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By F. H. Buttner, H. Udin, J. Wulff

Using a modified Udin, Shaler, and Wulff technique, the surface tension of gold Udin, purified helium was found to be 1400 ± 65 dynes per cm for the temperature range 1017° to 1042°C. IN the original Udin, Shaler, and Wulff technique for measuring the surface tension of copper: variously weighted wires were allowed to extend or contract in a copper cell held at elevated temperatures in vacuum. By plotting stress vs. strain for a wire array in one test, the stress at zero strain is obtained. This is the point where the contractile forces resulting from surface tension are balanced by the applied load, according to the expression: y = e=o r [1] where y is the surface tension in dynes per cm; a,,,, the stress at zero strain in dynes per cm; and T, the radius of the wire in cm. The assumption that the wires deform viscously permits the drawing of a straight line through the points on the stress-strain plot. Justification of the assumption has received further experimental support recently.'-' The presence of grain boundaries in the wires requires a correction to the original expression used." Thus: y = d4=T [l- (dl) (ar)Y1 [2] where, n/l is the number of grain boundaries per unit length, and a, the ratio of grain boundary tension to free surface tension. Alexander, Kuczynski, and Dawson in studying the creep of gold wire in vacuum were unable to obtain reproducible values of the surface tension of gold. In plotting stress vs. strain for progressively longer times, they found that the stress at zero strain drifted with time from positive stress values to negative values. Similarly, for the surface tension of silver, reproducible values were obtained only when a purified helium atmosphere was substituted.' Evidently the evaporation rate of silver in vacuum is too high at the temperatures employed to obtain solid-gas equilibrium even in a similar metal enclosure. Thus reproducibility of results is lost. Experimental Procedure The experimental procedure was much the same as that originally developed by Udin, Shaler, and Wulff with a few modifications and improvements. For greater accuracy in strain measurements, knots gave way to cut gage marks as shown in Fig. 1. These were made with a hand-driven lathe in which razor blades serv'ed as cutting tools. Also a more precise cathetometer with a screw accurate to 0.00015 cm was used. The tests were conducted in an atmosphere of purified .helium rather than in vacuum in order to avoid possible evaporation difficulties. Five mil wire of high purity gold (99.98 pct) was used. After cutting in the gage marks, each wire of a series of about 12 was differently loaded by welding a gold ball to one end. This was done by dipping the end of the wire in a cooling gold droplet, previously melted on a charcoal block with a No. 2 acetylene torch. The other end of the wire was strung through a hole in a gold lid and twisted over the edge to hold the wires fixed and in suspension from the lid. The lid and mounted wires were then dipped in pure ethyl alcohol to dissolve any skin oils and dirt on the surface of the wires due to handling. Finally the lid was put in place on an alundum crucible lined with gold so that the wires hung freely within the gold-lined chamber. This whole assembly was next heated in a quartz nichrome wound tube furnace and heated for a few minutes at 600°C to soften the wires. After this anneal the wires were easily straightened with tweezers. The wire assembly was finally annealed 10" to 25°C above the subsequent test temperature for 2 hr. This treatment allowed the grains to grow to equilibrium size and shape. After the anneal, the lid was mounted in front of the cathetometer. The gage length was measured by sighting the 40 power microscope on the upper lip of the lower gage mark for the first reading, then traveling up to the lower lip of the upper gage mark for the final reading. This procedure was repeated four times to give an average gage length value. In this manner the annealed gage length and the final gage length could be measured to determine the strains. During all measurements, grain counts were made.

Jan 1, 1952

By C. W. Sherman, N. J. Grant

The correlation of the high temperature chemical properties of slag-metal systems with some easily measured property of either slag or metal at room temperature has been the goal of both process metallurgists and melting operators for many years. There are several rapid methods for estimating various constituents in steel in addition to the conventional chemical methods which are quite fast, but these do not reveal the nature of the slag as a refining agent, which is of primary interest to the steelmaker. Furthermore, there are several methods for examining slag, the three principal ones being slag pancake, petrographic examination, and the previously mentioned chemical analysis. The main objection to the last two is the lime required to make a satisfactory estimate of the mineralogical or chemical components. The objection to the first is the inadequacy of the information obtained. A new technique has been developed by Philbrook, Jolly and Henry1 whereby the properties of slags are evaluated from an aqueous solution leached from a finely divided sample of slag. It is known that the pH or hydrogen ion concentration (of saturated solutions that have dissolved certain basic oxides, notably calcium oxide) will indicate a pronounced basicity. Philbrook, Jolly and Henry devised the pH measurement technique in order to supply open hearth operators with a fast, reasonably accurate method of estimating slag basicity. They offered the method as an empirical observation and made no claims as to its theoretical justification. The results were presented as an experi-metally observed relationship which applied over an important range of basic open hearth slags. They found that, in plotting the measured pH against the basicity, the best relationship existed between the pH and the log of the simple V ratio, CaO/SiO2. Extensive investigation also showed that there were several variables in the experimental technique that influenced the results and necessitated following a standard procedure to obtain reproducible pH readings. These variables were: 1. Particle size of the slag powder used. 2. Weight of sample used per given volume of water. 3. Time of shaking and standing allowed before the pH was measured. 4. Exclusion of free access of atmospheric carbon dioxide to the suspension. 5. Temperature of the extract at the time the pH was measured. In subsequent investigations of the pH method by Tenenbaum and Brown2 and by Smith, Monaghan and Hay3 the general conclusions of Philbrook's work were reaffirmed. It was the object of the present investigation to extend the technique to a point where it could be used to evaluate slags of all types. Experimental Results PARTICLE SIZK OF SLAG POWDER A large sample of commercial blast furnace slag of intermediate basicity (V-ratio 1.15) was selected for the study. The slag had been put through a jaw crusher until all of it passed through a 20 mesh screen. Five fractions of this crushed material were separated, -20 to +40, -40 to +60, -60 to +100, -100 to +200, and -200 mesh. A representative sample of 0.5 g was removed from each fraction and the pH determined using the method of Philbrook. Check pH analyses on the sample fractions varied due to the different amounts of shaking. To eliminate this variable, a mechanical shaker was employed. In order to know the exact time of contact between the slag and water, it was found necessary to filter the extract at the end of the shaking period. Using the mechanical shaker and a filtering apparatus, similar runs were made on the five fractions for contact times of 5, 10, 20, and 40 min. Random checks gave reproducible results within 0.02 pH. The data are plotted in Fig 1. It can be seen from the plot that each slag fraction is hydrolyzed to an extent that is roughly proportional to the surface area exposed to the water. The (—100 to +200) mesh material changed very little in pH after 10 min. shaking time. The curves are symmetrical and lie in proper relation to one another. The —200 mesh curve appears to be somewhat flatter than the others, but this can be attributed to the portion of very fine material that is not present in the other fractions. The closeness of the (-100 to +200) mesh curve to the —200 mesh curve and the fact that a —100 mesh sample would contain amounts of slag down to 1 or 2 microns in diam were considered sufficient reasons for selecting a —100 mesh sample as representative of the whole sample of slag for the purposes of this investigation.

Jan 1, 1950

By David A. Thomas, David N. French

The lrnrdness of WC crystals has been measured with the Knoop indenter at loads of 100 and 500 g on the (0001) and (1070) planes. The hardness as tneasitred on the basal plane is 2400 kg per sq mm and shows little anisotropy. The hardness on the prism plane, however, shows a marked orientation dependence, with a low value of 1000 kg -per sq mm when the long axis of the Knoop indenter is oriented parallel to the c axis and a high value of 2400 kg per sq mm when the indenter is perpendicular to the c axis. Slip lines (Ire observed surrounding the microhardness indentations and they show slip on (1010) planes, probably in [0001] and (1120) directions. This slip behavior can be explained by the crystal structure of TVC, which is simple hexagonal with a c/a ralio of 0.976. The hardness anisotropy call be explained by [0001]{1010} and (1130) {10l0] slii) and the resolved shear-stress analysis of Daniels and Dunn. HARDNESS anisotropy is a well-known phenomenon and has been reported for many metals, with both cubic and hexagonal structure.1-6 For hexagonal tungsten carbide, WC, a wide range of hardness values is reported in the literature. For example, Schwarzkopf and Kieffer7 give a hardness of 2400 kg per sq mm and report a value of 2500 kg per sq mm by Hinnüber. Foster and coworkerss give the average Knoop microhardness as 1307 kg per sq mm with a maximum value of 2105 kg per sq mm. Although these values and the structure of WC suggest the likelihood of hardness anisotropy, no such measurements have been made. We first detected a large apparent hardness anisotropy in WC crystals about 75 p large, in over-sintered cemented tungsten carbide. Prominent slip lines also occurred around many indentations. This report presents further observations and interpretations of hardness anisotropy and slip in WC crystals obtained from Kennametal, Inc. Both Knoop and diamond pyramid indenters were used on a Wilson microhardness tester with loads of 100 and 500 g. EXPERIMENTAL RESULTS The carbide crystals tended to be triangular plates parallel to the (0001) basal plane of the hexagonal structure. The side faces were parallel to the ( 1010) prism planes. Specimens were mounted approximately parallel to these two types of faces and metallographically polished. Laue back-reflection X-ray patterns were used to orient the specimens, which werethen ground to within ±1 deg of the (0001) and (1010) planes. The Knoop hardness values measured on the basal plane are plotted in Fig. 1. There is only a small anisotropy, with a hardness range of 2240 to 2510 kg per sq mm. The additional points at angles from 52.5 to 67.5 deg confirm the sharp minimum hardness at 60-deg intervals, consistent with the sixfold hexagonal symmetry. The average hardness of all values obtained on the basal plane is 2400 kg per sq mm. While the basal plane shows only slight anisotropy, the (1010) plane exhibits marked hardness anisotropy, from 1000 to 2400 kg per sq mm. Fig. 2 shows the hardness as a function of the angle between the long axis of the indenter and the hexagonal c axis, the [0001] direction. The minimum and maximum occur when the indenter is oriented parallel and perpendicular to the [0001] direction, respectively. The anisotropy of the prism plane is contrary to that reported for hexagonal zinc and hard- However, the basal-plane anisotropy is similar to these two metals.1'2 To check the accuracy and reproducibility of the measurements, a series of fifteen impressions was made at 100-g load in the same orientation in the same area of the specimen surface. The average for all was 2040 kg per sq mm, with a range of 1950 to 2130 kg per sq mm, giving an accuracy of about ± 5 pct. Thus the slight anisotropy on the basal plane is almost within experimental error. Fig. 3 shows slip lines around the Knoop indentations on the basal plane. The slip traces are in directions of the type (1130). The presence of slip steps on the basal plane indicates that the slip direction lies out of the (0001) plane. Because WC has a c/a ratio of 0.976,' the shortest slip vector is [0001], which suggests slip systems of the type [0001] (1010). Fig. 4 shows slip lines around the Knoop intentations on the (1010) plane. These slip lines are inconsistent with [0001] slip but can be

Jan 1, 1965

By E. J. Roberts

Prior to 1916, thickening was an art, and any accurate decision as to what size of machine to install to handle a given tonnage of a specific ore must have been one of those intuitive conclusions, based on both intimate and extensive acquaintance with thick-ners and ore pulps. Then in 1916 "knowledge of acquaintance," became "knowledge about" with the publication of the Coe and Clevenger paper.' The unit operation of thickening had graduated to the status of an engineering science. The fundamental similitude relationships for the two major phases of the operation were defined so clearly that batch tests on models as small as liter cylinders could serve to specify protypes as large as 325 ft in diameter. It is quite apparent from reading the literature that Coe and Clevenger's contribution is not generally appreciated. In so far as the basic engineering relationships are concerned, the only real advance which has occurred in the 30 odd years which have elapsed since the Coe and Clevenger paper is the recognition of the effect of the rakes on the thickening process. Bull and Darby2 noted this in 1926, and the extensive use of the "gluten type" thickener, in which the effect is magni-fied, bears witness to its importance. Comings3 further verified this effect of the rakes. As a matter of fact, a number of papers show an apparent regression from the Coe paper in that the area determinations are made on the basis of a single test from One concentration of solids. Coe and Clevenger amply demonstrated that this is unsafe, since the controlling zone may be one other than that of the feed dilution. Comings3 neatly demonstrated this without apparently realizing it. Of course there have been significant advances in the application of the operation to industry. Open tray thickeners were introduced to save area; balanced tray thickeners, washing thickeners, and multifeed clarifiers were developed with all of their special hydraulic and mechanical problems. Combinations of all kinds have been introduced, such as combination agitators and thickeners, combination flocculators and clarifiers, combination thickeners and filters. With the establishment of the operation on a firm engineering foundation, installation was facilitated and expansion proceeded. There are still problems, of course, functional as well as mechanical. Sometimes the moisture in the underflow obtained in practice is not as low as is expected on the basis of the test data. Sometimes the underflow is so "thick " that its discharge and subsequent handling requires special attention. Island formation plagues some operators. The use of the thickener as a surge basin and blending tank in the cement industry poses unusual problems. Design of rakes and the drive mechanism must be continually im-proved. Corrosion problems must he overcome. Power requirements for raking the settled solids occasionally is the controlling factor as it was in the case of the all American Canal desilting installation. Other similitude relationships and design problems come into the picture when we enter the field of clarification or nonline settlement. We have an energy dissipation problem in introducing the feed and any models must satisfy the Froude model relationships. Autoflocculation requires detention which involves the same similitude laws that we encounter in the compression zone. Approach to an Exact Science The next step beyond having control of the similitude relationships is to understand the why of these relationships right back up the line to first principles. The ultimate might be that, if given the mineralogical composition of the solids and their size distribution together with an analysis of the suspending liquid, we could calculate the entire thickening behavior of the system. Then we could say we had reduced the operation to an exact science. True it might be more trouble getting this basic analytical data than to make our empirical determinations for area and volume, and we would need an ENIAC to calculate the results, but that does not detract from the desirability of such understanding. Considerable work has been done by the chemical engineers with this end in view. Comings,3 Egolf,4 Work,5 Kam-mermeyer,6 Steinour,7 and others have studied the problem. The writer has no final answer to the thickening story but would like to propose a picture of the mechanics of the two phases of thickening which has been found useful in understanding the subject and which leads to some convenient relationship in treating the compression step and arriving at the compression depth.

Jan 1, 1950

By J. P. Morris

PREVIOUS investigations1,2 have shown that alloying elements in liquid iron influence the thermodynamic activity of sulphur and thereby affect the partition of sulphur between metal and slag in the desulphurization process. For example, the greater efficiency of desulphurization in the blast furnace as compared to the open hearth can be attributed in part to a higher level of sulphur activity in blast-furnace metal due to the higher concentration of carbon and silicon. In the present investigation, a short study was made of the influence of manganese on the activity of sulphur in liquid iron and iron-carbon alloys. In contrast to carbon and silicon, manganese was found to decrease the activity coefficient of sulphur; and in iron-carbon alloys it counteracts to some extent the influence of carbon. However, at manganese concentrations normally present in the blast furnace or open hearth, the effect of manganese is small. Since manganese sulphide has a limited solubility in iron, manganese can act, under certain conditions as a desulphurizing agent. Considerable data on the manganese-sulphur product in carbon-saturated melts were obtained in the investigation and have been included in this report. The experimental procedure was the same as that used in the earlier investigations on the effect of silicon' and carbon' on sulphur activity. Briefly, the method was as follows: The molten alloy, contained in a graphite or sintered alumina crucible, was brought to equilibrium at a constant temperature with a mixture of hydrogen and hydrogen sulphide of constant composition by bubbling the gas through the metal. Samples of' the melt were taken for analysis at regular intervals by suction through a 2 to 3 mm bore silica tube dipped into the metal. The experiments were run in a graphite spiral resistance furnace with melts weighing 50 to 60 g. The gas bubbling tubes were made of sintered alumina and were 5/16 in. OD, 1/16 in. ID, and 24 in. long. Equilibrium was assumed to have been attained when the sulphur content of the liquid metal reached a constant value. During an experiment there was a rapid loss of manganese from the melt by volatilization. To offset this loss, small additions of manganese were made periodically. The rate of manganese addition needed to maintain a constant manganese concentration was determined in preliminary tests. In all of the experiments, deposits of manganese sulphide formed above the melts in a cooler region of the furnace. Apparently, these deposits resulted from a reaction between manganese vapor and hydrogen sulphide in the gas. To prove that manganese sulphide did not volatilize from the melts to a measurable extent, an experiment was run in which helium was bubbled through liquid iron containing both manganese and sulphur. Although manganese volatilized rapidly in this test, there was no appreciable loss of sulphur. Volatilization of manganese sulphide from a melt would have led to an apparent equilibrium condition in which the sulphur content of the metal was lower than the true equilibrium value. The experimental results are shown in the first seven columns of Table I. The data in the last two columns were obtained from the previous work on the effect of carbon' and show what the results would have been in the absence of manganese but with temperature, gas composition, and carbon content of the metal remaining the same. Comparison of the last four columns show that, in the presence of manganese, the sulphur content of the metal increased at equilibrium and the activity coefficient of sulphur decreased. However, the results show that, for manganese concentrations below 3 pct, the effect of manganese is small. The values for activity coefficient of sulphur given in Table I were calculated from the following relations: S (in liquid metal) + H2 (gas) = H2S (gas) [l] K ph2s/?s X %S X phg = 0.00251 [2] where K is the equilibrium constant for the reaction, PH2S and ph2 are the partial pressures of hydrogen sulphide and hydrogen, respectively, and ?s is the activity coefficient of sulphur. The standard state for sulphur was taken to be a 1 pct solution of sulphur in pure iron. The numerical value for K at 1600°C was determined in the earlier work. For the purpose of showing graphically the results of the tests run at 1600°C, the activity coefficients of sulphur were recalculated so as to correspond to a manganese concentration in the metal of 2 pct in each case. In the calculation it was assumed that the increase in sulphur content of the metal at equilibrium caused by the presence of manganese

Jan 1, 1953

By O. G. Paasche, Yuan-Shou Shen

THERE is a disagreement among the various authors about the exact manner of transformation of ZrCr2. Rostokerl and others2 stated that ZrCr2 had a C-14 (MgZn2) type of structure below 1000°C and a C-15 (MgCu2) type of structure at temperatures above 1000°C. Alisova3 and others4 reached the opposite conclusion and stated that the transformation temperature is close to the melting point of ZrCr2. A literature survey shows that various investigators3'= who homogenized the specimens at a temperature higher than 1000°C have concluded that ZrCr2 had the C-15 structure at room temperature. Meanwhile, Jordan et al.4 reached similar conclusions without annealing the specimen. Other investigators1,2,6,7 who X-rayed the specimens in the as-cast condition without annealing reached different conclusions. The investigation reported herein was conducted with the aim of exploring the exact manner of transformation of ZrCr2 by various heat treatment tests. The alloys for this examination were prepared from iodide-reduced zirconium crystal bars, 99.9 pct purity, and electrolytic chromium, 99.9 pct purity. They were melted in a nonconsumable electrode arc furnace with water-cooled copper crucible in a helium atmosphere. The melting loss of each alloy was less than 1.5 pct by weight. Chemical analysis of a randomly selected specimen indicated that there was a very close agreement between calculated and analyzed compositions. Before being heat-treated each specimen was encapsulated in a vycor or quartz tube inside which an argon atmosphere was maintained at a pressure of lower than 1 atm. In determining the crystal structure of each specimen with a Debye-Scherrer camera, the standard procedure8 for X-ray quality analysis (Hanawalt method) was followed. The different series of heat treatment tests in this investigation are tabulated in Tables I and 11. The tests in Series I, specimens from 1-1 to 1-9, which were similar to Rostoker's experiment1 indicated that the transformation temperature seemed to fall between 870° and 900°C and that the crystal structure of ZrCr2 at lower temperature seemed to be of the C-14 type. However, once the compound is transformed to C-15 type, it is impossible to reverse the transformation back to the C-14 type by first heating the specimen above 900°C and then annealing it slowly below 900°C as shown in Experiments II-1 to II-3. Thus, it appears that the specimen of ZrCr2 will transform from C-14 to C-15 structure when heated above 900°C but will not transform from C-15 to C-14 when annealed slowly passing 900° C even after the extremely slow cooling process such as indicated in the experiment of Specimen II-3. As a valid transformation temperature is a temperature at which the transformation is reversible, therefore the temperature 900°C (or other temperature close to 900°C) is not the transformation temperature for ZrCr2 and the C-14 structure is not the stable structure of ZrCr2 at lower temperatures. The C-14 structure is retained at room temperature because the transformation to C-15 structure is very sluggish and the fast cooling after melting does not allow enough time for the transformation to take place. Additional energy is required to alter the metastable condition of the C-14 structure. The sluggishness of this transformation was again demonstrated through another series of experiments. Four specimens with C-14 structure were taken. Then they were annealed at 900°C but each specimen was soaked for a different period of time, Table 11. X-ray diffraction patterns of this group indicated that the C-14 structure gradually disappeared as the soaking period was lengthened. The figures listed under the column "C-14 Structure, pct" were estimated from the intensity of the d = 2.330 line of the diffraction pattern corresponding to the structure. Notice that the intensity of this line became weaker for longer soaking periods. To determine the transformation temperature of ZrCr2, specimens with C-14 structure (as-cast condition) were annealed at 1300°, 1400°, 1500°, 1550°, and 1600°C, respectively. A final specimen was first heat-treated to 1500°C in order to transform it to C-15 structure, then heat-treated at 1600°C again. From the X-ray analyses of this series of tests, Specimen Nos. III-1 to III-6, it is evident that a transition from C-15 structure at lower temperatures to the C-14 structure occurs at some temperature between 1550° and 1600°C.

Jan 1, 1969

By George B. Clark

Many basic engineering principles of all four phases of mining operations, namely, prospecting, exploration, development, and exploitation, can be analyzed better in terms of quantitative geology. Geological data from both field and laboratory will also complement scientific methods now being developed. THE geological data emphasized so successfully in prospecting for new deposits, that is, structural controls, strength of solutions, and type of mineralization, are basically those required for successful exploitation. In the mining of newly discovered deposits the most economical methods should be employed as early as possible to keep the overall cost per unit produced at a minimum and to permit maximum extraction of valuable minerals. A crucial question is: How can geological data be translated into useful quantitative results which will aid in achieving this end? H. E. McKinistry' has suggested that a solution may be reached in one of two ways: 1—the usual approach, use of judgment based on experience; or 2—mathematical calculations and tests on models, both subject to certain limitations. He also suggests that in addition to better use of geology more case data and theoretical data are needed on which to base sound judgment. Further research, therefore, is necessary. Perhaps in this field the emphasis should be on more specialization in mining methods and ground movement by men with thorough training in physics, engineering, geology, and underground mining. These specialists would be equipped to point out the most economical and scientific methods of exploitation. Selection of a stoping method is governed by the amount and type of support a deposit will require in the process of being mined, or by the possibility of employing the structure of the deposit to advantage in mining the ore by a caving method. In addition to these factors there are others which almost invariably influence the choice of an economical method of mining:' 1—strength of ore and wall rocks; 2—shape, horizontal area, volume, and regularity of the boundaries of the orebody, and thickness, dip and/or pitch of the deposit and individual ore shoots; 3—grade, distribution of minerals, and continuity of the ore within the boundaries of the deposit; 4—depth below surface and nature of the capping or overburden: and 5—position of the de- posit relative to surface improvements, drainage, and other mine openings. In the final analysis it is usually necessary to disregard the less important of these factors to satisfy the requirements of the more important. Because of the variation of geological conditions throughout and surrounding the deposit, no mining method will be everywhere ideally applicable to the conditions encountered in one ore deposit. The immediate problem is to interpret the above physical characteristics of deposits in terms of geological characteristics. Very few quantitative geological data are available on the factors related to a choice of mining methods. However, there are many descriptive data in mining and geological literature which collectively show how important an effect details of geology have upon all phases of mining operations. The following categories of basic mining methods were investigated to establish the geological factors that have affected their successful application: 1— open stopes with pillars; 2—sublevel stoping; 3— shrinkage stoping; 4—cut-and-fill stoping; 5— square-set mining; 6—top slicing and sublevel caving; and 7—block caving. It should be noted that the first five of these methods are listed in the order of increasing support requirements. Mines were selected as examples only where geological descriptions were complete enough to warrant their use. A study of the geological factors involved in mining operations led to a choice of the following classifications, employed in Table I: 1—structural type of orebody; 2—dimensions (geometry); 3— country rock (type); 4—faulting, folding, and fracturing; 5—alteration of ore and rock; 6—type of mineralization; and 7—geological factors determining mining method (summary). Of these factors only one yielded results that can be defined from available data in a quantitative manner, i.e., dimensions of the deposit. These are the most reliable guides that can be used in selection of suitable mining methods. They are, in general, the properties of geologic structure most difficult to evaluate by studies of models, pho-toelastic studies, and other laboratory methods, all of which are at present more limited in their applications than the geologic method. Application of geology has proved a reliable guide in other phases

Jan 1, 1955

By Norman Lamont

The evaluation of certain reser-voir properties, such as porosity and fluid saturation, from electrical well surveys has been widely accepted in petroleum engineering. Various investigators have established relationships between these properties and certain parameters which affect the response of the electrical log. Among these are the resistivities of the mud, its filtrate, and its filter cake. In 1949, Patnode1 established a relationship between the resistivities of the mud and filtrate. The well logging service companies have contributed relationships for the mud-mud cake resistivities2,3 These have been valuable since it was the practice to measure only resistivity of mud at the well site. During the mid-1940's the industry began drilling wells with oil-emulsion drilling fluids. These were conventional aqueous muds with a dispersed oil phase. Since 1950, oil-emulsion muds have been used on an increasing number of wells each year. However, the practice of measuring only the resistivity of the mud at the well site has continued, and the mud filtrate and mud cake resistivities have been determined by the above-mentioned relationships. Service companies are now equipped to measure all three resistivities at the well site. An investigation was conducted on the resistivities of oil-emulsion muds, mud filtrates, and mud cakes to determine if these values conformed to the relationships for aqueous muds. TYPES OF MUDS Fifty-one oil-emulsion mud samples were prepared in the laboratory following a standard manual' published by a leading mud company. The diesel oil in the samples varied from 5 to 50 per cent, the majority of the samples being in the 10 per cent region. The basic aqueous mud types which were converted to oil-emulsion muds were commercial clay and bentonite muds, low pH and high pH, caustic-quebracho treated muds, and lime treated muds. The emulsions were stabilized by dispersed solids, lignins, lignosulfo-nates, sodium carboxymethyl cellulose, or sulfonated petrolatum. It is worthy of note that after a quiescent period of two weeks at room temperature all samples, regardless of emulsifying agent, remained stable. The make-up water for the muds was from the laboratory tap. Resistivities were varied by the addition of table salt to the water. A range of mud resistivities from 0.44 to 3.9 ohm-m was obtained in this way. Twenty-three field muds were tested. These covered the same range of mud types as did laboratory muds. Oil provinces of the Gulf Coast, South Texas, West Texas, Oklahoma, Montana, and Canada were represented. MUD TEST PROCEDURE Each mud was tested for density, viscosity, pH, and filter loss by standard testing techniques. The resistivity measurements were obtained with a Schlumberger EMT meter. This meter required small volumes of sample, e.g., 2 mm. Filtrate was obtained from a Standard Baroid fil-ter press at the end of a 30-minute test. The filter cake from the same test was used for cake resistivity measurements. Mud, filtrate, and cake samples were heated to 100" F in a constant temperature water bath prior to measurement of resistivities. RESULTS The relation between mud resistivity (Rm) and mud filtrate resistivity (Rmf) is shown in Fig. 1. The solid line represents an average for the data. The equation of this line is Rmf =0.876 (Rm) 1.075 . . (1) Arbitrary limits, indicated by the dashed curves, have been set. The majority of the data falls within these limits, but some points do lie outside the limits. The approximate equation Rmt = 0.88 Rm , . . . . (2) will give satisfactory results within these limits. The data on mud cake resistivity Rmc is shown in Fig. 2. The solid line is an average for the data. The equation for the line is Rmc = 1.306 (Rm)0.88 The dashed lines are arbitrary limits on the data. Within these limits, Eq. 3 may be simplified to Rmc = 1.31 Rm . . . . (4) DISCUSSION The limiting curves in Figs. 1 and 2 represent maximum deviations of ±25 per cent. Thus the use of the average curves can introduce considerable error. There is no substitute for accurate measurements of mud, mud cake, and mud filtrate resistivities at the well site. The mud sample tested should be representative of the mud opposite the formation being logged. The average mud filtrate resistivity curve of Fig. 1 is reproduced in Fig. 3 with two curves which have been published for clay-base aqueous muds2,3. The latter curves were determined from average values of a large number of drilling fluids. The three curves have essentially the same slope and the differences between them are from 7 to 22 per cent. Comparison is made only to illustrate the possibility of error

Jan 1, 1958

The pressure of the gas in equilibrium with sphalerite has been determined in the temperature range of 680' to 825°C, using the Knudsen orifice method. A comparison of these experimental pressures with those calculated from thermal data and from other equilibrium measurements shows that the vapor above sphalerite is predominantly dissociated ZnS. Equations have been given for correctly calculating dissociation pressures using the Knudsen orifice method. It has been shown that the experimentally determined pressure is the same, whether the zinc sulphide is sphalerite or not, or a mixture of wurtzite and sphalerite. CONFLICTING points of view appear in the literature on the constitution of the vapor in equilibrium with solid zinc sulphide in the vicinity of 800°C. By comparing the dissociation pressure calculated from thermodynamic data and the vapor-pressure determination of ZnS by Veselovski,1 Lumsden2 has concluded that the vapor consists largely of dissociated ZnS. Sen Gupta,' however, concludes from his spectroscopic determinations that the vapor is largely ZnS molecules. In view of the fact that the thermodynamically calculated' dissociation pressure is higher than that experimentally measured by Veselovski, it seemed in order to repeat Veselovski's measurements. Experimental Procedure The method used for the determination of the pressures in this papel- is the Knudsen effusion cell. The apparatus and procedure were described in a previous paper- from this laboratory on the determination of the vapor pressure of silver. The only difference is that the Knudsen cell in this work is made from platinum and there is no external cover around the cell. The cell is an ordinary platinum crucible of 2.2 cm top diameter with a capsule cover. It was thought that platinum might stand up at these temperatures to the solid and gaseous ZnS, since it was found that the weight of the platinum cell itself did not change appreciably on heating ZnS in it at the working temperatures. To insure that reaction of the zinc sulphide with the cell was not giving' a false value, a stabilized zirconia cell was employed for check runs. Fig. 1 shows the comparison, which is satisfactory. Veselovski previously had measured the vapor pressure of ZnS using a silica Knudsen effusion cell. On repeating his experiment in this laboratory, it was found that ZnS at-tacked the silica cell, giving it a marked frosty appearance. This led to the belief that Veselovski's result:; may be in error. Also, he was operating at pressures above the range ordinarily considered safe for the Knudsen method. The effusion rate was measured by weighing the cell before and after each run. The weight loss during heating to temperature and cooling down was measured and subtracted from the weight loss during the actual run. The zinc sulphide used in this investigation was from two sources: Fisher cp grade, and a sample of pure sphalerite supplied by Mr. E. A. Anderson of the New Jersey Zinc Co. Before and after the series of runs with Fisher ZnS, X-ray analysis showed that both wurtzite and sphalerite were present. However, the ratio of sphalerite to wurtzite increased. All measurements were made below the transition temperature which has been reported" to be 1020°C. The data obtained in this investigation are tabulated in Table I. The pressure was calculated by the usual Knudsen formula" on the assumption that ZnS molecules were effusing. From these data, using pure sphalerite in the platinum Knudsen cell, the vapor pressure of ZnS, in mm of Hg, as a function of temperature is given by the solid line in Fig. 1. The best straight line, as determined by the method of least squares, is given by 14405 logpzns =-14405/T +11.032. A comparison of these results with Veselovski's shows that his results are about 50 pct lower. Discussion The vapor in equilibrium with solid zinc sulphide in the temperature range of this study will consist of Zn, S2, and ZnS mol, since other species of zinc and sulphur' are relatively unstable. The question to be settled is whether or not ZnS is largely dissociated. The derivation8 which follows gives the method of calculating the pressure of zinc and sulphur over solid ZnS, assuming complete dissociation, from Knudsen cell data. The free energy of the reaction 2 ZnS(solid) ? 2 Zn(gas) + S2(gas) is given by ?F?° = -RT In K = —RT In p12p2 where p1 is the zinc pressure and p is the sulphur pressure. If dissociation occurs in a closed system,

Jan 1, 1955

By V. Steve Reed, Ed L. Reed

The in situ leach mining process generates a waste stream that is high in sulfates, total dissolved solids, and radium 226. During the mining phase, the volume of the waste stream is relatively low and consists primarily of the bleed stream. During the restoration phase, larger volumes of waste water are generated. These waste streams require environrnentally sound disposal. The low net evaporation rate in the Coastal Bend area precludes pond evaporation as a feasible disposal alternative. Reverse osmosis is a practical method of reducing the volume of the waste water handled, but the concentrated waste stream from the reverse osmosis unit must be disposed properly. Deep well injection into highly saline reservoirs is considered a sound method of disposing of the liquid waste generated by in situ mining in the Gulf Coast uranium district. Thirteen injection wells have been permitted to serve the disposal needs of the leach mining industry in Texas. Of these 13, 11 have actually been drilled. Seven applications are pending. The injection zones for the permitted wells range from depths of 3050 to 6200 feet. Pressure limitations imposed on these wells range from 500 psi to 1350 psi. The following criteria are used to determine the desirability of a disposal well site: 1. A minimal number of nearby, improperly plugged borings which penetrate the disposal zone; 2. Minimal crustal disturbance; 3. Sufficient salinity of the water contained in the disposal zone; 4. Protection of oil and gas producing zones; and 5. Sand of sufficient permeability and areal extent to handle the desired volume without fracturing the reservoir. 1. Improperly plugged borings: During the early part of the century, oil wells, gas wells and test holes were drilled using cable tool equipment, often with a minimum amount of surface casing. Production casing, when it was set, was often partly removed when the holes were abandoned. Thus, wells drilled prior to 1940 frequently have less than 100 feet of surface casing and either no production casing or the upper part of the production casing removed. Additionally, these holes are often plugged only with mud. The close proximity of these holes to an injection well location are a concern in that they can provide an avenue for injection-depth fluids to migrate up the bore hole and jeopardize shallower fresh water reservoirs. Usually, where there are more than 6 or 8 poorly plugged borings in a 2 1/2 mile radius of the well site, it is preferable to examine deeper zones for disposal well potential. The deeper zones are especially attractive where the borings are not in a cluster, which renders monitoring more difficult. Often, even the deeper disposal zones are penetrated by a few improperly plugged borings. When this condition arises, the potential for leakage through the borings can be addressed in the following ways. a. Demonstration that the static head in the boring is higher than the anticipated increase in bottom hole pressure generated at the boring by the disposal well. A 100 psi differential between these two pressures is recommended. The calculated increased pressure at a boring caused by injection should be refined using annual bottom hole pressure measurements in the disposal well. Figure 1 illustrates an injection pressure map which can be overlain on the oil well map to determine the anticipated increase in pressure expected at each oil, gas or abandoned hole. b. Shallow ground water monitoring. A shallow monitor well is drilled next to the boring and both pressure and quality measurements are made periodically in the shallow well. c. Disposal zone monitoring. Recently there has been a tendency for regulators to require disposal depth monitor wells instead of shallow well monitoring. We consider disposal depth monitoring to be a less effective method of monitoring because it provides only indirect evidence of potential problems. Assumptions have to be made for the unplugged borings, such as mud weight, that are not addressed by the disposal zone monitoring program. There is little improvement with this system to that discussed in "a" above. A shallow zone monitoring program, however, yields direct evidence of a developing problem with an unplugged boring. Leakage by the boring will be detected quickly by an abnormal increase in pressure in the shallow well. Quality monitoring will detect upward migration of poor quality fluids. The pressure data provide an early warning of impending leakage; the quality monitoring will detect actual fluid migration.

Jan 1, 1980

By D. J. McAdam

In a series of papers the author and associates have discussed the influence of temperature on the tensile properties of metals.11-18 These papers present much information about the influence of temperature and the stress system on the conventional indices of mechanical properties, with special attention to the fracture stress. A recent study of the data, however, has revealed much additional information about the influence of temperature on the fundamental factors involved in the flow of metals. The present paper presents results of this study. Attention will be confined almost entirely to results derived from tension tests of unnotched cylindrical specimens at strain rates a little slower than those used in ordinary tension tests. According to a concept first presented by Ludwik and elaborated in recent papers by others,8,9,22,23 the mechanical state of a metal depends on the total plastic strain, but not on the temperature during straining, provided that the only structural changes are those essential to plastic deformation. In the summer of 1948, however, the author made the previously mentioned study of results of a general investigation by the author and associates and reached the conclusion that the mechanical state depends not only on the total strain, but also on the temperature during the straining. A number of diagrams were then prepared. These conclusions were presented without diagrams in a discussion last October of a paper by Dorn, Goldberg and Tietz.2 The metals used in the investigation on which this paper is based were Monel and oxygen-free copper. The Monel was supplied by the International Nickel Co. through the courtesy of Dr. W. A. Mudge. The copper was supplied by the Scomet Engineering Co. through the courtesy of Dr. Sidney Rolle. The data to be presented are based on results of tests at temperatures ranging between 165 and — 188°C. Description of the apparatus and methods of test are given in previous papers.1011'1"2 The present paper is the first part of the general discussion of the influence to temperature on the stress-strain-energy relationship for metals. The next paper will deal with metals that are subject to structural changes other than those induced solely by plastic deformation. Influence of Temperature and Plastic Strain on the Flow Stress of Monel and Copper For a study of the influence of temperature on the stress-strain relationship, flow-stress curves obtained with annealed metals at various temperatures will be compared with curves obtained with the same metals after cold drawing or cold rolling at room temperature. Diagrams thus obtained with Monel and copper are shown in Fig 1 to 8. Fig 1 to 7 show the variation of the flow stress with temperature and plastic strain; Fig 8 is a diagram of a different type, derived from Fig 4 to 7. In Fig 1 to 7 strain is expressed in terms of A0/A, in which A0, and A represent the initial and current areas of cross-section. Since values of Ao/A are represented on a logarithmic scale, abscissas are proportional to true strains; moreover, the true strains representing prior plastic deformation and those representing subsequent strain during a tension test are directly additive. Fig 1 shows flow-stress curves obtained with annealed Monel. Five of the curves are based on results of tension tests. Between yield and the maximum load, the flow was under longitudinal tensile stress; between the maximum load and fracture, the local contraction induced transverse radial tensile stress. The portions of curves designated F, therefore, represent flow with increasing radial stress ratio, the ratio of the transverse stress S3 to the longitudinal stress Si. Curve Fo is based on the ultimate stresses of specimens taken from bars that had been cold drawn various amounts.17 Since the tensile stress at the maximum load is unidirectional, curve Fo represents the course that a flow-stress curve would take if the stress during an entire tension test could be kept unidirectional. The flow-stress curve F obtained at room temperature (Fig 1) has been established accurately by numerous measurements of the diameter of the specimen during the extension from yield to fracture.17 At the time of the experiments, however, no apparatus was available for measuring the diameter during tension tests at low temperatures. Nevertheless, curves have been established to represent with sufficient accuracy the flow at low temperatures. Each flow-stress curve must be tangent to a curve U, which starts at a point representing the ultimate stress of annealed metal. Since the ultimate stress is based on the area of

Jan 1, 1950

By S. S. Arentz

PLANNED health and safety programs have become an essential part of American industry because such programs lead to increased operating efficiency, improved labor relations, better public relations, and to substantial savings in compensation insurance. Those of you who have had the unpleasant duty of informing the wife or widow of one of your men of his serious injury or death while on the job, know that all the benefits of a successful safety program do not show on the balance S. S. ARENTZ, Member AIME, is General Superintendent, Nevada Operations, Combined Metals Reduction Co., Pioche, Nevada. AIME San Francisco Meeting, February 1949. TP 2741 A. Discussion of this paper (2 copies) may be sent to Transactions AIME before March 31, 1950. Manuscript received Jan. 6, 1949. sheet. These programs are of particular importance to the mining ,industry because mining's reputation as an unusually hazardous industry and the commonly isolated location of mining operations tend to focus attention on these problems. Description of Operations: Before proceeding with a discussion of our health and safety programs at Pioche, it may be proper to give a brief description of Pioche and of our operations there. Pioche is one of the early Nevada mining camps. It was founded shortly after the discovery of high grade silver ore in 1863 and mining has continued with more or less regularity to the present day. In an era of lawlessness, Pioche was notorious. The story persists that 75 men died with their boots on before one died a natural death, and old payroll records show that nearly as many gunmen were employed to stand off claim jumpers as there were miners working the mine. That was probably as close to a safety program as the times permitted. Pioche is situated in southeastern Nevada on the main highway between Ely and Las Vegas. The camp is on the flank of "Treasure Hill," near the original silver discovery, at an elevation of about 6000 ft. The present day population of about 2000 is primarily dependent upon the mines of the area, although Pioche also serves as the county seat of Lincoln Couqty and as the center of the surrounding livestock industry. The camp is served by a branch of the Union Pacific Railroad and receives power from the generators at Hoover Dam. The Pioche operations of the Combined Metals Reduction Co. were started in 1923 when the first complex lead-zinc ore was shipped to the company's mill at Bauer, Utah. The modern mill at Pioche was completed in 1941. The operations are medium sized in the nonferrous field, employing an average of 350 men in the mine, mill, and related works. The complex lead-zinc ore is mined from replacement deposits in a comparatively flat, extensively faulted, limestone horizon. Mining methods vary from stull-supported open stopes to filled square-set stopes. The thin bedded limestone and shale overlying the ore is allowed to cave as areas are mined out and caving frequently follows closely upon ore extraction. The relatively heavy ground and the numerous faults add to the problems of safe mining. The mine is well mechanized and the mill and surface plant are modern and well equipped. Labor is organized in a C.I.O. union and labor-management relations have been unusually harmonious. During most of the period since 1923 a competent supervisory staff worked to reduce safety hazards but the primary responsibility for safety rested on the individual workman. Accidents happened and all too frequently they were regarded by all concerned as unavoidable. In October 1939, the late Robert L. Dean became superintendent at Pioche. Most of his previous experience had been in the fields of iron and coal mining and from that experience he brought the concept that no accident is unavoidable. Many of the features of our present health and safety programs were initiated by Mr. Dean during his term as superintendent. Health Program: Our health program centers in Dr. Q. E. Fortier and his new, well-equipped, and well-staffed, modern hospital in Pioche. The program starts with a thorough pre-employment physical examination and is followed by yearly re-examinations at the expense of the company. The Pioche Mutual Benefit Association, to which all Pioche mine operators and employees belong, pays benefits covering hospitalization and surgery expense incurred by employee members and their families. The Association is governed by a board of directors elected by its members. The mine operators of the district donated the original capital and pay the monthly dues of the employee members. The employees pay the dues covering members of their families. Though not strictly a part of the

Jan 1, 1951

By Bruce Chalmers, D. C. Larson

Aluminum crystals with longitudinal-axis orientations of (111) . (110), and (100) were deforined in tension and annealed. The conditions of deformation were controlled so that the re crystallization nuclei originated in either the heavily deformed regions at saw cuts {artificial nucleation) or in the lightly deformed matrix (spontaneous nucleation). The artificial-nucleatioln experiments showed that in lightly deformed (110) and (100) crystals low-angle twist boundaries are most mobile, while in (111> crystals and heavily deformed (110) and (100) crystals high-angle tilt boundaries with near (111) rotations are favored. The spontaneous-nucleation experiments showed the existence of preferred orientations in the (111) crystals. The nonrandomness of the grain orientations is quantitatively determined through a comparison with the results which would he obtained from a randowl set of grain ovientations. PREVIOUS recrystallization studies have been performed on single crystals deformed in tension.1 7 The crystals used in these studies usually had random tensile-axis orientations and the extent of deformation was not a primary consideration. The present study concerns the recrystallization of single crystals with tensile-axis orientations of (Ill), (110), and (100). The emphasis of this work is on the influence of the tensile-axis orientation and the degree of deformation on both the nucleation and growth processes. The multiple-slip orientations were chosen because secondary slip or slip intersection promotes nucleation.1,5,8 These crystals recrystallize at lower strains than the crystals which are oriented for single slip. Also, the greatest variation in deformation behavior is exhibited by the multiple-slip orientations. The stress-strain curves for crystals with tensile-axis orientations of (111) are higher than the stress-strain curves for poly-crystals, and the stress-strain curves for crystals with tensile-axis orientations of (100) are lower (at large strains) than the stress-strain curves for the crystals which deform initially in single slip.g The recrystallization nuclei originated in either 1) the homogeneously* deformed matrix of the crys- tals or 2) the heavily and inhomogeneously deformed regions at saw cuts. The nuclei will be referred to hereafter as spontaneous and artificial nuclei, respectively. The two terms do not imply a difference in the nature of the nuclei; they imply simply a difference in the mode of introduction of the nuclei. During spontaneous nucleation very few (always less than ten) grains nucleate, while during artificial nucleation large numbers of grains nucleate. Only a fraction of the artificially nucleated grains penetrate very far into the deformed matrix during annealing. The grains that penetrate the farthest into the deformed matrix will be referred to as the dominant grains. EXPERIMENTAL PROCEDURE The thirty-five crystals used in this investigation were grown from the melt in milled graphite boats at a rate of 1.6 cm per hr. The crystals had dimensions of approximately 6 by 12 by 80 or 6 by 6 by 80 mm and the aluminum was of 99.992 pet purity. The as-grown crystals were annealed at 610°C for 24 hr and furnace-cooled. They were then heavily etched and electropolished in a solution of five parts methanol to one part perchloric acid. The crystal orientations were obtained by back-reflection Laue photographs and were accurate to ±2 deg. The tensile-axis orientations were (loo), (110), and (111). Two of the side faces of the (111) crystals were (110) lanes. The (110) crystals had both {100) and {110) side faces and the (100) crystals had (100) side faces. The crystals were deformed at a strain rate of 0.003 per min. Shear stress and shear strain were obtained by multiplying and dividing the tensile stress and strain, respectively, by the Schmid factor, m. For the (111) crystals m = 0.272 and for the (110) and the (100) crystals m = 0.408. The Schmid factor is effectively constant during deformation for all orientations. The deformed crystals were sawed into 1-in.-long specimens while the crystals were totally enclosed in a graphite boat. The sawing was performed very carefully in order to limit the plastic deformation to the sawed regions. The specimens were electropolished in the solution mentioned above to remove the sawed-end deformation as well as controlled amounts of surface material. A special stainless-steel grip was used to hold the specimens during the electropolishing treatment. The gripping faces were flat, with no teeth, to prevent the introduction of extraneous de-

Jan 1, 1964

By N. A. Gokcen, J. Chipman

A revised treatment of the authors' published data eliminates the complex relation previously proposed between concentration of silicon and activity coefficient of oxygen in liquid iron. Revised values of the thermodynamic properties of the liquid solution are presented. IN a recent experimental study of the reaction SiO2 (s) = Si+ 2O; Kf, = [% Si] [% O]² [1] the authors' found a substantially constant equilibrium product in liquid iron at 1600°C of 2.8x10-5 They also reported extensive data on the reactions: SiO2 (s) + 2H2 (8) = Si- + 2H2O (g); K'2= [% Si] (H2O/H2 [2] and H, (g) + 0 = H2O (g); K'3 =( H2 O [31 (H2) [%O] From the results on reaction 3 and earlier data of Dastur² on this same reaction in the absence of silicon, they determined the activity coefficient of oxygen, f0, on the basis of the definition K3 = (H2O)/ (H2)f0 [% 0] where K, is the equilibrium constant and f0, is taken as unity in the pure Fe-0 system. Similarly values of fsi were deduced from results on reaction 2. In a more recent study" of analogous reactions in the system Fe-A1-0, it was found impossible to reconcile the results on reaction 3 with Dastur's data; accordingly the latter were ignored and the equilibrium results were extrapolated to find a value of K, at zero concentration of aluminum. This procedure failed to locate the cause of the discrepancy but it did yield reasonable values of activity coefficients. It also avoided introduction of the complex empirical relation between the oxygen activity coefficient and the concentration of the added element. The same type of discrepancy exists for system Fe-Si-0.' In the earlier paper an attempt was made to fit both sets of data by a single curved line (Fig. 6 of ref. l), the form of which is contrary to the theoretical requirement of a finite slope at infinite dilution. In the light of experience on the Fe-A1-0 system the discrepancy must be recognized as one which can be resolved only by more refined measurements. Accordingly Figs. 6 and 10 are retracted. It is pointed out also that until the discrepancy is resolved Figs. 7, 8, and 11 are subject to some uncertainty. Qualitatively the following conclusions still appear valid: 1—The activity coefficient of oxygen is reduced by addition of silicon. 2—In dilute solutions the activity coefficient of silicon increases with its concentration. 3—With respect to equilibrium in reaction 1, the above effects are approximately compensating. The discussion of K'1 in the previous paper requires no revision. It was pointed out that the constancy of the product [% Si] [% 0]² ndicated a compensating effect of the activity coefficients of silicon and oxygen. Therefore, as a very good approximation, K1 = K'1 and the following average values are suggested both for K, and K', at the temperatures 1550°, 1600°, and 1650", respectively, 1.0x10-", 2.8~10-" and 5.5 ~lo-'. Revision of the thermodynamic treatment is necessitated by the recent appearance of new data, based on a combination of combustion and solution calorimetry,' which yields for the heat of formation of low-cristobalite from the elements, the value —209,330 ±250 cal per mol at 25°C. This is about 4000 cal larger than the value previously accepted. The new value for cristobalite is used, together with Kelley's tables of high-temperature heat contents" and entropies and with Korber and Oelsen's' heat of fusion of silicon to obtain the following equation for the standard free energy of cristobalite in the temperature range 1700" to 2000°K: Si (1) + 02 (g) = Si02 (crist.); ?F° = -217,700 + 47.OT [4] The free energy of solution of 0, in liquid iron is:8 O2 (g) = 20 (in Fe); AF° = -55,860 - 1.14T [5] and these two equations are combined to give: Si (1) + 20 = SiO, (crist.); AF° = -161,840 + 48.14T [6] ?F°1873 = -71,700 cal. From the experimental value of K, = 2.8x10-5, Si + 20 = SiO2 (crist.); ?F°1873 = -39,000 cal. [7] The combination of Eqs. 6 and 7 yields the free energy change when liquid silicon dissolves in iron to form the dilute solution of unit activity (1 pct). Si (1) = Si; ?F°1873 = -32,700 cal. [8] The heat effect in this process according to Korber and Oelsen' is an evolution of 28,500 cal per gram

Jan 1, 1954

By G. W. Thomas

A mathematical tnodel of forward combustion in an oil reservoir is treated in this paper. The model describes a radial system having a vertical section of essentially infinite thickness, all of which is permeable to gas flow. Combustion, however, is presumed initiated over a limited thickness of the total vertical section. In the interval supporting cotnbustion, the mechanisms of radial conduction, convection and heat generation are taken into account. Above and below the burning interval, heat transport in the radial direction is by cottduction and convection. Vertical heat losses from the ignited interval are accounted for by conduction alone. A general solution is presented for the temperature distribution caused by radial movement of the combustion front. The results show that no feedback of heat occurs into the ignited interval when convection and conduction are acting in the bounding media. Peak temperatures are also 5 to 10 per cent higher than in the case where heat transport in the bounding media is by conduction alone. We arbitrarily define vertical coverage to be that fraction of the total ignited interval which is at 600F above atnbient, or greater, at any given time. The radial distance at which the vertical coverage becomes zero is the propagation range of the combustion front. It was found that an increase in vertical coverage results when the oxygen concentration, fuel concentration or gas-injection rate is increased. Moreover, the combustion front can be propagated 10 to 15 per cent further than in the case where only conduction is acting above and below the ignited interval. INTRODUCTION In the theoretical treatment of forward combustion in a radial system, one of the problems encountered is the determination of the transient temperature distributions caused by an expanding cylindrical heat source. Bailey and Larkin' and Ramey' simultaneously presented analytical solutions to the problem assuming heat transport by conduction alone. In a subsequent publication, Bailey and Larkin3 included the effects of both conduction and convection while treating linear and radial models. In this latter work, however, vertical heat losses were largely neglected. Selig and Couch' dealt with a radial model in which both conduction and convection were acting. Only a limiting case involving vertical heat losses was considered, however. Namely, temperatures on the boundary of the bed of interest were set equal to zero. Solutions thus obtained were representative of a system having a maximum vertical heat flux. Chu5 recently treated a more general case in which a permeable bed was considered bounded by impermeable media. Conduction and convection took place within the bed, and only conduction outside of the bed. The effects of vertical heat losses were included in his study. Solutions were obtained by numerical techniques. This paper is an extension of the theoretical work of other authors pertaining to forward combustion in a radial system. In particular, a mathematical model of the process is treated in which heat generation occurs over a small vertical interval of a larger permeable section. In the interval supporting heat generation, and above and below this interval, the mechanisms of radial conduction and convection are also presumed acting. Heat losses from the ignited interval are accounted for by vertical conduction. An analytical solution for the temperature distribution caused by radial movement of the burning front is presented. The effects of certain process variables are indicated and comparisons with Chu's results are made. THEORY To render the mechanism of forward combustion tractable to mathematical treatment, we idealize the problem to the extent of assuming continuous reservoir media possessing homogeneous and isotropic properties. The following additional assumptions are implicit in this analysis. 1. The thermal parameters, i.e., heat capacities, thermal conductivities and thermal diffusivities are invariant with temperature and pressure. Moreover, the bounding media possess the same thermal properties as the bed of interest. 2. The temperatures of the porous media and its contained fluids at any point and at any time are equal. 3. The reaction rate between the oxidant gas and the fuel is infinite. This assumption implies that the incoming oxygen concentration instantaneously goes to zero within an infinitesimal distance, i.e., the width of the combustion zone is negligible. 4. The rate of gas injection is constant and corresponds to the average rate throughout the lifetime of the project. 5. The fuel concentration is constant throughout the volume of rock swept out by the burning zone. 6. There is complete burnoff of fuel. This assumption demands that the rate of propagation of the burning front equals the rate of fuel burnoff. In a radial system, with a

By F. Wald, A. J. Rosenberg

The general features of the constitution of the B-Pt system were determined using standard rnetal-lograph~c, thermoanalytic, and X-ray diffraction techniques. Three compound were found. Two of these, Pt3B and Pt,B, are formed by peritectic reactions at 523° and 890°C, respectively. The third, Pt3B,, is congruently melting with a flat maximum at 940°C but decomposes eutectoidally in to Pt,B ant1 boron nt - 600° to 650°C. THE low-temperature allomorph of boron (red, simple rhombohedra1 a boron) is of scientific and technological interest as an elemental semiconductor.' However, the studies of this material have been hampered by its reported instability above 1200"~ which precludes crystal growth from the melt (mp - 2200°C). Crystallization from platinum solutions has been suggested as an alternative crystal-growth technique, but has met with only limited success.' The technique depends upon the existence of a significant difference between the eutectic temperature and the transformation temperature of boron. In order to clarify the conditions for further crystal-growth experiments, we found it desirable to redetermine the main features of the B-Pt phase diagram since previous reports on the system1'5'6'7 are in marked disagreement. EXPERIMENTAL The experimental methods used were thermal analysis, metallography, X-ray analysis, and, to a lesser extent, measurements of microhardness. Most of the alloys were prepared from spectrograph-ically standardized boron obtained from Johnson-Matthey &Co., Ltd. (212 ppm impurities, exclusive of carbon and oxygen) and platinum powder obtained from F. Bishop & Co. (200 ppm impurities, mainly of other platinum group metals). Some alloys were also prepared with very high-purity, float-zone refined boron (99.9999 pct obtained from "Wacker Chemie" and extrahigh-purity platinum (99.999 pct) obtained from Johnson-Matthey & Co., Ltd. The reported results did not depend on the choices of these starting materials. Five-gram alloy specimens containing 10, 20, 25, 27.5, 30, 33.3, 34, 35, 37, 37.5, 38, 39, 40, 41, 42, 43, 45, 50, 55, 60, 70, and 80 at. pct B were made by melting the elements together in boron nitride crucibles using rf heating of a graphite susceptor, either in vacuum or under high-purity argon. All alloys were heated to at least 1800°C for -5 to 15 min. Most of the alloys did not wet the crucibles when the latter were outgassed by preheating under vacuum. In any event, no weight loss was detected after melting, and the nominal composition was assumed for all specimens. Thermal analysis on 2.5-g samples were carried out in boron-nitride crucibles under a vacuum of 5 x X torr. The apparatus was heated in a "Kan-thal A 1" wound furnace, which limited the maximum temperature to about 1100°C. The output of the indicator thermocouple was fed to a dc recorder with a 1-mv full-scale span and an adjustable zero. The apparatus was calibrated repeatedly, using the freezing points of high-purity aluminum, silver, and gold. The results justified the use of the NBS voltage vs temperature tables for Pt/Pt 10 pct Rh thermocouples. All thermal analyses were run at least twice and both the heating and cooling effects were recorded. Most of the alloys had a very strong tendency to supercool. However, the use of mechanical vibration permitted reproducibility within *5°C for all alloys, except in the region around 40 at. pct B. Only the cooling effects are plotted in Fig. 2, since they appear to be more reliable. For metallography, the alloys were cut with a diamond cutting wheel, cast in a polymethacrylate resin, ground and polished with diamond paste, and etched with dilute aqua regia, a common etch for platinum alloys. Both copper and molybdenum radiation were employed to obtain X-ray diffraction data using Debye-Scherrer cameras and a "Norelco" diffractometer Diffractometry with high scanning speeds (1 deg per min) using nickel filtered CuK, radiation was used to identify the main regions of the diagram. However, molybdenum radiation was used for the detection of boron, since the latter showed very strong absorption and fluorescence effects with CuK, radiation. RESULTS AND DISCUSSION Three intermediate compounds, corresponding to the compositions Pt3B, Pt2B, and Pt3B2, were found in the system. Fig. 1 reproduces their X-ray diffraction spectra, together with those of pure boron and pure platinum. As can be seen from the thermal-analysis data in Fig. 2, Pt3B and Pt2B are formed by

Jan 1, 1965

By John L. Miles

John L. Miles (Arthur D. Little, 1nc.)—Pollack and orris" have reported measurements on electron tunneling through A1-A12O3-A1 sandwiches in which the oxide was formed by gaseous oxidation in a glow discharge. From these measurements they deduced the asymmetry of the barrier and, since this is small, conclude that the mechanism suggested by Mott19 for the growth of oxide in thin A12O3 films is inapplicable. In earlier papers20 Pollack and Morris report similar work for oxide films grown thermally. In this case they find a greater asymmetry and conclude that the Mott mechanism is valid. I would like to point out that both these conclusions are quite unjustified. Mott suggests that the growth of the oxide film on aluminum results from the passage of ions through the already present film of oxide under the action of an electric field. This field results from a constant voltage which is in effect a contact potential between metal on one side of the barrier and adsorbed oxygen ions on the other side of the barrier. The theory does not require that the oxide grown is nonuniform either in stoichiometry or structure. It does however specifically assume that the partial layer of ionized oxygen on the surface remains adsorbed on the surface of the growing oxide. In other words, the so-called "built-in field" remains in the oxide only as long as the ionized oxygen is present. When a counter electrode of aluminum is deposited on the oxide, it will react with the adsorbed oxygen on the surface of the oxide, thus forming a small additional amount of oxide. It is clear, then, that there is no requirement in the Mott theory of oxide growth which would necessitate tunneling currents through an Al-A1203-A1 sample to be different when the polarity is reversed. Neither does the theory eliminate the possibility that some additional mechanism could cause the tunneling barrier to be asymmetric and hence tunneling currents to be a function of polarity in such a sandwich. Thus these tunneling-currents measurements are not germane to the question of whether the Mott mechanism is the true method of growth of aluminum oxide films. In fact, it is not surprising that there should be a difference between the oxide properties at the two interfaces (with resulting asymmetry in the tunneling barrier) since the growth conditions and growth rates must have been quite different at these two positions. S. R. Pollack and C. E. Morris (authors' reply)— The point raised by Miles above is one has caused some confusion in the past. The following is an attempt to clarify this point. The built-in field which is responsible for the growth of the thermal oxide at low temperatures arises, according to Mott, because of the passage of electrons from the Fermi surface of the oxidizing metal to surface states introduced by the adsorbed oxygen. It is assumed that the energy of these surface states lies below the Fermi energy of the metal. Electrons therefore continue to flow from the metal to the surface until the built-in electric field raises the potential energy of the surface states to the value of the Fermi energy in the metal, at which time equilibrium is obtained between the surface states and the metal. That is in equilibrium as many excess electrons pass from the metal to the surface per unit time as vice versa. The surface of the oxide prior to deposition of a metallic counterelectrode can then be pictured as follows. The Fermi energy lies in the energy gap of the oxide and is essentially pinned at the energy of the oxygen surface states. The vacuum work function of the oxide is then given by the sum of the electron affinity of the oxide (i.e., the difference in energy between the vacuum and the conduction-band minimum) plus the energy difference between the conduction-band minimum and the Fermi energy. The deposition of a metal onto the surface of the oxide can result in a transfer of electrons across the extremely thin oxide only if there is a contact potential difference between the deposited metal and the parent metal or oxide. That is if the vacuum work function of the deposited metal differs from that of the parent metal, then charge can be redistributed across the oxide in order to equilibriate the Fermi energy across the structure. (It should be

Jan 1, 1965

By J. M. W. Mackenzie, C. J. Wood

In attempting to describe the size classification performance of a hydrocyclone, most workers have elected to use either an equilibrium orbit theory or an non-equilibrium orbit theory. The equilibrium orbit theory has been used by the majority of workers including Lilge,' Bradley; and Yoshioka and Hotta. In applying this theory, it is argued that particles in the body of a hydrocyclone attain an equilibrium radial position where the drag force on the particle resulting from the inward radial fluid velocity is balanced by the outward centrifugal force caused by the tangential component of fluid flow. When considered over the full height of the hydrocy-clone, attainment of this radial equilibrium orbit results in the particle following a conical equilibrium envelope. It is then argued that if this envelope lies outside the envelope of "zero vertical velocity," the particle will report to the underflow, while if the equilibrium envelope lies inside the envelope of "zero vertical velocity," the particle will report to the overflow or vortex finder product. The d50-sized particle which reports in equal quantities to the underflow and overflow is assumed to correspond to particles whose equilibrium envelope is coincident with the envelope of "zero vertical velocity." In considering the equilibrium orbit theory, it is apparent that the horizontal position of the particles in the feed inlet pipe should have no effect on their ultimate destination on the hydrocyclone. Each particle should attain an equilibrium position which depends on the density, size, and shape of the particle; the density and viscosity of the fluid; and the flow patterns within the hydrocy-clone. The nonequilibrium orbit or unsteady state theory has been largely developed by Rietema4 and Mizrahi.6 Mizrahi has listed four main objections to the equilibrium orbit theory. These objections center on the short residence time in the hydrocyclone, the fact that the experimental classification curve is much less sharp than is theoretically predicted, and the absence of negative efficiency conditions in hydro cyclones operating on a feed material which is much finer in size than d50. Proponents of the nonequilibrium orbit theory argue that for a particle to discharge with the underflow it must have sufficient outward radial velocity to reach the downward-flowing region close to the hydrocyclone wall in which the flow lines are parallel to the wall and the ratio of vertical to radial velocity is constant. It is then postulated that a d50 particle entering the cyclone at the center of the feed inlet will just reach this downward-flowing region as it reaches the apex. Thus for uniform distribution of particles across the feed inlet, half the d50 particles—that is, those injected in the half of the inlet area nearest the cyclone wall —will report to the underflow while those injected in the other half will not reach the downward-flowing region and will be carried inward to the center of the cyclone and thus report in the overflow. The exact thickness of the down-ward-flowing region of fluid adjacent to the outer wall of the hydrocyclone is uncertain but Mizrahi considers it to be equal to the apex radius minus the air core radius. Particles larger than d50 have a greater outward centrifugal force acting on them than the d50 particles and may reach the wall even if injected at a distance from the wall greater than Di/2 (Di is inlet diameter). Conversely, particles smaller than d50 may not reach the wall even if injected at a distance less than Di/2 from the cyclone wall. Since the equations put forward by the proponents of both theories yield approximately the same values of d50, it is not possible to decide between these theories by measurement of d50. It should be possible however to examine the theories by injecting a small stream of solids into the feed inlet of a hydrocyclone running on clear water. If the efficiency or classification curve is measured for various horizontal injection positions, then the curves should be coincident if the equilibrium orbit theory holds. If, however, the unsteady state theory describes the cyclone operation, then the classification curves should show finer d50 sizes for particles injected close to the cyclone wall. Experimental A 6-in.-diam hydrocyclone with geometry as in Figs. 1 and 2 was used. Quartz particles were injected as a 50% by wt pulp via an 1/8-in. steel probe. For each in-

Jan 1, 1971

By J. E. McDonald, J. G. Eberhart

A model is discussed from which the work of adhcslon .tor liquid transition metals on aluminum oxide surfaces can he calculated, A close-packed (00011 oxygen surface on A12O3 is assumed with two different types of surface sites: one type involving metal-oxygen bonds and the other man der Waals into actions. The work of adhesion is thus expressed as the sum of these two bonsding free energies. Calculated works of adhesion for nickel, titanium, chromium, and zirconium on sapphire agree well with the experimentally determined quantities. The model is extentled to the calculation of the work of adhesion and shear stress required to remove a thin metal film from a sapphire substrate and is in good agreement with experimental values. The obsewed dependence of the work of adhesion on the free energy of oxide formation of the metal is shown to also provide an interpretation of the tittle dependence of thin-film adhesion. THIS paper presents a model for the type of bonding which occurs across a metal-A12O3 interface. The model is used to explain the results of two types of experiments in which such an interface exists: 1) the adhesion of thin metal films on Al2O3 substrates and 2) the wetting of A12O3 by liquid metal drops. The adhesion of thin films to various substrates has been the subject of a variety of investigations.'-' Benjamin and weaver3 and Bowie,6 using the scratch test developed by Heavens,10 studied the adhesion of metallic films to glass substrates. Their observations for noble-metal film adhesion agree well with an adhesion model involving a van der Waals type of bonding between the film and the substrate. For films of metals whose free energy of oxide formation. ?F°f, has a negative value. Benjamin and weaver3 and Bowie6 found a time-dependent adhesion with an initial value that can be interpreted in terms of van der Waals interactions but a larger terminal value which was related to ?F°f, Karnow-sky and Estill7 deposited films on sapphire at elevated temperatures and noticed no time dependence of film adhesion but a similar correlation with ?F°f. Because of the kinetic problems associated with thin-film adhesion it is desirable to examine adhesion in an equilibrium system. The wetting behavior of liquid-metal drops on Al2O3 provides such a system. Systems of this metal-ceramic type have been studied extensively.11 Humenik and Kingeryl2 have measured the wetting of A12O3 (and other substrates) by several metals and have pointed out that the wetting ability of these metals increases with increasing values of -?F°f. It is thus seen that thin-film adhesion and metal wetting on A12O3 are both related to the tendency of the metal to react with the surface oxide ions of the Al2O3 substrate and, because of this, both phenomena should be explainable by an appropriate model for the metal-Al2O3 interfacial bonding. In the sections that follow, wetting and adhesion data on A2O3 are reviewed and a model is presented by which these phenomena can be interpreted. ANALYSIS OF WETTING EXPERIMENTS In an equilibrium system involving a liquid-metal drop on a solid Al2O3 substrate, the work of adhesion, WAD, is defined by the Dupre, equation as WAD = ?s + ?L -?sL [1] where ?s and ?L are the surface free energies of the solid substrate and the liquid drop, respectively, and ?sL is the interfacial free energy. The work of adhesion is the work required to separate a unit area of the solid-liquid interface into two surfaces. The work of adhesion is usually determined from a sessile-drop experiment in which yL and the contact angle, ?, are measured. The Young-Dupr6 equation is then used to calculate WAD Wad = ?L (1+ cos ?) [2] The literature of this subject has been examined and Table I shows work of adhesion data for various liquid metals as measured on A12O3 substrates. The standard free energy of oxide formation of the metal at the temperature of the wetting experiment, ?F°f . is also tabulated in kcal per g-atom of oxygen. The data is grouped according to the gaseous atmos-

Jan 1, 1965

By F. R. Jones

AT Tahawus, N. Y., National Lead Co. operates the MacIntyre development. Here the world's largest titanium mine produces 5200 long tons of ore per day and pours 8000 long tons of waste rock over its dumps. Concentrated ilmenite is sent by rail to National Lead Co. pigment plants, and a second product, magnetite, is sold to steel producers in raw form or is agglomerated and shipped as sinter. Several earlier attempts had been made to produce iron from the deposits, which have been known since 1826. These attempts failed, chiefly because of titanium impurity. In 1941 the present owners reestablished the operation for production of war-scarce ilmenite, and the impurity became the main product. The Ore: The MacIntyre ore zone is about 2400 ft long and 800 ft wide in horizontal measurements. Ore outcrops were found on the northwest side of Sanford Hill, 450 ft above Sanford Lake and 2500 ft southeast. The zone dips at about 45" toward the lake and plunges to the southwest. The ore minerals, ilmenite and magnetite, are unevenly distributed in bands roughly parallel to the long axis of the ore zone and are interspersed with bands and horses of waste. Hanging wall ores are fine grained and grade from rich ore to waste rock or gabbro. Footwall ores are coarse grained and are almost entirely ilmenite and magnetite. The foot-wall waste rock, anorthosite, is the common country rock. Several faults cut the ore zone. These faults have no great displacement but do contribute to the great physical variations in ore rock and surrounding waste. The Mine: The MacIntyre mine is an open pit operation, with benches at 35-ft intervals. The lowest bench is now 54 ft below lake level. Loading equipment consists of three electric-powered shovels (a P & H model 1400 with 4-yd dipper and two Bucyrus-Erie models 85-B with 2%-yd dippers) and one diesel-powered shovel (a Northwest model 80D with 2%-yd dipper). Ore and waste are transported to a 48x60-in. jaw crusher in ten 22-ton Euclid trucks with 300-hp diesel engines. Ordinarily the two Bucyrus-Erie 2 % -yd shovels load ore into a fleet of three or four trucks. This combination works two 8-hr shifts per day, moving 5200 long tons of ore to the crusher and removing a small portion of the waste rock. The P & H model 1400 shovel, with a fleet of four trucks, loads waste on three shifts per day. The mine operates on a 5-day week, with a small maintenance crew working Saturday. Oversize rock is broken by a dropball handled by an Osgood model 825 rubber-mounted crane.' Ore and waste are broken by drilling and blasting 9-in. diam vertical holes behind the benches. Bucyrus-Erie 42-T churn drills are used to drill the holes, which are extended 4 ft below the bench level on which the broken rock will fall. Drilling and Blasting History: In its early years the mine was equipped with Bucyrus-Erie 29-T churn drills, which drilled 6-in. holes. To keep up with production requirements the hole diameter was soon increased to 9 in., and by 1950 the three 42-T drills now in use had been acquired. Early blasting experiments with different kinds and grades of explosive led to adoption of 90 pct straight gelatin dynamite as standard. It was recognized that this explosive was expensive, and from the start of operations until 1950 extensive experiments were made using blasting agents of the ammonium nitrate family. Results were recorded as uniformly poor, with great build-up of oversize rock. The expense of these experiments, and the discouraging results, caused the abandonment of any expectation of breaking MacIntyre rock with anything but 90 pct straight gelatin dynamite. Further standardization led to 9-in. well drillhole spacings set at 16 ft in ore and 18 ft in waste, exceptions being permitted only for unusual conditions. The hole burdens were theoretically about 22 ft. Due to the extreme back-slope of bench faces, caused by blasting with heavy charges of dynamite, actual burdens were commonly well over 30 ft. Lack of precise control resulted in many holes having a burden as light as 15 ft. General practice was to stem 6 or 7 ft of hole with magnetite concentrate, the amount of stemming being left to the discretion of the pit foreman. Usually all holes in a row were fired instantaneously with Primacord detonating fuse. Millisecond delays were

Jan 1, 1957