in which this industry is a dominant one. The Federal Government last year passed a law, known as the Federal Vocational Education Act, that provides for the promotion of vocational&apos; education and for cooperation with the States in the promotion of such education in agriculture and trades and industries and in the preparation of teachers of vocational subjects. More than 1,000,000 men in the continental United States are employed in mines and quarries. The following table, from the-1910 Census Report, shows the relative importance of mining and manufacturing pursuits: Manufactures Mines and Quarries Employees 7,405,313 1,109,410. Capital $18,428,269,706 $3,380,525,841 Value of product $20,672,051,870 $1,238,410,322 Of the more than 1,000,000 men employed in mining, almost three-fourths are engaged in the, production of bituminous and anthracite coal; one-fifth, in the production of copper, iron, lead, zinc, and other metals; and the remainder in getting out structural and miscellaneous materials from quarries. The meeting of the needs of the coal-mining industry, therefore, is the problem of greatest importance in the point of numbers to be served. Moreover, the dangers in coal mining are greater than in metal mining. Practically all States have set up educational and technical standards , for the position of foreman, boss, superintendent, etc., and every applicant for these positions must have a certificate of competency. The qualifications, too, are becoming increasingly difficult in the field of technical information. In some States, every miner must have a certificate, which is issued by the mine examining board. only after an oral examination.

Jan 1, 1919

By Donald N. Rosenblatt

I would like to commend the author on his description of Pinto Valley&apos;s plant, and its interesting innovations. The author&apos;s discussion of 18 ft in diameter by 21 ft long primary ball mills (the largest wet ball mills made to date) is of considerable importance to those who have been engaged in the field of studying ball mill liner wear as it relates to various grinding installations. While it is stated that the mills have operated on a fixed tonnage and therefore it has been difficult to determine the best liner pattern, calculations from the data in the paper indicates that for mills 1, 2, 5, and 6, the tonnage per hour varies from a low of 223 for mill No. 1, up to a high of 314, for mill No. 2, while mills 5 and 6 calculate 285 and 246, respectively. Analysis of the wear patterns for the single wave liners used in mill No. 1 as compared to the double wave liners in mill No. 2, suggests that the author&apos;s data may tell more than he suspects about liner wear to be expected and thus what should be a best pattern selection from what may appear to be only limited hours of operation. [Fig. A] is a reproduction of author&apos;s [Fig. 10], to which wear forecast patterns [F-1 and F-2] have been added. Forecasts are based on many previous studies of large diameter ball mill operations over many years of field service. Calculations from the first-wear pattern provide an initial wear rate of only 13 lb of liner per 100 sq ft of shell liner surface per day, or 0.031 lb per to: ground. However, the second wear pattern is ac-

Jan 1, 1978

By John E. Willson, Ben L. Seegmiller

A portable electronic instrument was designed and constructed to detect unknown stress magnitudes in rocks. The principle used to detect stress is based on the propagation velocity method. This method allows the stress in a rock subjected to a load of unknown magnitude to be determined, provided that the wave velocity-stress relationship for similar rock is known. This is accomplished by measuring the travel time of longitudinal mechanical waves passing between two points a set distance apart in a rock. The velocity of the wave is calculated and the stress determined from the wave velocity-stress relationship. If the sending and receiving transducer spacing is constant, a time vs. stress relationship rather than the velocity-stress relationship may be used. The method is nondestructive and tests can be made without drilling or otherwise disturbing the rocks. The first studies undertaken in the United States to determine stress in rocks using propagation velocity techniques were reported by Obert1,2 in 1939 and 1940. The Soviet Union first reported using propagation methods to study rock pressures in 1951.8 Success of the method led to the development of a pulse-type ultrasonic seismoscope4 in 1953. Using this instrument, Ivanov and Betaneli5,6 in 1963 succeeded in devising and testing under field conditions a method of investigating coal pillar stresses. In 1967 Osipova7 reported results of similar studies in the Nakhichevan salt mines. In France, Tincelin8 has used the propagation velocity method to study the stress distribution in iron ore pillars. Uhlmann9 has investigated stresses in salt and potash pillars in Germany, using velocity techniques. Determination of uniaxial stress by propagation velocity methods is limited to rocks which have readily detectable wave velocity-stress variations. The present stress detection instrument is restricted in application to rocks which have a velocity change under stress of at least 300 fps. Examples of rocks which meet this requirement are sandstones, coal, and possibly some limestones. Testing of this instrument was limited to a laboratory study and the results may or may not be indicative of what would be found in a field test. A program of field experiments to study the feasibility of using this instrument to determine mine pillar and tunnel stresses is in progress. Instrument Design The instrument has two main components: A probe and a control-display unit. The probe is a hand-held device to which two identical rodlike transducers are rigidly mounted. Coaxial cable connects the probe to the control-display unit which is mounted in an enclosed carrying case measuring 7 % x 9 x 13 in. The instrument is designed so it may be carried and operated by one man. Weight of the probe and enclosed control-display unit is approximately 20.1b. The probe consists of two transducers identical in construction. One transducer is used to convert electronic pulses to mechanical waves and transmit these waves into a rock. The other detects the transmitted waves in the rock and converts them back into electronic signals. The basic element in both transducers is a piezoelectric crystal. The crystal is a disk made of lead titanate zirconate and has a natural resonant frequency of 400 kHz ±1% in the thickness mode. A schematic of one of the two identical transducing elements is shown in [Fig. 1.] A spacing of 6 in. between sending and receiving transducers has been found to be most satisfactory. The control-display unit consists of a pulse generator and a 1-in. oscilloscope. Various electronic devices are used for the power supply, amplification, and calibration. The amplitude of the square wave from the pulse generator can be continuously varied between 0 and +20 vdc. Pulse width can be set at 1, 2, 3, or 4 µsec. The repetition rate of the pulse can have the following values: 50, 75, 100, 150, 200, 250, 300, 400, 500, 750, and 1000 Hz. The oscilloscope delay system allows the travel time of the longitudinal wave passing through a rock to be measured to an accuracy ±0.1 µsec in the range from 0 to 995 µsec. Fig. 2 is a schematic of the various electronic sections in the control-display unit. Laboratory Testing The first step in using the instrument is to develop a velocity-stress or a time-stress curve for the par-

Jan 1, 1972

By Robert E. Hochscheid

The subject article appears in the February, 1978, edition of Mining Engineering, and on page 190 carries the statement, "As most operators know, in a Water-Only cyclone, ability to adjust the vortex finder is quite important due to the sensitivity of this type of operations to variance in feed conditions and having to take them apart to make this adjustment is always a nuisance." Our experience as a manufacturer and supplier of Water Only cyclones does not agree with the above statement, and we offer the following brief discussion. Our company finds that the sensitivity of the Water-Only cyclone is a function of the fineness of the coal being washed, with the sensitivity to vortex finder length being greatest for coals coarser than 3/4-in. and virtually insensitive to raw coal finer than 1 mm. For coals finer than 1 mm, we find that we can use the shortest standard vortex finder in either our 10- or 15- in. Water Only cyclone without exception. For coals 5/8-in. x 1 mm ceramic vortex finders of fixed length have been satisfactory in all cases following an initial testing period in which the length was determined. Coals coarser that 5/8-in. are more sensitive to the vortex finder length, and two installations handling coals coarser than 1-in. continued to experiment with vortex finders of various lengths long after the initial startup period. For coals of all sizes, final adjustment of product quality has been achieved by determining the optimum size of apex orifice, which is accessible and readily changed, if necessary. Comments above assume no unusual change in the coal being washed or the required quality of clean coal.

Jan 1, 1979

By H. A. Schwartz

FROM time to time, one of the authors has had occasion to investigate the graphitizing reaction and has published the results mainly as discussion1 of the work of other investigators. In view, therefore, of the diversity of both observations and interpretation, it seemed well to attempt a revision of the entire subject in the Research Laboratory of the National Malleable Castings Co. The work consisted of a careful study of a single, impure, iron-carbon alloy. The material, which may be considered typical low-carbon white cast iron of the malleable industry, was an air-furnace iron of the following composition: carbon 2.30 per cent., silicon 1.20 per cent., manganese 0.29 per cent., phosphorus 0.156 per cent., sulfur 0.048 per cent. The freezing range extended from 1213° to 1132° C., the latter figure agreeing with Gontermann&apos;s data .2 We have attempted to define the equilibrium diagram of this alloy, in the stable condition as related to the metastable diagram, from a temperature somewhat below the eutectic freezing point to a temperature below Ar1 and for a graphitic-carbon content from practically nothing to the original total carbon of the specimen (see Fig. 6). More especially we desired to locate the line in the stable diagram corresponding to the Acm line of the metastable system, which is designated by analogy, the Atm line (tm from temper carbon) ; also to determine the location of Aim with respect to Acm toward their lower extremities at Ar1 in order to determine whether or not graphitization is completed by the separation of an iron-carbon eutectoid analagous to pearlite in the iron-cementite system. This would involve further a comparison of the locations of the stable and metastable A1 points.

Jan 8, 1922

By Cyril Stanley Smith, Earl W. Palmer

A complete discussion of the literature on the subject of the influence of copper on iron and steel will be published elsewhere.&apos; The present paper is concerned especially with the precipitation-hardening of copper steels, and brief mention will be made here only of the work of Kin-near,2,3 Nehl4 and Buchholtz and Koster,5 who alone have studied this phase of the subject. Kinnear found that normalized low-carbon steel castings containing about 1.0 per cent copper on drawing at temperatures in the neighborhood of 530" C. (1000" F.) increased in tensile strength and yield point, with little decrease in ductility. The maximum effect was obtained with about 0.9 to 1.0 per cent copper, and when the copper exceeded 3 per cent no improvement occurred on reheating. Kinnear did not study the effect of time of reheating—an all important factor where precipitation is concerned. Neh1,6 who apparently was unaware of Kinnear&apos;s work, recognized the change as due to precipitation and studied the interdependence of time and temperature of reheating, the temperature of the initial solution heat treatment, and the speed of cooling necessary to retain the bulk of the copper in supersaturated solution and to give an increase in hardness and tensile strength on reheating. Nehl found that if the final rolling or forging temperatures exceeded 700" C. (1300" F.) it was not necessary even to normalize the steel, but full precipitation occurred on merely reheating the forged or rolled shapes at 500" C. (930" F.). Buchholtz and Koster7 studied in detail the precipitation-hardening of a mild steel containing 1.0 per cent copper and 0.5 per cent chromium. From data on the electrical and magnetic properties of copper steels, they drew the portion of the equilibrium diagram in Fig. 1, assuming the

Jan 1, 1933

By Meredith Johnson

THE year 1924 saw little change in the production of oil and gas it either Pennsylvania or New York. In the early months of the year the price of Pennsylvania crude oil, $4.00 to $4.50 a barrel, was sufficiently high to cause operators to lay plans for rather extensive drilling opera-tions. But when, in August, the price dropped to $2.85 a barrel, there was no inducement for new work and drilling was curtailed. In spite of the curtailment, however, the combined production of the two states for the year showed a slight increase over 1923, the production being 8;859,000 bbl. in 1923, and 8,901,000 bbl.1 in 1924. The increased production was due entirely to. the continued development of flooding operations in the Bradford field, all other oilfields showing a decline for the year. TABLE 1. Production of Crude Oil in Pennsylvania and New York for the Last Ten Years a Production, Barrels Production, Barrels Year Year Pennsylvania I New York Pennsylvania I New York 1915 7,838,000 888,000 1920- 7,438,000 906,000 1916 7,593,000 874,000 1921 7,418,000 988,000 1917 7,733,000 880,000 1922 7,425,000 1;000,000 1918 7,408,000 809,000 1923 7,609,000 1,250,000 1919 8,137,000 851,000 1924 7,535,0006 1;366,0006 a U. S. Geological Survey figures. b Estimated. Much of the drilling in the southwestern part of Pennsylvania was done in search of gas. Natural gas in Pennsylvania is fast being recognized at its true worth and the increased prices allowed &apos;the gas companies for their product induced quite a bit of new work. The relative importance of natural gas compared with manufactured gas is shown in Table 2.2

Jan 3, 1925

By Wilbur I. Duvall, James F. Devine

7.4-1. Introduction. Ever since explosives were discovered and developed for mining purposes, there has existed the problem of determining what effect the air and ground vibrations resulting from blasting have on structures of various types. This problem is especially acute for surface operations such as open pit mines, quarries, and construction projects that are in the proximity of residential districts, scientific laboratories, or other vibration-sensitive structures. Because the general public is directly involved in the blasting vibration problem, it has been of major concern to local, state, and federal governments, as well as to mining and construction companies, explosive manufacturers, insurance companies, and scientific investigators. Consequently, many investigations have been conducted, both in this country and abroad, on the effects of air and ground vibration on residential and other structures. E. H. Rockwell made some of the first studies in this country on the vibrations produced by blasting and the effect that these vibrations have on structures (18). In his publications, he stressed the need for measuring the vibrations from quarry blasting in order to establish the level of vibration as a function of charge size and distance. The concept of a damage criterion for structures was developed shortly thereafter, and a displacement of less than 0.03 in. was considered safe. During the period from 1939-1942, the Bureau of Mines conducted an extensive investigation on the problem of seismic effects of quarry blasting. This work, summarized in Bulletin 442 (20) was the first major effort for establishing damage criteria for residential structures based on experimental data. An acceleration of less than 0.1 g was considered safe, 0.1 to 1.0 g was caution, and greater than 1.0 g was damage. This work

Jan 1, 1968

By C. W. Tomlinson

INTRODUCTORY THE analyses which have been made by the writer according to the method described below were made as part of Professor Withey&apos;s investigation of the concrete aggregates&apos; of Wisconsin, in the hope of throwing light on the extent and nature of the influence of the mineralogical constitution of aggregates upon the properties of the concrete, and mortar made from them. This method was found most satisfactory after four weeks of experiment and modification. It attains an accuracy quite sufficient for the purpose of the analysis, and is fairly rapid. Mineralogical analyses may be thus made at a cost considerably less than that of chemical analyses of a similar degree of completeness. The results obtained by this method should not vary in any item by more than 1 per cent. of the total weight of the sample, and may be made very much better than this by the exercise of greater care in the microscopic counting (see below). The accuracy of the hand sorting &apos;is of course dependent upon the experience and carefulness of the analyst. . After a month&apos;s practice, the writer was enabled to turn out these analyses at the rate of four in 33 hr. of working time. It was found most economical to handle four samples at once. The samples analyzed weighed froth 290 to 340 g. each. The same method could be applied to larger or to smaller samples, within reason able limits. The following list covers the necessary apparatus and equipment: 1 good balance capable of weighing accurately to tenths (or, better, hundredths) of a gram, and of handling quantities up to 200 g. in weight. 1 Westphal balance, registering densities to two or three decimal places. 1 Meker or Bunsen burner, and gas supply. 1 nest of sieves, 10, 20, 40, and 100 meshes to the inch.

Jan 5, 1915

By P. Stark, B. S. Lement

ACCORDING to Hanemann1 the hardness of retained austenite in a 1.7 pct C steel is increased by subsequent tempering. He reported that the hardness of this constituent increases only slightly on tempering to 300°F, rises more rapidly above this temperature, and attains the same value as the tempered martensite present in steel at approximately 480°F. The variation of hardness of both retained austenite and martensite with tempering temperature as obtained by Hanemann1 is shown in Fig. 1. Hanemann attributes the increase in hardness of the retained austenite essentially to continued formation of fine needles of martensite on tempering. Although he did not explain the actual mechanism involved, work hardening of retained austenite by the plastic deformation associated with such martensite formation appeared to be a possibility. Recently Lement, Averbach, and Cohen2 reported the existence of subgrains, about l000Å in diameter, in retained austenite and suggested that they may be delineated by a subboundary carbide. This raised the possibility that carbide precipitation might be responsible for the hardening reported by Hanemann. In view of these conflicting possibilities, an investigation" was carried out on a high purity, Fe-C alloy containing 1.7 pct C in order to clarify the nature of the phenomenon involved in the reported hardening of retained austenite. This investigation involved correlation of changes in microhardness and microstructure of the austenite as a result of tempering. Microhardness tests were made with a Tukon tester using a 25 gram load and a Vickers diamond. It was found that two groups of austenite hardness values were obtained after tempering at a series of temperatures up to 450°F, which was the highest temperature at which retained austenite regions could be identified microscopically. One group represents a high hardness level which in-

Jan 1, 1955

By R. E. Miner, J. K. Jackson, T. L. Wilson

THE nature of the decomposition reactions producing hardening in aluminum-rich A1-Ag alloys is now fairly well understood. The silver is not uniformly dispersed even at temperatures above the solvus line: rather, it is clustered into zones and associated with dislocations.&apos; During quenching, nearly all of the silver segregates into zones&apos; and some is attracted to dislocations, producing stacking faults. Growth of these faults produces the semicoherent y precipitate.1 During aging, the larger zones and the y precipitates grow at the expense of the smaller zones. The large zones persist up to about 350°C; where they redissolve in favor of the y&apos; and y precipitates.&apos; The equilibrium hcp y precipitate occurs both by a cellular reaction at the grain boundaries1 and by growth of the Y&apos; plates to a critical size where incoherency results. Damask and Nowick 4 observed an internal friction peak in an A1-6 at. pct Ag alloy at about 140°C (0.25 cps). The peak was found to exhibit a systematic instability. After aging at successively higher temperatures between 155&apos; and 23Z°C, the peak first migrated to higher temperatures while increasing in height, then returned to lower temperatures and decreased in height. After aging at temperatures above about 200°C, the peak position remained constant while the height decreased very slowly. In quenched specimens, elastic after-effect measurements inhcated a rapid initial transient in peak position from low temperatures up to 155°C. The authors concluded that a stress induced growth and resolution of the clusters and the 7&apos; was the most reasonable mechanism to account for the behavior. It would seem, however, very unlikely that both the clusters and the y&apos; precipitate would provide damping contributions whose relaxation time and activation energies were identical, which is required for them to produce the same internal friction peak. The present investigation was motivated by this apparent anomaly and the development of a very low thermal inertia torsion pendulum5 capable of monitoring transient behavior. Low frequency internal friction measurements were made on four binary A1-Ag alloys nominally containing 5, 10: 15, and 20 at. pct Ag. The results of measurements made on a 10 at. pct Ag alloy in the quenched condition are shown in Fig. 1. There are two peaks apparent in the spectrum, one centered at about 160°C, the second at about 210°C. The two peak spectrum was observed in each alloy, but it was best resolved in the 10 pct Ag specimens. Because of the instability in the internal friction, 4 it seemed possible that the second peak was not a peak at all, but simply a manifestation of a single, sharply transient, peak. This effect has been observed in Au-Ni alloys.6 The existence of the higher tempera-

Jan 1, 1970

By A. H. Daane, B. J. Beaudry

The uranium-antimony system has been investigated by metal-lographic, therma1, and X-ray methods. There are four intermediate phases present: USbz and U,Sb, which undergo peritectic decomposition at 1355°and 1695OC, respectively, and USb and UPSbs which melt congruently at 1850"and 1800°C, respectively. There is a eutectic between U4Sbs and USb which melts at 1770°C. USb,, UsSb4, and USb are line compounds while U4Sbs has a solubility range of approximately 1.5 at. pct at 1770°C. Single crystal studies of the cmpo,und U.,SbS show it to have a hexagonal unit cell with a, = 9.268A and c, = 6.201b;. The solubility of uranium in liquid antimony increases from 0.1 at. pct at 650°C to 1.5 at. pct at 900°C. The solubility of antimony in a, P, and y uranium is approximately 0.02, 0.1, and 0.05 at. pct respectively. IN nuclear reactors, uranium has been used mainly in the form of solid-fuel elements of the pure-metal or uranium-rich alloys. However, in attempts to overcome some of the difficulties inherent in solid-fueled reactors, reactor concepts hive been proposed in which the fuel would be a liquid metal, such as a solution of uranium in another metal. The liquid-metal fueled reactor (LMFR) designed by the Brook-haven National Laboratory of the U.S. Atomic Energy Commission is an example. Fuel for this reactor is a solution of approximately 800 ppm enriched uranium in bismuth.&apos; To design such liquid-fueled reactors, a background of information on the solubility of uranium in various metals is obviously necessary. The present study was undertaken to pro- vide information for such purposes, and in addition, to study the uranium-antimony system as a whole. A preliminary literature search showed that some work had been done previously on uranium-antimony alloys, but no attempts had been made to make a complete study of this system. The solubility of uranium in antimony was studied by Cammack and Bridger&apos; up to 950°C and by Hayes and Gordon up to 900&apos;-C. Their data are plotted in Fig. 1. Ferro, in an X-ray structure study of some uranium-antimony compounds, found USg to be tetragonal with a, = 4.2724 and c, = 8.471A, U3S4 to be b.c.c. with aO = 9.095A and USb to be f.c.c. with a,= 6.191A. Colani reported the compound UsSba; however, work by Ferro and by this Laboratory showed that this compound does not exist. foote&apos; reported several intermediate phases in the uranium-antimony system, but did not identify them. Greninger&apos; re-ported finding an intermediate phase in a 5 at. pct Sb

Jan 1, 1960

By A. Gatti

A dispersion of AI2O3 was produced within chromium by the hydrogen reduction of a solid solution of Cr2U3-A1203. The yesulting powder. was compacted, extruded, and the material tested in tension at various temperatures and in creep at 1800°F. The material is ductile above 1200°F and is at least twice as strong as electrolytic chvomium at all temperatures. Its creep resistance, rupture strength, and ductility at 1800°F warrant further work to improve its Low -temperatutre properties. BECAUSE of its high melting point and excellent corrosion resistance, chromium has long been considered as a base for high-temperature alloys. Most chromium base alloys, however, are extremely brittle at low temperatures. Elements such as nickel and iron lower the melting point of chromium and thereby limit its high-temperature usefulness. Chromium-alumina materials made by othersl, 2 have shown good high-temperature creep resistance but poor ductility and thermal shock resistance. Pure ductile chromium3 is weak at high temperatures and reverts to a brittle state if exposed to nitrogenous atmospheres. In sintered aluminum powder (SAP)4 and internally oxidized alloys of silver, copper, and nickel, fine dispersions of oxide phases can control hardness, recrystallization, and yield stress. The possibility of formation of a ductile, high-strength material by precipitating dispersions of alumina in a chromium matrix appears worthy of investigation. MATERIAL Powder Production—Chromium oxide and alumina in amounts to produce a final alloy of Cr + 16 pct AI2O3 by weight were colloid milled, filtered, dried and high fired for 16 hr at 1600°C in air. The resulting solid solution of oxides was ball-milled to a fine powder and the chromium oxide reduced to chromium by heating in pure dry hydrogen for 64 hr at 1400°C. Final Materials Produced by Powder Metallurgy Techniques—-Chromium powder, containing a disper-sion of alumina, was pressed into 1 1/8-in. diam slugs at 25 tons per sq in. and sintered 16 hr at 1400° C in pure dry hydrogen. The sintered slugs were encased in a stainless steel can, 2 in. in diam, and extruded at 1200° C through a 1/2-in. die. The 16 to 1 extrusion ratio produced a final chromium-alumina rod 1/4 in. in diam about 30 in. long. Samples of the rod were plunge-ground to small rupture bars of 0.82-in. gage length with a 0.01-sq-in. cross section. RESULTS Microstructure Studies—Fig. 1 is the microstruc-ture of the material in the as-sintered condition while Fig. 2 is the microstructure of the material as-extruded. Comparison of the as-sintered and as-extruded micrographs show the redistribution and realignment of the alumina inclusions by the extrusion

Jan 1, 1961