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By Paul A. Beck, John Towers, W. D. Manly

Recent work on grain growth in high purity aluminum and in a solid solution type alloy of aluminum and magnesium1 showed that the isothermal increase of the average grain diameter D with time follows the equation D = K(tg + A)2 Eq 1 Here tg is that portion of the total annealing time 1, which is available for grain growth, that is the annealing time less the time necessary for complete recrystallization. The parameters A and n depend only on the temperature. A has the dimension of time, and is usually of the same order of magnitude as the time for recrystallization: R. In some instances R may be substituted for A in Eq I. This is the case for example at long periods of annealing or at high temperatures, where both A and R become small in comparison with t. In such instances Eq I is simplified to D = K, tn Eq 2 For high purity aluminum the exponent n increases in the ratio of about I to 5 in. the temperature range of 350' to 6w°C. 70-30 brass has been long known to exhibit grain growth of the continuous type, such as occurs in pure metals. For this reason, it was of interest to determine whether grain growth in brass also follows Eq I and 2 as it does in aluminum. Isotherma1 grain growth data for 70-30 brass were published in recent years by R. S. French2 and by H. L. Walker.' Fig I shows a log D vs. log t plot of French's data, together with the 500, 600 and 700°C data by Walker for 42.8 pct reduction by rolling. In such a log D vs. log t plot Eq 2 corresponds to a straight line with a slope of n. As seen in Fig I, the data of French and the 600°C data of Walker can be represented by straight lines. However, the 500 and 7c0°C data of Walker give curved lines and the average slope of his 500°C line is very different from the slope of the line representing French's data for the same temperature. Furthermore the curvature of the 700°C line suggests a tendency for grain growth to stop after 8 hr. If confirmed, this would constitute a definite deviation from Eq 2, not explainable on the basis of the specimen thickness effect found in high purity aluminum.' Walker's specimens were 0.125 in., or 3.18 mm thick, while the deviation occurs at a grain size of approximately 0.4 mm. A further point of divergence arose recently, when J. E. Burke suggested4 on the basis of Walker's data, that, instead of Eq 2, grain growth in 70-30 brass follows the equation D - D, = K . tgn Eq 3 Here D, is the grain size as recrystallized at the temperature in question. Although it was possible to show5 that, for pure aluminum, for which isothermal grain growth has been carefully investigated, Eq 3 is incompatible with the experimental facts, the situation remained somewhat obscure with respect to 70-30 brass. In view of the discrepancies between the published data on brass from different

Jan 1, 1949

By John Towers, Paul A. Beck, W. D. Manly

Recent work on grain growth in high purity aluminum and in a solid solution type alloy of aluminum and magnesium1 showed that the isothermal increase of the average grain diameter D with time follows the equation D = K(tg + A)2 Eq 1 Here tg is that portion of the total annealing time 1, which is available for grain growth, that is the annealing time less the time necessary for complete recrystallization. The parameters A and n depend only on the temperature. A has the dimension of time, and is usually of the same order of magnitude as the time for recrystallization: R. In some instances R may be substituted for A in Eq I. This is the case for example at long periods of annealing or at high temperatures, where both A and R become small in comparison with t. In such instances Eq I is simplified to D = K, tn Eq 2 For high purity aluminum the exponent n increases in the ratio of about I to 5 in. the temperature range of 350' to 6w°C. 70-30 brass has been long known to exhibit grain growth of the continuous type, such as occurs in pure metals. For this reason, it was of interest to determine whether grain growth in brass also follows Eq I and 2 as it does in aluminum. Isotherma1 grain growth data for 70-30 brass were published in recent years by R. S. French2 and by H. L. Walker.' Fig I shows a log D vs. log t plot of French's data, together with the 500, 600 and 700°C data by Walker for 42.8 pct reduction by rolling. In such a log D vs. log t plot Eq 2 corresponds to a straight line with a slope of n. As seen in Fig I, the data of French and the 600°C data of Walker can be represented by straight lines. However, the 500 and 7c0°C data of Walker give curved lines and the average slope of his 500°C line is very different from the slope of the line representing French's data for the same temperature. Furthermore the curvature of the 700°C line suggests a tendency for grain growth to stop after 8 hr. If confirmed, this would constitute a definite deviation from Eq 2, not explainable on the basis of the specimen thickness effect found in high purity aluminum.' Walker's specimens were 0.125 in., or 3.18 mm thick, while the deviation occurs at a grain size of approximately 0.4 mm. A further point of divergence arose recently, when J. E. Burke suggested4 on the basis of Walker's data, that, instead of Eq 2, grain growth in 70-30 brass follows the equation D - D, = K . tgn Eq 3 Here D, is the grain size as recrystallized at the temperature in question. Although it was possible to show5 that, for pure aluminum, for which isothermal grain growth has been carefully investigated, Eq 3 is incompatible with the experimental facts, the situation remained somewhat obscure with respect to 70-30 brass. In view of the discrepancies between the published data on brass from different

Jan 1, 1949

By William J. Gillich, Gerald L. Moss

PREVIOUS studies of the shock loading of randomly oriented polycrystalline aggregates have firmly established that, after rather short load duration, hydrostatic compression closely approximates the shocked state.' However, McQueen deduced from pressure measurements with good time resolution that, during shock loading, metals begin to undergo a stress relaxation at the yield stress. He suggested that the relaxation is due to plastic deformation accompanying the change from a one-dimensional compression to a three-dimensional, hydrostatic, compression.' Stress relaxation may occur faster as stress is increased above the yield stress, but the existence of unidirectionally compressed states for finite times suggests that the concept of a unidirectionally distorted lattice should furnish considerable insight into the mechanical behavior of materials subjected to shock loading. Here, a crystallographic analysis is presented for the hypothetical situation where atom motion may be considered as unidirectional which, as indicated above, is a useful extension of the macroscopic concept of shock loading to the microscopic domain. Only motion caused by the com-pressive forces is treated; i.e., lattice vibrations upon which this motion is superimposed, as discussed by Fitzgerald, are ignored.3 The purpose of the analysis to be presented is to develop a description for single crystals of the lattice rotations that occur for idealized uniaxial strain deformation. The usefulness of such an analysis in realistic deformation is to give a method of specification of those properties of an initially distorted lattice which would affect the subsequent deformation. Non-axial complications that arise experimentally, such as off-axis rarefaction waves, must be treated in proper context. Consider the material element shown in Fig. 1 and let the compression be uniform at all regions of the element. Then, any interior point of the element is displaced in proportion to its distance from the base of the element. Accordingly, if the top of a unit element is displaced by an amount, (-K), with respect to the base, a point P(X, Y) will be displaced by (-KY). Crystal planes are, therefore, rotated by the compression. Analysis of this rotation may be accomplished from the geometry shown in Fig. 1 where the line OP is the edge of the plane {hkl} and N is its normal. The orientation of the crystal plane with respect to the shock direction is specified by the angle, before deformation. After deformation the edge of plane {hkl}

Jan 1, 1969

By E. L. Kinsella, F. S. White

Solids fluidization utilized in two ways for the commercial production of lime is described. Crushed —6 mesh limestone is dried and dedusted in a single bed reactor, then calcined in a 5-stage reactor. Construction, operation, and results obtained with both processes are given. THE solids fluidization process brought out by the Standard Oil Development Co. in the early forties for catalytic cracking of petroleum enabled rapid transfer of large quantities of heat from gases to solids under closely controlled temperature conditions. As the technique developed, it was natural that applications for the process outside the petroleum industry would be sought. The calcination of limestone to quicklime requires the transfer of a large quantity of heat at high and closely controlled temperatures. For these reasons, this possible application was one of the first to be investigated. Laboratory work conducted by the Dorr Co. indicated that lime produced by solids fluidization would be of exceptional quality. Calculation of theoretical heat balances showed that savings in fuel could be expected. Accordingly, a pilot plant of 10 tons per day capacity was erected by the New England Lime Co. at Adams, Mass., to substantiate the process. Operation of the pilot plant, previously described,' showed that excellent lime could be produced with a saving in fuel. The work also indicated the necessity of preparing feed to remove the finer fractions if dust losses were to be held to a tolerable level. A single-stage fluidized solids drier-stripper was developed to accomplish this sizing.' Subsequently a commercial lime reactor was erected at Adams, beginning operation in June 1949, followed by air sizer facilities completed in November 1951. This paper will describe these two applications of solids fluidization. The raw material for preparation of lime by the fluidizing process is a high calcium crystalline lime- stone having the approximate chemical composition shown in Table I. The stone is first prepared by crushing and grinding in conventional equipment to pass a 6-mesh sieve. After crushing, the material is sent to the FluoDry unit for drying and dedusting. Fluidized Drying and Sizing As shown in Fig. 1, the FluoDry unit consists of a -in. cylindrical steel shell, 9-ft ID by 22-ft 6-in. overall height. Attached horizontallv at the bottom is a 6-ft diam combustion chamber, approximately 7 ft 6 in. long. A refractory constriction dome, consisting of first-quality fire brick shapes, partitions the inside into two sections. The lower section is called the windbox and the upper the bed compartment. The vertical wall of the shell is lined with 6-in. first-quality rotary kiln blocks. In the bottom of the windbox a hoppered section is formed with castable refractory to permit the windbox to be readily cleaned via a 6-in. screw conveyor. The combustion chamber is similarly lined with refractory shapes. On the 6-in. diam section, 4 in. of insulating brick, HW-26, are used. The inner combustion chamber is lined with 4-in. Korundal brick to withstand the high flame temperatures encountered. An 8-in. steel feed pipe, entering vertically through the top and fitted with an air-operated cone valve,

Jan 1, 1953

By F. S. White, E. L. Kinsella

Solids fluidization utilized in two ways for the commercial production of lime is described. Crushed —6 mesh limestone is dried and dedusted in a single bed reactor, then calcined in a 5-stage reactor. Construction, operation, and results obtained with both processes are given. THE solids fluidization process brought out by the Standard Oil Development Co. in the early forties for catalytic cracking of petroleum enabled rapid transfer of large quantities of heat from gases to solids under closely controlled temperature conditions. As the technique developed, it was natural that applications for the process outside the petroleum industry would be sought. The calcination of limestone to quicklime requires the transfer of a large quantity of heat at high and closely controlled temperatures. For these reasons, this possible application was one of the first to be investigated. Laboratory work conducted by the Dorr Co. indicated that lime produced by solids fluidization would be of exceptional quality. Calculation of theoretical heat balances showed that savings in fuel could be expected. Accordingly, a pilot plant of 10 tons per day capacity was erected by the New England Lime Co. at Adams, Mass., to substantiate the process. Operation of the pilot plant, previously described,' showed that excellent lime could be produced with a saving in fuel. The work also indicated the necessity of preparing feed to remove the finer fractions if dust losses were to be held to a tolerable level. A single-stage fluidized solids drier-stripper was developed to accomplish this sizing.' Subsequently a commercial lime reactor was erected at Adams, beginning operation in June 1949, followed by air sizer facilities completed in November 1951. This paper will describe these two applications of solids fluidization. The raw material for preparation of lime by the fluidizing process is a high calcium crystalline lime- stone having the approximate chemical composition shown in Table I. The stone is first prepared by crushing and grinding in conventional equipment to pass a 6-mesh sieve. After crushing, the material is sent to the FluoDry unit for drying and dedusting. Fluidized Drying and Sizing As shown in Fig. 1, the FluoDry unit consists of a -in. cylindrical steel shell, 9-ft ID by 22-ft 6-in. overall height. Attached horizontallv at the bottom is a 6-ft diam combustion chamber, approximately 7 ft 6 in. long. A refractory constriction dome, consisting of first-quality fire brick shapes, partitions the inside into two sections. The lower section is called the windbox and the upper the bed compartment. The vertical wall of the shell is lined with 6-in. first-quality rotary kiln blocks. In the bottom of the windbox a hoppered section is formed with castable refractory to permit the windbox to be readily cleaned via a 6-in. screw conveyor. The combustion chamber is similarly lined with refractory shapes. On the 6-in. diam section, 4 in. of insulating brick, HW-26, are used. The inner combustion chamber is lined with 4-in. Korundal brick to withstand the high flame temperatures encountered. An 8-in. steel feed pipe, entering vertically through the top and fitted with an air-operated cone valve,

Jan 1, 1953

By J. G. Savins

Certain types of macromolecules added to water and salt solutions flowing in turbulent motion can reduce the pressure gradient. Alternatively, the volumetric capacity of a pipe for these fluids is increased by the presence of these materials. Examples presented show that the drag reduction can become significant. Thus, the presence of 0.28 per cent of a gum derivative in a solution of sodium chloride flowing at 200 gal/min in a 1.89-in. pipe yields a pressure drop which is 0.44 of the single-phase drop measured under the same conditions of turbulent flow; the addition of 0.1 per cent of a vinyl derivative to a I-in. water line yields a throughput capacity which is 1.78 of the single-phase capacity at the same pressure drop. It is further shown that these phenomena are distinctly different from previous observations with other classes of non-Newtonian systems. There a simple lowering of friction factors below the levels predicted from the resistance laws for Newtonian fluids is associated with a suppression of turbulent motion. A rational physical explanation for drag reduction is advanced. Briefly, the proposed mechanism is a storage by the molecular elastic elements of the macromolecules in solution of the kinetic energy of the turbulent motion. INTRODUCTION This study was inspired by a recent review of some paradoxical drag reduction phenomena in turbulent pipe f1ow.I Under very moderate conditions of turbulent flow, the pressure gradient necessary to pump solutions containing certain specific kinds of polymers, fibers and metallic soaps may become appreciably lower than that required* to pump the solvent, i.e., water or a low-viscosity hydrocarbon, under identical flow rates in the same conduit. As shown by our review, this phenomenon of drag reduction in turbulent duct flow was first noted during the second world war, apparently arising in connection with the development of flame warfare weapons. Since that time several papers illustrating this phenomenon have appeared: Toms, Oldroyd,3 Agoston et al., Bundrant and Matthews, Robertson and Mason,6 Ousterhout and Hall,7 Daily and Bugliarello~ Lummus, Anderson, and Fox. 9 That there are practical applications for techniques which increase discharge or decrease the pressure necessary to transport a liquid through a pipeline is illustrated in the patents which have issued which take advantage of this peculiar phenomenon, e.g., Mysels,l0 Dever, Harbour, and Seifert.11 One also finds fragmentary evidence of this effect in the data pertaining to a few of the polymeric solutions studied by Shaver12 and Dodge.13 However, these investigators were concerned with the development of friction factor vs Reynolds number correlations for a variety of non-Newtonian solutions and suspensions, rather than in a study of drag reduction-. A similar kind of drag reduction effect has been observed in gases. Sproull, for example, reports that adding dust to air flowing in turbulent motion through a pipe results in a lowering of the pressure gradient at identical flow rates. There are also military applications for reducing the drag on hydrodynamic vehicles. For example, the possibility of injecting a rheologically complex fluid into the boundary layers of bodies to reduce the skin friction has been investigated by Fabulal5 and Granville.16 Along somewhat different lines are the drag reduction studies of Kramer.17118 He has shown that skin friction can be reduced by covering the surface of a vehicle with a flexible skin. The effect is apparently due to the boundary layer being stabilized by the presence of the skin. Drag reduction by means of coexisting gas and liquid boundary layers, e.g., film boiling and continuous gas injection, has been proposed by Bradfield, Barkdoll, and Byrne,19 Cess and Sparrow,20'21 Sparrow, Jonsson, and Eckert.22 Here the skin friction occurs between a vapor and a surface rather than between a liquid and a surface. There are several references in the literature to friction-factor correlations for non-Newtonian solutions and suspensions: Shaver and Merri11,23 Dodge and Metzner, Clapp,25 and Thomas.26

Jan 1, 1965

Bethlehem Steel Co. has just brought an all-new iron ore mine into production at Marmora, Canada, about 120 miles east of Toronto. High grade pellets produced at the mine from open-pit magnetite ore travel 64 miles by rail to the port at Picton, then go by Lake boat to help supply Bethlehem's Lackawanna, N. Y., plant. At full production about half a million tons of pellets a year will be shipped, and mining will be at a scale of more than a million tons per year of ore. Aeromagnetic surveys carried out in 1949 by the Ontario Dept. of Mines and the Canadian Geological Survey revealed an anomaly that interested Bethlehem geologists. By 1951 the company had put down some 40 diamond drillholes ranging from 460 to 1800 ft to block out the magnetite replacement deposit which lies in pre-Cambrian limestone below about 130 ft of Paleozoic limestone. The orebody, roughly 2100 ft long and up to 400 ft wide, averages about 37 pct iron, too low for direct shipment. Not only construction of concentrating and pelletizing facilities, but 20 million tons of stripping lay ahead before production could start

Jul 1, 1955

By Ralph W. Marsden

The natural iron ores of the Lake Superior Region in the United States are being replaced by iron-ore concentrates produced from magnetite- or hematite-rich horizons in the Precambrian cherty iron formations. The production of natural ores will soon be less than one-half of the total ore shipped annually. The tonnage of natural ore produced will continue to decline until it is a minor percentage of the ore shipped. The natural ores include (1) soft ore, (2) hard ore, ( 3) conglomerate ore and ( 4) siliceous ore. The soft ores are the important ore type and are sometimes termed "Lake Superior type ore". These ores are in situ deposits of hematite and limonite formed by the leaching of silica and other gangue materials in the iron formations. The leaching is believed to have been done by downward moving surface waters. In deep ore deposits, generally over 1000 feet below the present surface, circulation and leaching action may have been assisted by hydrothermal activity. Hard ore deposits occur on the Vermilion and Marquette ranges where hematite or magnetite has replaced the silica (and possibly some silicates) in cherty iron formation. These ores commonly occur in structural situations favorable for the movement of hydrothermal waters. The ores are believed to be the result of hydrothermal action with iron, possibly from the iron formation, transported by hydrothermal waters to areas favorable for the replacement of silica by iron oxides. Certain stratigraphic horizons of the Lake Superior Precambrian cherty iron formations contain sufficient iron as magnetite or hematite, with a grain size and texture that will give adequate mineral liberation, to yield an ore-quality product after grinding and concentration. The magnetite-rich deposits are termed magnetite taconites. These ores occur on the Mesabi, Gogebic, and Marquette ranges. Magnetite taconite ores are unenriched cherty iron formation and are the result of the sedimentary deposition of the iron formation and the metamorphic history of the rock. The commercial quality hematite-rich iron formations in Michigan are termed jasper ores. The jasper ores are being mined on the Marquette and Menominee ranges. These ores are composed of specular or granular hematite associated with quartz and iron silicates. The jasper ores appear to be recrystallized iron formations in which iron, prior to recrystallization, occurred as hematite.

Jan 1, 1968

By David P. Shoemaker, Clara Brink Shoemaker

The I phase was discovered by Bardos, Malik, Spiegel, and Beck1 in the V-Ni-Si system at 1100°C and in the Mn-Co-Si and Mn-Ni-Si systems at 1000°C. Kuzma and Hladyshevskii2 and Kuzma, Hladyshevskii, and cherkashin3 have reported a new phase, which they named the S phase, in the Mn-Co-Si and the Mn-Ni-Si systems at 800°C at approximately the I phase compositions. The powder diffraction diagrams of the I and S phases, compared by Bardos el al.,1 are related but show distinct differences in relative intensities. Professor Paul A. Beck kindly sent us a specimen of the (v-Ni-Si) I phase for single-crystal study. Precession and Weissenberg diagrams taken with a tiny single-crystal fragment showed that the structure is monoclinic with a large unit cell. The cell dimensions were refined from a powder diffraction diagram taken with CrKa radiation in a 57.3-mm camera with our computer program LSCELP4 using only the indices of those reflections that were observed on Weissenberg diagrams (hnl, with n = 0 through 9 and n = 12) and precession diagrams (Okl, lkl, Izk0, hkl, hk2). Table I shows the agreement obtained between the l/d values calculated with the final values for the cell dimensions and the l/d values observed by Bardos et al .1 for I(V42Ni33Si25), by US for I(V41Ni36Si23), and by Kuzma and Hladyshevskii2 for S(MnpCoSi). The powder patterns contain many broad and diffuse lines. In fact practically all powder lines observed are near-superpositions of several reflections. The refinement program is not set up to handle several reflections contributing to one observed line: the position of each line is refined with the aid of one set of indices only. The final values for the cell dimensions were obtained from the lines of strong or medium intensity with the indices indicated by asterisks in Table I: a0 = 13.403(6), bo = 23.336(15), co = 9.129(6)A ß = 99.11(6) deg The standard deviations (given in parentheses) are those calculated by the refinement program from least-squares theory and do not take into account that the lines included in the refinement may not have been correctly indexed in every case or that the position of a line may also be partly determined by the overlapping reflections. The agreement between the calculated and observed I phase patterns is fairly good, although some discrepancies occur. Not all observed lines could be indexed with the incomplete set of sin-

Jan 1, 1968

By AIME AIME

AT a joint meeting of the Tulsa Geological Society and the Mid-Continent Section of the A; I. M. E., held at Tulsa on March 21, the history of unit development in the Kettle- man Hills field was discussed. Robert C. Patterson? Super- visor for the U. S. Geological Survey at Taft, Calif., was the speaker. An abstract of some of his remarks follows: On July 1, 1923, curtailment agreement for two years was made between the operators of North Dome Kettleman and the U. S. Department of the Interior. In this it was agreed to consider a plan of unit development, or other cooperate method of developing the structure. As a result, Kettleman North Dome Association was incorporated in December, 1930, as a non-profit cooperative corporation under the laws of California. Approved in January, 1931, by California Oil and Gas Supervisor and by Secretary of Interior. On Feb. 10, 1931, it became an active organization through election of executive officers and eleven- directors. General manager chosen. Association started to function April 1, 1931. Assumed control of 10,800 acres of proved and semi-proved land, as shown in Table 1.

Jan 1, 1932

By Herbert K. Kutter

This paper is concerned with the physical phenomena in the fracture process of presplitting and only indirectly with the establishment of the optimum presplitting parameters. Its nature is therefore qualitative, rather than quantitative.

Jan 1, 1968

By J. M. Steininger

The mechanism of flotation of sphalerite with sulthydryl compounds of different acidities has been interpreted in terms of chemisorption of unionized collector molecules on hydroxylated zinc surface sites. At a constant collector concentration of 1.5 x 10 -4 M, the upper pH limit of flotation has been found to be directly and closely related to the negative logarithm of the collector acidity constant, pKa The experimental results indicate that flotation is prevented when about 95% of the collector is ionized. At low pH, collector adsorption decreases because of the reduction in number of hydroxylated zinc surface sites. These two effects explain why sphalerite does not float with the more acidic sulfhydryl compounds. Typical anionic collectors for sulfide minerals are short chain xanthates or dithiophosphates which are nearly completely ionized in the generally alka- line flotation circuits.1,2 Flotation mechanisms have therefore been usually interpreted in terms of ionic collector adsorption and competition with ionic depressants of the same sign as the collector. A few attempts have been made, however, to consider adsorption of neutral collector molecules. Contact angle measurements obtained by Wark and cox3 have been interpreted by Cook and Nixon4 as representing adsorption of the free acid molecules on sulfide minerals such as pyrite, galena, chalcopyrite, and sphalerite. Cook and Nixon tried to show that the free acid concentration of the collector was substantially constant along the curves relating collector addition and minimum pH for bubble adherence. However, in a review of the subject,5 Sutherland presented new data on the flotation of pyrite with mercaptobenzothiazole which did not support the hypothesis of constant concentration of either collector ions or collector free molecules. He further showed that, with the assumption of constant occupation of surface sites, both theories lead to relations between pH and collector addition having similar mathematical expressions. In fact, with the more likely assumption of variable occupation of surface sites which results in more complex relationships, it appears impossible to

Jan 1, 1968

I-Mine Ventilation A C. CALLEN, Chairman E A. HOLBROOK, Vice-chairman H I. SMITH, Secretary CLARENCE E. ABBOTT ALFRED W. HESSE FRANK H. PROBERT GRAHAM BRIGHT G R. JACKSON GEORGE S RICE THOMAS H. CLAGTETT JAMES F. MCCARTHY E. J RISTEDT OSCAR A GLAESER W. B. PLANK GERALD F. SBERMAN, Coal Mining M D COOPER, Chairman HOWARD N EAVENSON JOHN A GARCIA THOMAS D THOMAS CHARLES ENZIAN JAMES W PAUL C M YOUNG JOHN J. RUTLEDGE Physiological Studies R R. SAYERS, Chairman C. L. BERRIEN W. D. BRENNAN A. J. LANZA

Jan 1, 1923

By M. J. Sinnott

WITH the advent of nodular iron as an engineering material, considerable interest has been shown in developing this material for various applications. Generally the strength, hardness, or ductility are the more common properties used for design purposes but occasionally applications involving a knowledge of the thermal conductivity of the material is required. A search of the literature showed, rather surprisingly, that relatively little work has been done on this property for all types of cast iron. Insufficient data are available on thermal and electrical conductivities to develop the Weidemann-Franz ratio, or the Lorentz relationship, for estimation purposes. There are no data on nodular irons and the published data on malleable irons are quite sketchy. For these reasons the conductivity of the group of nodular irons listed in Table I were measured. Irons I to IV were made in such a fashion as to maintain the carbon and alloy content essentially the same while varying the silicon content. Test specimens were cut from the mid-section of 4-in. keel castings. Iron V is a commercial casting of nodular iron. The test specimen of this iron was cut from the center of a large cast billet 3x10x26 in. All alloys were tested in the as-cast condition. A microstructural examination showed that each iron had spheroidal graphite but the size and distribution varied with the silicon content of the iron. The spheroid size decreased while the number of graphite nodules increased as the silicon content of the iron was increased. The matrix structures of each iron consisted of pearlite and ferrite with the proportions varying from predominantly pearlite in the low-silicon iron to mostly ferrite in the high-silicon iron. The commercial heat, iron V, was predominantly pearlite probably because of its higher manganese and phosphorus content. Two methods for measuring the thermal conductivity were used. The first was a modification of the techniques described by Powell and Hickman and Van Dusen and Shelton,² the second method was a simpler technique described by Jakob and Hawkins." Since the agreement of both methods as tested on two sets of duplicate samples was excellent, the simpler experimental technique was used. The results of the measurements are given in Table II along with a summary of the microstructural constituents. These conductivity data indicate that as the percentage of silicon increases there is a corresponding decrease in the thermal conductivity even though the structure is changing from pearlitic to ferritic. Such a change in microstructure is usually accompanied by an increase in conductivity. A third factor involved in the behavior of these irons is the size and distribution of the graphite nodules. The relative amounts of graphite do not change greatly but the size and distribution do. Generally graphite improves the conductivity but in the present case such an effect is probably masked by the alloying effect from the silicon and its dispersion through the matrix. The thermal conductivity of the nodular irons studied is less than that of regular foundry irons by some 10 to 30 pct and is less than that of regular malleable irons.' The cause of the decreased conductivity is probably the more highly alloyed fer-rites obtained in these irons and the changed mode of distribution of the graphite as compared to the regular foundry irons. References R. W. Powell and M. J. Hickman: Journal Iron and Steel Inst. (1946) 154, No. 2, p. 99. 'M. S. Van Dusen and S. M. Shelton: U. S. Bur. Standards Journal of Research, RP668 (1934) 12, p. 429. ³M. Jakob and G. A. Hawkins: Elements of Heat Transfer and Insulation. 2nd Ed. (1950) New York. John Wiley and Sons. ' J. W. Donaldson: Journal Iron and Steel Inst. (1933) 128, No. 2, p. 255.

Jan 1, 1954

M. G. CORSON.'-I am not going to discuss any of the papers presented but merely the question of season cracking of brass in general. Alpha brass has been known for nearly 500 yr. and from the very beginning a number of articles were made from sheets of this alloy by the process of spinning. Metal turning lathes were not yet known but wood turning lathes already existed and the wood turners used to make the mandrils and do the brass spinning themselves. Cases of season cracking were known even then. In 1915 I was connected with a small company doing a certain amount of brass spinning. Before the war its products never gave cause for complaint but their war-time products were cracking too frequently. The reason was, that due to the exigencies of war an excessive speed-up was attempted and the material simply could not take it. Today corrosion cracking of brass is most important in the case of cartridges for small arms ammunition and shell cases for large guns. The cracking takes place in storage and too frequently it is explained by the hypothetical presence of vapors of ammonia. However, I had the experience of purchasing some 20,000 rounds for the old Vetterli rifle from the Swiss government in 1905. The rifle and the cartridges were obsolete for many years and those which I bought were stored for at least 10 yr. None of them was found to be cracked or to crack in use. And the reason again was in the amount of work used in manufacturing the cartridges between anneals. The machinery today is more powerful and the amount of

Jan 1, 1945

By Shiou-Chuan Sun

THE fact that coals of different ranks and even of the same rank differ greatly in their amenability to iroth flotation is well known. In recognition of the need for an explanation of this phenomenon, two hypotheses have been suggested. Wilkinsl reported that the floatability of coals increased with an increase of the carbon content or rank. This postulate is handicapped by the fact that bituminous coals that possess moderate carbon contents are actually more floatable than anthracite coals that have high carbon contents, as shown in columns 6 and 9 of Table I. Taggart and his associates' implied that the difference of floatability between bituminous and anthracite coal was caused by the variation of carbon-hydrogen ratio. This is not applicable to the relative floatability of other coals and carbons. For example, column 11 of Table I shows that the carbon-hydrogen ratios of low-floating lignitic coal and non-floating animal charcoal are not only smaller than the moderate-floating anthracite coal, but are also similar to the high-floating bituminous coal. Furthermore, according to this hypothesis, high temperature coke-A (464), Ceylon graphite (1238), and lamp-black (357), all possessing extremely high carbon-hydrogen ratios, should be less floatable than other substances having much lower carbon-hydrogen ratios such as high volatile-B bituminous coal (11.9 to 22), anthracite coal (35.7 to 60.5), lignitic coal (15.6 to 33.6), and charcoal (13 to 26.2). However the former group is actually more floatable than the latter group. In this paper, a surface components hypothesis is Proposed to explain the different floatabilities of coals, carbons, and hydrocarbon minerals. The validity of the hypothesis is experimentally supported by the actual floatability, natural floatability, wettability, and adsorbability for neutral oils of coals, carbons, and hydrocarbon minerals tested. The combustible recovery of the flotation results, as used in this paper. was calculated from Eq. 1: P (100-Ep) 100 RWCP Rc= [1] F (100-E,) C, where R, is the percent combustible recovery; F and P are, respectively, the weight of feed and the weight of concentrate or product; E, and Ep are, respectively, the total percent of ash plus moisture in feed and in concentrate; Ru. is the percent weight recovery: and C, and C, are, respectively, the percent of combustible in feed and in concentrate. Except for ash and moisture content, all chemical components of a coal are assumed combustible. The experimental work included studies on flotation, ultimate and proximate analyses, contact angle tests, extractions of bitumen-A with benzene, adsorptions for liquid hydrocarbons, and wetting tests. Most of the flotation experiments were performed in a laboratory Fagergren machine; others were tested in a small Denver machine. The solid feed for the former was 300 g and for the latter was 30 g. The solid materials used for flotation were crushed to —48 mesh. After the mineral pulp in the flotation cell was agitated for 6 min and the pH was adjusted to 7.5 & 0.2 with sodium hydroxide or hydrochloric acid, a petroleum light oil having a viscosity of 5.73 centipoises at 77 °F was added and conditioned for 2 min. Finally, pine oil was introduced and the froth was collected for exactly 3 min. The weight ratio of petroleum light oil to pine oil was kept constant at 1.5 to 1. Tap water was used for all flotation tests. Contact angles were measured with a captive bubble machine. For each coal sample, three specimens were mounted in transoptic mounts and polished with levigated alumina, first on a sheet glass, then on a cloth-covered metal polishing wheel. The polished specimen was first washed with distilled water and wiped thoroughly on a cleaned linen pad, then transferred into the pyrex cell of the captive bubble machine and conditioned for 6 min., and finally measured for contact angles at three or more points. Except where otherwise stated, the induction time for each measurement was 1 min. The contact angle representing each material was obtained by averaging the measurements of three specimens. The linen pad was first washed with warm distilled water, then boiled 30 min in a 2N sodium hydroxide solution, and finally washed with distilled water until no trace of sodium hydroxide could be detected in the decanted solution. The cleaned linen pad was stored under distilled water. Immediately before using, the pad was rewashed and transferred into a clean pyrex petri dish partly filled with distilled water. The glassware and rubber gloves used were cleaned by soaking in sulphuric acid-potassium dichromate cleaning solution, followed by rinsing with distilled water. The polished specimens were handled only by glass forceps. The ultimate and proximate analyses were made in accordance with the ASTM standard procedures for coal and coke. The extractable bitumen-A was determined by weighing 1 g of —100 mesh sample and placing it in a desiccated and weighed ASTM aluminum-extraction thimble. The thimble was placed in condenser hooks and inserted into an extraction flask containing 100 cu cm of benzene. The flask was heated and the benzene vapor was condensed by water coils. At the end of 24 hr of percolation, the thimble was removed, desiccated, and weighed. Loss in weight of sample was taken as bitumen-A and calculated to dry and ash-free basis.

Jan 1, 1955

By Charles Schwab

WHEN I leave this life, as an employer of labor there is no one thing that I want so much to be engraven upon my monument as the fact that I have been one of the men who have worked, whether with my brains or my hands, for the advance-ment of industry in the United States, and to be regarded, as I should like to be regarded by labor with love and kindly remembrance for our years of association. In my relationship with my employees, I have always contended that every man should be paid for what he does. I will not employ or be associated with any kind of labor of which the organization insists that a man shall receive a fixed sum of money for his day's wages, however poor or good his day's labor may be. What applies to the common laborer applies equally well to the management, the brains of an industrial establishment. I have always thought of industry as a three-legged stool, with capital, management and labor as the legs. All three legs of that stool must be there before that stool can stand up. If you take any one of those legs away, it is going to topple over. The. three must stand together for successful business accomplishment. I believe that a man who has his labor to sell, to put it plainly, has just as good a right to bargain for the best price for that labor as those who employ him. It is a perfect right for him to do so. If the government of our United States is right, our method of handling labor in our Bethlehem plants is right because it is representative and democratic. We have our labor elect delegates; elect representa-tives from their own ranks, who shall meet frequently. That constitutes, as. it were, the House of our great Congress in Washington, the industrial house. We have our management, who represent the Senate, who meet independently. Working men have learned to express their views freely as to whether they are treated fairly and whether conditions are right or wrong.

Jan 12, 1922

Discussion of the paper of M. E. Lormbardi,. presented at the New York meeting, February, 1915, and printed in Bulletin No. 9S, February, 1915, pp. 209 to 215. I. N. KNAPP, Ardmore, Pa.-Some years ago I made a swinging spider to carry a string of casing in a way somewhat similar to that described in this paper. The ground I tried it in was too soft for its successful operation. The spider had a central one-flange bushing to carry the slips. Under this flange and on top of the spider was a properly protected ball runway, or ball bearing, which permitted rotating the casing when it was in place and supported by the slips.

Jan 5, 1915

By G. P. Chatterjee

HARDNESS is one of the most common properties of metals and alloys but is one of the most difficult to define or express in terms of the fundamental units. In the Brinell or Vicker's hardness test, the hardness number N is defined as N = L/A where L is the load in kilograms and A is the area in mm'. This does not necessarily imply that a pressure of N kilograms per mm2 actually acts on the specimen.' It is essential that for a true comparison of hardness values the angle of indentation be kept constant,' a which is not done in the Brinell test. For this and other reasons, the Brinell hardness value is not based on a satisfactory physical concept.' The concept of hardness based on the work done per unit volume of indentation3-6 offers a simple and scientific definition. An attempt has been made here to develop this concept (at constant angle) and to compare the values so obtained with the Bhn for different metals and alloys. In Fig. 1 a vertical load L is applied to a compara- tively hard ball of diameter D creating an indentation of diameter 6 and depth h on the specimen S. The work done, W, is then given by W = So Ldh. [I] Since L is not independent of h, Honda and Takahasip and Matsumura" assumed different functions leading to equations which do not apply equallye to a wide variety of metals and alloys. A better relation is obtained from Meyer's Law L = ad" [2] where A and n are constants depending on the nature of the material but independent of the ball diameter D. From Fig. 1 it is obvious that h = R(l — cosq) and d = 2R sin 8.

Jan 1, 1957

By J. L. Brimhall, R. A. Huggins, M. J. Klein

IN a recent paper1 it was shown that small additions of magnesium, copper, and oxygen decrease the stacking fault probability in plastically deformed silver. Correlation of :X-ray data with measurements of the twinning frequency and metallographic observation of slip lines indicated that this effect may be due to an increase in the relative stacking fault energy. It is the purpose of the present note to report experiments on these alloys of known stacking fault probability that relate to the stacking fault contribution to electron scattering in silver. There has been a considerable amount of disagreement in the literature concerning the contribution of dislocations to electrical resistivity in metals. Calculations relating to the scattering of conduction electrons by the dislocation core, its strain fields, and related effects have been made by several authors.2-5 In addition, experimental work on Cu,3,6-10 Ni,11 and Ni-Co alloys1:! has shown that dislocations in those materials exert an anomalously large effect on resistivity, this effect being related to the presence of extended dislocations. The problem of the scattering due to the stacking fault region between partial dislocations has been treated theoretically using several different procedures.13-'0 The results of these calculations show a wide range of predicted values, from rather large to completely negligible, due to the models and the approximations involved. Almost all of this work has been applied to the specific problem of scattering in copper. The more complete theoretical treatments depend rather critically. on the shape of the Fermi surfaces and the effective mass parameters. The scattering from stacking faults might therefore be expected to vary significantly from one metal to another. Unfortunately, sufficiently accurate information is not available .which would allow one to have any confidence in a theoretical prediction of the scattering in the case of silver. Resistivity measurements were made on small wire specimens of vacuum melted high purity silver and several dilute alloys. These measurements were made at approximately 20°C, with great care being given to temperature measurement and corrections. At such temperatures phonon scattering is significantly greater than imperfection scattering. However, the experimental technique was such that measurements on similarly treated samples were reproducible to better than 2.4 X 10-9 ohm-cm, which was satisfactory for this work. Values of the stacking fault probability and the increase in resistivity after severe plastic deformation are given in Table I. The increases in resistivity resulting from plastic deformation agree quite well with previous work on silver.14,21-26 In addition, it has been previously found that interetitials and vacancies in silver anneal out well below room temperature, so that the residual imperfection scattering in these experiments could be expected to be primarily associated with dislocations;. It is seen from the data in Table I that large changes in stacking fault probability are accompanied by differences in resistivity due to imperfection scattering of only a few percent. If the scattering of conduction electrons by extended dislocations is primarily due to the area of the stacking fault region independent of the stacking fault energy, and not the core and strain fields associated with partial dislocations, the resistivity increase should be proportional to the product of the dislocation density and the stacking fault width, as is the stacking fault probability. Therefore, the imperfection scattering caused by plastic deformation of the alloys should be considerably smaller than that found in the pure silver. Since this was not observed, these results sugegest the possibility that stacking faults do not themselves contribute appreciably to the electrical resistivity of silver deformed at room temperature. An alternative explanation of these observations is that the scattering by the stacking fault is roughly proportional to the stacking fault energy, rather than being determined by the total area of such faults. Sufficient evidence is not yet available to separate these two possibilities. This work was supported by the Office of Naval Research.

Jan 1, 1962