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By W. E. Stiles

A method is presented for predicting the performance of water flooding operations in depleted, or nearly depleted, petroleum reservoirs. The method makes use of permeability variations and the vertical distribution of productive capacity. From these two parameters can be calculated the produced water cut versus the oil recovery. Derivations of the mathematical analogy is shown and sample calculations and curves of prediction are presented. Comparison is made of the predicted and actual performance of a typical 5-spot in an Illinois water flood. INTRODUCTION The use of water as a flooding medium in both depleted and "flush" oil reservoirs is gaining greater recognition and acceptance. Many of the shallower fields, depleted by primary production, have been and are being subjected to water injection in order to obtain some part of the large volume of oil remaining after primary production. Some of the earlier water flood installation proved highly discouraging and the value of water flooding was often questioned. Many of these earlier floods were haphazardly selected and developed as little was known of the physical characteristics and contents of the producing formations. The prior evaluation of the flood performance was impossible. During the past decade the development of the required reservoir engineering tools-—core analysis, reservoir fluid analysis, electric logs, fluid flow formulae, etc.—has allowed the engineer to construct and apply the methods which are presently being used to evaluate the economic and mechanical susceptibility of a reservoir to flooding. This discussion will present a method for taking into account the effect of permeability variations in predicting the performance of water floods in depleted reservoirs. PERMEABILITY AND CAPACITY DISTRIBUTION It is generally agreed by most investigators that in a single phase system fluid will flow in a porous and permeable medium in proportion to the permeability of the medium. Producing formations are usually highly irregular in permeability, both vertically and horizontally. However, zones of higher or lower permeability are often found to exhibit lateral continuity. Thus, while structurally comparable stringers in adjacent wells may differ several fold in permeability values, they usually bear resemblance as being part of a general continuous higher or lower permeability section. It is generally agreed that where such stratification of permeability exists, injected water sweeps first the zones of higher permeability, and it is in these zones that "break-through" first occurs in the producing well. It is a basic assurnption of the presently described method that penetration of a water front follows the individual permeability variations as if such variations were continuous from input to producing well. This is admittedly not rigorously true, but can be justified as making possible a simplifying mathematical approach to an otherwise extremely complicated three dimensional flow problem. As a basis for study of the lateral flow of fluids in formations of irregular permeability, the irregularities may be conveniently represented by a permeability distribution curve and a capacity distribution curve. In obtaining these curves, the permeability values, regardless of their structural position in the formation, are rearranged in order of decreasing permeability. If these permeability values so arranged are plotted against the cumulative thickness, a permeability distribu- tion curve is obtained. This curve may then be likened to a "smoothed" permeability profile of the formation. In making comparison between different distribution curves it is convenient to state the permeabilities in terms of the ratio of the actual permeability values to the average permeability of the formation. These ratios termed "dimensionless permeabilities", are used in this paper rather than the permeabilities in terms of millidarcys. The capacity distribution curve is a lot of the cumulative capacity (starting with the highest permeabilities) versus the cumulative thickness. The capacity and thickness are given as fractions of the total capacity and thickness. Mathematically, the capacity distribution is the intergration of the permeability distribution curve. In practice it is convenient to first obtain the capacity distribution curve and derive from it a smoothed dimen-sionaless permeability curve. The method of obtaining the capacity distribution curve is illustrated in the successive column of Table 1, in which capacity and thickness are derived as fractions of their respective totals. If only a small number of permeability values are available, it is generally desirable to smooth the resultant curve. This has been done to give the capacity distribution curve shown in Figure 1. The differentiation of the capacity distribution curve to obtain the permeability distribution curve is shown in Table 2. Here, the capacity values are read from the smoothed curve at intervals of cumulative thickness, and the increments of capacity are divided by the increments of thickness to obtain the dimensionless permeability, K. Due to this stepwise procedure of calculation these premeability values must be plotted at the midpoints of the successive increments of thickness. The curve so obtained from these data is shown in Figure 1. The total area under the K' curve is equal to unity.

Jan 1, 1949

By W. E. Stiles

A method is presented for predicting the performance of water flooding operations in depleted, or nearly depleted, petroleum reservoirs. The method makes use of permeability variations and the vertical distribution of productive capacity. From these two parameters can be calculated the produced water cut versus the oil recovery. Derivations of the mathematical analogy is shown and sample calculations and curves of prediction are presented. Comparison is made of the predicted and actual performance of a typical 5-spot in an Illinois water flood. INTRODUCTION The use of water as a flooding medium in both depleted and "flush" oil reservoirs is gaining greater recognition and acceptance. Many of the shallower fields, depleted by primary production, have been and are being subjected to water injection in order to obtain some part of the large volume of oil remaining after primary production. Some of the earlier water flood installation proved highly discouraging and the value of water flooding was often questioned. Many of these earlier floods were haphazardly selected and developed as little was known of the physical characteristics and contents of the producing formations. The prior evaluation of the flood performance was impossible. During the past decade the development of the required reservoir engineering tools-—core analysis, reservoir fluid analysis, electric logs, fluid flow formulae, etc.—has allowed the engineer to construct and apply the methods which are presently being used to evaluate the economic and mechanical susceptibility of a reservoir to flooding. This discussion will present a method for taking into account the effect of permeability variations in predicting the performance of water floods in depleted reservoirs. PERMEABILITY AND CAPACITY DISTRIBUTION It is generally agreed by most investigators that in a single phase system fluid will flow in a porous and permeable medium in proportion to the permeability of the medium. Producing formations are usually highly irregular in permeability, both vertically and horizontally. However, zones of higher or lower permeability are often found to exhibit lateral continuity. Thus, while structurally comparable stringers in adjacent wells may differ several fold in permeability values, they usually bear resemblance as being part of a general continuous higher or lower permeability section. It is generally agreed that where such stratification of permeability exists, injected water sweeps first the zones of higher permeability, and it is in these zones that "break-through" first occurs in the producing well. It is a basic assurnption of the presently described method that penetration of a water front follows the individual permeability variations as if such variations were continuous from input to producing well. This is admittedly not rigorously true, but can be justified as making possible a simplifying mathematical approach to an otherwise extremely complicated three dimensional flow problem. As a basis for study of the lateral flow of fluids in formations of irregular permeability, the irregularities may be conveniently represented by a permeability distribution curve and a capacity distribution curve. In obtaining these curves, the permeability values, regardless of their structural position in the formation, are rearranged in order of decreasing permeability. If these permeability values so arranged are plotted against the cumulative thickness, a permeability distribu- tion curve is obtained. This curve may then be likened to a "smoothed" permeability profile of the formation. In making comparison between different distribution curves it is convenient to state the permeabilities in terms of the ratio of the actual permeability values to the average permeability of the formation. These ratios termed "dimensionless permeabilities", are used in this paper rather than the permeabilities in terms of millidarcys. The capacity distribution curve is a lot of the cumulative capacity (starting with the highest permeabilities) versus the cumulative thickness. The capacity and thickness are given as fractions of the total capacity and thickness. Mathematically, the capacity distribution is the intergration of the permeability distribution curve. In practice it is convenient to first obtain the capacity distribution curve and derive from it a smoothed dimen-sionaless permeability curve. The method of obtaining the capacity distribution curve is illustrated in the successive column of Table 1, in which capacity and thickness are derived as fractions of their respective totals. If only a small number of permeability values are available, it is generally desirable to smooth the resultant curve. This has been done to give the capacity distribution curve shown in Figure 1. The differentiation of the capacity distribution curve to obtain the permeability distribution curve is shown in Table 2. Here, the capacity values are read from the smoothed curve at intervals of cumulative thickness, and the increments of capacity are divided by the increments of thickness to obtain the dimensionless permeability, K. Due to this stepwise procedure of calculation these premeability values must be plotted at the midpoints of the successive increments of thickness. The curve so obtained from these data is shown in Figure 1. The total area under the K' curve is equal to unity.

Jan 1, 1949

By B. C. Allen

The interface energies of chronzium, molybdenunz. hugsten, and their solid-solution alloys Cv-35Re, MO-33Re, and UJ-25Re were studied at 0.6 to 1.0 of the absolllte liquidus ter)zpe,vature using fiz&apos;e )izethods. Liquid surface tension, yv , was deter mined clsing the pendant-drop and drop-weight methods. Results are, respectizlely, 1700, 2370, and 2480 +100 dynes per ct for the rhernium -containing alloys and essentially the same as tlwse reported for liquid chro)riln, trolybdenum, and tungsten. Average solid slrjace energy, rsv< xias ))zeasured using tlre fiber-extetlsion method. The ratio of ysS, the acerage high-angle grain-boundary energy, to ySV cclas jolnd fronz grain-bolzdary grooue angles fort)zed at the surface in an inert atrfizosphere. Absolllte iute?:face energies were deterawined using ?nultip/rase equilibria involzing suitable liquids of known surface tension (tin, silver). Interpretation of the experimented results in view of pvobable tenzperatzcre, orientation, and purity effects giz,e the follouling approximations in ergs per sq ctn: ysv (i2lo. Mo-33Re) - 2100, ySS (Mo, Mo-33Re) - 800, rr (defornzation twins in MO-33Re at 1200"C) - 800. ysV (Cr. Cr-35Re) - 2400. YSS (CY, Cr-35Re) - 1000. Probably Ylv- YSV- 2500 for tungsten and W-25Re, giving yss (It&apos;, W-25Re) - 900. The interface energies of solid and liqid ch?&apos;omiu?.z, nolybdenu?rr, and tungsten are not geatly aff&apos;ected by rhenium and therefore are not a ttlajor factor in the ductili zing rhenium effect in Croup VI-A metals. THE interface energies of the refractory Group VI-A metals, chromium, molybdenum, and tungsten, are not well-established. The objective of this investigation was to study the liquid surface tension, solid surface energy, and grain-boundary energy of these metals and compare them to those found for the bcc solid-solution alloys, Cr-35e,&apos; 0-33e,&apos; and -25e. Five techniques were used to measure interface energies in high-purity polycrystal rod, wire, and sheet at 0.6 to 1.0 of the absolute liquidus temperature. The alloys were chosen to see if there was any connection between interface-energy behavior and the ductilizing rhenium effecL4j5 EXPERIMENTAL WORK Materials. A description of the materials used is presented in Table I. Chromium rod was prepared by arc melting iodide process crystals supplied by Chromallo Cor., hot extruding, and warm swaging to 0.63-cm-diam rod.6 The sheet was prepared by rolling as-extruded rod to 95 pct reduction in area from a hydrogen furnace at 800" to 900°C and surface grinding off 0.02 cm from each face. Cr-35Re rod was prepared by arc melting sintered rhenium powder and iodide chromium crystals, warm rod rolling to 50 pct reduction in area in cans, and swaging to 60 pct reduction in area at 1100" to 1200°C. Some of the rod was warm-rolled to sheet and then surface-ground. Portions of swaged chromium and Cr-35Re were further reduced by swaging and drawing to 0.013-cm-diam wire by the General Electric Co. Mo-33Re and W-25Re rod, sheet, and wire were provided by Chase Brass and Copper Co. The molybdenum sheet consisted of two lots, both essentially the same except for the carbon content. Liquid Surface Tension. The liquid surface tension of Cr-35Re, Mo-33Re, and W-25Re was measured by a combination of pendant-drop and drop-weight methods using techniques already decribed." Following out-gassing, molten drops were formed on the ends of centerless-ground Mo-33Re and W-25Re rods by electron bombardment at 5 x 106 mm. Similar drops were formed on outgassed Cr-35Re rods by induction heating under 1 atm of 99.995 pct Ar. Solid Surface Energy. Solid surface energy was measured by conducting microcreep experiments on molybdenum, Mo-33Re, chromium, and Cr-35Re wires at 2350°, 2306, 1550°, and 180O°C, respectively. In preparation, gage marks -2.5 cm apart and -0.001 cm deep were circumferentially scribed on the wire with a razor blade. Weights of the wire material were then attached. Five to seven reasonably straight wires were hung in a container made out of the wire material. The free end was placed through a small hole in the removable top and secured by bending a small portion 90 deg. The containers not only tended to provide vapor-solid equilibrium for the wires but also protected them from gaseous impurities. They were nominally 2.5 cm in diam by 5 cm high and were made from extruded chromium rod, Cr-35Re arc casting, molybdenum bar stock, or welded Mo-33Re sheet. After deg re as ing, the assembly was outgassed at a relatively low temperature to 2 x 10"5 mm and then recrystallized 2 to 8 hr at the creep temperature in a rhenium-element resistance furnace. The static argon atmosphere was gettered by tantalum radiation shielding. Specimen temperature was measured optically to 25"C using calibration with known melting points and blackbody conditions. The wires generally developed a stable bamboo-type structure according to Fig, l(b), (c), and (d) and retained their gage marks [upper portion of Fig. l(d)]. One or two of the weights were clipped off to provide a low load for the creep anneal. To minimize the possibility of bending or breakage, the wires remained attached to the top of the annealing container which was held to keep the wires vertical. The distance between gage marks was

Jan 1, 1967

By M. Gell, G. R. Leveran

The high- and low-cycle fatigue properties of the nickel-base superalloy, Mar-MBOO, in columnar-grained and single-crystal forms were determined at room temperature. It was found that the fatigue lives of these materials were greatly affected by the size of preexisting cracks in MC-type carbides contained in the micro structure. Most of the data falls on two curves given by: (zN)&apos;/A€= K, where Nf is the number of cycles to failure, Af is the total strain range, and K is a function of carbide size. No difference was observed in the fatigue behavior of the columnar-grained and single-crystal materials for the same MC carbide size. Matrix slip and crack initiation occurred at precracked MC carbides and, to a lesser extent, at micropores. Fatigue crack propagation was mainly in the Stage I mode, i.e., on cry stallo graPhic slip planes. The Stage I fracture in these materials was unusual in that distinct features were observed on the fracture surfaces. In high-cycle fatigue, these features resembled those commonly observed on the cleavage fracture surfaces of bcc and hcp materials. Yet, in this study, the cracks propagated slowly in a cyclic manner. In low-cycle fatigue, the Stage I facets contained equiaxed dimples, similar to those observed on the tensile fracture surfaces of ductile materials. These observations indicate that both local normal and shear stresses are involved in these Stage I fractures. A model is proposed to explain these results based on the weakening of the cohesive energy of the active slip planes by reversed shear deformation and the fracture of the bonds across the weakened planes by the local normal stress. RECENT developments in casting technology have produced cast nickel-base superalloys in columnar -grained and single-crystal forms.1&apos;2 The tensile and creep properties of the nickel-base superalloy, Mar-M200, cast in these forms have been shown to be superior to the corresponding properties of the conventionally cast polycrystalline material.lp2 This improvement in properties results, in part, from the elimination of grain boundaries in the single crystals and the alignment of the grain boundaries parallel to the stress axis in the columnar-grained castings. As part of a program to evaluate the fatigue properties of nickel-base superalloys cast in single-crystal and columnar-grained forms, a study has been made of the cyclic deformation and fracture of Mar-M2OO at room temperature. M. fiFl I .hininr Mpmher AIMF ic ^pninr Rocoarrh Accn^iata anH I) EXPERIMENTAL PROCEDURE The composition range of Mar-Ma00 in weight percent is: 8 to 10 Cr, 9 to 11 Co, 11.5 to 13.5 W, 0.75 to 1.25 Cb, 1.75 to 2.25 Ti, 4.75 to 5.25 Al, 0.01 to 0.02 B, 0.03 to 0.08 Zr, 0.07 to 0.12 C, bal. Ni. All of the castings met the above specifications. The castings were solutionized for 1 to 4 hr at 2250°F followed by aging at 1600°F for 32 hr which resulted in a 0.2 pct offset yield stress of 150,000 psi at room temperature. The microstructure of the material consisted of cuboidal, coherent particles of ordered, fcc Ni3(A1,Ti) (commonly designated y&apos;), approximately 0.3 p on edge, distributed in an fcc y solid-solution matrix. MC carbides together with shrinkage and gas micropores were also distributed throughout the materials. The MC carbides and micropores were located preferentially in the interdendritic interstices, as well as in the grain boundaries in the columnar-grained castings. The (100) direction of all the single crystals and the common (100) axis of the grains in the columnar materials were aligned within about 5 deg of the specimen axis. Fatigue testing was carried out in the high-cycle (HCF) and low-cycle (LCF) fatigue regions, with the major difference being gross yielding of the specimen occurred during the first cycle in the LCF region. This division also corresponded with the more usual one in which the life of a specimen in LCF is less than lo4 cycles and that in HCF is greater than lo4 cycles. The designs of the high-cycle fatigue and low-cycle fatigue specimens are shown in Figs. l(a) and (c), respectively. The gage sections of both HCF and LCF specimens were electropolished prior to testing. The HCF specimens were tested in an MTS, closed-loop, hydraulic fatigue machine at 10 cps in air. The specimens were cycled between a tensile stress of 5000 psi and a maximum tensile stress which ranged from 35,000 to 125,000 psi, Fig. l(b). The LCF specimens were cycled under strain control from zero to a maximum tensile strain, Fig. l(d), in a Wiedemann-Baldwin testing machine. The experimental procedure has been described elsewhere.3&apos;4 Both HCF and LCF tests were interrupted periodically in order to replicate the development of slip and cracking at the specimen surface. This was accomplished by placing plastic replicating tape around the gage section of the specimen while the specimen was in the mahine. The size of the MC carbides for all specimens was measured on a polished longitudinal section through the gage section after fatigue testing. The method of measurement consisted of carefully scanning the entire polished section in order to locate the largest MC carbides. Photographs were then taken of the six longest carbides oriented approximately normal to the

Jan 1, 1969

THOSE MARKED THUS * ARE MEMBERS, MARKED THUS t ARE ASSOCIATES THESE SIGNS DOUBLED INDICATE LIFE MEMBERS AND ASSOCIATES RESPECTIVELY THE FIGURES AT THE END OF THE ADDRESS INDICATE THE YEAR OF ELECTION HEAVY FACED TYPE SIGNIFIES HONORARY MEMBERSHIP ` *AARONS, JULIAN B , Mine Supt, The Boulder Deep Levels, Ltd , Boulder, Western Australia, &apos;05 *ABADIE, EMILE R, Cons and Min Engr, 1004 James Flood Bldg, San Francisco, Cal &apos;76 *ABBE, CLEVELAND, JR, GeOl Aid, U S Geological Survey, Washington, D C &apos;00 *ABBOTT, AI ARTHUR, Cerro de Pasco Mining CO , Cerro de Pasco, Peru, So America &apos;95 *ABBOTT, ARTHUR V , Electrical Engr , Westinghouse, Church, Kerr & CO , 8 Bridge St, New York, N Y &apos;82 *ABBOTT, CLARENCE E, Chief Min Engr, Care Fayal Mine, Eveleth, Minn &apos;04

Jan 1, 1910

By Carl Henrich

Jan 1, 1893

By C. E. Nelson, T. E. Leontis

THE properties and casting characteristics of sand-cast Mg-Al-Zn alloys, used commercially in this country and abroad, have been discussed in a number of articles during the past few years.&apos;-" In addition, broad surveys of many compositions in the magnesium corner of the Mg-Al-Zn diagram between 0 and 12 pct A1 and 0 and 6 pct Zn have been presented by Busk and Marande7 and by Fox." The purpose of this paper is to furnish a comprehensive survey of the changes in tensile properties, micro-structure, and dimensional stability of two magnesium alloys, AZ92A and AZ63A,+ as a function of aging time and temperature after solution heat treatment. Rates of precipitation in Mg-Al-Zn alloys are slow enough so that a substantially homogeneous, supersaturated solid solution can be maintained even when these alloys are air-cooled from the heat-treating temperature. Consequently, even though it has been demonstrated9 10 that significantly higher strengths can be developed in some alloys in the aged condition by quenching rather than air cooling from the heat-treating temperature, general practice is to air-cool magnesium alloys. All the work discussed in this paper is based, therefore, on aging studies performed on material air-cooled from the heat-treating temperature. This investigation was intended to be a study of commercially practical aging cycles; the tensile properties were determined, therefore, up to aging times of 25-hr duration. The alloys, used in this investigation were prepared in the laboratory by remelting commercial ingots of the appropriate composition according to the procedures described in a previous paper" as the "crucible method." All melts were given the standard superheating treatment in order to insure the maintenance of a uniform grain size of 0.003 to 0.004 in. Standard test bars, 61/2 in. long with a 2½ in. long reduced section having a diameter of Y2 in., were cast from these melts in green-sand molds using a four-bar pattern. The alloys were heat treated according to the following schedules: AZ92A alloy: 500" to 780°F in 2 hr + 780°F (24 hr) AZ63Aalloy: 500" to740°Fin2 hr + 740°F (16hr) In commercial practice," AZ92A alloy is heat treated at 770°F for 18 hr and AZ63A alloy is heat treated at 730°F for 10 hr. The slightly higher temperatures and longer periods of heat treatment were used in order to obtain as homogeneous a solid solution as possible and thus to minimize inconsistencies in the

Jan 1, 1952

By D. I. Bardos, F. X. Spiegel, R. K. Malik, Paul A. Beck

Study of the occurrence of elongated ß-Mn phase fields and of their direction of extension in fifteen ternary systems indicates that the ß-,mn phases formed by manganese with other transition elements and with either silicon, germanium, or tin may be considered as electron phases. In these phases the alloying behavior of silicon on the one hand and of germanium and tin on the other shows a significant difference . Formally all observations can be satisfactorily described by attributing to germanium and tin an apparent electron concentration of nine electrons per atom. This behavior may be interpreted by assuming that germanium and tin exclude itinerant electvons because of a strong repulsive potential arising fromthe requirement that their wave junctions be orthogonal to those of- the closed d shells of these atoms. The absence of a d shell in silicon allows this atom to admit itinerant d electrons. The following new ternary phases of unknown structure were identified by means of their X-vay difjraction patterns: (Mn,Co,Si)I , (Mn,Co ,Si)Y, (Mn.Co,Si)Q , (Mn ,Ni ,Si)W , and (Mn, Ti ,Si)A. The X-ray data suggest that these phases have large unit cells and that to some extent they may he related structurally to the 0 phase. The a phase and certain other phases formed by transition elements are characterized by their preferred occurrence at a constant electron concentration. AS a result, these "electron phases" have relation to electron concentration. The tendency of the B phase to form elongated phase fields in ternary systems along the directions of constant electron concentration has been previously demonstrated. The occurrence of an elongated 0-Mn phase field in the ternary system Mn-Co-Si was schematically indicated by Cherkashin et al.l0 The present study includes fifteen ternary systems, each containing in addition to manganese a second transition element from the group Ti, V, Cr, Fe, Co, Ni and one of the nontransition elements silicon, germanium, or tin. In each system the boundaries of the B-Mn phase were determined in the 900°, 950°, or 1000°C isothermal section, the temperature being dictated on the one hand by the requirement of sufficient rates of diffusion to approach equilibrium conditions in reasonably short annealing periods and on the other hand by the need to avoid as much as possible the partial melting of alloys during annealing.

Jan 1, 1967

By Carl Zapffe

The object of the Bradley process is to free manganese oxide from its associated gangue and separate the contained iron oxide by dissolving the manganese and precipitating it from the solution. This results in producing two oxides as concentrates, the one to be converted into a manganese sinter and the other into an iron sinter. The manganese sinter may be made to contain as high as 70 per cent. inanganese and as low as 1 per cent. silica, with phos&apos;phorus not to exceed 0.02 per cent, and from that figure downward to about 0.005 per cent., and with sulfur at 0.007 per cent. The iron sinter contains about 2.50 per cent. manganese, making the total recovery of the manganese 98 per cent. The method involves selective roasting, which converts the manganese into a form in which it is soluble in an aqueous solution of ammonium sulfate, then precipitating the manganese as a hydrate by using ammonia gas, then stabilizing and washing the hydrate, followed by sintering; concurrently most of the iron oxide is converted by roasting into an insoluble form and becomes a residue with magnetic qualities, which makes possible the recovery of much of the iron in the ore. Leaching manganese by using sulfur dioxide gas was shown possiblc Inany years ago. Mr. Bradley tried it for a year, working in conjunction with the original experimenters of that method, and found that while relatively pure inanganese oxide could be produced, the likelihood of developing a commercial process along this line was not very encouraging. He also attempted leaching with nitric and sulfur acids, but the cost of recovery mounted and the quality of the product was not as good Mr. Bradley tried nearly every leaching method proposed, and the expen-ments were made on a large scale in the spacious laboratories of John C. Wiarda & Co., in Brooklyn, N. Y. Insufficient encouragement was discernible in the results obtained. Upon the suggestion of N. Arthur Laury, of John C. Wiarda & Co., Bradley undertook using arnrnoniurn sulfate. Because a high rate of recovery appeared feasible, Bradley planned the erection of a 4-ton pilot plant at the State of Minnesota Mines Experiment Station, at Minneapolis. The results of that work are described in this treatise.

Jan 1, 1929

By Raymond Mancha

TO determine the optimum inside dimension of an airshaft, it is necessary to strike the proper balance between the cost of power for air friction and turbulence losses within the airshaft, on the one hand, and the cost of sinking the airshaft, on the other. An airshaft larger than necessary is extravagant because the increased amortization rate exceeds the reduction in power cost. Conversely, an airshaft smaller than is needed wastes more power than can be justified by the corresponding reduction in amortization rate. The airshaft power is a function of the air volume Q cu. ft. per min. and the air-pressure drop p in. water gauge across the airshaft. † The airshaft pressure drop is calculated by the formula: R X l X O X Q2 P= 5-2 X A3 wherein R is the air-friction coefficient for the portion of the airshaft 1 ft. deep, 0 ft. in perimeter and A sq. ft. in area, when passing an air volume of Q cu. ft. per minute. The airpower w kw. dissipated within the airshaft is computed by the formula: w= 0.746Q X P 6345 With a combined over-all unit efficiency of the fan and drive of E per cent, the air-shaft power requirements become W = 100w There are 8760 hours per year, so with energy at I cent per kilowatt-hour the annual cost per kilowatt for continuous operation is $87.60; therefore, with energy at C¢ per kw-hr. the annual airshaft power cost MI dollars per year becomes: M1 = 87.60W X C which, expressed by the component parts, becomes: M1 V 5.2X4&apos; / XQ Z 0.746 X 100 X 87.60 X C V 60 X E ) The amortization rate a dollars per year per dollar invested includes both depreciation and interest on the investment. These in turn involve the expected useful life of the airshaft and available molletary interest rate. For the purpose of this discussion the cost of airshaft sinking is assumed to comprise a fixed move-in charge Y dollars, which covers the cost of transporting, erecting and removing the shaft-sinking equipment and is independent of the air-shaft size. Also, it is assumed that the cost of shaft lining, if any, is fixed at $X per cu. yd. of lining material and that excava-

Jan 1, 1947

By Raymond Mancha

TO determine the optimum inside dimension of an airshaft, it is necessary to strike the proper balance between the cost of power for air friction and turbulence losses within the airshaft, on the one hand, and the cost of sinking the airshaft, on the other. An airshaft larger than necessary is extravagant because the increased amortization rate exceeds the reduction in power cost. Conversely, an airshaft smaller than is needed wastes more power than can be justified by the corresponding reduction in amortization rate. The airshaft power is a function of the air volume Q cu. ft. per min. and the air-pressure drop p in. water gauge across the airshaft. † The airshaft pressure drop is calculated by the formula: R X l X O X Q2 P= 5-2 X A3 wherein R is the air-friction coefficient for the portion of the airshaft 1 ft. deep, 0 ft. in perimeter and A sq. ft. in area, when passing an air volume of Q cu. ft. per minute. The airpower w kw. dissipated within the airshaft is computed by the formula: w= 0.746Q X P 6345 With a combined over-all unit efficiency of the fan and drive of E per cent, the air-shaft power requirements become W = 100w There are 8760 hours per year, so with energy at I cent per kilowatt-hour the annual cost per kilowatt for continuous operation is $87.60; therefore, with energy at C¢ per kw-hr. the annual airshaft power cost MI dollars per year becomes: M1 = 87.60W X C which, expressed by the component parts, becomes: M1 V 5.2X4&apos; / XQ Z 0.746 X 100 X 87.60 X C V 60 X E ) The amortization rate a dollars per year per dollar invested includes both depreciation and interest on the investment. These in turn involve the expected useful life of the airshaft and available molletary interest rate. For the purpose of this discussion the cost of airshaft sinking is assumed to comprise a fixed move-in charge Y dollars, which covers the cost of transporting, erecting and removing the shaft-sinking equipment and is independent of the air-shaft size. Also, it is assumed that the cost of shaft lining, if any, is fixed at $X per cu. yd. of lining material and that excava-

Jan 1, 1947

By Franklin H. Sharp, Kenneth L. Clifford

During the last few years the New Lead Belt of Southeastern Missouri has become the main source of lead in the United States. It also produces significant amounts of zinc, copper and silver. The mines are located on a narrow, mineralized trend extending southward from Viburnum, MG. for about 35 miles. The ore occurs as replacement deposits in the Bonneterre lime- stone formation which is about 1000 ft below the surface. In most of the mines, enough water is produced to supply the milling requirements without any reuse. There are dry areas, however, and at least two mills must rely on water reclaimed from their tailings lake to augment the mine water. These mills have no choice but to utilize recycled mill water.

Jan 7, 1973

By Dogan E. Gücer

In the following note, with the help of a new parameter of strain hardening, the attention is drawn to the close relationship between the relative performances of steels under plastic range cycling and their strain hardening characteristics. If a cantilever beam of rectangular cross section is tested with the arrangement shown in Fig. 1 in reversed bending between constant deflection limits at the lever head, the longitudinal strain-deflection hysteresis shown in Fig. 2 is obtained. There is no lateral strain due to constraint at the rectangular test section.&apos; If the absolute value of strain undergone between a maximum state of tension and maximum state of compression is called A, it will be seen that A decreases with cycling rapidly due to the strain hardening in the test section and the ensuing change in the bent beam profile. It will also be seen that A reaches an approximately stable value around 10th cycle. If we call the first cycle strain range, A,, and that of 10th cycle, A,,, then is an index of the strain hardening capacity at a particular deflection level. A plot of n vs ? is reproduced in Fig. 3 for two ASTM steels: A-302 and A-201. A-302 is a low alloy structural steel, A-201 a low-carbon steel. The specimens were first normalized at 1650°F for 75 min then grooved with a milling cutter; and the notches were ground to 50G finish. The final operation before testing was a stress relief at 1150°F for 75 min. The behavior of the two steels changes around ?1 = 0.8 pct or A,, = A, -n = 0.7. If these curves are compared with the endurance life curves given in Fig. 4, it will be seen that the relative performance of the two steels also change around A,, =. 0.70 pct. The steel with lower strain hardening capacity (7) at a certain strain level shows a longer endurance life, irrespective of their strength levels. Since the exponent governs strain hardening at lower strain levels than the Modulus, the relative values of both parameters also coincide with the change in n values and the relative performances of the two steels at different strain levels. These two observations indicate to the close relationship be-

Jan 1, 1962

By G. B. Gould

FROM the very inception of the price-control experiment in the bituminous-coal industry, the problem of price differentials was of major importance. In fact, assuming that there will be no legal or Govern- mental obstacles to price control, the successful maintenance of the present or any other method designed to control bituminous-coal prices will depend upon the discovery of some satisfactory way to determine these differentials in price, among: individual mines and groups of mines. Being of organic origin, bituminous coal is a commodity that presents an almost infinite variety of qualities. It is inevitable, therefore, that a work- able price-control system must attempt

Jan 1, 1935

By R. M. Goldhoff, J. W. Spretnak, R. Speiser

IT has been observed by Grange and Garvey&apos; that the homogenization of boron-treated steels could lead to complete elimination of the hardenability effect caused by boron. The experimental conditions leading to this conclusion involved the heating of steel specimens with and without boron at 2350°F for 24 hr. To prevent oxidation and decarburization, the specimens were encapsulated under vacuum in silica tubes. This observation of permanent deterioration of boron effect was noted in both a 0.63 pct plain carbon steel and a 0.25 pct C low alloy steel, both of which had been treated with ferroboron and homogenized at 2350°F for 24 hr. In another series of boron-treated steels which had been homogenized for shorter times, the effect was to reduce, but not completely eliminate, the boron effect. Under these circumstances, it was presumed that longer heating times would cause permanent deterioration of the boron effect in this series of steels also. Since the characteristic grain boundary constituent could not be formed in any of these steels after homogenization, it was concluded that possibly boron was being converted into an ineffective form through chemical combination resulting from prolonged heating at high temperatures. Assuming the validity of this observation, a series of experiments to establish the relationship between temperature of homogenization and kinetics of deterioration was contemplated. Initially an attempt was made to reproduce the results just described. For this purpose commercial lots of AISI 8640 and AISI 86B40 steels were used. The analyses of these steels appear in Table I. Hardenability evaluation was on the basis of the hardness of normalized sections. Grange and Garvey&apos; pointed out that these steels are air hardening to some extent and such a technique can be used. For these tests the specimens were held for 20 min in a salt bath at 1600°F and then cooled in still air. Several evacuated silica capsules containing pairs of the test steels were made up and heated for 24 hr at 2350°F. A number of failures occurred among these capsules, but evaluation of the specimens successfully treated indicated a complete deterioration of the boron effect. However, examination of the surface of the treated specimens showed that some reaction, probably with the silica tube, had taken place, since none of the specimens was bright. One further test was made at 2350°F for 12 hr, using a pair of specimens wrapped in tantalum foil prior to encapsulation. These specimens appeared bright after treatment, and their hardenability evaluation showed that no deterioration had taken place. Since these steels are extremely sensitive to de-boronization at high temperatures and particularly at long times of holding, the experimental observations raised some doubt as to the real occurrence of the permanent deterioration effect. To resolve this question an experiment was undertaken in which the high temperature portion of the treatment was completed in vacuum. For this purpose a molybdenum wound resistance furnace capable of maintaining good vacuum was used. All the pertinent accessories to this furnace were machined from molybdenum stock in order to withstand the high temperature involved in the experiment. With this equipment a temperature of 2400 °F was obtained and held under a pressure of 5x10-" mm Hg for 24 hr. A specimen of each of the two test steels was suspended in the furnace and both were heated simultaneously. Temperature was measured with a standardized Pt—Pt-13 pct Rh thermocouple, whose millivoltage at 2400°F was extrapolated from the standardization equation, good up to 2200°F. The complete evaluation of these specimens is shown in Table 11. The hardness values attained by each steel after normalizing, both before and after the prolonged high temperature treatment, remain the same. A microscopic study of the decomposition products present in the microstructures of specimens at various points in the heat treatment correlated with the hardness values observed. It seems that no permanent loss of boron hardenability effect occurs, at least in these Grainal-treated steels, when deboronization is circumvented. Acknowledgments This work was conducted under the sponsorship of the Aeronautical Research Laboratory, Wright Air Development Center, under Contract No. AF 18(600)-94, and their support and permission to publish is gratefully acknowledged. A note of thanks is also extended to the Republic Steel Corp. for supplying the materials used in the work. References &apos; R. A. Grange and T. M. Garvey: Factors Affecting the Harden-ability of Boron Treated Steels. Tro.ns. ASM, 1946, vol. 37, p. 136.

Jan 1, 1957

By A. W. Schlechten, W. J. Kroll

As this paper is written, the only method for the commercial production of lithium metal is by the fusion electrolysis of LiC1-KC1 mixtures, as first proposed by Gunkz.2 The details of the industrial process have not been made public hut Osborg8 stated that an efficiency of more than 90 pct is obtained with a lithium metal recovery above 95 pct and that the metal is 99.5 pct pure. Pletenev9 gave a power consumption of 66 kw-hr per kilogram of lithium metal produced with a salt consumption of 9 kg of LiCl and 0.4 kg of KCI. Pletenev&apos;s figures indicate a somewhat lower recovery than that given by Osborg. In a recent report on&apos; German lithium plants, Motock7 gives figures that are also less favorable than those reported by Osborg. At the German plants the lithium recovery in electrolysis was only 83.4 pct, with a power consumption of 140 kw-hr per kilogram of lithium. The metal averaged only 97 pct pure, with sodium and potassium as the chief impurities. The presence of SO2, SO3, SiO2, Ba, Ca, and appreciable amounts of Na and Fe2O3 were especially disturbing in the electrolysis, and it was necessary to run the cell for some time to eliminate most of these impurities and condition the bath. It may be noted here that most of these impurities do not interfere with a thermal reduction process. These facts show that the present method of producing lithium is fairly efficient and would suggest that other methods could not compete. However, fusion electrolysis has certain drawbacks, such as the necessity of supplying costly, low-voltage direct current and of collecting and processing the anodic chlorine if it is not to be wasted. An even greater disadvantage is that lithium chloride must be used for an electrolyte. It is one of the most expensive salts of lithium, since it cannot be produced in anhydrous form by precipitation from aqueous solution, as is done with LiF or Li2CO3. The reduction of lithium by a vacuum process similar to the ferrosilicon method for the production of magnesium would have certain advantages. Small units such as Pidgeon retorts, would give flexibility of production to meet demand, and a small plant would require moderate capital outlay. There is a possibility of using idle Government-owned magnesium plants for such purposes. The raw materials needed for vacuum thermal reduction would not have to be extremely pure, Director, with respect to the iron and calcium Content. This paper describes a series of experiments to determine the effectiveness of various agents for the reduction of lithium compounds. The investigation is part of a program of vacuum metallurgy con-

Jan 1, 1949

By Louise Reber

THE town of Jerome is located in Yavapai County in north central Arizona. It has a population of over 6000 people and the two important mines of the district, the United Verde and the United Verde Extension, lie almost within its limits. Perched high on the southwest side of the Verde Valley, at elevations of 4800 and 5500 ft.. (1463 and 1676 m.), both mines have long tunnels from their lower levels through which ore is hauled to their respective smelters. The two smelter towns of Clarkdale and Verde, out in. the center of the valley and about 2000 ft. (609 m.) below, form part of the wide panorama that meets the eye at Jerome. This panorama and the topographic and surface features and the geologic formations are shown in a separate pamphlet. In the vicinity of Jerome, the great southern escarpment of the Arizona plateau forms the northeast side of the Verde valley. The general level of this escarpment, or "the rim" as it is called, is about 4000 ft. above the Verde River. A corresponding elevation is reached only by the, highest parts of the Black Hills, which bound the valley to the southwest. The two lava-capped mesas, known as Mingus and Woodschute Mountains, form the culmination of the Black Hills. That part of the old Black Hills mining district which lies between their crests -and the Verde River is now known as the Jerome or Verde copper district. However, if the smelters are not included, a strip about 7 mi. long and 3 mi. wide (11 by 5 km.) embraces practically all the properties belonging to the Jerome district. This strip lies parallel to and along the side of the Verde valley from about First View, 2 mi. northwest of Jerome, to the neighborhood of Allen Spring, near the south end of Mingus. Mountain. The rank of the district in copper production, as well as features of peculiar geologic interest in connection with the ore occurrences, entitle the district to much more consideration than it has received in the literature of copper mining. Arizona stands preeminent as a copper-producing state. In 1918, it produced four-tenths of the total amount produced in the United States, and nearly equaled the, combined production of its three chief competitors, Montana, Michigan, and Utah. The production of the Jerome district places it in the sixth place among the copper

Jan 8, 1920

Jan 1, 1963

NEW MEMBERS The following list comprises the names of those persons who became members during the period Oct. 10 to Nov. 10, 1914: Members ADDAMS, CHARLES EDWARD, Asst. Supt.. . Ray Consolidated Copper Co., Ray, Ariz. ALEXANDER, HARRY H., Asst. Mgr. and Mgr., American Smelting & Refining Co., Maurer, N. J. ALLING, MARK,N., Min Engr Downieville, Sierra Co., Cal. ANDERSON, LESLIE DOUGLASS, Supt U. S. Smelting Co., Midvale, Utah. ATKINS, WILLIAM, JR., Sec&apos;y Kingwood Coal Co., Howesville, W. Va. BLAKE, TRUE WALTER, Min. Engr Rombauer Coal Co., Novinger, Mo. BRACKETT, GEORGE SYLVESTER, Supt... Pittsvein Coal Co., Flemington, W. Va. BRAINERD, HOWARD STANFORD, Met Ingersoll-Rand Co., Phillipsburg, N. J. BROWN, THOMAS J., Genl. Supt., Nova Scotia Steel & Coal Co., Sydney Mines, N. S., Canada.

Jan 12, 1914

R. W. Raymond, New York, N. Y.: Mr. Bain&apos;s introductory statement that there is a growing disposition on the part of the public " to change emphasis " in government from political equality " to " social justice " is somewhat vague; but I gather from his argument that he approves the growing disposition to which he refers, and that the " emphasis " which it demands involves a radical change in the policy hitherto pursued by the United States—the policy under which the Pacific States and Territories have been developed thus far. I do not understand how "social justice" is to be furthered by the sacrifice of " political equality," or how the change is to be effected by "emphasis." But I fancy that the beneficent and necessary

Jan 1, 1913