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By David Hoadley, Kenneth R. Maser, Ashok B. Boghani, James E. Billar, D. Randolph Berry, Mackenzie Burnett, Robert H. Trent

B 5 Typical Input - Ouput The input data and the printed output for case 8 of the simulations (Section A.4) are shown in the following. The branch, level, node, miner and hoist numbers mentioned in the input and output refer to the corresponding numbers in Figures A-3, A-4 and A -5, The plotted ouput, already shown in Figures A,11 and A-12 is not repeated. The input data shown does not include the following simulation parameters: DTIME = 0, 05 FTIME = 130 JPLT(K) = 1 for K = 1, 2, 3 0 for other K TPLSTR = 0.0 TPLST P = 130

Jan 1, 1976

By E. G. King

Enthalpy increments above 298 K were determined by the Bureau of Mines for magnesium hydroxide (brucite) between 298 and 900 K by the method of drop calorimetry. Enthalpy and entropy increments from this investigation were combined with other data from the literature to calculate absolute entropies, free energy functions, enthalpies of formation, Gibbs energies of formation, and equilibrium constants of formation from 298 to 900 K. By the use of published PVT data for water, a table of Gibbs energy change was given for the reaction MgO (periclase) + H2O = Mg(OH)2 (brucite) as a function of temperature and pressure.

Jan 1, 1975

By R. Schuhmann

Distribution of material in the fine sizes of a comminution product generally is well represented by the empirical equation [ ] in which y = cumulative percent finer, x = particle size, = distribution modulus, and k = size modulus. Charles found that the energy consumption in comminution is usefully expressed by another empirical relation, [ ] in which E = energy input per unit volume of material, A = a constant, k = size modulus based on Eq. 1, and n = a constant; ( 1-n) is the slope of a plot of log E vs log k. Holmes has presented energy equations similar to Eq. 2.

Jan 2, 1960

By Gregory B. Poole

Data collected with 16K sampling at a variety of blasting sites (construction, quarry, trench, mine) and monitoring locations was analyzed to effectively show the 16 different 1K readings all based on the microsecond that the unit went into monitoring.

Feb 1, 2020

By D. Duff, R. A. Cameron, R. E. S. Whitehead

"Although CO2 content indicates the quantity of carbonate in altered basaltic rocks, it does not measure the intensity of carbonate alteration. CO/CaO molar ratios, which correlate with the relative proportions of individual carbonate species, measure the intensity of carbonatization. The Timmins gold area is characterized by pervasive carbonatization, with large parts of the area having CO2/Ca0 ratios greater than 1, representing C02 supersaturation relative to CaO. The mean CO/CaO in the Timmins area is0.67, compared to 0.13 in two unmineralized control areas. Regional arsenic abundance in the gold-bearing area averages 4 ppm, compared to about 0.35 ppm for the non-mineralized areas. The average K20 and Rb abundances are similar in both mineralized and non-mineralized areas (about 0.30% K20 and 7 ppm Rb). Although there appears to have been no wholesale addition of K20 or Rb to the mineralized area from external sources, these alkalies have been redistributed internally, resulting in patterns of local enrichments and depletions which are characteristic of the mineralized region.Local target anomalies associated with known mineralization consist of CO2/Ca0 > 1.5, As > 10 ppm and K2O > 0. 75%. Not all three values are required to be above these thresholds in order to identify local targets. Combinations such as C02/ Ca0 > 1.5 and As < 10 ppm, CO/CaO 1-1.5 and K20 > 0. 75 or K20 > 0. 75 and As > 10 ppm are equally indicative of proximity to gold occurrences. Gold-bearing veins and associated anomalies occur throughout the entire volcanic succession and are not restricted to any particular stratigraphic horizons. The levels and patterns of alteration and occurrences of gold appear to be structurally controlled."

Jan 1, 1982

By Franz Xaver Spachtholz

K+S Potash Canada GP (KSPC) is part of the K+S Group, one of the world&apos;s leading suppliers of standard and speciality fertilizers and the world?s leading salt producer. Globally, the K+S Group employs more than 14,000 people. The K+S share ? the only commodities stock on the German DAX index ? is listed on all German stock exchanges. Legacy, K+S Group?s most recent project, is a greenfield solution potash mine and production plant in the south of the Canadian province of Saskatchewan. Therefore K+S will invest 4.1 billion CAD. With the expected commissioning in summer 2016 the company projects to reach the two-million-tonnesmark of production capacity at the new location at the end of 2017. The gradual extension of the annual capacity to 2.86 million tonnes of potash products will occur afterwards. The Legacy Project will be Saskatchewan?s first new potash mine in nearly 40 years. It is one of the world?s most economically attractive potash projects. The project will contribute to Saskatchewan?s growth through employment, tax revenues, utilities, and local partnerships. Located in the southern part of Saskatchewan?s potash belt, the deposit is too deep to be mined conventionally due to rock mechanics and high temperatures. Legacy will therefore be a solution mine. Proximity to the urban centers of Regina and Moose Jaw is imperative, with their workforces and economy, as well as access to highways, railway carriers and, most importantly in the prairies, water. With this unique project setting and mining technique, we are convinced that Legacy will be one of the most efficient, economic, sustainable and safe potash facilities in the world.

Aug 1, 2013

By W. M. LONIE

Address by the President of the Institute (Mr W M Lonie) in making the presentation of the Certificate of Honorary Membership to Sir Ian McLellan, K.C.G.M, K.B.E. Address by Sir Ian McLellan after accepting the award.

Jan 1, 1979

By W. D. Robertson, H. Conrad

Incremental loading creep tests were conducted on magnesium single crystals in tension and direct shear at 78" to 364°K. The creep curves were transient in character and exhibited waves or periodic behavior, especially at small strain rates. A sigmoidal curve was frequently observed following the first occurrence of creep. Repeating of segments of a master curve was observed at all temperatures indicating that identical conditions for flow existed after various stress and strain histories. A logarithmic time law with a independent of temperature whs observed at 78°and 203°K, while a parabolic (power) time law was approximated at 298" and 364°K. The departure from logarithmic creep was attributed. to recovery, which was observed at 298" and 364°K. The temperature independence of a represents a serious objection to the theory of low-temperature creep based on the intersection of dislocations. IN a previous paper1 it was found that the initial creep rate of magnesium single crystals, y, following the addition of a stress increment in the range of 78° to 364°K, was given by the expression ? = C exp - {?H/kT} exp {B(T- h ?)} [1] where B and C are constants, AH is an activation energy of 10,300 cal per mole, T is the applied stress, ? is the total strain at the time of the stress increment, and h is the strain-hardening coefficient. The present paper is concerned with certain other features of creep behavior observed during this investigation; namely, the form of the creep curves, time laws for creep, recovery, and microscopic and X-ray observations of the structural manifestations of the creep flow. To the authors&apos; knowledge the temperature range of the present investigation (0.08 Tm to 0.40 Tm) has not been covered previ-

Jan 1, 1959

By Yang Cui

The partition ratios of lead and antimony between liquid copper and MgO saturated Cu2O-CaO-SiO2 flux and their removal rate by using CuCl-based flux were measured between 1423 K and 1573 K. For Cu2O-CaO-MgOsatd. system, the partition ratio of Pb was not affected by the CaO content at 1573 K, while that of Sb increased with increasing CaO content. The maximum value was 160 at the CaO content of 15.4 mass%. For Cu2O-CaO-SiO2-MgOsatd. system, two liquid phases of calcium silicate rich phase and Cu2O rich phase were observed at 1573 K. The partition ratios of Pb and Sb between calcium silicate phase and liquid copper depended on the (mass%CaO)/(mass%SiO2) ratio. When the (mass%CaO)/(mass%SiO2) ratio increased, the partition ratio of Pb decreased and that of Sb increased. For CuCl flux, the removal of Pb from liquid copper at 1473 K was possible, and a relatively low concentration could be reached. Moreover, most of lead as metal chloride vaporized from flux. For the removal of antimony from liquid copper at 1423 K, it was difficult by CuCl flux, and the removal ratio was approximately 20% for 75 minutes. However, the addition of CaO or Na2CO3 to CuCl flux was relatively effective to remove antimony, and the concentration of antimony reached a relatively low value after 15 minutes by using CuCl-CaO or CuCl-Na2CO3fluxes. Moreover, a small part of antimony as chloride vaporized from flux.

Jan 1, 2006

By Kenji Monden

"The creep properties of tin-based lead-free solders Sn3.0AgO.5Cu and SnO.7CuO.1Ni and lead solders 10Sn90Pb and 50Sn50Pb were investigated at temperatures between 298 K and 398 K. The creep-rupture time decreases with increasing initial stress and temperature. The analysis of the solder creep curves uses the Omega method. The creep rate E is expressed by following formula: 1n ? = 1n ?0 + O?, where ?0 and O are experimentally determined. The parameter Eo increases with increasing initial stress and temperature. The activation energy for ?0 is 108 kJ/mol for Sn3.0Ag0.5Cu and 85 kJ/mol for SnO.7CuO.INi. These energies suggest that the lattice diffusion of tin is dominant. The energy of 1OSn90Pb is 37 kJ/mol and that of 50Sn50Pb is 67 kJ/mol. The creep-rupture time is calculated using the parameters, ?0 and O. The calculated creep-rupture time is in good agreement with the measured creep-rupture time.IntroductionLead-free solders are replacing traditional tin-lead solders as part of a major trend toward protecting the earth's environment. However, tin-lead solders are currently needed for a high temperature range. Since creep is the main deformation mechanism in solder joints, understanding a constitutive model for creep and examining creep data for solders are important for reliability analysis of solder joints. A steady-state creep equation is generally used for solders [1,2], and it supposes that the secondary-, or minimum-, creep rate occupies the entire creep life. However, secondary creep accounts for only about 15 % to 40 % of the entire creep life for solders. Thus, the equation may not represent the entire creep life."

Jan 1, 2012

By J. D. Mooney, D. L. Darnell

The accompanying chart has been drawn for the convenicncc of engineers as&apos; a means of quickly determining the correct number of plies of conveyor belts operating under specific conditions. The calculations are based on the average safe strength (factor of safety, 15) of the various standard rubber conveyor belts. The calculations assume maximum loading conditions; that is, the belt is considered as carrying the greatest load that it will handle without spillage at ordinary belt speeds. This not only produces the most economical operating conditions, but also the maximum tension in the belt. The chart is a graphical representation of the formula: Where, p = kg W(L + 10H) p = the correct number of plies k = a constant, depending on the type of drive g = the weight &apos;in pounds per cubic foot of material handled W = the width of the belt in inches L = the length of the belt in feet (approximately twice the center distance). H = the difference in elevation between the head and tail pulleys, in feet. For a simple drive, with a bare pulley, k = ----1 For a simple drive, with a rubber-lagged pulley, k = --- 1 For a, tandem drivej with bare pulleys, k = 1/375.000 For a tandem drive, with rubber-lagged pulleys, k = ---1 The chart is drawn for a simple drive, with a bare pulley (k — =e)1 ; therefore, the number of plies obtained from the chart should be multi- &apos; plied by the factor 0.83 or 5/6 for simple, lagged drive; the factor 0.67 or 2/3 for tandem, bare; and the factor 0.55 or 11/20, for tandem, lagged. The formula p = kgW(L + 10H) has been developed mathematically

Jan 1, 1916

By Mariana H. Staia

Thermodynamic calculations based on the free energy minimization method were performed for the system Ti-Fe-Mn-0-C1-C in order to evaluate the chemistry associated with the chlorination of an ilmenitic concentrate using CCl4. In the calculations various weight ratios of FeTi03/ CC14, a , were considered, which varied between 0.32 to 0.98 at temperatures ranging from 673 K to 1273 K , at atmospheric pressure. The results obtained indicated that the operating conditions for a process in a packed bed reactor, which implies the chlorination of both Fe and Ti, are 0.39 < a < .49 at T 1173 K.

Jan 1, 1993

By T. M Morris

THE rate of flotation of solid particles determines the percentage recovery of these particles which can be obtained during a given time interval. It is an established fact that the recovery is greatest in the fine sand size range and falls off on either side of this size range. It is also a fact that the history of the development of concentrating devices has been marked by a desire to recover the small particles: tables were developed to treat material too fine to recover in jigs; flotation recovers material too fine to be recovered on tables. As our ores become more complex and require finer grinding, a greater proportion is in the slow-floating size range. It is therefore of economic importance that the effect of size of particle upon rate of flotation be studied. Many studies have been made along this line. The present investigation attempts to quantify the rate of flotation by means of a rate equation, and then compares the effect of size upon the rate of flotation. It was found that the flotation rate constant increases with increase in diameter of particle in the size range smaller than the optimum size particle. The flotation rate constant decreases with an increase in diameter of particle in the size range coarser than the optimum size particle. This suggested two different actions taking place, and these are discussed. Suggestions for further work in order to improve the flotation rate of both fine and coarse particles are made. The usual method of expressing rate is to plot cumulative percent recovery vs time. If we knew the equation of this kind of curve we could replot it as a straight line, and the slope of the line would be an index of the magnitude of the flotation rate. It is easier to compare straight lines than curves. The basis used for obtaining the equation of the recovery vs time curve is the law of mass action. This law states that the "rate of a chemical reaction is proportional to the &apos;active masses&apos; of the reacting substances present at any given time." This, of course, implies a constant temperature. We can modify this for our use by using the term flotation reaction instead of chemical reaction and the term floatable particle instead of active masses. dC Thus----— kc [1] where C denotes concentration of particles, t denotes time, k denotes a proportionality constant. If the exponent of C is 1, then we have a first order rate reaction. If the exponent of C is 2, then we have a second order rate reaction. For a first order rate reaction, we obtain, after integration, where k is again a proportionality constant and has the dimension of the reciprocal of time. C, is the original concentration or weight or percentage of floatable particles. C, is the concentration or weight or percentage of floatable particles remaining in the flotation cell after time t. If all particles of the mineral to be floated were capable of floating we could take C, as 100 and C, would be 100 - R, where R is the cumulative percentage recovery of mineral at time t. But it is not usual for all of the particles to be capable of floating, because of surface oxidation, etc. Therefore, we must introduce a factor x which is the percentage of un-floatable mineral. Hence the first order equation is It has been proposed by various workers1-&apos; that this equation describes the flotation rate. More recently N. Arbiter4 states that a second order rate equation fits the recovery vs time curve more closely than does a first order rate equation. Morris- as discussed the evidence presented for this statement. It has been concluded that the first

Jan 1, 1953

By T. M. Morris

THE rate of flotation of solid particles determines the percentage recovery of these particles which can be obtained during a given time interval. It is an established fact that the recovery is greatest in the fine sand size range and falls off on either side of this size range. It is also a fact that the history of the development of concentrating devices has been marked by a desire to recover the small particles: tables were developed to treat material too fine to recover in jigs; flotation recovers material too fine to be recovered on tables. As our ores become more complex and require finer grinding, a greater proportion is in the slow-floating size range. It is therefore of economic importance that the effect of size of particle upon rate of flotation be studied. Many studies have been made along this line. The present investigation attempts to quantify the rate of flotation by means of a rate equation, and then compares the effect of size upon the rate of flotation. It was found that the flotation rate constant increases with increase in diameter of particle in the size range smaller than the optimum size particle. The flotation rate constant decreases with an increase in diameter of particle in the size range coarser than the optimum size particle. This suggested two different actions taking place, and these are discussed. Suggestions for further work in order to improve the flotation rate of both fine and coarse particles are made. The usual method of expressing rate is to plot cumulative percent recovery vs time. If we knew the equation of this kind of curve we could replot it as a straight line, and the slope of the line would be an index of the magnitude of the flotation rate. It is easier to compare straight lines than curves. The basis used for obtaining the equation of the recovery vs time curve is the law of mass action. This law states that the "rate of a chemical reaction is proportional to the &apos;active masses&apos; of the reacting substances present at any given time." This, of course, implies a constant temperature. We can modify this for our use by using the term flotation reaction instead of chemical reaction and the term floatable particle instead of active masses. dC Thus----— kc [1] where C denotes concentration of particles, t denotes time, k denotes a proportionality constant. If the exponent of C is 1, then we have a first order rate reaction. If the exponent of C is 2, then we have a second order rate reaction. For a first order rate reaction, we obtain, after integration, where k is again a proportionality constant and has the dimension of the reciprocal of time. C, is the original concentration or weight or percentage of floatable particles. C, is the concentration or weight or percentage of floatable particles remaining in the flotation cell after time t. If all particles of the mineral to be floated were capable of floating we could take C, as 100 and C, would be 100 - R, where R is the cumulative percentage recovery of mineral at time t. But it is not usual for all of the particles to be capable of floating, because of surface oxidation, etc. Therefore, we must introduce a factor x which is the percentage of un-floatable mineral. Hence the first order equation is It has been proposed by various workers1-&apos; that this equation describes the flotation rate. More recently N. Arbiter4 states that a second order rate equation fits the recovery vs time curve more closely than does a first order rate equation. Morris- as discussed the evidence presented for this statement. It has been concluded that the first

Jan 1, 1953

By T. M. Morris

THE rate of flotation of solid particles determines the percentage recovery of these particles which can be obtained during a given time interval. It is an established fact that the recovery is greatest in the fine sand size range and falls off on either side of this size range. It is also a fact that the history of the development of concentrating devices has been marked by a desire to recover the small particles: tables were developed to treat material too fine to recover in jigs; flotation recovers material too fine to be recovered on tables. As our ores become more complex and require finer grinding, a greater proportion is in the slow-floating size range. It is therefore of economic importance that the effect of size of particle upon rate of flotation be studied. Many studies have been made along this line. The present investigation attempts to quantify the rate of flotation by means of a rate equation, and then compares the effect of size upon the rate of flotation. It was found that the flotation rate constant increases with increase in diameter of particle in the size range smaller than the optimum size particle. The flotation rate constant decreases with an increase in diameter of particle in the size range coarser than the optimum size particle. This suggested two different actions taking place, and these are discussed. Suggestions for further work in order to improve the flotation rate of both fine and coarse particles are made. The usual method of expressing rate is to plot cumulative percent recovery vs time. If we knew the equation of this kind of curve we could replot it as a straight line, and the slope of the line would be an index of the magnitude of the flotation rate. It is easier to compare straight lines than curves. The basis used for obtaining the equation of the recovery vs time curve is the law of mass action. This law states that the "rate of a chemical reaction is proportional to the `active masses&apos; of the reacting substances present at any given time." This, of course, implies a constant temperature. We can modify this for our use by using the term flotation reaction instead of chemical reaction and the term floatable particle instead of active masses. [ ] where C denotes concentration of particles, t denotes time, k denotes a proportionality constant. If the exponent of C is 1, then we have a first order rate reaction. If the exponent of C is 2, then we have a second order rate reaction. For a first order rate reaction, we obtain, after integration, [ ] where k is again a proportionality constant and has the dimension of the reciprocal of time. C, is the original concentration or weight or percentage of floatable particles. C, is the concentration or weight or percentage of floatable particles remaining in the flotation cell after time t. If all particles of the mineral to be floated were capable of floating we could take C, as 100 and C, would be 100 - R, where R is the cumulative percentage recovery of mineral at time t. But it is not usual for all of the particles to be capable of floating, because of surface oxidation, etc. Therefore, we must introduce a factor x which is the percentage of unfloatable mineral. Hence the first order equation is [ ] It has been proposed by various workers1-3 that this equation describes the flotation rate. More recently N. Arbiter4 states that a second order rate equation fits the recovery vs time curve more closely than does a first order rate equation. Morris&apos; has discussed the evidence presented for this statement. It has been concluded that the first

Jan 1, 1952

By W. R. Hibbard

It has been proposed1 on the basis of slip and flow that the ideal deformation texture of drawn wire for face-centered cubic metals is a (111) direction parallel to the wire axis. Under these considerations, an intermediate [loo] texture is permissible but should eventually change to the stable [Ill] orientation. This theory does not explain why copper and its alloys containing up to about 30 pct nickel, 5 pct zinc and 1 pct aluminum were found to have a double [Ill], [100] texture, while alloys with at least 8 pct zinc and 2 pct aluminum developed essentially a single [Ill] texture.1,2 It does suggest, however, that the [loo] component of the former alloys should disappear with additional deformation.3 Some evidence exists to support this suggestion. Howald4 reported that annealed copper wire previously drawn 93.8 pct and 98.5 pct developed a single [Ill] texture. Since, in wire, the drawn and annealed textures are similar, these wires probably had a single [Ill] deformation texture before annealing. In addition, although cold rolled polycrystalline copper and 80-20 brass were found to form different types of pole figures,5,6 single crystals of these metals when rolled from the stable (110) [112] orientation were found to produce essentially the same polo figure.7.8 It may be, therefore, that while the addition of solute elements changes the path by which the final texture is obtained, the final texture is not changed. Experimental Results The wires remaining from the previous investigation1 already drawn about 50 pct reduction in diameter were further drawn to 86 pct and then to 96.4 pct total reduction in diameter. The textures were determined after these stages by transmission X ray photograms taken with a copper K-alpha beam rotated 0 degrees toward the wire axis. It was found that the ratio of the [111].] and [100] intensities of the double texture in copper and its alloys containing 0.94 pct aluminum, 30.70 pct nickel, 2.61 pct zinc and 4.99 pct zinc increased from about 3:l for the 50 pct reduction to about 9:l for the 86 pct reduction. After 96.4 pct reduction, these wires exhibited essentially a single [lll] texture. Summary Thus, the drawn wire deformation texture of copper and its alpha solid solution alloys containing aluminum, nickel or zinc is essentially a single [Ill] texture as predicted by theoretical considerations of slip and flow.&apos; For copper and its alloys containing up to about 1 pct aluminum, 5 pct zinc and at least 30 pct nickel, an intermediate [loo] component is found after about 50 pct reduction in diameter, but disappears after about 95 pct reduction in diameter. For alpha solid solution alloys containing at least about 2 pct aluminum and 8 pct zinc the single [Ill] texture is developed considerably more easily and is found after about 50 pct reduction in diameter. No explanation of this compositional effect is apparent. Similar work is contemplated for cold rolled and cold compressed alloys. References 1. w. R. Hibbard, Jr. and M. K. Yen: Trans. AIME 175, 126. Metals Tech., Feb. 1948, TP 2334. 2. W. R. Hibbard, Jr. and D. E. Trout, 11: Trans. AIME 185, 620. Jnl. of Melals, Sept. 1949, TN 18. 3. W. R. Hibbard. Jr.: Disc. of R. M. Brick&apos;s paper on Deformation Textures, ASM Symp. on Cold Working of Metals, Oct. 1948. 4. T. S. Howald: Disc. of Ref. 1. Trans. AIME, 175, 126. Melals Tech. Aug. 1948, TP 2449. 5. R. M. Brick, D. L. Martin and R. P. Angier: Trans. ASM (1943) 31, 675. 6. 0. Dahl and F. Pawlek: Zlsch. Melall. (1936) 28, 266. 7. C. S. Barrett and F. W. Steadman: Trans. AIME (1942) 147, 57. 8. J. E. Burke and C. S. Barrett: Trans. AIME 175, 106. Melals Tech., Feb. 1948. TP 2327.

Jan 1, 1950

By D. A. Elgillani, M. C. Fuerstenau

In this paper, oxidation potentials measured in the presence of various concentrations of cyanide, ferro-cyanide, and ferricyanide and ethyl xanthate at various values of pH are related to flotation response. Eh-pH diagrams are presented to show that the formation of surface ferric ferrocyanide is probably responsible for depression when cyanide is added. The influence of cyanide on the depression of pyrite with xanthates as collector has been the subject of a number of investigations,&apos;-6 and several theories on the mechanism of depression have evolved from these studies. Wark and Cox7 and Gaudin8 have suggested that the depressing effect is due to a competition of cyanide ion with xanthate ion for the surface. Cook and his colleagues9-11 have explained this phenomenon in terms of competition between hydrocyanic acid and xanthic acid. Sutherland 12 has shown that although both of these theories accurately describe the relation between pH value and cyanide addition at constant collector addition, they fail to describe the relation between pH value and the amount of collector required to cause flotation. Taggart 13 suggested that depression in these systems is due to the formation of a reaction product between ferric ion at the pyrite surface and ferrocyanide ion derived from solution. Majumdar4,6 has attempted to prove this hypothesis by measuring the contact angles of pyrite in the presence of 25 mg per liter ethyl xanthate and different concentrations of potassium ferrocyanide and ferricyanide. In all cases the contact angles were quite high up to pH 10. These results indicate that pyrite should not be depressed by either potassium ferrocyanide or ferricyanide. In view of these facts, Majumdar has assumed that the compound Fe(CN)2 forms at the surface. Gründer and Bornl4 have stated that depression may be due to the formation of the compound K2Fe(II)Fe(CN)6 at the pyrite-solution interface. This compound is thought to be an interaction product between the K2Fe(CN)6-2 ion from solution and the Fe++ ion at the pyrite surface and, accordingly, K4Fe(CN)6 should depress pyrite at least as effectively as KCN. This was proven experimentally, but there was no simple relation between the depression of pyrite and the concentration of either KCN or K4Fe(CN)6 in solution. In view of the many mechanisms that have been proposed for pyrite depression by cyanide, it is apparent that a clear understanding of the phenomena occurring in these systems is lacking. One reason for this may be the fact that the species responsible for pyrite flotation in the presence of xanthate is not the xan-thate ion but rather dixanthogen.15 Since the oxidation of xanthate to dixanthogen is dependent on the oxidation potential of the solution, it would seem that knowledge of these potentials would be a requisite to understanding the pyrite-xanthate-cyanide system. It is the object of this paper to measure both the oxidation potential and pH of the pyrite systems in the presence of various concentrations of cyanide, ferrocyanide, and ferricyanide and xanthate and to relate these values to flotation response. EXPERIMENTAL MATERIALS AND &apos;TECHNIQUES In the experiments discussed here, pure potassium ethyl xanthate was used as collector, and reagent grade potassium cyanide, potassium ferrocyanide, and potassium ferricyanide were used as depressants. Reagent grade HC1 and KOH were added for pH adjustment. Conductivity water, made by passing distilled water through an ion exchange column, was used in all experimental work. Two natural samples of pyrite were used in the investigation. Sample preparation for flotation included dry grinding with a mortar and pestle and sizing the product to 100 x 200 mesh. Prior to flotation, a 0.75-gm sample of pyrite was added to a solution containing a known amount of depressant at the desired pH value, and the system was conditioned for 4 min. Following this, a known amount of collector was added and the system was conditioned for another 4 min. The pH — termed flotation pH - was measured; the pulp was transferred to a Hallimond cell, and flo-

Jan 1, 1969

By William A. Rhoades

INTRODUCTION Ever since miners were faced with the task of moving ore, some form of scraper has been in use. At first, men and beasts of burden supplied the power to move the scraper, and later, machines were developed for this purpose. In the early 1900s, a few mining properties used small pneumatic single-drum winches to pull loaded scrapers to a raise or ore pocket, and the empty scraper then was dragged back to the muck pile by a miner, as shown in Fig. 1. Just prior to 1920, an improvement was introduced, using two single-drum hoists. As illustrated by Fig. 2, the second hoist was used to return the empty scraper to the muck pile. However, this arrangement still re¬quired two men, one operating each hoist. The next developmental step was to eliminate the second man by locating the hoists side-by-side and using one man to control both hoists. As illustrated by Fig. 3, this involved the use of a tail rope over a sheave to pull the scraper back. The greatest progress in development and the great¬est increase in the use of slusher haulers occurred between 1920 and 1930. In 1921, the Sullivan Machinery Co. designed and built the first two-drum scraper oper¬ating on the principle illustrated by Fig. 4. It was powered by a 4.5-kW (6-hp) Turbinair(r) motor and would pull a 450-kg (1000-1b) load at 0.61 m/s (120 fpm). In 1922, this unit was shipped to the Verona Mining Co. of Caspian, MI, and it experienced immediate success in the Lake Superior iron ore district. Since the two-drum slusher was much less expensive and more efficient than hand mucking, the Lake Superior mines were saved from financial disaster when iron-ore prices fell 25% between 1923 and 1925. Immediately there¬after, a demand developed for slushers that would oper¬ate with electric power, which was considerably cheaper than compressed-air power. In 1923, the Sullivan Machinery Co. responded with the first electric-powered double-drum hoist. During subsequent years, design improvements in¬cluded separate tail-drum gearing to increase the tail¬drum speed, as welt as a number of safety features such as rope guides. One result of these improvements and their utilization by the Michigan iron mines was an in¬crease of 100% in the tons of ore per miner per day in those mines between 1924 and 1929. Since the two-drum slusher was capable only of straight-line mucking, it was not a practical machine for use in open stopes. In 1929, the Sullivan Machinery Co. introduced the first three-drum slusher. As illustrated by Fig. 5, two tail drums and one hauling drum were provided. A tail sheave could be placed at each side of the stope, and the ore then could be loaded and hauled to a central point from the entire width of the stope. During the 1930s, progressively larger slushers were demanded. By 1940, two- and three-drum units were available with motor power as high as 45 kW (60 hp), and slusher power continued to increase after 1940. Be¬tween 1951 and 1952, Joy Manufacturing Co. designed a 112-kW (150-hp) two-drum Blusher for the Climax Molydenum Co. Although there has been no demand for a slusher more powerful than this, 150- to 225-kW (200- to 300-hp) slushers are quite feasible at the pres¬ent time. During the last 30 years, many slusher improvements have been made to the operating life, operational safety, and ease of operation and maintenance. Increased tail-drum speeds have decreased overall scraping times. Rope guards, totally enclosed drums, and operator shields have reduced the hazards of injury due to wire¬rope breakage. Improvements in lubrication have made the slushers relatively maintenance-free, with long operating lives. The introduction of spring-actuated drag brakes prevented uncontrolled unreeling of dis¬engaged drums, allowing the development of practical remote-control slusher operation. Remote control now is available in a choice of all-air, all-electric, or air¬electric slushers. APPLICATIONS Quite simply, slushers are used to load and transport material (ore), generally over a short distance of from

Jan 1, 1982

By R. E. Ogilvie, E. T. Peters

The X-ray Kossel-line method has been used preaioz~sly for measuring lattice parameters to accuracies of 1 part in 100,000.5 A second application of this method is described for determining the crystallographic orientation of a randomly positioned single-crysta1 spherical volume that can he as small as 50 µ in diameter, within accuracy limits of ±1/2 deg. The theory, experimental procedure, and interpretation of Kossel-line patterns and an experimenta1 verification of the predicted orientation relationship of the cph k and fee a Cu-Si phases are presented. This orientation relationship can be described as (111)a, (00.1)k; [110]a [11.0]k. In addition, the lattice parameter of the a phase was found to he a. = 3.62154 ± 0.00014Å. The Kossel-line method when used in conjunction with electron microanalysis is shown to he capable of providing a complete chemical and structural analysis of a given crystal. THE most easily accomplished methods for determining the orientation of a single crystal are variations of the Laue X-ray diffraction method. Although certain materials can be oriented to within ± 1 deg accuracy by observation of exterior macroscopic features, such as etch pits, cleavage faces, or growth features, the Laue method is generally preferred for routine laboratory application. Both back-reflection and transmission patterns are coordinated by appropriate reference charts and are plotted in terms of reflection plane poles (normals) on a stereographic projection. Orientation is deduced by relating the pole distribution (which is fixed by the crystallographic symmetry of the specimen) to two specified external reference directions. The Laue method has several limitations. 1) Orientation can rarely be determined to better than ±l deg accuracy. Principal errors involve inaccuracy in specimen-to-film distance, measurement confidence of individual Laue spots, and stereographic plotting. 2) Because of the indeterminacy of the X-ray wave length diffracted to a given spot, it is not possible to determine supplementary crystallographic data, such as interplanar spacings and lattice parameters. 3) The method is generally limited to specimens of cubic symmetry or to specimens of high symmetry and known structure. 4) The relatively large cross-sectional area of the incident X-ray beam generally precludes the measurement of relative grain orientation in a polycrystalline material.* _____ Several of these limitations can be overcome by application of the Kossel-line method, which has been previously employed for precision lattice-parameter determinations.&apos;-= For this method, a point source of divergent monochromatic X-rays is generated within the crystal by means of an incident electron beam. The divergent X-rays which fulfill the Bragg law are diffracted by the specimen and are recorded on a film placed either in transmission or back reflection. Analysis of the film yields a direct measurement of orientation and lattice-spacing values to an accuracy of ±1/2 pct. As analyses can be obtained from spherical specimen volumes as small as 50 µ in diameter, the method provides a means for structural analysis of second-phase or impurity precipitates within a given matrix. The primary limitation of the Kossel-line method is the requirement for an electron microanalyzer or similar apparatus capable of producing a finely focused electron beam. This paper is designed to present the theory, experimental procedure, and geometrical interpretation of Kossel patterns. The experimentally determined orientation relationship between the k and a phases occurring in the Cu-Si system and a precision measurement of the a lattice parameter are presented as a practical application of the method. THEORY OF KOSSEL LINES Divergent X-ray beam photography utilizes an effective point source of characteristic X-rays which, when diffracted from a single crystal, form numerous diffraction and absorption cones that are recorded on film.7 The cones generated from a source lying within the crystal are called Kossel lines.&apos; Although the X-ray scattering from a divergent point source contained within a crystal is described in terms of Laue dynamical theory,9 the directions of the diffracted spectra can be ade-

Jan 1, 1965

By Owen Thornton

Recently equations have been presented by Rose and Bruce&apos; and by Rose², showing how the relative permeability of a reservoir rock may be determined from the capillary character of the rock. In particular, equations were developed to show the relationship between capillary pressure and the effective permeability to the wetting phase in a poly-phase flow system. The equations are as follows (nomenclature the same as in the Rose and Bruce paper) Rose and Bruce assume that the average length of path (L,), which the wetting fluid follows in flowing through a porous media is independent of the saturation to the wetting phase, so that 1 is a constant dependent only on the k rock, and the permeability to the wetting phase is a function of the saturation and of the capillnry pressure: In measuring the flow of ekctric current through partially saturated core samples, it has been observed that the electrical resistance usually is not a simple linear function of saturation. This fact indicates that the average length of path in the flow of electric current is not independent of saturation, but rather that the tortuosity of the path depends upon the average saturation to the conducting fluid as well as upon the characteristics of the rock. itself. Inasmuch as the flow of fluid and the flow of electric current in many respects are analogous, it may be indicated further that the length of path for fluid also may not be independent of saturation. The following relationship can be derived to express the resistance to flow of electric current through a core sample partially saturated with conductive wetting phase, the non-wetting phase being non-conductive: where L is the average length of path followed by the current in flowing through a length L of partially saturated core, where L is the length of current path when the core is fully saturated, and where Rs and R, are the specific resistivities of the partially saturated and saturated core sample, respectively. It will be noted that R., the specific resistivity of the core sample when S, = 1.0, is equal to the "formation resistivity factor"&apos; multiplied by the specific resistivity of the saturating fluid. If it be assumed that the average lengths of path for fluid flow and for the flow of electric current are the same, then equation (5) may be combined with equations (l), (2) and (3) to give the following relationship: The above equation suggests including a correction for tortuosity when calculating the relative permeability of a reservoir rock to a wetting phase. The correction factor may be obtained from the resistivity-saturation relationship. For instance, the following calcu- lations are obtained for the unconsoli-dated sands described by Leverett Kw Kw Sw R°/Rs Pt/Pc¹ (Calc.) (Obs.) 0.30 0.08 0.88 0.02 0.0 0.40 0.16 0.94 0.06 0.04 0.50 0.27 0.97 0.14 0.11 0.60 0.40 0.98 0.26 0.21 0.70 0.53 0.98 0.39 0.35 0.80 0.68 0.99 0.57 0.54 0.90 0.84 0.99 0.77 0.76 1.00 1.00 1.00 1.00 1.00 It will be noted that the agreement between the calculated points and the observed data is rather good. Further investigation of the above method for obtaining relative permeabilities may be merited. For many sands within specified ranges of saturations the following relationship has been found to hold approximately&apos;: — =SW"......(7) Rs The exponent n in the above equation has been found to equal two for many sands so that equation (6) reduces to the following: The writer acknowledges the permission of The Texas Co. to submit this note for publication. 1. Walter Rose and W. A. Bruce—Jnl. of Pet. Tech., Vol. 1, No. 5, p. 127 (1949). 2. Walter Rose—Jnl. of Petr. Tech., Vol. 1, No. 5, p. 111 (1949). 3. G. E. Archie— -Trans. AIME, 146, 54 (1942). 4. M. C. Leverett— Trans. AIME, 142, 152 (1941). 5. M. C. Leverett—Trans. AIME, 132, 149 (1939).

Jan 1, 1949