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By Charles Legrand

The problem of mine pumping is so much affected by local conditions, and those conditions are so liable to changes during the life of a mine, that the best system to use is difficult to determine. The experience of the writer has been that, in general, for copper mines, electric pumping is most satisfactory unless the quantity of water to be pumped is great or the mine does not use electric power for any other purpose. As a rule, however, figuring on the cost of necessary power plant, the total cost of installation is greater with electric pumps than steam pumps. As the water is liable to be gritty, outside-packed plunger pumps are to be recommended, and if the water is not acid, or only slightly so, chilled cast-iron plungers pay for their extra cost very quickly in diminished cost of packing. With chilled plungers it is possible to use metallic packing instead of soft packing, if the water is not too gritty or the lift too high. In vertical pumps this packing is satisfactory for lifts up to 400 ft., in the experience of the writer, and may be found satisfactory for higher lifts. For water slightly acid, cement-lined pump bodies have given good satisfaction. For acid water, both plungers and puhp bodies should be made of acid-resisting bronze. For high lifts, especially with gritty water, a satisfactory pump valve is difficult to find. To reduce the unbalanced pressure on the valve at time of opening it is advisable to have narrow seats; this, however, brings high pressures on the seats when the valve is closed. On large valves for lifts of 600 to 1,000 ft. we have used leather, hard rubber, vulcanized fiber, and soft brass; for clear water the latter has given best satisfaction. The difficulties due to grit make it advisable to have large sumps where water can settle, and if possible, to have two of them so that they can be cleaned alternately. At the mines of the Old Dominion Copper Mining & Smelting Co., Globe, it was found that a small air ejector discharging into a mine car provides a very convenient way to clean mud and sand from a sump.

Jan 1, 1916

By N. F. Fiore, M. C. Mangalick

DISLOCATION damping in single crystals is orientation dependent.&apos; If dislocations are distributed uniformly over all the slip systems in a crystal, and if each dislocation contributes identically to the damping process, then a damping loss or decrement 6 arises. In fact, however, for a crystal of axial direction (hkl), dislocations on certain slip systems are oriented to make a larger contribution to the damping loss than are dislocations on other slip systems. Thus the measuring damping, 6 (hkl) is proportional to an orientation factor (hkl), which weights the ability of each slip system in the crystal to contribute to the total damping. If it is assumed that dislocations are distributed uniformly on all slip systems, then for a given orientation of an fcc crystal this factor is determined2,3 by the three nonzero elastic constants Cll, C12, and C44 and the density p. In this communication measured values of elastic constants for two Fe-18Cr-Ni alloys are reported, and the dislocation damping results on two different orientations of each alloy are analyzed for orientation dependence. Elastic constants were determined by measuring the ultrasonic velocities of one longitudinal and two shear waves in (110) direction of the fcc crystals.4 Variation of velocities with the change in orientation is very small for (110) orientation,3 and so this direction is ideal for orientation factor and comparative damping measurements. Experimental Procedures. The nominal compositions of the alloys studied are given in Table I. These alloys were vacuum-melted and vacuum-cast from high-purity materials at U.S. Steel Applied Research Laboratories. The single crystals were grown by the modified Bridgeman technique. The specimens were homogenized at 1350°C for 20 hr, cut to shape and size with a spark cutter, and then annealed at 1100°C for 20 hr prior to use. In the KHz frequency range, the Marx composite piezoelectric oscillator was used to obtain the damping. The ultrasonic pulse technique was used to measure the ultrasonic velocities and the attenuation in the MHz frequency range. M.c.mANGALICK,-- StudentMemberAIME,formerly Research Assistant, University of Notre Dame, is now with Casting Laboratory, Chase Brass and Copper Co., Cleveland, Ohio. N. F. FIORE, Associate Member AIME, is Assistant Professor, Department of Metallurgical Engineering and Materials Science. University of Notre Dame, Notre Dame, Ind. This communication is part of a Ph.D. thesis presented by M. C. MANGALICK to the Graduate School of the University of Notre Dame. Manuscript submitted May 20. 1968. IMD Elastic Constants. Table I gives the values of the ultrasonic velocities. Values of C11, C12, and C44 calculated from these velocities are also given in Table I, and it is evident that there is very little variation in C11, C12, and C44 with nickel content. This slight variation may be attributed to inaccuracies in velocity measurements and practical differences in the orientations of the specimens. For example, in B the time interval for vl was 2.4 ± 0.05 sec giving an error in v1 of 0.235 x or 0.005 in. per sec. The elastic constants obtained by pulse technique were tested by calculating E<110>, Young&apos;s modulus for the (110) direction, from the equation: This value was compared with the value computed from the resonance bar measurements in which where fr is the resonance frequency and I, is the length of the specimen. In all cases, the values calculated from the ultrasonic velocity measurements

Jan 1, 1969

By H. Ries

SPECIAL sands are those that are employed for special purposes. They have a limited use, as compared with sands for concrete and plaster. The sands discussed in this chapter are those used for foundry purposes, glass manufacture, filter beds, abrasive purposes, sandblasting, engine sand, sand for sheet asphalt, and furnace sand. These sands are almost invariably high in silica, consisting chiefly of quartz grains. In some-for instance, glass sands-freedom from ingredients other than silica, such as iron oxide, is important. Certain grades, like filter sand, require rather careful sizing, and in foundry sands the bond, usually clay, is important. One deposit may sometimes be used to supply several grades or types of sand by appropriate methods of screening and washing. The special sands discussed here are usually obtained from deposits of unconsolidated material but sands high in silica are derived occasionally from soft sand- stones that can be disintegrated easily. Requirements are not the same for all uses, and, while certain properties may be demanded, there are no officially adopted standard tests except for foundry sands. FOUNDRY SANDS Definition-Foundry sands include those siliceous sands that are used to make the forms for casting metals. If these are employed for making molds, they are known as molding sands; if for the cores that form the hollow spaces in the casting, they are called core sands. Naturally bonded sands are those containing a variable, though not usually large, amount of clay, which serves as a bond. Synthetic sands, now much used, represent an artificial mixture of sand with usually fire clay or bentonite. Core sands may be bonded with oil, cereal binders, resin, pitch, or other materials. The foundrymen apply the term "sharp sand" to one that is free from bond. Properties-The important properties to be tested at room temperature are texture or fineness, permeability, strength, deformation, and flowability. Those tested at elevated temperatures are compressive strength, expansion and contraction, sintering point, and durability. More attention probably has been given to research on the properties

Jan 1, 1949

By R. V. Smith

The theory for unsteady radial flow of gas, as simplified by Aronofsky and Jenkins, has been reviewed and compared with actual well performance. This study indicated that the theory should be modified by the introduction of an empirical "rate of flow" function "Y(q)". The expansion of the theory on the flow of gas to include a Y(q) term bridges the gap between the theory of unsteady-state flow of gas and actual gas-well behavior. Apparently, the Y(q) term is a function only of the rate of flow for a given well. The completion factor or skin effect is associated with the Y(q) function in such a manner that at least two or more sets of drawdown or pressure build-up tests are needed to separate the Y(q) function and the completion factor. Since the Y(q) function used in this report represents an energy loss dependent on rate of flow near the well-bore of a gas well which is in addition to the loss required by Darcy&apos;s law, the Y(q) function is related to the exponent of the back-pressure curve for a gas well. Techniques and equations are presented which permit the estimation of stabilized and "short-time" deliver-abilities. INTRODUCTION The published theory of radial unsteady-state flow of gas through porous media may be divided into papers1-" which use Darcy&apos;s law as a premise and other papers4-7,16-10 which recognize that Darcy&apos;s law may not hold for flow into gas wells. Miller8 compared theory based on Darcy&apos;s law with actual gas-well behavior and concluded that flow into gas wells deviates from Darcy&apos;s law. Houpeurt4,5 in a thorough investigation of flow into gas wells concluded that the flow rate into a gas well is proportional to the difference between the squares of upstream and downstream pressures elevated to a power between 1.0 and 0.5. Houpeurt also concluded that the deviation from Darcy&apos;s law is caused by irreversible, kinetic energy exchange between the flowing gas and the porous media. Tek9 presented a complete evaluation of the flow problem in gas wells. Several investigators state that the inclusion of gas compressibility into flow equations based on Darcy&apos;s law does not begin to explain the flow problems suggested by gas-well behavior. For many years, natural-gas engineers have realized that gas wells by actual test exhibit behavior which cannot be explained entirely by Darcy&apos;s law. Many tests on gas wells have indicated that a value of 0.85 for the exponent in the back-pressure equation is a fair average for many wells. However, Darcy&apos;s law either for steady or unsteady-state conditions indicates that the exponent should be 1.0. The purpose of this paper has been to examine this deviation from Darcy&apos;s law and to present a reasonable mathematical means of evaluating the deviation. The method presented in this paper combines previous theory with an empirical function to provide a description of gas-well behavior. All equations in this report are for constant terminal-rate-of-flow conditions. While present methods of testing wells usually are not based on this concept, they are believed to approach the constant terminal-rate conditions closely in many tests. However, several gas wells were tested by holding the rates of flow constant in order to approach more closely the constant terminal rate of flow; the results confirm the concepts presented in this paper. REVIEW OF THEORY FOR UNSTEADY-STATE FLOW OF GAS The basic partial differential equation for the flow of natural gas has been given in technical literature; the equation for a cylindrical reservoir with the well in the center is

By O. D. Sherby, J. E. Dorn, C. D. Wiseman

MCLEAN&apos; has shown that the total strain obtained during creep of aluminum polycrystals arises exclusively from the mechanisms of 1) microscopically observable slip, 2) subgrain tilting, and 3) grain boundary shearing. Extensive analyses, however, suggest that the high temperature creep of metal polycrystals is controlled by a single thermal activation process."" These somewhat contradictory observations might be reconciled if it could be shown that the activation energies are identical for each of the mechanisms isolated by McLean. Such identity of the activation energies, of course, would demand that the same unit process controls the functioning of each individual mechanism. Under these conditions the activation energies for creep of single metal crystals and polycrystalline aggregates and for grain boundary shearing should be identical. This paper is primarily concerned with the possible identity between the activation energies for high temperature creep of single crystal and polycrystalline metals. In addition, comparisons will also be made between the activation energies for creep of single crystals and polycrystals with previously reported activation energies for grain boundary shearing.0"8 Materials and Techniques The same high purity aluminum, lead, and tin were used in the creep tests for single crystals and polycrystalline aggregates. Single crystals of aluminum and tin were prepared by the strain-anneal technique, whereas lead single crystals were prepared by growth from the melt. The tensile creep stress was maintained constant by Chalmers-Andrade parabolic-type lever arms. Strains were measured to &0.0002 in. per in. with rack and pinion type extensometers, and temperatures were maintained to within & 0.2"C of the reported values by mercury-thermoregulated resistance-heated silicone oil baths. The previously reported unambiguous technique for determination of the activation energy for creep was employed throughout this investigation.&apos; ." It involves periodic abrupt small changes in temperature during the course of creep, accomplished by rapidly replacing one creep bath at temperature TI by a second bath at T,. The abrupt change in creep rate from i, to & attending the abrupt change in temperature is due exclusively to the change in temperature, since both the stress and the substructure are identical the instant before and the instant after a change of temperature. Therefore, all other factors being identical where AH is the activation energy for creep and R is the gas constant. Results Typical examples of the creep rate vs creep strain data during temperature cycled creep tests of single

Jan 1, 1958

By D. Mizer, J. B. Clark

In a recent investigation of the entire magnesium-yttrium phase diagram, Gibson and Carlsonl report the maximum solid solubility of yttrium in magnesium as 8.0 wt pct Y at 1050°F (565°C). However, independent studies of the precipitation in magnesium-yttrium alloys, 2 conducted prior to the publication of the above phase diagram, indicated a greater solid solubility of yttrium in magnesium. Accordingly, the magnesium-rich region of the diagram was re-investigated to determine more accurately the liqui-dus, solidus, and solvus of the magnesium solid solution in this system. Metallography and thermal analysis were chosen as the best corroboratory methods for accurately determining this region of the diagram. A magnesium-yttrium master alloy was prepared from yttrium of 99.9 pct purity and singly sublimed magnesium of 99.9 pct purity. Subsequently, the twelve magnesium-yttrium alloys shown in Fig. 1 were prepared using standard alloying procedures for magnesium. Metallographic samples of the alloys were heat treated under an SO, atmosphere in an electric furnace capable of control within ±3oF of the desired temperature. Alloy samples for the metallographic determination of the solvus, and for the precipitation studies were solution treated at 1000° F for 16 hr. Subsequent aging times for the solvus determination varied from a minimum of 6 hr for the higher temperatures up to 816 hr for lower temperatures. Periods of 30 min at temperature were used to generate liquation for the metallographic placement of the solidus and eutectic horizontal. The liquidus and the eutectic horizontal were determined by thermal analysis. Simple time-temper- ature heating and cooling curves were obtained from 10-g alloy samples heated and cooled in a well-insulated electric furnace. An Iron-Constantan thermocouple was placed in direct contact with the sample. The thermocouple did not contaminate the alloy sample since metallographic examination showed that the alloy melt did not wet or attack the thermocouple. The heating and cooling rates were approximately 5 to 8OF per min. The standard metallographic procedures for magnesium were used in polishing and etching the magnesium-yttrium alloys studied. Glycol etchant* was etchant was used to reveal precipitate in the solvus determination. The redetermined magnesium-rich region of the magnesium-yttrium diagram is shown in Fig. 1. The

Jan 1, 1962

By AIME

TO say that Milton Fies has been active in promoting the Southern Research Institute is a masterpiece of understatement. He is a director and trustee who was in on the ground floor when plans were first made for the Institute and since it became an actuality early in 1945 his interest and efforts continue. He talked about it at the last regional meeting of the Birmingham Section of the A.F.A. and has talked all over the South-and when he makes a speech he has really prepared it. He does not try to rival a political spellbinder, but his careful, concise English is always well thought out and to the point. Southern Research might be called his present hobby as it has been for the past two or three years. Another major interest has been his encouragement and training of engineering students. Active in the formation of the Student Chapter of the A.I.M.E. at the University of Alabama, for several years he saw to it that the students had a prominent engineer at each monthly meeting, usually taking them down to Tuscaloosa from Birmingham in his own automobile. He has also been keenly interested in the mining engineering activities at Tuscaloosa, both in the Bureau of Mines and the University.

Jan 1, 1946

By Regis M. N. Pelloux, N. J. Grant

Five or six alloys each in the systems Ni-C.v, Ni-Mo, and NL-W, spaced to cover the single phase areas as well as a part of the adjacent two-phase field, were prepared as uacuum-melted alloys. Tensile tests at room temperature and stress rupture tests at 1200° and 1500°F we.ve run for time periods to give fracture in 0.1 to 500 hr. Observations mere made of solid-solution and second-phase strengthening or uleakening, coupled with studies of ductility and the role of structure on the noted behavior patterns. THE development of stable creep-resistant alloys while broad and systematic is also highly empirical. A number of simple rules have been suggestedL-" but these are based on scant data. Whereas rules for low-temperature alloy strengthening can be applied with some success in more simple systems, the extension to behavior at high temperatures (greater than about 0.4 of the absolute melting temperature) is very poorly understood and has encountered important exceptions and deviations in terms of atom size or valency effects. In particular, the occurrence of both solid-solution weakening4,5 and second-phase weakening2 at high temperatures is of interest. MATERIALS AND PROCEDURE Nickel was chosen as the base material because of its important role in current high-temperature alloy applications, because of its freedom from a phase transformation, and because of wide solubilities for many elements of interest. The selected alloys were single- and two-phase structures in the Ni-Cr,6 Ni-Mo,7 and Ni-w 8 systems. The alloys were vacuum melted and cast, utilizing high-purity raw materials, in the form of 15 lb ingots. They were readily forged into 1/2-in. bar stock (except the two highest chromium alloys). The alloy compositions are shown in Table I. Typical impurity values are: Carbon: Max 0.03 pct (except for 30 and 35 pct Cr alloys which had 0.07) Nitrogen: Max 0.005 pct (except for two of the Cr alloys which had 0.05) Sulfur: Max 0.004 pct Silicon: Max 0.05 pct Each alloy was solution treated (air quenched) to produce comparable grain size. The two-phase alloys were then overaged in an effort to produce a stable structure for testing at 1200° and 1500°F, The results are summarized in Table 11. Following the heat treatments outlined in Table 11, the bars were machined to a gage section of 1.25-in. long by 0.250-in. diam. In addition to room-temperature tensile tests, stress-rupture tests were run at 1200" and 1500°F for time periods from 0.1 to about 500 hr. RESULTS Room-Temperature Mechanical Properties—The ultimate tensile strength, total elongation, and reduction of area values for all the alloys are plotted vs atomic percent of the solute element in Fig. 1. It can be seen that for the same amount of solute element the ultimate tensile strength increases in the order chromium, molybdenum, tungsten. Elongation increases (reduction of area decreases) with increasing amounts of solute for all the solid solution alloys, the values ranging from 50 to 70 pct. The ultimate tensile strength values for the two-

Jan 1, 1961

By Michael Epstein, Daniel E. Rosner

Fume nucleation sufficiently close to vaporizing suvfaces can augment net vaporization rates into cooler environments. Environmental conditions favoring large vaporization rate enhancements are briefly discussed and a previous theoretical treatment of this nucleation phenomenon is generalized to account for the self-regulating effect of condensalion-heat release within the boundary layer. Despite kinetic limitations on homogeneous nucleation, and latent heat release, non-diffusive condensate removal processes appear to make possible large enhancements in steady-state vaporization rates, provided surface temperatures are well below the boiling point. When condensed phases vaporize (or dissolve) into cooler media, the diffusion-limited mass loss rate can be strongly influenced by the process of nucleation/con-densation (or precipitation) within the thermal boundary layer. This condensation process (which typically leads to mists or fumes in the case of evaporation into cooler gases) has the effect of steepening the vapor pressure profiles near the evaporating surface, since the condensation zone acts as a vapor sink. However. the resulting enhancement in the diffusion-limited evaporation rate can be estimated (as first done by Turkdogan1 for the case of molten iron/nickel alloys evaporating into helium) only if one has independent knowledge of the critical supersaturation, sCrit(T), required to homogeneously nucleate the vapor.* In a recent reformulation and generalization of the theoretical model of Ref. 1 it has been shown that, when log sCrit is approximately linear in reciprocal temperature, rather simple expressions can be derived4 for the ratio of the actual rate of vaporization j" to either the minimum (no condensation) rate j"min, or the maximum (equilibrium condensation) rate j"max In the present communication we wish to briefly report on further developments and implications of the formulation of Ref. 4, with emphasis on i) environmental conditions favoring large enhancements in vaporization rate, and ii) the self-regulatory influence of condensation heat release (neglected in Refs. 1 to 4) on predicted vaporization rates. Additionally, we take this opportunity to correct several misprints appear- ing in Ref. 4, and comment on Elenbaas&apos;s recent criticism5 of Ref. 1. MAXIMUM POSSIBLE VAPORIZATION RATE IN PRESENCE OF CONDENSATION A nonequilibrium theory is of interest because of the very large difference between the minimum (no condensation) and maximum (equilibrium condensation) vaporization rate. The magnitude of this maximum possible enhancement can be shown quite clearly by combining a result of Refs. 3 and 4 with the fact that for most liquids there is a simple relationship between the molar heat of evaporation, A, and its boiling point, i.e., A/(RTBp) = C, where the constant C, often called the Trouton ratio, takes on values not very different from 11.* More generally, for any substance (including The Trouton ratio (which for water is 13, for methane, 10, and so forth) will be recognized as the ratio of the molar entropy change upon vaporization (at TBP or Ttransf) to the unlversal gas constant R. Its near constancy reflects the fact that the change in atomic order upon vaporization depends only weakly on the kinds of molecules involved. those that sublime under ordinary conditions) we can define a characteristic transformation temperature. Ttransf, by a relation of the form Ttransf =A/(CR), and then examine the maximum possible evaporation rates as a function of how far removed from Ttransf are the surface temperature, Tw, and ambient temperature, T. Subject to the assumptions: 1) equilibrium vapor pressure, pv,eq, everywhere small compared to prevailing total pressure, p, and 2) negligible effect of condensation heat on temperature profile, the maximum enhancement ratio was found (Eq. [17], Ref. 4) to be: where, for most vapors, Nu/NuD (the ratio of heat transfer coefficient to mass transfer coefficient for the same configuration) is a number near unity.* Ex- *An alternative derivation of the Nu = NuD special case of this equation. revealing its validity for arbitrary velocity/temperature profiles in a laminar boundary layer, is given in Ref. 3. amining this result for a "Trouton substance", one obtains the results shown in Fig. 1, constructed for C = 11. Since we are concerned with vaporization enhancements (j&apos;max/J"min > 1) at surface temperatures below Ttransf, this area of interest is shown unshaded. One notes that at a fixed ambient temperature (hence, T/TtranSf) there is a unique surface temperature, 2T , at which j"max/j"min attains its peak value; moreover, the peak enhancement ratio, see dashed locus. Fig. 1, is: (NuA/NuD)(C/4)(Ttransf/T,). Hence, if Nu = NuD, when the ambient temperature is less than 1/4 of TtranSf the peak enhancement exceeds the Trouton

Jan 1, 1969

By E. W. Filer, L. S. Darker

ONE of the primary purposes of this investigation was to determine how far blast-furnace metal and slag depart from equilibrium, particularly with respect to sulphur distribution. In studying the equilibrium between blast-furnace metal and slag, there are two approaches that can be used. One method is to use synthetic slags, as was done by Hatch and Chipman;&apos; the other is to equilibrate the metal and slag from the blast furnace by remelting in the laboratory. In the set of experiments here reported, metal and slag tapped simultaneously from the same blast furnace were used for all the runs. The experiments were divided into two groups: 1—a time series at each of three different temperatures to determine the t.ime required for metal and slag to equilibrate in various respects under the experimental conditions of remelting, and 2—an addition series to determine the effect of additions to the slag on the equilibrium between the metal and slag. An atmosphere of carbon monoxide was used to simulate blastfurnace conditions. The furnace used for this investigation was a vertically mounted tubular Globar type with two concentric porcelain tubes inside the heating element. The control couple was located between the two porcelain tubes. The carbon monoxide atmosphere was introduced through a mercury seal at the bottom of the inner tube. On top, a glass head (with ground joint) provided access for samples and a long outlet tube prevented air from sucking back into the furnace. The charge used was iron 6 g, slag 5 g for the time series, or iron 9 g, slag 7 % g for the addition series. This slag-to-metal ratio of 0.83 approximates the average for blast-furnace practice, which commonly ranges from about 0.6 to 1.1. A crucible of AUC graphite containing the above charge was suspended by a molybdenum wire in the head and, after flush, was lowered to the center of the furnace as shown in Fig. 1. The cylindrical crucible was 2 in. long x % in. OD. The furnace was held within &3"C of the desired temperature for all the runs. The temperature was checked after the end of each run by flushing the inner tube with air and placing a platinum-platinum-10 pct rhodium thermocouple in the position previously occupied by the crucible; the temperature of the majority of the runs was much closer than the deviation specified above. The couple was checked against a standard couple which had been calibrated at the gold and palladium points, and against a Bureau of Standards couple. The carbon monoxide atmosphere was prepared by passing COz over granular graphite at about 1200°C. It was purified by bubbling through a 30 pct aqueous solution of potassium hydroxide and passing through ascarite and phosphorus pentoxide. The train and connections were all glass except for a few butt joints where rubber tubing was used for flexibility. The rate of gas flow was 25 to 40 cc per min. As atmospheric pressure prevailed in the furnace, the pressure of carbon monoxide was only slightly higher than the partial pressure thereof in the bosh and hearth zones of a blast furnace—by virtue of the elevated total pressure therein. Simultaneous samples of blast-furnace metal and slag were taken for these remelting experiments. The composition of each is given in the first line of Table I. There is considerable uncertainty as to the significant temperature in a blast furnace at which to compare experimental results. This uncertainty arises not only from lack of temperature measurements in the furnace, but also from lack of knowledge of the zone where the slag-metal reactions occur. (Do they occur principally at the slag-metal interface in the crucible, or as the metal is descending through the slag, or even higher as slag and metal are splashing over the coke?) The known temperatures are those of the metal at cast, which averages about 2600°F, and of the cast or flush slag, which is usually about 100°F hotter. To bridge this uncertainty, remelting temperatures were chosen as 1400°, 1500" (2732°F), and 1600°C. For the time series the duration of remelt was 1, 2, 4, 8, 17, or 66 hr; crucible and contents were quenched in brine. The addition series were quenched by rapidly transferring the crucible and contents from the furnace to a close-fitting copper "mold." Of incidental interest here is the fact that the slag wet the crucible

Jan 1, 1953

By Pol Duwez, Spencer R. Baen

A LTHOUGH the practical importance of Fe-Cr--iV Mo alloys has long been recognized, constitution studies have been limited to a few alloys within rather narrow ranges of composition. The purpose of the present study was to establish the phase boundaries in the ternary diagram at a temperature of 1200°F (650°C).* The identification of phases and the determination of phase boundaries were based primarily on X-ray diffraction measurements and occasionally on microscopic observation. The methods used for locating the phase boundaries were essentially those described in detail by Andersen and Jette,&apos; and by others.&apos;, The compositions of 180 alloys, prepared by powder metallurgy techniques, are shown in Fig. 1. The chemical analysis and mesh size of the powders which were used for the preparation of the alloys are given in Table I. The powders were mixed for 48 hr or longer, then compacted under 80,000 psi in a %-in. diam steel die. Each specimen weighed approximately 5 g. The compacts were sintered for 4 hr at 2500°F (1371°C) in an atmosphere of pure dry hydrogen (dew point near the temperature of liquid nitrogen). An open tray of chromium powder was placed upstream of the compacts to act as a "getter" for the last traces of oxygen or water vapor. In many cases this chromium powder was bright after sintering. Occasionally the upstream side of the powder was slightly greenish, but in all cases the downstream side was metallic. Some of the alloys contracted during sintering and some expanded as much as 20 pct. The latter alloys, which in all cases were brittle, were crushed in a tool steel mortar, repressed, and resintered for 4 hr at 2500°F. After sintering, the alloys were aged for ten days at 1200°F (650°C) in evacuated (105 mm Hg or better) fused silica tubes and were cooled rapidly to room temperature following the aging treatment. The use of the powder metallurgy technique in preparing alloys for structural investigation offers a simple method &apos;of avoiding contamination or non-homogeniety which could result from a melting operation. It is very often possible to dispense with chemical analysis of the alloys which have been prepared by the powder metallurgy method. In lieu of chemical analysis, it is necessary to determine the weight of the compact before and after sintering. The chemical composition of the alloy may be considered as that of the green compact, within the limits of accuracy determined by the loss of weight during sintering. In phase-diagram studies it is important to achieve homogeniety in structure, whether the alloys are prepared by powder metallurgy or by melting. Homogeniety means complete diffusion during the sintering of the alloys which are prepared by powder metallurgy. It has been found from experience that the best criterion for establishing the completeness of diffusion is the sharpness of the X-ray diffraction patterns. This criterion has been used throughout the present study. The X-ray diffraction data were obtained using powder cameras of either 214.86 or 71.62 mm diam. Copper, cobalt, or chromium Ka. radiation was used, the choice being based on the alloy composition. The powder samples were obtained by filing or crushing, depending on whether the alloy was ductile or brittle. In some cases, the powder was annealed for 4 to 6 hr at 1200°F in vacuum. The measurements were corrected for film shrinkage and absorption. The large camera, in which the film is mounted

Jan 1, 1952

By David Cornell, Yusuf K. Sukkar

The fundamental differential equation for fluid flow has been rearranged, integrated numerically for natural gases, and the results presented in tabular and graphical form suitable for the direct calculation of vertical gas flow problems. The integration is rigorous except for the use of a constant average temperature. Both static and flowing bottom-hole pressures may be calculated by (his method without resort to trial and error procedures. INTRODUCTION Bottom-hole pressures can be determined either by direct measurement with a bottom-hole pressure gauge or by calculation. Since measurement with a bottom-hole pressure gauge is costly, calculation of bottom-hole pressures is to be preferred when possible. Calculation involves knowledge of the wellhead pressure, the properties of the gas, the depth of the well, the flow rate, the temperature of the gas, and the size of the flow line. These quantities must be combined in a suitable equation for vertical flow. Several equations have been developed for the purpose of calculating gas flow through both horizontal and vertical pipes1,2,4,5,6,7,8,9 These equations all have the basic differential equation for flow as their starting point. They differ only in the assumptions which are made in order to simplify the integration step. The bottom-hole pressures calculated using any of the better known equations are in good agreement with bottom-hole pressures calculated by a more rigorous but considerably more tedious stepwise integration method. All of these methods, however. require a trial and error solution for the calculation of flowing bottom-hole pressures. The purpose of this paper is to present a method which provides a direct solution of static and flowing bottom-hole pressures without use of a trial and error solution. At the same time the method involves fewer simplifying assumptions than any of the previous methods used, except lor the lengthy stepwise method. The derivation of any practical flow equation begins with the basic differential equation which is rigorous for almost all flow problems. The five terms in Equation 1 represent the potential energy, the potential energy of position, the kinetic energy, the shaft work done, and the energy losses due to motion of the fluid respectively. Equation 1 reduces to Bernoulli&apos;s theorem for the case of horizontal flow of an incompressible fluid, no work done, and no energy losses due to friction. In the present case, the term (v) (dv)/(g,.) can be shown to be quite negligible compared to the other terms, and is therefore neglected in the derivation. The term — dws includes work done by a pump or turbine in the system and is zero in this case. The frictional losses in the pipe are included in the term — d(lw) which is expressed in terms of the Moody friction factor as in Equation 2. (f) (v2) (dL) d(lW) = (f)(V2) (dl)/(2) (g) (d) The specific volume of a gas can be expressed in terms of the temperature, pressure, molecular weight, and compressibility factor. The velocity of the gas can be expressed in terms of the weight rate of flow, the specific volume, and the dimensions of the pipe.The reduced pressure Pr is to be used because z can be expressed as a function of Pr and Tr alone for most natural gases:

Jan 1, 1956

By T. W. Watson, D. R. Flynn

,Measurements are reported for the thermal conductivity of four heat-resistant alloys which were measured at the National Bureau of Standards. Data are given for two samples of AISI Type 316 stainless steel covering a combined temperature range from -150° to + 800°C. The thermal conductivities of three samples of AISI Type 347 were measured over the temperature range from -150° to +150C. Thermal conductivity values are given for Carpenter 20 stainless steel from 100" to 800" C. Data are given for Inconel 702, in both the solution-annealed and the age-hardened condition, over the temperature range from -150" to +800C. SINCE 1958 the National Bureau of Standards, at the request of other government agencies or private laboratories, has carried out measurements of the thermal conductivity of a number of metals and alloys. This paper presents available thermal conductivity data obtained in this way on samples of four heat-resistant alloys. The data presented for Carpenter 20* stainless steel are, to the best of our knowledge, the only thermal conductivity values which have been reported for this alloy to date. Although thermal conductivity values for AISI stainless steels 316 and 347 have been reported by previous investigators, the present data are believed to be of particular interest because, for each alloy, the effect on thermal conductivity of small changes in chemical composition can be observed. The data on Inconel 702&apos; are of interest, not only for

Jan 1, 1969

By A. S. Nowick, C. Y. Li

It is deMonstrated that quenching from an elevated temperataupe accelerates the grain boundary relaxation in two solid solutions (aAg-Zn and a Cu-Al). This result is consistent with the proposal that, in solid solutions, grain boundary relaxation occurs by a mechanism of&apos; self diffusion. Nevertheless, an alternative possibilitg, that quenching introduces vacancies into the boundary itself, must also be considered. THE phcnomenon of grain boundary relaxation has been well known for many years,1,2 yet the mechanism of this process is very poorly understood. One of the most interesting suggestions which relates to the mechanism of grain boundary relaxation was that of Ke,3 who claimed that the activation energy for grain boundary relaxation and for lattice self diffusion were essentially the same. The implication is therefore that the elementary step in the two processes is the same. This suggestion is particularly startling in view of the fact that activation energy for self diffusion along a grain boundary is very significantly lower than that for volume self-diffusion. Later evidence5-7 showed that there really are two grain boundary peaks, one which appears in high-purity metals, and the other (which develops at a higher temperature than the first) which appears in solid solutions beginning at solute concentrations in the range of 0.1 pct. Data for silver6 show that Kg&apos;s hypothesis is surely incorrect for the grain boundary peak in the high-purity metal, since it has an activation energy of only 22 kcal per mole, but that the hypothesis may still be correct for the grain boundary peak in various silver solid solutions, for which activation energies in the range 40 to 50 kcal per mole are observed. If the elementary step in the grain boundary relaxation process were the same as that for self-diffusion, it would be expected that the relaxation process could be hastened by quenching, 2.c. by introducing a non-equilibrium excess of lattice vacancies. Such a quenching effect has already been demonstrated in the case of another anelastic relaxation process, viz., the Zener relaxation effect. The Zener effect, which occurs in essentially all solid solutions, may be attributed to the reorientation of pairs of solute atoms in the presence of an applied shear stress,&apos; and therefore must take place by means of a volume diffusion mechanism. The hastening of this process through quenching9 has been one way of demonstrating that atom movements in the lattice take place through a defect mechanism, presumably single vacancies. In order to see if the grain boundary relaxation is affected by quenching, it is particularly convenient to compare the grain boundary relaxation with the Zener effect, by choosing a specimen for which both relaxation effects appear. Specifically, a fine-grained sample of a solid solution shows in the curve of internal friction vs temperature, first a peak due to the Zener effect, then a second rise (and eventually a peak at substantially higher temperatures) due to the grain boundary relaxation. The same phenomena are also observable in static anelastic measurements, such as creep at very low stress levels. Thus, for the same fine-grained solid solution, the creepstrain, when plotted against log time, falls on a sigmodial curve with a sharp inflection point, due to the Zener effect, which is followed by a second rise and inflection resulting from the grain boundary relaxation. To look for a quenching effect, static measurements are preferable to the dynamic internal friction measurements, due to the fact that quenching effects tend to anneal out too rapidly at the temperatures at which the internal friction is measured.9 RESULTS AND DISCUSSION Creep experiments in torsion were carried out in an apparatus similar to that described by Ke1, whereby a wire is held under constant torque and its angular displacement is observed as a function of time. The alloy Ag-30 at. pct Znwas selected because of the large Zener relaxation that it displays. The two samples used were a "coarse grained" wire with a mean grain size about twice the diameter of the wire (diam = 0.032 in.), and a "fine-grained" wire which had several grains across the diameter. In Fig. 1 a comparison is made of the creep curves at 160°C of these two samples after they had been cooled slowly from 400°C. Curve A, which represents the coarsegrained sample, shows a unique relaxation process due to the Zener relaxation, with a relaxation time, T , in the vicinity of 100 sec. Curve B, which represents the behavior of the fine-grained sample, on the other hand, shows first the same relaxation process as that in A, followed by a turning up of the curve which corresponds to the onset of a second overlap-

Jan 1, 1962

By C. E. Armantrout, J. A. Rowland, D. F. Walsh

THE close-packed-hexagonal structure of mag-J- nesium is converted to a ductile and malleable body-centered-cubic lattice by the addition of lithium in excess of 10 pct. Further, the density of magnesium or magnesium-base alloys is decreased by additions of lithium. The practical possibilities of such alloys as a basis for uniquely light, malleable, and ductile structural materials were pointed out by Dean in 1944&apos; and by Hume-Rothery in 1945.2 It was apparent to these investigators, however, that more complex compositions would be required if strengths sufficient for structural applications were to be developed in these alloys. In a search for strengthening additions, various investigators w have examined a number of the ternary and more complex alloys containing magnesium and lithium. An investigation of the fundamental characteristics of these alloys was undertaken by the Bureau of Mines. The investigation was initiated with a study of the magnesium-rich corner of the equilibrium diagram for the ternary system, Mg-Li-Al. The following data from published investigations of Mg-Li-A1 alloys were available: 1—a description of isothermal sections at 20" and 400°C through the Mg-Li-A1 constitution diagram by F. I. Shamrai;&apos; 2—a diagram by P. D. Frost et al." showing approximate phase relationships at 700°F for a number of the Mg-Li-A1 alloys; and 3—diagrams showing the constitution at 500" and 700°F for the Mg-Li-A1 alloy system published by A. Jones et al.&apos; Where compositions and temperatures permit comparison, these diagrams show disagreement. The 700°F isotherms of Frost and Jones differ only in the placement of the phase boundaries. But Sham-rai&apos;s 400°C (752°F) isotherm shows a variation in phases as well as in phase boundaries. Although rigid comparison of these different isothermal sections might not be justifiable, it seems impossible to reconcile Shamrai&apos;s construction with the isotherms of Frost or Jones. The isothermal sections presented in this paper were prepared to determine compositions which might be suitable for age hardening and to develop the general slope and placement of the various phase boundaries in the magnesium-rich corner of the diagram. Sections at 375", 200°, and 100°C were selected for investigation. In constructing these sections, the solubility of aluminum in magnesium, as reported by W. L. Fink and L. A. Willey Vn 1948, was used at the binary Mg-A1 boundary and the solubility of lithium in magnesium was obtained from the equilibrium diagram for that system as reported by G. F. Sager and B. J. Nelson" in the same year. The solubility of magnesium in lithium was determined experimentally and conforms in general to data reported by P. Saldau and F. Shamrai." Parameters for AlLi and MgI7A1, were taken from American Society for Testing Materials X-ray diffraction data cards. Experimental Procedures Although the isothermal sections presented in this paper are not unusually complex, the experimental techniques involved in their construction are made extremely difficult by the relatively high vapor pressure of lithium and the great chemical activity of both magnesium and lithium. Because of these characteristics, which make precise control of the composition of equilibrium-treated filings practically impossible, the disappearing phase method was used in preference to the parametric method in conjunction with metallographic studies. The alloys used in this investigation were melted and cast in an atmosphere of helium using a tilting-type furnace which enclosed a steel crucible and mold in a single unit. Each portion of the charge (500 to 600 g) was cleaned carefully just before placing it in the crucible; and the charge, crucible, and entire melting apparatus were evacuated and then washed with grade A helium while preheating to approximately 100°C. The alloys were melted and chill cast in an atmosphere of helium. Alloys prepared in this way were relatively free from inclusions and a fluxing treatment was considered unnecessary. The cylindrical ingots obtained were scalped and then reduced 96 pct in area by direct extrusion, yielding % in. diam rod. Sections of the rod, approximately 3 in. long, were given equilibrium heat treatments and then sampled for metallographic examination, X-ray diffraction study, and chemical analysis. The surface of each equilibrium-treated rod was machined to a depth sufficient to insure removal of contaminated material before samples for chemical analysis or X-ray diffraction study were obtained, and all decisions on microstructure were based on the examination of the central portion of the metallographic specimen. All specimens homogenized at 375°C were analyzed after this equilibrium heat treatment. When the composition of an alloy placed it in a critical area of the 200" or 100°C isothermal section, a check chemical analysis was made on a sample taken from the alloy specimen as-heat-treated at the particular temperature. Standard chemical procedures of gravimetric analysis were used in the determination of magnesium and aluminum; lithium, potassium, and sodium were determined by flame photometer methods

Jan 1, 1956

By Strathmore R. B. Cooke, Emert W. Lindroos, Norman F. Schulz

DURING the course of an investigation of the effects of various starch products on hematite and quartz in regard to their separation by&apos; flotation, it was found that whereas most starches flocculated suspensions of hematite in water, they did not flocculate similar suspensions of quartz. However, a derivative of whole corn starch containing approximately one tenth aminoethyl group (--CH2CH2NH2) per glucose unit proved to be an exception. This aminoethyl starch was a very good flocculant for both hematite and quartz, and also possessed the property of depressing both minerals in cationic flotation tests. In view of the widespread use of starch as a flocculating agent of solids in tailing water, pulps, and for coal slimes, ore and various other purposes in sulphide and nonsulphide flotation, attention is called to the effective surface activity of aminoethyl starch in flocculating suspensions of the two minerals investigated. Test Materials The starch products employed included Globe Pearl starch No. 144 and fresh aminoethyl starch, both supplied by the Corn Products Refining Co. Dr. A. L. Elder[4] has kindly furnished a description The minerals used in the various tests are listed in Table I. Their specific surfaces were determined by the air permeability method, which gave results reproducible to within plus or minus 3 pct. Surface areas obtained by krypton adsorption were greater than the air permeability values by a factor of 2.0 for quartz and 62.4 for hematite. Starch Adsorption on Mineral Surfaces Starch adsorption data were obtained by agitating for 1 hr measured volumes of neutral starch solutions of known concentrations with weighed quantities of finely divided mineral of measured specific surface. The mineral then was separated from the pulp liquor by centrifugation and the residual starch content determined by chemical analysis. The analytical procedure consisted of oxidizing the starch with an excess of dichromate in acid medium at elevated temperature and titrating the excess dichromate by the ferrous ammonium sulphate-potassium permanganate method. Analytical results were reproducible to within less than 1 pct. Adsorption data obtained in unbuffered solutions near the neutral point are plotted in Fig. 1. The adsorption on either mineral was far greater for the aminoethyl starch than for the Pearl starch. Flocculation The degree of flocculation was determined for a given pulp by measuring the time required for the supernatant liquid to reach a certain degree of clarity. This was accomplished by passing a standard light beam through the pulp, contained in a glass settling tube, and measuring the time required after agitation ceased for the intensity of the transmitted beam to reach a predetermined value. Starch reagent was added to each pulp in small uniform increments, and the settling time was determined after each such addition. Increase in pulp liquor volume due to the added starch reagent was ignored in these tests. Pulps were maintained at or near neutrality.

Jan 1, 1952

By Vaughn Moyer

A method for incorporating well drawdown effect into reservoir calculations is presented in detail, together with examples of its use for widely divergent conditions that could be normally encountered in oil reservoir sands. Tentative conclusions as to the effect such factors as permeability, producing interval, well spacing, or well damage caused by completion or repair practice could have upon economic ultimate recovery are reached. It is carefully pointed out that the conclusions reached are subject to limitations indicated in the text, and placed upon all calculations made in the paper. Introduction Papers have been presentedl-3 in which theoretical depletion histories for reservoirs under equilibrium conditions and no appreciable well-bore drawdown pressure are determined. However, except for a limited number of cases, appreciable drawdown Pressures in a well bore exist, and Pressure and saturation gradients consequently vary throughout a reservoir. This Paper is an attempt to incorporate well-pressure drawdown into reservoir history, and to analyze certain factors, such as those involving recovery, well spacing, and productivity index, from the standpoint of permeability and drainage radius. BY incorporation of drawdown into reservoir history, it is believed that theoretical considerations will more closely approxi- mate actual reservoir behavior experienced in an oil field. For purposes of simplification, a solution gas-drive type field1 without primary gas cap, has been selected. Basic Data and Assumptions Details of equilibrium saturation conditions for no drawdown production for various pressures in a reservoir are deter-minable by methods described by several are permitting the construction of a curve of oil saturation vs Pressure for any selected combination of initial interstitial water and hydrocarbon saturation (Fig I). This curve is then a description of equilibrium saturation conditions at any given pressure in a reservoir. Developing the above, details of saturation and flow permeability may thus be determined for any increment or portion of a well drainage area or reservoir. In this manner the effect of pressure differential (drawdown) may be incorporated into reservoir and well drainage considerations. It is pointed out that this method thus makes use of no-drawdown equilibrium production data, superimposing a draw-hwn effect upon them. It also assumes that net flow conditions in any considered increment of reservoir are not materially altered by the entrance of fluids from other increments into them, but approach equilibrium for the existing pressure and saturation (i.e., each increment behaves in more or less independent fashion, dependent largely upon the equilibrium considerations mentioned above)

Jan 1, 1948

By Vaughn Moyer

A method for incorporating well drawdown effect into reservoir calculations is presented in detail, together with examples of its use for widely divergent conditions that could be normally encountered in oil reservoir sands. Tentative conclusions as to the effect such factors as permeability, producing interval, well spacing, or well damage caused by completion or repair practice could have upon economic ultimate recovery are reached. It is carefully pointed out that the conclusions reached are subject to limitations indicated in the text, and placed upon all calculations made in the paper. Introduction Papers have been presentedl-3 in which theoretical depletion histories for reservoirs under equilibrium conditions and no appreciable well-bore drawdown pressure are determined. However, except for a limited number of cases, appreciable drawdown Pressures in a well bore exist, and Pressure and saturation gradients consequently vary throughout a reservoir. This Paper is an attempt to incorporate well-pressure drawdown into reservoir history, and to analyze certain factors, such as those involving recovery, well spacing, and productivity index, from the standpoint of permeability and drainage radius. BY incorporation of drawdown into reservoir history, it is believed that theoretical considerations will more closely approxi- mate actual reservoir behavior experienced in an oil field. For purposes of simplification, a solution gas-drive type field1 without primary gas cap, has been selected. Basic Data and Assumptions Details of equilibrium saturation conditions for no drawdown production for various pressures in a reservoir are deter-minable by methods described by several are permitting the construction of a curve of oil saturation vs Pressure for any selected combination of initial interstitial water and hydrocarbon saturation (Fig I). This curve is then a description of equilibrium saturation conditions at any given pressure in a reservoir. Developing the above, details of saturation and flow permeability may thus be determined for any increment or portion of a well drainage area or reservoir. In this manner the effect of pressure differential (drawdown) may be incorporated into reservoir and well drainage considerations. It is pointed out that this method thus makes use of no-drawdown equilibrium production data, superimposing a draw-hwn effect upon them. It also assumes that net flow conditions in any considered increment of reservoir are not materially altered by the entrance of fluids from other increments into them, but approach equilibrium for the existing pressure and saturation (i.e., each increment behaves in more or less independent fashion, dependent largely upon the equilibrium considerations mentioned above)

Jan 1, 1948

By Stanley D. Michaelson

The selection of the proper size of grinding mill for a given installation has long been a subject of discussion by mill manufacturers, consulting engineers, and their clients. It would be presumptuous for the author to propose a simple cure-all. Instead this paper describes a method that has been used over a period of years to determine mill size with a reasonable degree of accuracy. Furthermore, this method claims no theoretical background, for ball milling, as practiced today, is still an art. The selection of the proper size and type of mill must, therefore, be made on an empirical basis. A ball mill, in order to grind, must deliver power through its grinding media to the material to be ground. Assuming that a mill is properly supplied with grinding media, ore pulp, and power, the amount of grinding done is closely proportional to the power expended. With this as a basic premise we can, for the moment, forget about type of mill (i.e., overflow, low-level mill, or other design) and base our calculation of mill size on the total power required to grind a ton of a particular ore from a given size of feed to a given size of product. When this is determined the type and size of mill, its speed, and the weight of the load of grinding media can be selected to absorb this amount of power. Development of Technique Some years ago, W. L. Maxson, F. Cadena, and F. C. Bond, working in the Minerals Research Laboratory of Allis-Chalmers Co., Milwaukee, developed and described a laboratory technique for comparing grindability (or relative ease of grinding) of various ores at different mesh sizes of grind.&apos; Additional papers2 have described the work done since that time, and the most recent paper3 summarizes results for hundreds of tests on many ores and materials. The essentials of the testing method are: a 12 by 12-in. laboratory mill operating at a constant speed of rotation, using a constant volume of ore, and a constant ball load having a relatively constant grinding surface. Tests are run at 250 per cent circulating load with the feed, circulating load and products screened on calibrated Tyler standard screens. Certain ores from operating mills, representing various mineralogical types, were selected for reference and designated "standard ores." In most cases, detailed studies of the grinding circuits were made when samples of mill feed were taken for grindability tests. Each analysis formed a basis for comparison of the grindability in the laboratory with actual power requirements for a particular grinding circuit. In addition, a critical study was made to evaluate the differences in the various grinding circuits so that two or more field installations having different layouts (single-stage grinding, two-stage grinding, or others), and using mills of various diameters operating at various speeds with differing circulating loads, etc., could be compared. Large-scale tests by mill operators complemented by laboratory investiga-

Jan 1, 1947

By Stanley D. Michaelson

The selection of the proper size of grinding mill for a given installation has long been a subject of discussion by mill manufacturers, consulting engineers, and their clients. It would be presumptuous for the author to propose a simple cure-all. Instead this paper describes a method that has been used over a period of years to determine mill size with a reasonable degree of accuracy. Furthermore, this method claims no theoretical background, for ball milling, as practiced today, is still an art. The selection of the proper size and type of mill must, therefore, be made on an empirical basis. A ball mill, in order to grind, must deliver power through its grinding media to the material to be ground. Assuming that a mill is properly supplied with grinding media, ore pulp, and power, the amount of grinding done is closely proportional to the power expended. With this as a basic premise we can, for the moment, forget about type of mill (i.e., overflow, low-level mill, or other design) and base our calculation of mill size on the total power required to grind a ton of a particular ore from a given size of feed to a given size of product. When this is determined the type and size of mill, its speed, and the weight of the load of grinding media can be selected to absorb this amount of power. Development of Technique Some years ago, W. L. Maxson, F. Cadena, and F. C. Bond, working in the Minerals Research Laboratory of Allis-Chalmers Co., Milwaukee, developed and described a laboratory technique for comparing grindability (or relative ease of grinding) of various ores at different mesh sizes of grind.&apos; Additional papers2 have described the work done since that time, and the most recent paper3 summarizes results for hundreds of tests on many ores and materials. The essentials of the testing method are: a 12 by 12-in. laboratory mill operating at a constant speed of rotation, using a constant volume of ore, and a constant ball load having a relatively constant grinding surface. Tests are run at 250 per cent circulating load with the feed, circulating load and products screened on calibrated Tyler standard screens. Certain ores from operating mills, representing various mineralogical types, were selected for reference and designated "standard ores." In most cases, detailed studies of the grinding circuits were made when samples of mill feed were taken for grindability tests. Each analysis formed a basis for comparison of the grindability in the laboratory with actual power requirements for a particular grinding circuit. In addition, a critical study was made to evaluate the differences in the various grinding circuits so that two or more field installations having different layouts (single-stage grinding, two-stage grinding, or others), and using mills of various diameters operating at various speeds with differing circulating loads, etc., could be compared. Large-scale tests by mill operators complemented by laboratory investiga-

Jan 1, 1947