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By J. T. Norton, J. G. McMullin

E. R. JETTE*—The way this ternary was developed there are two directly determined points on each of the iso-thermals except the 700° isothermal, where I believe there is only one. How were the end points determined, and what was the basis for it? Whose work formed the basis for the silver copper line ? M. B. BEVER †—At most temperatures the authors determined three points of the isothermal line. For example it can be seen from Fig 3 and 4 how three points for the 600°C contour line were obtained. The authors took two additional points for each temperature from the copper-silver binary diagram contained in Hansen's volume on constitution diagrams. It is correct that the 700°C contour line is based only on one point in the ternary field which was found by a slight extrapolation of Fig 4. R. I. JAFFEE*—Are the tie lines representing phases in equilibrium parallel to the silver-copper boundary ? M. B. BEVER—Are you asking that as a matter of principle ? R. I. JAFFEE—I assumed that would follow from principle. I wonder if that is in accordance with the authors' ideas too. M. B. BEVER—I do not believe the authors obtained any information on the tie lines in the two-phase field. Of course, in the limiting case of very low gold contents the tie-lines approach to the silver-copper boundary, but generally in the two-phase field I do not think that the tie-lines have to be parallel to the boundary. E. R. JETTE—I am quite sure they do not have to be parallel. You cannot tell from this work in which direction they go. E. A. GULBRANSEN*—Were any pycnometric measurements made to check the lattice parameters of the alloys ? M. B. BEVER—I did not hear any-

Jan 1, 1950

By Carl Benedicks

IT is well known by every one who has had to deal with boiler tubes that these are often seriously affected by a sort of corrosion, occurring as a local pitting, that frequently causes a perforation of the tube in a rather short time. The characteristic appearance of this corrosion is shown in Fig. 1. A very similar corrosion occurs in condenser tubes (Fig. 2).1 A study of the microstructure or local chemical composition of the metal (iron; 70 :30 brass) fails to reveal any cause for this pitting. Of course one might possibly assume the starting point-to be some oxide particle, causing an electrolytic local action, but this in no way explains the continued corrosion, progressing even when the assumed oxide particle must have disappeared. By the excellent work done by the Corrosion Research Committee of the Institute of Metals, and communicated, in October, 1923, to the Institution of Engineers and Shipbuilders in Newcastle-on-tyne, stress was laid on the fact that a content of entangled air has a very obnoxious effect. In the following correspondence, Sir George Goodwin stated that "engineers have, for a long time past, regarded air as the big enemy to be fought against in all questions of corrosion"2 and Kenneth Fraser pointed out that the most vital point is "the avoidance or removal of possibly the chief cause of present day corrosion, i. e. entrapped air and gases brought out of solution."' In their reply, however, the authors state that they had "never for a moment believed that entangled air was responsible for the beginnings of corrosion, and express the opinion that the gradual liberation of air could not account for localized corrosion."4 On consulting the many textbooks on corrosion, as that

Jan 4, 1925

E. R. JETTE*—The way this ternary was developed there are two directly determined points on each of the iso-thermals except the 700° isothermal, where I believe there is only one. How were the end points determined, and what was the basis for it? Whose work formed the basis for the silver copper line ? M. B. BEVER †—At most temperatures the authors determined three points of the isothermal line. For example it can be seen from Fig 3 and 4 how three points for the 600°C contour line were obtained. The authors took two additional points for each temperature from the copper-silver binary diagram contained in Hansen's volume on constitution diagrams. It is correct that the 700°C contour line is based only on one point in the ternary field which was found by a slight extrapolation of Fig 4. R. I. JAFFEE*—Are the tie lines representing phases in equilibrium parallel to the silver-copper boundary ? M. B. BEVER—Are you asking that as a matter of principle ? R. I. JAFFEE—I assumed that would follow from principle. I wonder if that is in accordance with the authors' ideas too. M. B. BEVER—I do not believe the authors obtained any information on the tie lines in the two-phase field. Of course, in the limiting case of very low gold contents the tie-lines approach to the silver-copper boundary, but generally in the two-phase field I do not think that the tie-lines have to be parallel to the boundary. E. R. JETTE—1 am quite sure they do not have to be parallel. You cannot tell from this work in which direction they go. E. A. GULBRANSEN*—Were any pycnometric measurements made to check the lattice parameters of the alloys ? M. B. BEVER—I did not hear any-

Jan 1, 1950

At the meeting of the Board of Directors the following resolution was passed: "The members of the Board of Directors of the American Institute of Mining Engineers desire: to place on record their appreciation of the very great generosity of one of their colleagues and a member of the Board, George D. Barron, expressed in his gift of $5000 to relieve the financial stringency caused by the unusual expenditures of the Institute for war work and the remission of dues of those members on active service." By-Law VIII. was amended to make a slight change in District No. 5. The report of the Auditing Committee was accepted. Tellers were appointed to: count the ballots on officers and on the proposed amendment. The proposed budget for the year 1919 was presented. The following resolution re Secretary Emeritus Rossiter W. Raymond, which had been prepared, by request, by Dr. A. R. Ledoux was then unanimously adopted. "The Board of Directors of the American Institute of Mining Engineers would place upon its minutes its profound sense of loss and sorrow in the death of Rossiter Worthington Raymond, Ph. D., LL. D., Secretary Emeritus of the Institute. Both as one of its founders and as its Sec-retary for 27 years, his was the guiding spirit of the Institute for more than a generation. "During the greater part of this long period, it might almost have been said that the Institute was Dr. Raymond-and Dr. Raymond the Institute. When with the progress of growth and development, great changes were introduced, Dr. Raymond acquiesced in these in spite of some misgivings, such as those with which a father might contemplate the emergence of his child from the careful supervision of the home; but as Secretary Emeritus for the past eight years, lie was always ready with valuable advice and helpful suggestion.

Jan 3, 1919

JUST as I have told you that it is impossible to give an exact rule for the bell scale, so I say the same concerning the clappers. Yet, if one wishes the bell to sound well, it is necessary that it have a proportioned blow; for if it is struck with a clapper that is lighter than it should be, it will not give out all the sound that it contains, and if it is struck with a heavier one than it can support, it is forced and the sound is made crude and harsh, in addition to the danger of breaking the bell. Therefore-the clappers must be made so that they bear a certain proportion to the bell, and in order not to proceed by chance, I have diligently tried to find the method as well as possible. I have found that experienced masters [99] have given to bells weighing ten pounds a clapper of one and a half pounds [and, to larger ones, clappers as tabulated on the next page].* Larger numbers than these I have not found because, in truth, few of the extraordinarily large ones are made in our regions. In my opinion, the masters should not be limited to these rules for the differences in clappers but should proceed according to the kind of bell, using a certain discretion for it is not possible to give a definite rule.

Jan 1, 1942

By W. van der Knapp

A theoretical and experimental ana1ysis is given of the change in volume of a porous medium due to changes in external and internal pressures. The result enable one to deduce directly the effect of large increlents in stresses. It is shown that the three-dimensional representation of the volume of a nonlinear porous system is a cylindrical rules surface, the generating lines of Which have a slope according to the solid rock compressibility. Analytical expression for such a ,surface is very simple. It follows that the bulk and pore compressihilities depend only on the difference between pore fluid and external hydrostatic pressure, i.e., the effective rock frame pressure. From measurements on a number of sandstones as well as limestones it can be conduded that for practical values of this effective frame pressure the pore compressibility falls in the range between the compressibility of water and that of undersaturated crude oil. In general, pore compressibility is found to be higher the lower the porosity. For a particular limestone reservoir pore compressibility and porosity could he related by means of a simple empirical formula. Results from the described study find a direct applicxtion in material balance calculations and in prob1ems of liquid flow through porous media. Since the elastic constants of the rock bulk material enter into equatiorls for velocity of acoustic waves, the results are also of importance for calculating rates of wave propagation in these media. It is shown for sandstones with 1.5 to -10 prr cent porosity that laboratory measured porosities may differ from those lender reservoir conditions by about I per cent of their value. The difference for low porosity limestotnes can be of the order of 10 per cent. The subsidence caused by elastic deformation of reservoir rock is shown to he very small. General formulas relating pore and bulk volume to external and internal pressure can be derived, as has been shown by Biot1, Gassmann2 and Geertsma.2 Their theories describe the differential change in volume caused by a variation in pore fluid tension and external hydrostatic tension. It has been known for a long timei that the stress-strain relations of a packing of spheres are essentially noninear. Our experiments show that for almost all reservoir rocks which arc elastically deformable the relation between volume and loading is also a nonlinear one, resembling that of a packing of spheres very closely. The behavior of these media is often described by using the linear stress-strain relations for small incremental stresses and strains in the vicinity of a pre-stressed condition." The present study differs from these considerations in that it enables one to deduce the effect of large increments in the stresses directly. In analogy to the stress tensor for a homogeneous body the stress tensor for a porous body can be written in the conventional way: where the tensor components, o, and t, represent the total forces acting on the faces of a unit cube of porous material. This tensor can be split into two parts. a hydrostatic and a deviatoric part, as follows2,3,1 where the component of the hydrostatic stress system, a. equals The hydrostatic stress causes a change in volume whereas the deviatoric stress system gives rise to a change in shape only. One of the difficulties in describing changes of volume occurring in reservoir rocks is that the relations connecting these volume changes with external and internal stresses are essentially non-linear. A complete description requires the bulk volume, V,,, as well as the pore volume,V,,. to be known as a function of both the changes in pore fluid tension (v) and the hydrostatic tension on the outer boundary.

By Robert E. Green, Edmund G. Henneke

Expressions are derived for the orientation factors associated with dislocation damping due to dislocation motion in the basal slip plane. The results of the calculation of these factors for zinc crystals aye pvesepzted and the relevance of these results to ultvasonic-attenuation measuvements is discussed. In this paper expressions are derived for the orientation factors associated with damping of sound waves due to dislocation motion in the basal slip plane of hexagonal crystals. For zinc crystals, the orientation factors for dislocation motion in one slip direction only, as well as for simultaneous motion in all three basal slip directions, have been calculated at 1-deg intervals. The results of these calculations are presented in the form of contour diagrams in a sector of a standard polar coordinate graph. The calculations have been carried out for both the traveling- and the standing-wave conditions. The relevance of these results to ultra sonic-attenuation measurements is discussed. Associated with the calculation of the orientation factors, the modes of elastic-wave calculation were also evaluated, according to the procedure previously employed by Hinton and Green.' The results of these calculations are not presented here since similar results have been reported earlier by Mus-grave.2 However, it should be mentioned that, although the qualitative features of Musgrave's results are preserved, the quantitative values are somewhat altered since different values for the elastic constants of zinc were used. ORIENTATION FACTORS It is convenient to refer to three different sets of coordinate axes, Fig. 1, in the following calculations. First, there are the Miller-Bravais axes which are used in describing the Miller-Bravais indices of planes and directions in a hexagonal crystal. These axes are denoted in the conventional manner by a,, a,, a,, and c. The second set of axes used is a rectangular Cartesian coordinate system located as shown in Fig. 1. Note that the x axis coincides with the close-packed direction al [2110], the z axis coincides with the c axis [0001], and the y axis is orthogonal to these two lying in the [0110 ] direction. Finally, a form of Eulerian angles is used to locate the position of the specimen axis. The angle between the specimen axis and the c axis is denoted by 0 and the angle between the projection of the specimen axis on the basal plane and the a, direction is denoted by ø. The majority of observations on plastic deformation of metal crystals indicates that such deformation is caused by the motion of dislocations on a limited number of crystalline planes and in a close-packed direction. In crystals with hexagonal symmetry the most prominent slip plane is the basal plane (0001) and the slip directiops- are one of the close-packed directions, [2110], [ 12101, or [ 1120]. To properly set the stage for the following work, several points must be clarified. First, we want to consider the contribution to the attenuation of elastic waves by dislocations only. While the Granato-Lucke, model for dislocation damping is presently the one most quoted, it is not necessary to refer the present work to it. The only assumption that need be made is that the motion of dislocations absorbs energy from the elastic wave and therefore attenuates it. It is immaterial here how this energy absorption occurs. However, because dislocations have been observed to move most freely on the (0001) plane for hexagonal crystals undergoing plastic deformation, it will be assumed that it is the dislocation motion on this plane which attributes to the observed attenuation. The basic orientation factor, Ok, is essentially the same as that used by Green and Hinton for fee crystals,4 defined by the equation

Jan 1, 1968

By R. E. Holmes, T. W. Leland

The solution to the unsteady-state mass- and heat-transfer equations describing the adsorption of a dilute component from a gas stream flowing through a packed bed is readily applicable to the design of hydrocarbon recovery units. The solution to these equations depends on the assumption of a linear equilibrium equation to relate concentrations on the surface to concentrations in the gas. This assumption is shown to be justifiable for components which are very dilute in the gas stream. This approach permits the prediction of the effects of bed shape and size, cycle time, and gas mass velocity on the recovery of each component. It also allows the time and gas flow rate needed for regeneration to be determined. The equations involve a constant for each component and another constant characterized by the type of adsorbent. These constants must be evaluated from operating data. Values of these constants determined for silica gel are presented. This paper points out the type of experitrrental data needed for a thorough design of hydrocarbon adsorption units. INTRODUCTION The growing importance of adsorption processes to recover hydrocarbons from natural-gas streams has created an urgent need for better methods to design and predict the performance of these processes. Since the original use of solid adsorbents in natural-gas processing was for dehydration, most of the current design methods for hydrocarbon recovery re extensions of those developed earlier for dehydration. These methods are based on an adsorptivity or bed loading factor defined as the amount of a particular component adsorbed per unit weight or per unit volume of the packed bed. The recovery of a given component is then expressed as an empirical function of its adsorptivity on a given adsorbent. In spite of the simplicity and initial utility of these methods, they are not completely satisfactory in that they do not predict the effect of the dimensions of the bed in relation to flow rate and cycle time, and do not predict the distribution of various components along the length of the bed. It is possible, however, to make a considerable improvement in the design of these processes by applying the basic equations of mass and heat transfer applicable to gases containing small concentrations of adsorbing components flowing through granular beds. These equations have been developed and solved by Anzelius,' Furnas,% nd Hougen and Marshall." where G = mass velocity of the inert constituents (assumed to be methane), lb/hr-sq ft, y = weight of an adsorbing component in the gas per unit weight of inert, lb of a given constituent per lb of methane, x = bed length, ft, p, = packed bed density, lb/cu ft, P, = gas density, Ib/cu ft, f = void fraction in the bed, t = time, hours, and w = weight of a given component adsorbed per unit weight of adsorbate free solid, Ib/lb pure adsorbent. The term on the left side of Eq. 1 represents the flow rate per unit bed volume of a component entering less the flow rate per unit bed volume of that component leaving the differential bed length. The terms on the right side represent, respectively, the accumulation rates per unit bed volume on the solid and in the gas phase within the differential bed section. The last term is negligible for this application since pgfv is quite small in comparison with p, and G. The partial differential is determined by the mass-transfer rate between gas and solid in the differential section.

By John Lamont, Walter Crafts

The principle formulated by Grossmannl for calculating hardenability of steel by multiplying the ideal diameter of "pure" iron-carbon alloys by factors for grain size and alloying elements has been confirmed on many steels and has proved useful in designing substitute alloy steels. Grossmann found that the typical behavior of the multiplying factors was to increase directly in proportion to the amount of the alloy present in a steel. However, the multiplying factor for the effect of silicon on hardenability was found by Grossmann to increase to a lesser degree as the amount of silicon was increased. In view of the increasing use of silicon in heat-treating steels, the effect of silicon on hardenability is becoming more important. Calculation of the hardenability of high-silicon steels made at the Union Carbide and Carbon Research Laboratories, Inc., indicated that the multiplying factor for silicon increased directly in proportion to the amount of silicon present. A more extended study of high-silicon steels covering fairly wide ranges of grain size, carbon, manganese, and aluminum has confirmed that up to 2 per cent silicon the multiplying factor is directly proportional to the silicon content. Manganese and aluminum factors have also been! determined. The results of the study offer strong confirmation of the general validity of Grossmann's multiplying principle. Testing Procedure Steels on which the study of the effect of silicon on hardenability was based were made in high-frequency induction furnaces using Armco iron and standard alloying materials. The ingots were forged to bars, and sections were normalized prior to machining Jominy hardenability test specimens. The Jominy test was carried out under standard conditions.2 The depth of hardness penetration for determination of ideal diameter at "half hardness" was taken at the hardness level indicated in Fig. 29 in Grossmann's paper by the solid line for 50 per cent martensite in plain carbon steel. The depth of penetration of hardness on the Jominy specimen was converted to ideal diameter by the relation given in Fig. 28 of Grossmann's paper. Chemical analysis was determined on a sample taken adjacent to the Jominy specimen. Aluminum was determined as "acid-soluble aluminum" rather than as total aluminum, as it is probably more representative of the aluminum that is effective as an alloying agent. In addition to the elements nominally present in the steels, analyses were made for residual phosphorus, sulphur, nickel, copper, molybdenum and chromium on one heat of those made with the same materials during the same period. Actual grain size was determined by microscopic examination of small samples quenched in water after receiving the same thermal cycle as the Jominy hardenability specimens. Chemical analyses, heat-treatment, grain size, and hardenability are shown in Table I representing three ranges of composition.

Jan 1, 1943

By John Lamont, Walter Crafts

The principle formulated by Grossmannl for calculating hardenability of steel by multiplying the ideal diameter of "pure" iron-carbon alloys by factors for grain size and alloying elements has been confirmed on many steels and has proved useful in designing substitute alloy steels. Grossmann found that the typical behavior of the multiplying factors was to increase directly in proportion to the amount of the alloy present in a steel. However, the multiplying factor for the effect of silicon on hardenability was found by Grossmann to increase to a lesser degree as the amount of silicon was increased. In view of the increasing use of silicon in heat-treating steels, the effect of silicon on hardenability is becoming more important. Calculation of the hardenability of high-silicon steels made at the Union Carbide and Carbon Research Laboratories, Inc., indicated that the multiplying factor for silicon increased directly in proportion to the amount of silicon present. A more extended study of high-silicon steels covering fairly wide ranges of grain size, carbon, manganese, and aluminum has confirmed that up to 2 per cent silicon the multiplying factor is directly proportional to the silicon content. Manganese and aluminum factors have also been! determined. The results of the study offer strong confirmation of the general validity of Grossmann's multiplying principle. Testing Procedure Steels on which the study of the effect of silicon on hardenability was based were made in high-frequency induction furnaces using Armco iron and standard alloying materials. The ingots were forged to bars, and sections were normalized prior to machining Jominy hardenability test specimens. The Jominy test was carried out under standard conditions.2 The depth of hardness penetration for determination of ideal diameter at "half hardness" was taken at the hardness level indicated in Fig. 29 in Grossmann's paper by the solid line for 50 per cent martensite in plain carbon steel. The depth of penetration of hardness on the Jominy specimen was converted to ideal diameter by the relation given in Fig. 28 of Grossmann's paper. Chemical analysis was determined on a sample taken adjacent to the Jominy specimen. Aluminum was determined as "acid-soluble aluminum" rather than as total aluminum, as it is probably more representative of the aluminum that is effective as an alloying agent. In addition to the elements nominally present in the steels, analyses were made for residual phosphorus, sulphur, nickel, copper, molybdenum and chromium on one heat of those made with the same materials during the same period. Actual grain size was determined by microscopic examination of small samples quenched in water after receiving the same thermal cycle as the Jominy hardenability specimens. Chemical analyses, heat-treatment, grain size, and hardenability are shown in Table I representing three ranges of composition.

Jan 1, 1943

By John L. Lamont, Walter Crafts

THE principle formulated by Grossmann1 for calculating hardenability of steel by multiplying the ideal diameter of "pure" iron-carbon alloys by factors for grain size and alloying elements has been confirmed on many steels and has proved useful in designing substitute alloy steels. Grossmann found that the typical behavior of the multiplying factors was to increase directly in proportion to the amount of the alloy present in a steel. However, the multiplying factor for the effect of silicon on hardenability was found by Grossmann to increase to a lesser degree as the amount of silicon was increased. In view of the in¬creasing use of silicon in heat-treating steels, the effect of silicon on hardenability is becoming more important. Calculation of the hardenability of high-silicon steels made at the Union Carbide and Carbon Research Laboratories, Inc., indicated that the multiplying factor for silicon increased directly in proportion to the amount of silicon present. A more extended study of high-silicon steels covering fairly wide ranges of grain size, carbon, manganese, and aluminum has confirmed that up to 2 per cent silicon the multiplying factor is directly proportional to the silicon content. Manganese and aluminum factors have also been determined. The results of the study offer strong confirmation of the general validity of Grossmann's multiplying principle. TESTING PROCEDURE Steels on which the study of the effect of silicon on hardenability was based were made in high-frequency induction furnaces using Armco iron and standard alloying materials. The ingots were forged to bars, and sections were normalized prior to machining Jominy hardenability test specimens. The Jominy test was carried out under standard conditions.2 The depth of hardness penetration for determination of ideal diameter at "half hardness" was taken at the hardness level indicated in Fig. 20 in Grossmann's paper by the solid line for 50 per cent martensite in plain carbon steel. The depth of penetration of hardness on the Jominy specimen was converted to ideal diameter by the relation given in Fig. 28 of Grossmann's paper. Chemical analysis was determined on a sample taken adjacent to the Jominy specimen. Aluminum was determined as "acid-soluble aluminum" rather than as total aluminum, as it is probably more representative of the aluminum that is effective as an alloying agent. In addition to the elements nominally present in the steels, analyses were made for residual phosphorus, sulphur, nickel, copper, molybdenum and chromium on one heat of those made with the same materials during the same period. Actual grain size was determined by microscopic examination of small samples quenched in water after receiving the same thermal cycle as the Jominy hardenability specimens. Chemical analyses, heat-treatment, grain size, and hardenability are shown in Table I representing three ranges of composition.

Jan 1, 1943

By R. L. Featherly, J. R. James

The replacement or reconditioning of oil storage tank bottoms due to external corrosion is an expensive maintenance item to the oil industry. Cathodic protection as a means of mitigating this problem has proved very successful in the past few years. One pipe-line company recently used magnesium anodes to cathodically protect two 55,000 bbl and two 20,000 bbl tanks. The tank bottoms were protected as individual units with considerable factual data being taken on one of the 20,000 bbl tanks. Twelve 60-lb anodes were installed about 2½ ft from the tank on approximately 17-½ft centers around the tank. Previous installations and tests have shown that sufficient protection is provided when the tank-to-soil potential using a copper-copper sulphate half-cell as a reference electrode is — 0.90 volts at the edge of tank. Potential measurements taken a few months after completing the installation indicated that the potentials were higher than were needed for protection. Resistors were then inserted in the individual anode circuits which increased the life of the anode installation and still provided adequate protection for the structure. All connecting pipe lines were then insulated from the tank. This limited the anode current strictly to the tank resulting in a greater degree of protection with a still longer anticipated anode installation life. The total installation cost for this tank was less than $300 which amounts to $20 per year The corrosion of metallic structures in contact with soil or water has been a very serious and costly problem for centuries. The annual loss to our nation's pipelines alone is estimated at $200,000,000. Mitigation of this problem by various methods has been carried on quite successfully for a number of years. An electrical method called cathodic protection has been used with considerable success to control corrosive influences on pipelines, tank bottoms, cable sheathing and other buried metal structures. Corrosion of oil storage tank bottoms has been a serious problem among the oil industries for years. In addition to damage costs and loss of oil, the replacement of one tank alone represents several thousand dollars. A reconditioning program at periodic intervals was instituted by several organizations and found to be a very expensive process. A few years ago two of the pipeline companies installed cathodic protection units at their tank farms.3'4 The average current required for protection of the tank bottoms varied from 0.7 to 1.0 milliampere per square foot. The results indicated that adequate protection was obtained at a cost of approximately 10 pct of the previous reconditioning program. The Republic Pipe Line Company experienced severe corrosion on their oil storage tanks. One of their 55,000-bbl tanks

Jan 1, 1948

By R. L. Featherly, J. R. James

The replacement or reconditioning of oil storage tank bottoms due to external corrosion is an expensive maintenance item to the oil industry. Cathodic protection as a means of mitigating this problem has proved very successful in the past few years. One pipe-line company recently used magnesium anodes to cathodically protect two 55,000 bbl and two 20,000 bbl tanks. The tank bottoms were protected as individual units with considerable factual data being taken on one of the 20,000 bbl tanks. Twelve 60-lb anodes were installed about 2½ ft from the tank on approximately 17-½ft centers around the tank. Previous installations and tests have shown that sufficient protection is provided when the tank-to-soil potential using a copper-copper sulphate half-cell as a reference electrode is — 0.90 volts at the edge of tank. Potential measurements taken a few months after completing the installation indicated that the potentials were higher than were needed for protection. Resistors were then inserted in the individual anode circuits which increased the life of the anode installation and still provided adequate protection for the structure. All connecting pipe lines were then insulated from the tank. This limited the anode current strictly to the tank resulting in a greater degree of protection with a still longer anticipated anode installation life. The total installation cost for this tank was less than $300 which amounts to $20 per year The corrosion of metallic structures in contact with soil or water has been a very serious and costly problem for centuries. The annual loss to our nation's pipelines alone is estimated at $200,000,000. Mitigation of this problem by various methods has been carried on quite successfully for a number of years. An electrical method called cathodic protection has been used with considerable success to control corrosive influences on pipelines, tank bottoms, cable sheathing and other buried metal structures. Corrosion of oil storage tank bottoms has been a serious problem among the oil industries for years. In addition to damage costs and loss of oil, the replacement of one tank alone represents several thousand dollars. A reconditioning program at periodic intervals was instituted by several organizations and found to be a very expensive process. A few years ago two of the pipeline companies installed cathodic protection units at their tank farms.3'4 The average current required for protection of the tank bottoms varied from 0.7 to 1.0 milliampere per square foot. The results indicated that adequate protection was obtained at a cost of approximately 10 pct of the previous reconditioning program. The Republic Pipe Line Company experienced severe corrosion on their oil storage tanks. One of their 55,000-bbl tanks

Jan 1, 1948

By E. S. Machlin

The wzodel of Blandin and Diplunt;, generalized to include a phase factor, is applied to the carbon-carbon interaction in iron. Darken&apos;s "energetic" model is generalized to include not only first neighbor interactions but further neighbor interactions as well. On the bases of these generalized models relations are derived for the activity of carbon in both austenite and ferrite in terms of the carbon-carbon Pair interaction energies. A single function then yields the pair interaction energies consistent with the experimental activities of carbon in both ferrite and austenite. Thus, a simple explanation is given for the observation that the nearest-neighbor interaction between carbon is repulsive in austenite and attractive in ferrite. Certain consequences of this approach are explored. OnE object of the present paper is to attempt to take into account the consequences of electrostatic contributions to the carbon-carbon pair interaction energy for carbon as a solute in iron. Friedel&apos; has shown that oscillations in electrostatic potential are to be expected about a solute atom in a metallic solution. Blandin and 6lant6&apos; have shown that such oscillations yield an interaction energy between pairs of solute atoms that obeys the relation: W{ = A cos(2ftFri + 4>)/(kFri)3 [l] where kF = Fermi wave vector, ri = distance between solute atoms comprising the pair7 <p = phase factor dependent only on electronic nature of solute and solvent, A = coefficient dependent only on electronic nature of solute and solvent. Machlin3 found that Eq. [I] accurately described the pair interaction energy derived from short-range order measurements based on field ion microscope observations of dilute alloys of platinum. He also found that the value of the phase factor $ derived from residual resistivity measurements agreed well with that obtained from the analysis of the short-range order data. Harrison and paskin4 were able to predict the long-range ordering energy of 0 brass using Relation [I] and residual resistivity values to predict the value of the phase factor $. Machlin5 has repeated their analysis and applied it to the prediction of the long-range ordering energy in AgZn and AgCd with excellent agreement between prediction and experiment. Both A and $ are independent of the crystal structure. The Fermi wave vector depends uniquely upon the conduction electron concentration per unit volume in the spherical approximation of the Fermi surface. Thus, Eq. [I] is expected to apply to both fer- rite and austenite with only one set of values of A and $. Mossbauer studies6 yield the result that iron has one 4s electron. We shall make an assumption found to hold previously for platinum3&apos;7 and nikel, which is that only the s electrons are involved in shielding the perturbing potential of carbon. With this assumption, kF = 1.35 A-l. Although A and $ may be obtained from certain mdels&apos;&apos;&apos; we shall take A and $ to be empirical constants in the spirit of Kohn and osko.&apos; Thus, Eq. [I] involves two adjustible parameters. Consequently, two independent relations in A and $I are required in order to evaluate them for carbon as a solute in iron. We may use a recent analysis of Aaronson, Domain, and poundg who showed that Darken&apos;s energetic model,1° as well as others, can be used to describe the activity-temperature data for carbon in iron in both the aus-tenitic and ferritic phases. Darken&apos;s model takes into account only first neighbor pair interactions. For our needs, all neighbor pairs need to be taken into account. It is convenient to generalize Darken&apos;s model. The result for the partition function for austenite is: over the temperature range 800" to 1200°C and where the uncertainty corresponds to one standard deviation. Eq. [4] effectively yields only one relation. Another relation is required to obtain unique values for A and $. One property of Eq. [4] is that it is independent of crystal structure. Hence, data for a iron can be used to obtain another relation. To arrive at this relation we must generalize Eqs. [2] and [3] so that they may be applied to the bcc a iron. The result is that:

Jan 1, 1969

By J. A. Putnam, A. D. K. Laird

In laboratory research on the behavior of oil, gas and water in porous materials, no direct method has been devised to measure saturation without disturbing the flow. Indirect methods involving various forms of radiation have been developed for measuring one or two components. The X-ray absorption method for two components has been described previously1-3. The possibility of three component measurement has also been mentioned1. Since two component measurement still requires considerable care to insure accuracy, the fact that three components can be measured satisfactorily is not generally appreciated. The fraction of the energy of an X-ray beam of one wavelength transmitted through a material is given by the transmission factor. in which 1 is the beam&apos;s path length through the fluid. p is the fluid density, and 1,. is its mass absorption coefficient. Eq. 1 implies that all parts of the beam pass through equal amounts of the fluid. Practical applications are, therefore, limited to systems in which the fluid is distributed so that I varies negligibly over the cross-section of the beam. Typical variations of the transmission factor with X-ray tube potential arc shown in Fig. 2 from which experimental points have been omitted for clarity. The X-ray absorption of low pressure gases in a core is practically zero so it cannot he distinguished from that a vacuum by instruments that will differentiate between liquids. The transmission factor of a gas in a core is. therefore, unity. Consequently, the saturation of only one gaseous component can be found. If two fluids completely fill a core, their two saturations add to unity and, therefore, only one value of the transmission factor, T, is needed to determine both of them. When two liquids do not completely fill the core, their saturations do not add to unity and transmission factors at two wavelengths must be measured. At one of these wavelengths, 7 for one liquid must be different from T or the other liquid. The change of T for one liquid us the wavelength is changed between the two values must be different from the corresponding change of T in the other liquid. The two values of fluid saturation will also give that of the gas because the three saturations must add to unity. In Fig. 2 it can be seen that there are many large differences of T between various oil solutions and water solutions at any one potential, and between potentials for any one solution. Consequently, conditions for optimum accuracy of measurement of the saturations of a gas and two liquids can be chosen. Once the operating potentials have been selected, however. it is more convenient to cross plot the data from Fig. 2 on transmission factor-solution concentration coordinates for the chosen potentials, as shown in Fig. 3. Fig. 3 was used to plan the experimental procedure for Core N-38-1. Since the voltage control at 45 kv was good, the slopes of the curves in Fig. 2 at this potential would cause little error. At low potentials, however, the voltage control was not good, so 33 kv was chosen because the small slopes of the curves made voltage control less critical. The lowest strength beam that could be measured reasonably at 33 kv corresponded to a minimum T factor of 0.385. Fig. 3 shows that both 4.80 weight per cent of cadmium chloride in water and 24.0 per cent of iodobenzene in crystal oil had this T value, and, consequently, were indistinguishable to the X-rays at 33 kv. Thus, at 33 kv, the liquid saturation was given by the same curve regardless of the relative amounts of oil and water solutions present. This simple calibration curve from which the gas and liquid saturation can be read, is shown in Fig. 4. Fig. 3 also shows that at 45 kv, T for the aqueous solution is 0.610. and that T for the oil solution is 0.220. These values were transferred to Fig. 4 to give the triangular calibration plot from which the brine and oil saturations can be found, after the gas saturation has been determined at 33 kv.

By V. P. Yashin

Modern trends in the use of autogenous grinding of minerals have been investigated by Mekhanobr Institute. The conclusions drawn are based on experimental and theoretical investigations by the author and his coworkers. A method is suggested for determining the efficiency of converting from ball milling to autogenous grinding. A qualitative interpretation is given for comminution theory using both methods. This interpretation is used as a basis for considering both the principal trends in increasing use of primary autogenous grinding and the results of such in industrial application. Data are given from testing some ores from USSR deposits for the effect of the ore Preparation step on subsequent metallurgical treatment. Autogenous grinding is shown to be widely used in various USSR industries.

Jan 1, 1980

By Cyril Smith

FOR several years an investigation has been in progress in the research laboratory of The American Brass Co. to determine the thermal and electrical conductivities of most copper alloys of commercial importance and several of scientific interest. This paper contains such data on the binary alloys of copper with silicon, aluminum, manganese and nickel, and on a large number of ternary or more complex alloys. Some inter-esting general relationships between the thermal and electrical conductivi-ties for all copper alloys at all temperatures are pointed out. The data for the copper-zinc, copper-phosphorus and copper-tin binary systems have already been published1,2 ? except for the electrical conductivity at 200° C., which is given in the appendix of this paper.

Jan 1, 1935

The North American Coal Corp. Increasing demand for coal, stimulated by the national energy shortage, and complicating and often conflicting, demands by government and ecology groups for better land reclamation methods continued to propel advances in surface mining technology during 1974. To meet current demands for coal (estimated at more than 700 million t for 1975 and 1.5 billion t in 19851, equipment manufacturers and coal operators concentrated on increasing the productivity of all surface mining equipment. The trend towards larger equipment, particularly in overburden removal, loading, and haulage machinery, prevailed during the year, notwithstanding the perennial debate over the economics and practicality of using large, inflexible equipment. Increasingly, the industry called on computers to monitor productivity and help design more efficient production systems

Jan 2, 1975

By E. P. Klier, Jane Jellison

The course of the transformations on cooling in a series of Cu-A1 alloys has been followed by means of thermal analysis, X-ray diffraction, and optical metallogrAphy. Specimen size was found to have a pronounced effect on the time of completion of the pearlite transformation due to modification in the unmber of nuclei. The ordeving transformation zvas observed for all cooling rates for all alloys studied, and is considered insuppvessible. At intermediate cooling&apos; rates, a transitional struchire for 011 alloy compositions and based ox the bainite reaction was observed. The alloys of aluminum and copper possess a eu-tectoidal region at from about 9.4 to 16 pct Al. Above the critical temperature, the disordered bcc phase, 0, is stable. At subcritical temperatures, the 0 phase decomposes under equilibrium conditions into the stable phases, a and y,. At the eutectoid composition, these phases develop a lamellar pearlite structure, in appearance much like that found in steels. During rapid quenching from supercritical temperatures, the diffusion processes required for the formation of the equilibrium phases are suppressed, and the 0 transforms to an acicular martensitic phase, designated B or y&apos; depending on the aluminum content, or to a mixture of the two. Early investigators of this system concentrated on determining the equilibrium phase diagram, &apos;&apos;&apos; the structure of the equilibrium phases,3 and later the isothermal decomposition of the 0 phase at subcritical temperatures, or the decomposition of the martensite on tempering. Other work which treated the nonequilibrium transformations concentrated largely on the martensite transformations and the structure of the b and y&apos; phases.15"17 Since the pearlite and martensite transformations in this system resemble those in steels, the one being formed by nucleation and growth of adjacent lamellae of the equilibrium phases, and the other by a diffusionless, shear-type transformation, the present study has been undertaken to investigate the structures formed on continuous cooling, with the objective of constructing cooling transformation diagrams analogous to those derived for alloy steel systems. In particular it was desired to determine whether parallels between the A1-Cu and iron-iron carbide systems on continuous cooling extended to encompass the intermediate phase, namely the bainite structures, which form by a combined martensitic-diffusion mechanism,&apos;&apos; but which are supersaturated with respect to the equilibrium phase upon which their crystal structure is based. To this end the transformations in representative alloys of the Cu-A1 B eutectoid system have been followed, using the methods of thermal analysis, X-ray diffraction, and optical microscopy. EXPERIMENTAL PROCEDURE Materials. The materials tested were those used by Klier and Grmko 5 in their study of the isothermal decomposition of these alloys. The alloys were melted in graphite crucibles and chill-cast into 5 by 2 by 1/4 in. ingots. The ingots were then hot-rolled to sheets 1/8, 1/16, and 1/32 in. thick. The compositions are given in Table I. Specimen Preparation. Specimens for thermal analysis were in the form of small squares approximately 1/4 in. on a side. For the 11.9 and 13.5 Al-Cu alloys, the sheet thickness was 1/16 in. and for the 10.5 A1-Cu alloy the 1/32-in. sheet was used. One lead of a chromel-alumel thermocouple was spot-welded to either face of the square with the specimen itself thus forming the hot junction of the thermocouple. Specimens for X-ray analysis were prepared as filings from martensitic sheet stock and were sealed in glass ampoules under 1/2 atm of argon. Metallographic examination of the powder specimens was carried out on a portion of the heat-treated filings mounted in lucite. All specimens were examined metallographic ally after etching

Jan 1, 1965

By R. W. Armstrong, P. C. Jindal

WE should like to point out that Hu and cline.1 in their study of the recrystallization of titanium, have obtained data which appear very important for accurately determining the dependence of hardness on grain size for this material. It has been proposed2,3 that the macrohardness. H. of a poly crystalline material should depend on its average grain diameter, l. in accordance with the relationship: H =H0 + kHl-1/2 [1] where Ho and kH are experimental constants. This relationship follows from the hardness dependence on yield stress or flow stress4 and. thence. the Hall-Petch stress-grain size relations.5&apos;6 Fig. 1 shows the Diamond Pyramid Hardness. Dph. values vs the reciprocal square root of the average diameter of the recrystallized grains, taken from the titanium data reported by Hu and Cline. The data in this figure give a linear relationship for the hardness over a wider range of grain diameters than anyone has previously measured. The data are not linear, for example. on an I-1 or l-1/3 scale. At the very large grain sizes the data are in reasonable agreement with measurements earlier reported by Holden, Ogden, and Jaffee.7 For the grain sizes smaller than 1 µ, which correspond in Fig. 1 to hardness values greater than 150, the material was reported as not being completely recrystallized and, therefore, these data appear less meaningful as far as the hardness dependence on grain size is concerned. However. it has been suggested by other investigators that deformed materials exhibit strengthening due to an effective grain size refinement even in the unrecrys-tallized condition.8-10 Also. Hu and Cline have shown that a change in texture occurs for titanium during the recrystallization process and. especially. following it. The influence of this texture change on the hardness-grain size relation would be expected to be less than that reflected in the unidirectional stress-strain behavior, even for an hcp material. because of the actual combined stress system which applies for the hardness test. For, at least, certain combined stresses, average values of the orientation factors should still occur in the stress-grain size analysis of a material exhibiting a preferred orientation.&apos; We previously suggested3 that the experimental constants. H0 and kH. from the Meyer hardness-grain size relation should be correlated with the Hall-Petch constants, uoy and ky.5,6 Our recent study" of the Hall-Petch relation and the hardness-grain size results which obtain for a number of polycrystalline materials shows that: The inequality [2] is explained by noting that the hardness should be correlated. perhaps, with the flow stress at some value of strain, as suggested by Dugdale,12 and. also, that the influence of strain on the Hall-Petch relation is to increase uo and not change k very much." Normally. Dph values and Brinell hardness values are nearly equivalent to each other and to the Meyer hardness number. For the case of titanium, then. the inequality [2] may be evaluated using the experimental values of ao = 8.0 kg per sq mm and ky = 1.3 kg per mm3/2 obtained from the experiment of Guard.6,13 Thus, (Ho/uo) = 12 and (kH/k3) = 1.3. This research has been supported at the University of Maryland through the Center of Materials Research by the Advanced Research Projects Agency of the U.S. Government.

Jan 1, 1969