CN1491740A - Device and method for heat synthesis - Google Patents
Device and method for heat synthesis Download PDFInfo
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- CN1491740A CN1491740A CNA021469369A CN02146936A CN1491740A CN 1491740 A CN1491740 A CN 1491740A CN A021469369 A CNA021469369 A CN A021469369A CN 02146936 A CN02146936 A CN 02146936A CN 1491740 A CN1491740 A CN 1491740A
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Abstract
The apparatus for converting one or several reactants into required ultimate product via heat conversion includes insulated reactor with high temperature heater, such as plasma torch, in the inlet end, and limited convergent-divergent nozzle in the outlet end. During the heat conversion process, reactants are injected before the reactor, so that the reactants are made to mix with plasma flow thoroughly. The reactor has reaction region with temperature maintained constant radically. Hot gas flow is cooled quickly via the nozzle to 'freeze' ultimate product(s) in heat balanced reaction stage. Or, the hot gas flow may be drained via outlet pipe without passing through a convergent-divergent nozzle. Then, the required ultimate product is separated from the gas flow.
Description
Technical Field
The present invention relates generally to a thermal synthesis process. The present invention relates specifically to a method and apparatus for thermally converting reactants in a thermodynamically stable high temperature gas stream to a desired end product, such as a gas or ultra-fine solid particles.
Background
In the united states, natural gas (wheremethane is the predominant hydrocarbon) is low in price and an underutilized energy source. It is known that large reservoirs of natural gas exist in remote areas of the continental united states, but that this energy source cannot be economically and safely transported from these areas. In recent decades, the conversion of natural gas into more valuable hydrocarbons has been studied, but such research has met with very limited success in modern economies. In recent years, some studies have been conducted to evaluate the technology of converting natural gas (now used for direct combustion) into acetylene as a raw material for commodity chemicals. The utilization of existing large natural gas reserves in connection with oil fields and inexpensive labor has led to the conversion of natural gas to acetylene for the manufacture of commodity chemicals that are particularly attractive in this part of the world.
Acetylene can be used as a feedstock for plastics manufacture or conversion to liquid hydrocarbon fuels by proven catalytic reactions. The versatility of acetylene as a starting material is well known and recognized. The current raw materials for plastics are petrochemical-based. The declining supply of petroleum reserves from both domestic and foreign sources to produce these petrochemical-based feedstocks has placed pressure on finding alternatives to petrochemical-based feedstocks. Thus, interest in acetylene-based feedstocks is now regained again.
The thermal conversion of methane to liquid hydrocarbons includes indirect or direct processes. The conventional Methanol To Gasoline (MTG) process and the Fischer-tropsch (ft) process are two major examples of such indirect conversion processes, which reform (reform) methane to synthesis gas, which is then converted to final product. These expensive endothermic processes are operated at high temperatures and pressures.
A direct catalytic conversion of methane to light olefins (e.g. C) is sought2H4) The process of conversion to liquid hydrocarbons has then become the focus of current natural gas conversion technology. Oxidative coupling, oxychlorination (oxychlorination) and partial oxidation are examples of direct conversion processes. These techniques require operation at higher pressures, moderate temperatures and the use of catalysts. The development of specific catalysts for the direct conversion of natural gas is the greatest challenge facing the development of these technologies. The method has low conversion rate and is indirectThe cost of the process is high and these techniques have not been proven.
Light olefins can be made by extracting hydrogen from methane at very high temperatures (>1800 ℃) and then coupling the hydrocarbon radicals. The high temperature conversion of methane to acetylene can be achieved by the reaction scheme This is shown as an example. This process is known for a long time.
The conversion of methane to acetylene is currently prevented by the use of cold liquid hydrocarbon quenchers. Perhaps the best known of these processes is the hull process, which has been used in commercial germany applications for many years. The Huels arc reactor is capable of transferring electrical energy by "direct" contact between a high temperature arc (15000-. The product gas was quenched with water and liquefied propane to prevent back reaction. For the Huels process, the single pass yield of acetylene is less than 40%. By recycling all hydrocarbons other than acetylene and ethylene, C2H2The total yield of (A) can be increased to 58%.
Although used commercially, the Huels process is only marginally cost effective economically due to the low single pass efficiency and the need to separate the product gas from the quench gas. The german government's allowance has helped to use this method in production.
A similar process using a 9MW reactor was established by dupont operating in 1963 to 1968 and which sent acetylene produced from a liquid hydrocarbon source to a neoprene mill. It has also been reported that this process was carried out on a pilot plant scale using a methane feedstock. Whereas plant-scale operations are limited to liquefied petroleum gas or liquid hydrocarbon distillates. The scale of the DuPont pilot plant method is not reported. In the dupont method, the arc is magnetically rotated, whereas in the original hues method, the arc is "swirled" with a tangentially injected gas. In the dupont process, all of the feedstock diluted with hydrogen is passed through an arc column. In the hull process, a portion of the reactants is injected at a location downstream of the arc.
Westinghouse uses a hydrogen plasma reactor to crack natural gas to produce acetylene. In a plasma reactor, hydrogen is fed into the arc zone and heated to a plasma state. Blazing H at temperatures above 5000K2The plasma exit stream is rapidly mixed with natural gas below the arc zone and electrical energy is indirectly transferred to the feedstock. The hot product gas is quenched with liquefied propane and water as in the hull process to prevent back reactions. However, the Huels methodLikewise, it is desirable to separate the product gas from the quench gas. It has been reported that recycling all hydrocarbons other than acetylene and ethylene can increase the overall yield to 67%. H for natural gas conversion2The plasma process has been extensively tested on a small scale, but further development and validation has been required on a pilot scale.
The Norwegian Scientific and Industrial research Foundation of Norway developed a reactor consisting of two concentric, heat-resistant graphite tubes. At temperatures of 1900-. During operation, carbon formation in the annulus can lead to significant operational problems. Liquefied quenchers are also used to quench the reaction products to prevent back reactions. As with the previous two acetylene production processes described above, it is necessary to separate the product gas from the quench gas. The overall yield of acetylene from this refractory reactor was about 80%, which was tested in pilot plants.
Thus, there is a need for improvements in the moderate methane conversion efficiencies, acetylene yields, selectivities, and energy consumption rates observed in those processes described above.
The high corrosion resistance and strength properties of titanium combined with its relatively low density make titanium alloys ideally suited for many high tech applications, particularly in aerospace systems. The use of titanium in chemical plants and power plants is also attractive.
Unfortunately, the widespread use of titanium is severely limited by its high cost. This high cost is a direct result of the batch nature of the conventional Kroll and Hunter processes used to manufacture titanium metal and the high energy consumption rates required for its application.
The large scale manufacturing processes used in the titanium industry have been relatively constant for many years. They comprise the following main steps: (1) chlorination of impure oxide ores, (2) purification of TiCl4Reducing with sodium or magnesium to obtain titanium sponge, (4) removing the titanium sponge, (5) leaching, distilling and vacuum remelting to remove Cl, Na and Mg impurities. The inherent cost of these processesThe resultant effects, difficulties associated with forging and machining titanium, and recent shortages in titanium sponge utilization have led to the use of less titanium.
One of the most promising methods currently being developed to prevent the high cost of titanium alloy components is the powder metallurgy process that produces shapes that approach the final component shape. For example, for each kilogram of titanium currently used in aircraft, it is estimated that 8 kilograms of scrap will be produced. Powder metallurgy can greatly reduce this waste rate. Although this technique mainly involves simple steps of manufacturing the powder and then compacting it into a compact article, considerable research is currently being carried out to optimize this process, so that the final product has properties at least equivalent to those of forged or cast materials and lower in cost than the latter.
One possible powder metallurgical route to titanium alloy parts is to mix elemental metal powders directly and then compact. Currently, titanium sponge fines from the Kroll process are used, but one major disadvantage is their high residual impurity content (mainly chlorides), which can form pores in the final material. Alternative powder metallurgy methods include hot isostatic pressing using titanium alloy powders directly.
Several solutions are currently involved in optimizing such titanium alloy powders. The results are very promising, but all solutions involve the use of Kroll titanium as starting material. The use of such existing powders requires a number of expensive purification and alloying steps as described above.
In the literature of the last 30 years, methods for producing titanium under plasma conditions have been of continuous interest. Reports generally relate to the reduction of titanium tetrachloride or titanium dioxide with hydrogen, and some individual references relate to the reduction with sodium or magnesium.
The reduction of titanium tetrachloride using hydrogen was investigated in an electric arc furnace. At 2100K only partial reduction was performed. The same reaction system in plasma flame has been studied more extensively and patented on the preparation of titanium subchlorides (German patent 1,142,159, 1 month 10 days 1963) and titanium metals (Japanese patent 6854, 5 months 23 days 1963; 7408, 10 months 15 days 1955; U.S. Pat. No.3,123,464, 3 months 3 days 1964).
Although early thermodynamic calculations indicated that the reduction of titanium tetrachloride to metallic titanium with hydrogen should begin at 2500K, this system is not a simple system. Calculations show that the formation of titanium subchlorides in this temperature range is thermodynamically more favorable.
U.S. Pat. No.3,123,464 discloses heating the reactants (TiCl) at temperatures to a low temperature, preferably above the boiling point (3535K) of titanium4And H2) The titanium tetrachloride can be successfully reduced to liquid titanium. At such high temperatures, it is disclosed that while titanium tetrachloride vapors are effectively reduced by atomic hydrogen, H2The tendency to dissolve or react with Ti is not significant, the HCl formed dissociates only about 10%, and the formation of titanium subchlorides is not favored. The titanium vapor product is then either condensed to a liquid in a water-cooled steel condenser at about 3000K, from which it overflows into the mold, or quenched with hydrogen to a powder, which is collected in a bin. Since the liquid titanium condensed from the gas is only accompanied by gaseous by-products or impurities, its purity (in addition to hydrogen impurities) is expected to be high.
Japanese patent 7408 describes the following reaction conditions: making TiCl4Gas and H2(50% excess) of the mixture at 4X 10-3m3The flow rate/min was discharged (3720V and 533mA) through a tungsten electrode nozzle with an internal diameter of 5mm onto another electrode at a distance of 15 mm.The resulting powdered crystals were heated in vacuo to yield 99.4% pure titanium.
None of the above patents clearly mentions energy consumption. Attempts have been made to develop industrial scale hydrogen reduction processes using autogenous crucible furnaces, but this effort has been discontinued afterwards. Recently, it has been claimed that small amounts of titanium are produced in the hydrogen plasma, but this statement is subsequently withdrawn when it is confirmed that the product is indeed titanium carbide.
In summary, attempts have been made to treat TiCl in a hydrogen plasma4It seems that only partial reduction to titanium and its subchlorides is possible unless very high temperatures are reached(>4000K). Previous researchers concluded that it was necessary to first condense vapor phase titanium to overcome the thermodynamic instability of the system.
Accordingly, there is a need for methods and apparatus that overcome or avoid the above-described problems and limitations.
Disclosure of Invention
It is an object of the present invention to provide a method and apparatus for thermally converting one or more reactants in a thermodynamically stable, high temperature gas stream into a desired end product.
It is a further object of the present invention to provide an improved process conversion efficiency and product yield for thermally converting one or more reactants in a thermodynamically stable high temperature gas stream to a desired end product.
It is a further object of the present invention to provide a process and apparatus for increasing the yield of thermal conversion of natural gas to acetylene.
To achieve these objects, and in accordance with the invention as embodied and broadly described herein, there is provided a method and apparatus for thermally converting one or more reactants in a thermodynamically stable, high temperature gas stream into a desired end product in the form of a gas or ultra-fine solid particles. Generally, the method comprises the following steps. First, a reactant stream is introduced at one end of an axial reactor. The reactant stream is then heated as it flows axially through the injection channel of reduced diameter relative to the axial reactor to create turbulence, thereby thoroughly mixing the reactant stream with the hot gases. The thoroughly mixed reactant stream is then passed axially through a reaction zone of an axial reactor, which reaction zone is maintained at a substantially uniform temperature over its length. The axial reactor has a length and temperature and is operated under conditions sufficient to heat the reactant stream to a selected reaction temperature at which a desired product stream is produced at a location proximate to the outlet end of the axial reactor.
Specifically, the method of the present invention comprises the following steps. First, a plasma arc gas flow is introduced between plasma torch electrodes comprising at least one pair of electrodes located near the inlet end of the axial reactor chamber. The plasma arc gas stream is introduced at a selected flow rate while the electrode receives a selected plasma input power level to generate a plasma in a diametrically confined injection passage extending into the reactor chamber and toward the outlet end of the reactor chamber. Next, at least one reactant is injected into the injection channel, the introduced reactant stream is thoroughly mixed into the plasma, thorough mixing is achieved, and then introduced into the reactor chamber. The reactor chamber is maintained at a substantially uniform temperature throughout the flow field to allow the reaction to reach equilibrium. The gas stream exitingthe nozzle is cooled at the outlet end of the reactor chamber by reducing its velocity while removing heat at a rate sufficient to prevent its motive temperature from increasing. The desired end product is then separated from the gas remaining in the cooled gas stream.
The present invention provides an increase in process conversion efficiency and acetylene yield over existing conventional processes. This improvement is achieved primarily by more efficient injection and mixing of the reactants with the plasma gas and minimizing temperature gradients and cold boundary layers in the reactor. Improved mixing and thermal control also increases this feature by reducing the yield of hydrocarbons other than acetylene. The quench rate achieved by heat transfer on the walls in a small reactor is sufficient to inhibit acetylene decomposition and formation of the soot. The formation of hydrocarbons other than ethylene is not affected by significantly increasing the quench rate.
These and other objects, features and advantages of the present invention will be more fully described by the following description and appended claims, or may be learned by the practice of the invention as set forth hereinafter.
Drawings
In order to describe the manner in which the above-recited and other advantages and objects of the invention are obtained, a more particular description of the invention briefly described above will be rendered by reference to specific embodiments thereof which are illustrated in the appended drawings. It is appreciated that these drawings depict only typical embodiments of the invention and are therefore not to be considered limiting of its scope. The invention will be described with additional specificity and detail through the use of the accompanying drawings.
FIG. 1 is a schematic cross-sectional view of a reactor system according to one embodiment of the present invention;
FIG. 2 is a schematic cross-sectional view of a reactor system according to another embodiment of the present invention;
FIG. 3 is a schematic cross-sectional view of a reactor system according to yet another embodiment of the present invention;
FIG. 4 is a graph depicting the theoretical maximum amount of methane that can be converted in the reactor system of the present invention;
FIG. 5 is a graph depicting conversion efficiency versus methane feed flow;
FIG. 6 is a graph depicting estimated reactor temperature and residence time in the reactor versus methane injection flow;
FIG. 7 is a graph depicting methane conversion efficiency versus reactor pressure;
FIG. 8 is a graph depicting acetylene yield versus methane injection flow;
FIG. 9 is a graph depicting acetylene yield versus methane injection flow;
FIG. 10 is a graph depicting hydrocarbon yield versus methane injection flow;
FIG. 11 is a graph depicting hydrocarbon (no methane) yield versus methane injection flow;
FIG. 12 is a graph depicting conversion efficiency versus methane feed flow;
FIG. 13 is a graph depicting acetylene yield versus methane feed flow;
FIG. 14 is a graph depicting yield efficiency versus methane feed flow;
FIG. 15 is a graph depicting hydrocarbon (no methane) yield versus methane feed flow;
FIG. 16 is a graph depicting acetylene yield versus pressure;
FIG. 17 is a graph depicting hydrocarbon yield versus reactor pressure;
FIG. 18 is a graph depicting the acetylene and energy consumption rates produced versus methane feed flow.
Detailed Description
The present invention relates generally to a method and apparatus for thermally converting reactants into a desired end product, such as a gas or ultra-fine solid particles. The apparatus of the invention generally comprises a multi-channel injector for injecting reactants into a chamber in front of the reactor, thereby mixing the reactants with the plasma stream. The reactor is insulated using, for example, a carbon lining, which provides residence time for the reaction to proceed while minimizing radial temperature gradients. The reactor may also be equipped with a converging-diverging nozzle at the outlet end, which produces ultrasonic expansion and thus greatly increases the quench rate.
There are various reactors and processes for high temperature reactions in which rapid cooling is again required to freeze the reaction products to prevent back reactions or decomposition into unwanted products. For example, U.S. patent No.5,749,937 to Detering (hereinafter "Detering") uses adiabatic and isentropic expansion of a gas in a converging-diverging nozzle for rapid quenching. This expansion results in a cooling rate of over 1010K/s, which preserves the reaction product that is only equilibrated at high temperatures.
However, there is still a continuing need for values that improve conversion efficiency and yield. It was found that temperature gradients and poor mixing lead to an uneven distribution of temperature in the reactor. The composition of the product stream varies with the kinetics or rate of the reaction and also with temperature non-uniformity. These effects can result in significant changes in the composition of the product stream. If the quenching process is too slow or delayed, the product produced, such as acetylene, will thermally decompose to solid carbon soot or may further react to form benzene and heavier hydrocarbons primarily on the hot surface.
The present invention addresses these issues. The problem of temperature non-uniformity can be overcome by using an insulation-lined "hot-wall" reactor configuration, such as a carbon-lined reactor, which minimizes radial temperature gradients and heat loss from portions of the reactor. The problem of poor mixing can be solved by using a restricted passage injector design that provides good mixing of the reactants into the plasma stream. The effect of the quench rate can be addressed by placing the ultrasonic quench nozzle in the apparatus just behind the reactor section. In the ultrasonic nozzle, the hot gas from the reactor section is rapidly expanded to a lower pressure. In this way, the thermal energy is converted into kinetic energy, so that the mixture is rapidly cooled. This process is sometimes referred to as pneumatic quenching. A quench rate of 107-108 c/s can be achieved which is one to two orders of magnitude greater than the quench rate reported in the original hues process. Without this converging-diverging nozzle, the quench rate achieved by wall heat transfer is only about 0.1X 106 deg.C/s.
Although this reactor concept was originally developed in the study of acetylene formation from natural gas, those skilled in the art will appreciate that the method and apparatus of the present invention may be used in other processes that require rapid quenching, including titanium production.
The rapid quench reactor and method of operation described herein utilizes the high temperatures (5,000-20,000 ℃) produced by high temperature heating devices, such as thermal plasma devices, toproduce materials that are thermodynamically stable at such high temperatures. These materials include metals, alloys, intermetallics, composites, gases, and ceramics.
The rapid quench reactor and method of the present invention will be described and illustrated with respect to a rapid heating device comprising a plasma torch and a plasma arc gas flow. However, it should be understood that the rapid heating means may also include other rapid heating means such as a laser and a flame produced by oxidation of a suitable fuel such as an oxygen/hydrogen flame.
It should be understood that various features described below in the various embodiments may be interchanged to provide further embodiments encompassed by the present invention. For example, various embodiments may or may not include a converging-diverging nozzle, a mixing chamber, a downstream injector, or an anode injector.
Referring now to the drawings, FIG. 1 is a schematic diagram of an overspeed quench apparatus 10. The apparatus 10 generally includes a flare portion 12, an injector portion 14, a closed axial reactor chamber 16, a converging-diverging nozzle 18, and a cooling portion 20.
In one embodiment of the present invention, apparatus 10 is a rapid quench axial reactor for thermally converting one or more reactants in a thermodynamically stable high temperature gas stream into a desired end product in the form of a gas or ultrafine solid particles. The apparatus 10 includes means for introducing a reactant stream at or in front of the inlet end of the axial reactor, such as a multi-channel injector or an anode injector as will be described later, located in the injector section 14 or flare section 12. The apparatus 10 further comprises heating means for generating a hot gas stream in front of the inlet end of the axial reactor, which hot gas stream flows axially towards the outlet end of the axial reactor. Such heating means may be selected from a torch for generating a plasma, a laser, a flame produced by oxidation of a suitable fuel such as an oxygen/hydrogen flame and equivalents thereof. In addition, the apparatus 10 also includes means for causing turbulence in the reactant stream and the hot gas stream through an injection passage of reduced diameter, such as injection passage 23 in the injector portion 14 or the flare portion 12 or their equivalents, to thoroughly mix the reactant stream with the hot gas stream. The apparatus 10 also includes means for minimizing radial temperature gradients in the axial reactor, such as an insulating layer or equivalent on the interior of the reactor chamber 16. The axial reactor is preferably operated under conditions sufficient to heat the reactant stream to a selected reaction temperature that produces the desired end product at a location near the outlet end of the axial reactor. The various components of the device 10 will be described in further detail below.
The torch portion 12 includes a plasma torch for thermally decomposing the incoming gas stream in the resulting plasma as it is fed through the inlet of the reactor chamber.
The plasma is a high temperature luminescent gas, which is at least partially (1-100%) ionized. The plasma is composed of gas atoms, gas ions, and electrons. The plasma is electrically neutral throughout the phase. A hot plasma can be obtained by passing a gas through an electric arc generated between two electrodes (anode and cathode). The arc rapidly heats the gas to very high temperatures by resistive and radiant heat during the several microseconds of the gas passing through the arc. The plasma typically emits light at temperatures above 9000K.
The plasma can be generated in this manner with any gas. Since the gas can be neutral (argon, helium, neon), reducing (hydrogen, methane, ammonia, carbon monoxide) or oxidizing (oxygen, nitrogen, carbon dioxide), it is possible to have a good control of the chemical reactions in the plasma. Oxygen or oxygen/argon mixtures can be used to make metal oxide ceramics and composites. Other nitride, boride and carbide ceramic materials require gases such as nitrogen, ammonia, hydrogen, methane or carbon monoxide to obtain the proper chemical environment required for the synthesis of these materials.
The details of torches used to generate plasma are well known and will be understood by those skilled in the art from the disclosure herein and need not be described in further detail.
The incoming plasma gas flow is indicated by arrows 31. The plasma gas may also be a reactant or inert. The one or more reactant gas flows (arrows 30) are typically separately injected into the plasma, which flows toward the outlet of the reactor chamber 16. The gas flow axially through the reactor chamber 16 comprises reactants injected into a plasma arc or in a carrier gas.
Gases and liquids are preferred forms of injecting reactants. A solid can be injected but it typically evaporates too slowly to allow chemical reactions to occur in the fast flowing plasma before the gas cools. If a solid is used as a reactant, it is typically heated to become gaseous or liquid before injection into the plasma.
In free-flowing plasma, typical residence times of the material are on the order of milliseconds. To maximize mixing of the reactants with the plasma gas, the reactants (liquid or gas) are injected through small holes under pressure (10-100 atmospheres) to achieve a velocity sufficient to penetrate and mix with the plasma. It is preferred to use gaseous or vaporized reactants whenever possible, since this eliminates the need for phase changes in the plasma, thereby improving the dynamic performance of the reactor. In addition, the injected reactant stream is preferably injected at right angles (90 °) to the flow of plasma gas. However, in some cases, a positive or negative deviation from this 90 ° angle of up to 30 ° may be optimal.
The high temperature of the plasma causes the injected liquid material to rapidly vaporize and cause its gaseous molecular species to fragment into its atomic constituents. Various metals (titanium, vanadium, antimony, silicon, aluminium, uranium, tungsten), metal alloys (titanium/vanadium, titanium/aluminium/vanadium), intermetallic compounds (nickel aluminide), titanium aluminide) and ceramics (metal oxides, nitrides, borides and carbides) may be synthesized by injecting metal halides (chlorides, bromides, iodides and fluorides) in liquid or gaseous form into a plasma of a suitable gas from behind the anode arc connection in the torch outlet or along the length of the reactor chamber.
The closed axial reactor chamber 16 is connected at one inlet end to the injector portion 14 and at one outlet end to a nozzle 18. The reactor chamber 16 of one embodiment of the present invention has an insulating liner 34 on its interior surface. The insulated reactor provides residence time for the reaction to proceed while minimizing radial temperature gradients. A cooling water jacket (not shown) is typically provided outside of the reactor chamber 16.
The reactor chamber 16 is where the chemical reaction takes place. The reactor chamber 16 begins at the inlet behind the plasma arc and ends at the nozzle throat 26. The reactor chamber 16 includes a converging portion 24 that forms the main reactor zone for the product and a portion of the nozzle 18.
The temperature requirements in the reactor chamber and its geometry depend on the temperature required to reach an equilibrium state for each desired end product content.
Since the reaction chamber is a hot and chemically active area, the material of the reactor chamber must be made compatible with temperature and chemical activity to minimize chemical corrosion from the reactants and to minimize melting degradation and ablation due to the resulting intense plasma radiation. The reactor chamber is typically made of water-cooled stainless steel, nickel, titanium or other suitable material. The reactor chamber may also be made of a ceramic material to withstand the intense chemical and thermal environment.
As described above, the walls of the reactor chamber are lined with an insulator, such as carbon, to maintain a constant temperature gradient in the reactor chamber 16. The purpose of the insulation is to provide a barrier to heat transfer to reduce the heat loss of the process to the cold water jacket outside the reactor chamber. Various insulating materials may be used as long as the material selected does not react with the reactants in the reactor chamber and has a sufficiently low coefficient of expansion to prevent swelling and bursting of the outer wall of the reactor chamber. Thus, preferred materials such as carbon are good insulators, are chemically inert to the process reactants, have a low coefficient of expansion, and withstand high temperatures, for example, of about 2000-3000K. Other suitable materials include, for example, boron nitride, zirconia, silicon carbide, and the like. However, any insulating material that reduces the heat transfer from the reactor chamber 16 to the outer wall may be used, provided that the above criteria are met.
The walls ofthe reaction chamber are heated from the inside by a combination of radiation, convection and conduction. Cooling the walls of the reaction chamber prevents undesirable melting and/or corrosion on the surfaces thereof. The system for controlling this cooling should maintain the walls at a high temperature allowed by the selected wall material, which must be inert to the reactants in the reactor chamber 16 at the desired wall temperature. The same applies to nozzle walls that are heated only by convection and conduction.
The dimensions of the reactor chamber 16 are selected to minimize recirculation of the plasma and reactant gases to maintain sufficient heat (enthalpy) into the nozzle throat to prevent degradation (unwanted back or side reactions).
The length of the reactor chamber 16 must first be determined experimentally using an elongated tube that allows the user to reach the target reaction threshold temperature. The reactor chamber 16 can then be designed to be of sufficient length that the reactants have sufficient residence time at the high reaction temperature to reach equilibrium and complete formation of the desired end product. Such reaction temperatures may range from a minimum of about 1700 ℃ to 4000 ℃.
The inner diameter of the reactor chamber 16 is determined by the fluid properties of the plasma and the moving gas stream. The internal diameter of the reactor 16 must be large enough to allow the desired gas flow therethrough, but not so large that undesirable back-flow eddies or stagnation zones form in the walls of the reactor. This detrimental flow pattern can prematurely cool the gas and precipitate unwanted products such as low chlorides or carbon. Typically, the inner diameter of the reactor chamber 16 should be in the range of about 100-150% of the plasma diameter at the inlet end of the reactor chamber 16.
A converging-divergingnozzle 18 is located coaxially behind the reactor chamber 16. The effect of the converging-diverging nozzle is to cause a rapid drop in motive temperature in the flowing gas stream. This effectively "freezes" or stops all chemical reactions. When the gas is rapidly cooled, the desired end product can be efficiently collected without reaching equilibrium conditions. The final product, which is produced in the plasma at high temperatures but is thermodynamically unstable or unavailable at lower temperatures, can then be collected due to the resulting phase transition (gas-solid) or stabilization by cooling to a lower equilibrium state (gas-gas).
The converging or upstream portion of the nozzle 18 restricts the passage of gases and controls the residence time of the hot gas stream in the reactor chamber 16 to bring the contents of the reactor into thermodynamic equilibrium. The constriction that occurs in the cross-sectional dimension of the gas stream, as it passes through the converging portion of the nozzle 18, changes the movement of the gas molecules in random directions (including rotational and vibrational movements) to a linear movement parallel to the axis of the reactor chamber. The dimensions of the reactor chamber 16 and the flow rate of the gas introduced are selected to achieve sonic velocities within a limited nozzle throat.
As the restricted gas flow enters the diverging or downstream portion of the nozzle 18, its pressure drops at an excessive rate due to the increasing volume along the conical wall of the nozzle outlet. The resulting pressure change immediately lowers the temperature of the gas stream to the new equilibrium condition.
By appropriately selecting the dimensions of the nozzle, the reactor chamber 16 can be operated at atmospheric pressure or under pressurized conditions, while the cooling section 20, which is located behind the nozzle 18, is kept under vacuum pressure by the action ofa pump. The sudden change in pressure as the gas stream passes through the nozzle 18 immediately places the gas stream in a lower equilibrium condition and prevents unwanted back reactions that may occur under prolonged cooling conditions.
The converging portion 24 of the nozzle 18, the purpose of which is to rapidly compress the hot gases into a restricted nozzle throat 26, minimizes heat loss through the wall while maintaining laminar flow and minimal turbulence. This requires that the aspect ratio of the nozzle diameter vary widely to maintain a smooth transition to the first steep angle (>45 °) and then to a smaller angle (<45 °) into the nozzle throat.
The purpose of the nozzle throat 26 is to compress the gas to achieve sonic velocity in the flowing hot gas stream. This process converts the random energy of the hot gas into translational energy (velocity) in the axial direction of the gas stream. This effectively lowers the kinetic temperature of the gas and further limits the chemical reaction almost immediately. The pressure differential between the reactor chamber 16 and downstream of the diverging section 28 of the nozzle 18 controls the velocity achieved in the throat and downstream diverging section of the nozzle. Negative pressure may be applied downstream or positive pressure may be applied upstream for this purpose.
The purpose of the diverging portion 28 of the nozzle 18 is to smoothly expand and accelerate the gas exiting the nozzle from sonic to supersonic speeds, which can further reduce the motive temperature of the gas.
In practice the term "smooth acceleration" requires the use of a small divergence angle of less than 35 ° for expanding the gas without the detrimental consequences of separation from the converging walls and induced turbulence. The separation of the expanding gas from the diverging wall causes some backflowof gas between the wall and the gas jet exiting the throat of the nozzle. This backflow, in turn, leads to local reheating of the expanding gas and to poor degradation reactions, which reduce the yield of the desired end product.
The overspeed quenching phenomenon observed in the embodiment of the present invention comprising a converging-diverging nozzle in the plant is achieved by rapidly converting the thermal energy of the gas into kinetic energy through improved adiabatic and isentropic expansion of the converging-diverging nozzle. For a detailed description of the physical function of the nozzle, see Detering.
Another reactant at ambient temperature, such as hydrogen, may be injected tangentially into the diverging portion of the nozzle 18 to complete the reaction or to prevent a reverse reaction as the gas cools.
A cooling section 20 is located coaxially behind the nozzle 18 and serves to further cool the gas stream and quench the reaction. The walls of the reactor chamber 16, the nozzle 18 and the coaxial cooling section 20 are cooled with a flow of cooling water.
The end product of the reaction may be collected in a cyclone (not shown). A downstream liquid trap, such as a liquid nitrogen trap, may be used to condense and collect reaction products such as hydrogen chloride and ultra-fine powders in the gas stream before the gas stream enters the vacuum pump.
Referring now to fig. 2, an apparatus 50 according to another embodiment of the invention generally has a torch/injection portion 52, an injector portion 54, an insulated reactor chamber 56, and a cooling portion 58. It should be understood that most of the components are as described above for the device 10.
The various improvements described above were found to render the nozzle 18 unnecessary for obtaining the positive results of this embodiment ofthe invention. Thus, the nozzle assembly is absent and instead a cooling straight tube portion of the same internal diameter as the following tube is provided. Although not shown, the present invention may also be used with a converging portion similar to the converging portion 24 of the converging-diverging nozzle 18 in place of the nozzle assembly 18, when it is desired to avoid clogging with solid materials. Without this requirement, an abrupt transition from the reactor chamber 56 to the cooling section 58 is also suitable.
Still another feature of this embodiment of the invention is the presence of an anode injector 64 as part of the torch/injection portion 52. Placing the anode injector 64 closer to the plasma arc reduces heat loss and results in higher mixing efficiency. However, this embodiment also illustrates that one or more injector holes 60 may also be provided behind the torch/injection portion, providing separate reactant injection holes connected to injection channel 62.
The reactor chamber 56 of this embodiment of the invention has an insulating liner 66 on its inner surface.
Referring now to fig. 3, an apparatus 100 according to another embodiment of the present invention generally comprises a torch/injection portion 102, an insulated reactor chamber 104, and a cooling portion 106. This embodiment specifically eliminates the syringe portion used in the above-described embodiment. It should be understood that most of the components remain as described for the device 10.
The apparatus 100 includes an anode injector 110, but no injector behind it. Thus, the anode injector 110 is close to the plasma arc, has space to provide thorough mixing before entering the reaction chamber, and is also closer to the reaction chamber to avoid excessive cooling. However, it can be seen that the torch portion 102 still includes an injection channel 112 ofreduced diameter to create turbulence to ensure thorough mixing of the reactants with the plasma prior to entering the reactor chamber 104.
The apparatus 100 achieves the desired reaction with high efficiency, even without a converging-diverging nozzle, due to the insulating layer 108 on the inner wall of the reactor chamber 104, the injection channel 112 and the judicious placement of the anode injector 110 in closer proximity to the plasma arc.
Although the disclosure herein focuses primarily on the production of acetylene from methane, it will be apparent to those skilled in the art that other materials may be produced using the method and apparatus of the present invention. These materials include, by way of example only, titanium, vanadium, aluminum, and titanium/vanadium alloys.
The following examples are given to illustrate the invention, but they do not limit the scope of the invention.
Examples
During the experiment, all instruments used in the examples described below were connected directly to the data acquisition system which continuously recorded the system parameters, except for Gas Chromatography (GC). Once the specified process power, pressure and gas flow were reached, the gas flow was continuously sampled by gas chromatography for a period of 7 minutes to ensure that a representative sample was obtained, and then taken formally for gas chromatography analysis. This 7 minute sampling time is approximately three times the time required to completely clean the sample tube. The pressure behind the quench nozzle was controlled with a mechanical vacuum pump and a flow control valve. The test pressure may be independently adjusted to between atmospheric pressure and about 100 torr, depending on the test conditions. After 1 minute or less, the test reached a steady state. The arrival of a steady state operating condition is determined with a continuously reading Residual Gas Analyzer (RGA). All cooling water flows and inlet and outlet temperatures are monitored and recorded to calculate the energy balance of the whole system.
The plasma gases used were Ar and H2A mixture of (a); methane or natural gas is injected after the lateral injection injector of the restricted passage. The DC plasma torch used could not be operated for a long time without severe anodic corrosion if pure hydrogen was used, so all experimental data were obtained using a plasma gas containing at least some amount of Ar. The use of inert Ar which does not participate in the process chemistry has the advantage that it provides a reference within the system for confirming the mass balance of the overall process. The severe corrosion of the tungsten cathode by the formation of volatile tungsten carbides avoids the direct handling of methane during the discharge process. The two most critical aspects of the experiment are the chemical analysis of the product stream and the overall mass balance.
Two experimental series of examples were conducted, the primary difference between the two groups being the use or absence of a convergent-divergent quench nozzle. When the quench nozzle is installed, the downstream valve is open and the capacity of the vacuum pump determines the downstream pressure. The downstream pressure for this type of test is typically 100 and 200 torr. In this configuration, the flow is restricted at the location of the nozzle throat, and the reactor pressure is determined by the nozzle throat diameter, the reactor temperature, and the mass flow rates of the plasma gas and the reactant gas. Under these conditions, the reactor pressure is typically 600 and 800 torr at an upstream to downstream pressure ratio of 4 to 6. The corresponding mach number is 1.6-1.8. Assuming a reactor temperature of 2000 ℃, pneumatic quenching rapidly reduces the temperature to1100 ℃ and 1300 ℃. The thermal efficiency of the plasma torch was measured to be 80-90% depending on the type and flow of the gas mixture. The power of the plasma torch was adjusted to a constant 60kW into the plasma gas. Since the plasma torch voltage is determined primarily by the ratio of argon to hydrogen, the power can be adjusted by adjusting the current to achieve the 60kW required to enter the plasma. The energy balance of the injector ring, reactor section and nozzle assembly showed that about 14.6kW of power was lost to the cooling water in these components. These distributions of energy losses are listed in table 1 below:
TABLE 1
Position of | Torch | Syringe with a needle | Reactor with a reactor shell | Nozzle with a nozzle body |
Energy loss: | +60kW enters etc Plasma body | -7.3kW into Cooling water | -3.9kW into Cooling water | -3.4kW into Cooling water |
As a result, approximately 45kW of power is available for conversion of natural gas to acetylene. Again care is taken with the system, including placing the components of the injector function within the flare and optimizing (shortening) the length of the reactor, these losses may be reduced by 70% or more. FIG. 4 is a graph of the theoretical maximum amount of methane that can be converted to acetylene in the present configuration under nominal operating conditions. The nominal operating conditions are defined as: 160 standard liters per minute (slm) Ar, 100slm H for plasma torch gas2The power into the plasma was 60 kW. The target temperature for the reactor was 2000 ℃. Under this nominal condition, the maximum theoretical amount of methane that can be converted to acetylene is about 145 slm.
The conversion efficiency at pure methane injection was defined as:
wherein [ CH4]Is the mole fraction of methane, Q, in the product stream obtained from GCCH4Is the injection flow rate of methane.
The yield of acetylene (for pure methane injection) is defined as:
wherein [ C2H2]Is the mole fraction of acetylene in the product stream as measured by GC, and the actual measured gas flow converted to standard conditions (1 atmosphere and 0 ℃ C.) is QSTP。
Example 1
In this example, the results of an experiment conducted in the presence of a converging-diverging nozzle are given. The conversion efficiency as a function of the methane injection flow rate is shown in fig. 5. The power into the plasma was kept constant at 60kW during the acquisition of this set ofdata, the plasmaThe flow rate of the sub-gas was constant at 160slm of Ar and 100slm of H2. The measured reactor pressure was relatively constant and varied between about 670 and 730 torr depending on the flow rate of the methane injection. The conversion of methane is substantially complete, i.e., the conversion efficiency is 100% at feed streams of methane up to about 100 slm. At feed flows above 100slm the conversion efficiency begins to drop and at feed flows of about 120slm the conversion efficiency drops to below about 95%. The estimated bulk gas temperature in the reactor and the corresponding residence time in the reactor are plotted in fig. 6. This estimate was obtained from the measured system energy balance, the flow of plasma gas and methane, and assuming 100% acetylene yield. The target temperature achieved at a methane flow rate of about 145slm was about 2000 ℃. For methane injection flow rates less than 145slm, the estimated reactor temperature is above 2000 ℃. If the reactor temperature is not uniform (as it is not), and if the process is run according to the equilibrium diagram in the graph shown in fig. 4, it is expected that the conversion efficiency will begin to drop at temperatures much lower than 2000 ℃ (methane feed flow greater than about 145 slm).
Another illustration of this situation is shown in fig. 4, where the theoretically possible maximum amount of methane conversion is plotted as a function of available energy (power). For nominal operating conditions of about 60kW minus about 15kW loss into the plasma gas in the injector, reactor and nozzle assembly, a net amount of about 45kW, the maximum amount of methane that can be converted is about 145 slm. For injection flow rates in excess of 145slm, the energy available is insufficient to decompose the injected methane and convert it to acetylene with 100% efficiency, resulting in a product stream temperature of 2000 ℃. The presence of an inevitable cold boundary layer in the injection ring and the reactor also allows some amount of gas to pass through the reactor without decomposing. At lower flow rates and correspondingly higher temperatures, the methane is almost completely converted, with conversion efficiencies reaching 100%. At 120slm (a value slightly below the expected value of 145slm), a decrease in conversion efficiency was observed. This may be due to the presence of cold boundary layer flow or due to insufficient residence time upon decomposition. Looking at the residence time curve in fig. 6, it can be seen that the residence time in the reactor is relatively constant and independent of the methane injection flow. The increase in mass flow and expected velocity with increasing methane injection flow is offset by the cooling of the gas mixture that also occurs with increasing methane injection flow.
The possible effect of residence time on conversion efficiency was evaluated by replacing the converging-diverging nozzle with a straight pipe of constant diameter section matching the internal diameter of the downstream pipe. The pressure of the reactor can be controlled independently of the flow rate by removing the converging-diverging nozzle. When a convergent-divergent nozzle is installed, flow is restricted (sonic velocity is achieved) at the throat of the nozzle. Under such conditions, the reactor pressure is independent of the downstream pressure, which is determined only by mass flow and temperature. With the convergent-divergent nozzle removed, the reactor pressure is controlled by the position of the downstream valve. Reducing the reactor pressure increases the velocity in the reactor and shortens the residence time. For this series of experiments, the pressure was varied between 300 and 700 torr and the residence time in the reactor was reduced by a factor of 2.3, i.e., from about 3.25ms to 1.4 ms. As shown in fig. 7, at lower pressures, a slight decrease in conversion efficiency of about 2 percentage points was observed. Although this indicates a slight dependence on residence time, this is not sufficient to account for the reduction in conversion efficiency seen in FIG. 5. This indicates that the presence of the cold boundary layer actually plays a major role in the observed reduction in conversion efficiency. The residence time in the reactor should be sufficient to decompose the methane and, as described below, also sufficient to form acetylene.
The acetylene yield as a function of the methane injection flow is shown in fig. 8. The observed trend of the yield data was similar to that of the conversion efficiency shown in fig. 5. At methane injection flow rates less than about 100slm, acetylene yield is about 95%. Further increases in the methane injection flow result in reduced yields. When the injection flow was 145slm (i.e. the theoretical maximum feed flow that could be processed to bring the conversion efficiency and yield close to 100%), the actual measured yield dropped to 75%. In fig. 9, the measured yields are normalized to illustrate the measured reduction in conversion efficiency.
This normalization is simple.
This normalized yield is a measure of the selectivity of conversion to acetylene. As shown in fig. 9, this normalization accounts for the major portion of the observed acetylene yield reduction. Increasing the conversion efficiency results in a significant flattening of the yield curve. This shows that minimizing the cold boundary layer by improving the thermal design results in an improvement in the overall performance of the system and an increase in the intrinsic yield of acetylene over a wide range of reactant flow rates (reactor temperatures).
The measured decrease in yield cannot be used to account for the decrease in conversion efficiency due to the formation of other carbonaceous material. These materials include other hydrocarbons and soot. The observed carbon-based yields of other hydrocarbons are plotted as a function of methane injection flow in fig. 10 and 11. The yields given are expressed as a percentage of carbon introduced into the system with methane. The two figures are identical except that fig. 11 has a different coordinate scale and no methane is present. In FIG. 10, the conversion efficiency is clearly seen as a function of the methane injection flow rate during the methane yield improvementIncreasing and decreasing. Fig. 11 is the same data depicted on an enlarged scale. The observed reduction in normalized yield is due to the increased yield of other carbonaceous material. The species represented on the figure are only those that can be observed by gas chromatography. Of sufficient interest, in olefin + C6And heavy (C)5=/C6+) After the hydrocarbon yield begins to increase, the relative amount decreases slightly at higher methane injection flow rates. These heavier hydrocarbons were then analyzed by GCMS to determine that almost all were benzene. The other carbonaceous species seen, ethylene, propadiene and t-2-butene, steadily increased with increasing injection rate of methane. Ar and H are reacted2After decreasing the ratio by one-half, the conversion, yield and yield of other hydrocarbons were examined for Ar and H2The ratio of the two components is dependent on the relative amount. As a result, there is no significant effect on the conversion efficiency or composition of the product stream. Since the product stream is always enriched in H by its characteristics2So the presence of Ar appears to have little or no effect on the kinetics and does not alter the equilibrium of the product stream.
In addition to the tests using methane as a reactant gas, limited tests were conducted using pipeline natural gas. The observed conversion efficiencies, acetylene yields and yields of other hydrocarbons were the same as previously observed when pure methane was used as the feedstock. The results of the analysis of the natural gas and product streams are listed in table 2.
Except that there is substantially inert N2And CO2The results are virtually identical to those in the pure methane test, except for the conversion to CO. In carbon-rich systems, CO-CO at moderate temperatures of about 1000 deg.C2The equilibrium of (c) tends towards CO.
Table 2.Gas chromatographic analysis of the natural gas reactant and product streams, expressed as mole percentages, 60kW plasma power, 140slm of Ar, 100slm of H2And 98.5slm of natural gas
H2 | C5=/C6+ | Propane C3H8 | Acetylene C2H2 | Isobutaneous kang C4H10 | N-butane C4H10 | CO | Ethylene C2H4 | Ethane (III) C2H6 | Ar | N2 | Methane CH4 | CO | Solid carbon Yield of | |
Natural substance (such as natural gas) Qi (Qi) | 0 | 0.04 | 0.74 | 0 | 0.1 | 0.11 | 0.5 | 0 | 3.79 | 0 | 1.18 | 93.47 | 0 | - |
Product of Flow of | 53. 5 | 0.34 | 0 | 11.8 | 0 | 0 | 0 | 0.182 | 0 | 33 | 0.44 | 0.21 | 0.2 | 3.2% |
Example 2
In this example, the results of the test without the use of the convergent-divergent nozzle are given. The resulting yield conversion efficiency (-100% vs. 70%) and selectivity (95% vs. 51%) of this process appeared to be slightly better than the original Huels process. This may be due to better mixing, temperature uniformity, or faster quenching. The analysis of the hull product stream and the laboratory scale results are summarized in table 2.
To evaluate the effect of rapid pneumatic quenching, the same series of tests as described in example 1 were performed, but without the use of a converging-diverging nozzle. In these runs, the system pressure was maintained at 700 and 900 torr, which was substantially the same as the reactor pressure in the series of runs of example 1. The results of conversion efficiency, yield and normalized yield are summarized in FIGS. 12-14. The yields of other hydrocarbons are summarized in fig. 15.
The yield and conversion efficiency results obtained without the convergent-divergent nozzle and ultrasonic pneumatic quenching are virtually identical to the previous results obtained in the presence of the nozzle. When nozzles are present, there appears to be some minor improvement in yield at higher methane injection rates, but the deviation is within the uncertainty estimation range and the error ranges overlap significantly. As shown in fig. 15, the results of the examination of the yields of other hydrocarbons showed that there was a statistically significant difference in the yield of ethylene between the results obtained without the use of the nozzle and with the use of the nozzle (fig. 11). It is clear that the high quench rate provided by the nozzle has the effect of inhibiting ethylene formation, cutting the ethylene yield by about 30%. The precise mechanism of ethylene formation is not yet clear; however, CH2The kinetics and number of free radicals may play an important role.
In the case of no convergent-divergent nozzle, the reactor pressure can be controlled independently by valves. This structure enables the study of the effect of pressure on yield. As shown in fig. 7, it has been previously demonstrated that pressure and residence time have only a small effect on conversion efficiency. The measured acetylene yield as a function of reactor pressure is depicted in fig. 16. The power of the inlet gas was constant at 60kW, the flow rate was maintained at 160slm of Ar, 100slm of H2And CH of 98.5slm4. The acetylene yield seems to decrease slightly with increasing pressure, but this effect is not as great. As shown in fig. 17, while the yield of acetylene was decreased, the yield of benzene was also slightly decreased, and the yield of ethylene was increased. Typically, pressure variations in the range of 300 torr to 700 torr are not significantly affected.
TABLE 2 analysis of the product stream in mole percent
Acetylene | Third two Alkene(s) | Diethyl (diethylene) Alkynes | Ethylene | Methane | Ethane (III) | Propane | Chinese medicinal composition Alkane (I) and its preparation method | N-butyl Alkane (I) and its preparation method | Benzene and its derivatives | |||||||
C2H2 | C3H4 | C4H2 | C4H4 | C2H4 | CH4 | C2H6 | C3H8 | C4H10 | C4H10 | C6H6 | H2 | CO | N2 | Carbon yield | ||
Huels Method of | Raw materials | - | - | - | - | 92.3 | 1.4 | 0.5 | - | 0.4 | - | - | - | 5.4 | - | |
Product of | 14.5 | 0.6 | 0.1 | 0.9 | 16.3 | 0.04 | 0.03 | 0.01 | - | 0.3 | 63.4 | 0.6 | 2.7 | 2.7% | ||
Current methods W/quench | Raw materials | - | - | - | - | 93.47 | 3.79 | 0.74 | 0.10 | 0.11 | 0.04 | - | - | 1.18 | - | |
Product of* | 11.8 | - | - | - | 0.21 | - | - | - | - | 0.34 | 53.47 | 0.25 | 0.44 | 3.2% |
Also contains 33.3 mol% argon
Example 3
The overall data for one experiment is given in tables 3-5 below. This test was performed without a convergent-divergent nozzle. A summary of the flow, power, measured flare thermal efficiency, nozzle geometry, and reactor and outlet pressures is presented in Table 3.
TABLE 3
Test SEPT13A-4P | |
Ar(slm) | 160.4 |
H2(slm) | 100.1 |
CH4(slm) | 75.9 |
Net power (kW) | 60.2 |
Efficiency (%) | 85 |
Geometric shape | Straight |
Outlet pressure (torr) | 409 |
Reactor pressure (torr) | 550 |
Table 4 lists theacetylene and hydrogen yields, volumetric flow rates, energy consumption rates (SER) and relative yields for this experiment. The column and row headings are generally sufficient to illustrate the data included. Specifically, columns labeled yield%: qt meas. provides the yield of acetylene and hydrogen from the methane feed and the measured conversion efficiency. The acetylene yield is the percentage of carbon in the methane feedstock to form acetylene and the hydrogen yield is the percentage of hydrogen in the methane feedstock to form elemental hydrogen. Qt represents the measurement based on the downstream turbine flow meter. Columns labeled yield%: qt At std. provides the yield of acetylene and hydrogen from the methane feed and the measured conversion efficiency. Q Ar std represents the measurement based on argon input flow and gas chromatography data.
The column labeled volumetric flow (slm) provides the volumetric flow of acetylene and hydrogen produced from the methane feedstock. Qt represents the measurement based on the downstream turbine flow meter. Q Ar std represents the measurement based on argon input flow and gas chromatography data.
The next four columns provide the energy consumption rate. Qt represents the measurement based on the downstream turbine flow meter. Q Ar std represents the measurement based on argon input flow and gas chromatography data.
The columns labeled R yield provide relative yields. The yield figures have normalized conversion efficiency.
TABLE 4
Yield% | Volume flow (slm) | SER Qt meas | SER Q Ar std | SER Qt meas | SER Q Ar std | R product Rate of change | |||
Qt meas | Q Ar std | Qt meas | Q Ar std | kW- hr/kg | kW- hr/kg | kW- hr/Mscf | kW- hr/Mscf | ||
Ar std | |||||||||
C2H2 | 98.2 | 96.14 | 37.27 | 36.48 | 23.19 | 23.69 | 765.60 | 782.14 | 0.97 |
Qt | |||||||||
Is obtained from CH1 H of (A) to (B)2 | 47.9 0 | 45.49 | 72.70 | 69.1 | 154.56 | 162.74 | 392.49 | 413.26 | 0.99 |
Transformation of Efficiency of % | 99.3 | 99.30 |
Table 5 lists the flow and mass balance data for the various materials in this test. The column and row headings are generally sufficient to describe the data contained therein to those skilled in the art. Specifically, the column labeled [ conc%]provides concentrations in mole percent as determined by gas chromatography.
Table 5 is presented in two parts due to its bulkiness.
TABLE 5 part A
Substance(s) | [conc%] | Qin slm | mdot in | Qt meas | mt meas | mdot H in | mdot H out |
H2(Hydrogen) Gas) | 45.825 | 100.1 | 8.94 | 377.1 | 15.43 | 8.94 | 15.43 |
C5=/C6+ (C6H6) | 0.217 | 0.0 | 0.00 | 377.1 | 2.85 | 0.00 | 0.22 |
C3H8(III) Alkane) | 0.000 | 0.0 | 0.00 | 377.1 | 0.00 | 0.00 | 0.00 |
C2H2(B) Alkyne) | 9.884 | 0.0 | 0.00 | 377.1 | 43.26 | 0.00 | 3.33 |
C3H6(III) Alkene) | 0.000 | 0.0 | 0.00 | 377.1 | 0.00 | 0.00 | 0.00 |
C4H10(different from each other Butane) | 0.000 | 0.0 | 0.00 | 377.1 | 0.00 | 0.00 | 0.00 |
C3H4(III) Diene) | 0.000 | 0.0 | 0.00 | 377.1 | 0.00 | 0.00 | 0.00 |
C1H10(n-type) Butane) | 0.000 | 0.0 | 0.00 | 377.1 | 0.00 | 0.00 | 0.00 |
C1H8(1- Butylene) | 0.000 | 0.0 | 0.00 | 377.1 | 0.00 | 0.00 | 0.00 |
C4H8(different from each other Butane iso Butylene) | 0.000 | 0.0 | 0.00 | 377.1 | 0.00 | 0.00 | 0.00 |
C1H8(t- 2-butene) | 0.000 | 0.0 | 0.00 | 377.1 | 0.08 | 0.00 | 0.01 |
C1H8(c- 2-butene) | 0.008 | 0.0 | 0.00 | 377.1 | 0.00 | 0.00 | 0.00 |
C1H6(1,3 -butylene glycol Alkene) | 0.000 | 0.0 | 0.00 | 377.1 | 0.00 | 0.00 | 0.00 |
C5H12(different from each other Pentane) | 0.000 | 0.0 | 0.00 | 377.1 | 0.00 | 0.00 | 0.00 |
C5H12(n-type) Pentane) | 0.000 | 0.0 | 0.00 | 377.1 | 0.00 | 0.00 | 0.00 |
CO2(II) Oxidation by oxygen Carbon) | 0.000 | 0.0 | 0.00 | 377.1 | 0.00 | 0.00 | 0.00 |
C2H4(B) Alkene) | 0.459 | 0.0 | 0.00 | 377.1 | 2.16 | 0.00 | 0.31 |
C2H6(B) Alkane) | 0.000 | 0.0 | 0.00 | 377.1 | 0.00 | 0.00 | 0.00 |
Ar (argon) Gas) | 43.454 | 160.4 | 286.43 | 377.1 | 292.62 | 0.00 | 0.00 |
N2(Nitrogen) | 0.000 | 0.0 | 0.00 | 377.1 | 0.00 | 0.00 | 0.00 |
CH4(first) Alkane) | 0.145 | 75.9 | 54.21 | 377.1 | 0.39 | 13.55 | 0.10 |
CO (oxygen monoxide) Chemical carbon) | 0.000 | 0.0 | 0.00 | 377.1 | 0.00 | 0.00 | 0.00 |
Total of g/min | 349.58 | 356.79 | 22.49 | 19.39 | |||
% difference | -2.06 | 13.77 | |||||
Soot produced on carbon balance |
TABLE 5 part B
Substance(s) | mdot C in | mdot C out | Yield Qt | Q Ar std | mt Ar std | mdot H Ar | mdot C Ar std | Yield Ar std | |
H2(Hydrogen) Gas) | 0.00 | 0.00 | 0.0 | 369.13 | 15.10 | 15.10 | 0.00 | 0.0 | |
C5=/C6+ (C6H6) | 0.00 | 2.63 | 6.5 | 369.13 | 2.79 | 0.21 | 2.58 | 6.34 | |
C3H8(III) Alkane) | 0.00 | 0.00 | 0.0 | 369.13 | 0.00 | 0.00 | 0.00 | 0.00 | |
C2H2(B) Alkyne) | 0.00 | 39.93 | 98.2 | 369.13 | 42.35 | 3.26 | 39.09 | 96.14 | |
C3H6(III) Alkene) | 0.00 | 0.00 | 0.0 | 369.13 | 0.00 | 0.00 | 0.00 | 0.00 | |
C1H10(different from each other Butane) | 0.00 | 0.00 | 0.0 | 369.13 | 0.00 | 0.00 | 0.00 | 0.00 | |
C3H4(III) Diene) | 0.00 | 0.00 | 0.0 | 369.13 | 0.00 | 0.00 | 0.00 | 0.00 | |
C1H10(n-type) Butane) | 0.00 | 0.00 | 0.0 | 369.13 | 0.00 | 0.00 | 0.00 | 0.00 | |
C1H8(1- Butylene) | 0.00 | 0.00 | 0.0 | 369.13 | 0.00 | 0.00 | 0.00 | 0.00 | |
C4H8(different from each other Butane iso Butylene) | 0.00 | 0.00 | 0.0 | 369.13 | 0.00 | 0.00 | 0.00 | 0.00 | |
C4H8(t- 2-butene) | 0.00 | 0.07 | 0.2 | 369.13 | 0.08 | 0.01 | 0.07 | 0.16 | |
C1H8(c- 2-butene) | 0.00 | 0.00 | 0.0 | 369.13 | 0.00 | 0.00 | 0.00 | 0.00 | |
C1H6(1,3 -butylene glycol Alkene) | 0.00 | 0.00 | 0.0 | 369.13 | 0.00 | 0.00 | 0.00 | 0.00 |
C5H12(different from each other Pentane) | 0.00 | 0.00 | 0.0 | 369.1 3 | 0.00 | 0.00 | 0.00 | 0.00 |
C5H12(n-type) Pentane) | 0.00 | 0.00 | 0.0 | 369.13 | 0.00 | 0.00 | 0.00 | 0.00 |
CO2(II) Oxidation by oxygen Carbon) | 0.00 | 0.00 | 0.0 | 369.13 | 0.00 | 0.00 | 0.00 | 0.00 |
C2H4(B) Alkene) | 0.00 | 1.85 | 4.6 | 369.13 | 2.12 | 0.30 | 1.82 | 4.47 |
C2H6(B) Alkane) | 0.00 | 0.00 | 0.0 | 369.13 | 0.00 | 0.00 | 0.00 | 0.00 |
Ar (argon) Gas) | 0.00 | 0.00 | 0.0 | 369.13 | 286.43 | 0.00 | 0.00 | 0.00 |
N2(Nitrogen) | 0.00 | 0.00 | 0.0 | 369.13 | 0.00 | 0.00 | 0.00 | 0.00 |
CH1(first) Alkane) | 40.66 | 0.29 | 0.7 | 369.13 | 0.38 | 0.10 | 0.29 | 0.70 |
CO (oxygen monoxide) Chemical carbon) | 0.00 | 0.00 | 0.0 | 369.13 | 0.00 | 0.00 | 0.00 | 0.00 |
Total of g/min | 40.66 | 44.78 | 110.1 | 349.25 | 18.98 | 43.84 | 107.8 | |
% difference | -10.14 | 0.09 | 15.59 | -7.81 | ||||
Based on carbon Balance of birth Raw soot % | -10.14 | -7.81 |
The conversion efficiency and acetylene yield measured in the laboratory reactor system described in this example 1 and 2 are generally slightly better than reported in the literature; conversion Efficiency (CE) was confirmed to be close to 100% with yields in the range of 90-95% and soot produced in the range of 2-4%. This seems to be comparable to the Huels method (CE 70.5%, C)2H2Yield 51.4% carbon soot 2.7%) and DuPont method (CE-not reported, C)2H2Yield 70%) was somewhat improved. The method reported here appears to have a slightly better specialization also for acetyleneAnd (4) uniformity. At the conversion efficiency,The improvements in yield and specificity are mainly due to improvements in injector design and mixing (better "stirred" reactors) and minimization of temperature gradients and cold boundary layers. The cooling rate by heat transfer on the wall appears to be sufficient to quench the product stream, preventing further decomposition of the acetylene into soot or preventing further reaction to heavier hydrocarbon products. The quench rate was significantly increased by rapidly expanding the product stream through a converging-diverging nozzle, with only marginal improvement in the composition of the product stream, primarily a reduction in ethylene yield.
The energy consumed (kW-hr) per unit quantity (kg) of acetylene produced ultimately determines the economic efficiency of the process. The Huels process is reported to consume 12.1kW-hr/kg of C produced2H2. Although not measured, the DuPont method is estimated to have an energy consumption rate of 8.8kW-hr/kg C produced2H2. This latter value is about 7.9kW-hr/kg-C, compared to the theoretical minimum for a product stream produced at a temperature of 2000 ℃, conversion efficiency and yield of 100%, and without loss of electricity or heat2H2Preferably. The energy consumption rates measured in the laboratory-scale process of this study are depicted in fig. 18. The minimum energy consumption rate measured was about 16kW-hr/kg of C produced2H2. It is estimated that this value can be improved to about 13kW-hr/kg-C by improving the thermal design2H2. This improvement involves moving the injection into the torch body, which avoids heat loss in the injector ring and reduces heat loss in the reactor section. The energy consumption rate can be further reduced by about 20% by recovering heat in the process to about 10kW-hr/kg-C2H2The range of (1). These values are better than the energy consumption rates reported by the Huels and DuPont methods, while at the same time the conversion efficiency and yield are higher.
The present invention may be embodied in other specific forms without departing from its spirit or essential characteristics. The described embodiments are to be considered in all respects only as illustrative and not restrictive. The scope of the invention is, therefore, indicated by the appended claims rather than by the foregoing description. All changes which come within the meaning and range of equivalency of the claims are to be embraced within their scope.
Claims (52)
1. A process for converting one or more reactants to at least one desired product comprising:
contacting the reactant stream with a hot gas to create a turbulent flow regime, thereby substantially thoroughly mixing the reactant stream with the hot gas;
axially passing a substantially thoroughly mixed reactant stream through a reaction zone of an axial reactor while maintaining the reaction zone at a substantially uniform temperature throughout its length, the axial reactor being operated under conditions sufficient to heat the reactant stream to a selected reaction temperature such that a stream comprising at least one desired product is produced at least proximate to the outlet end of the axial reactor.
2. The method of claim 1, wherein the reactant stream comprises methane and the at least one desired product comprises acetylene.
3. The method of claim 1, wherein the reactant stream comprises methane or carbon monoxide and at least one desired product comprises hydrogen.
4. The method of claim 1, wherein the reactant stream comprises a titanium compound and the at least one desired product comprises titanium or titanium dioxide.
5. The process as set forth in claim 1 wherein the temperature of the reaction zone is maintained at about 1500-4000 ℃.
6. The process as set forth in claim 5 wherein the temperature of the reaction zone is maintained at about 1700-2000 ℃.
7. The process of claim 1 wherein the hot gas comprises a plasma, the process further comprising passing a plasma arc gas stream between plasma torch electrodes comprising at least one pair of electrodes positioned proximate the inlet end of the axial reaction zone while subjecting the electrodes to a plasma input power selected to generate a plasma in an injection passage connected to the reaction zone.
8. The method of claim 7, further comprising injecting at least one reactant into the injection channel to thoroughly mix the reactant stream into the plasma, thereby achieving thorough mixing, and then introducing into the reaction zone while maintaining a substantially uniform temperature throughout the length of the reaction zone, thereby allowing the at least one reaction to reach equilibrium within the reaction zone.
9. The method of claim 1, further comprising:
cooling the stream comprising at least one desired product after it exits the outlet end of the axial reaction zone by reducing the velocity of the stream while removing heat energy at a rate sufficient to prevent an increase in its motive temperature;
at least one desired product is separated from the gas remaining in the cooling stream.
10. A process according to claim 9 wherein the stream comprising the desired product and residual gases is cooled and slowed down as it leaves the axial reaction zone by the addition of a quench gas to the stream at a rate which condenses at least one of the desired products and which suppresses the formation of other equilibrium products as the stream leaves the reaction zone.
11. The process of claim 1 further comprising rapidly cooling the stream containing at least one desired product produced in the reaction zone by passing it through a coaxial converging-diverging nozzle located proximate the outlet end of the axial reactor.
12. The method of claim 1, further comprising selecting the size of the restricted opening throat in the nozzle to control the residence time and reaction pressure of the substantially thoroughly mixed reactant stream in the axial reactor.
13. The method of claim 11, further comprising reducing the pressure of the stream containing at least one desired product ultra-rapidly by smoothly accelerating and expanding the stream along the diverging portion of the nozzle to further reduce its motive temperature and prevent unwanted side or reverse reactions from occurring.
14. The method of claim 1, further comprising selecting the reactant stream to comprise at least one reactant selected from the group consisting of titanium tetrachloride, vanadium tetrachloride, aluminum trichloride, methane, and natural gas prior to reacting or thermally decomposing the reactant stream.
15. The method of claim 1, wherein:
the hot gas comprises a plasma generated from a gas comprising an inert gas, hydrogen, or a mixture thereof;
the reactant stream comprises gaseous or volatile compounds of the selected metals:
operating the axial reactor, including forming an equilibrium mixture comprising the desired product in the form of at least one selected metal or oxide or alloy thereof, which selected metal, metal oxide or metal alloy is thermodynamically stable at a selected reaction temperature; the method further comprises the following steps:
cooling the stream comprising at least one desired product as it exits the outlet end of the axial reactor by reducing the velocity of the stream while removing heat energy at a rate sufficient to prevent an increase in its motive temperature;
at least one desired product is separated from the gas remaining in the cooling stream.
16. The method of claim 15 wherein the gaseous or volatile compound of the selected metal is a gaseous and volatile halide.
17. The method of claim 15, wherein the selected metal is titanium, vanadium, or aluminum.
18. The method of claim 15, wherein the compound of the selected metal is titanium tetrachloride, vanadium tetrachloride, or aluminum trichloride.
19. The method of claim 15, wherein the reactant stream further comprises at least one other material capable of reacting at the reaction temperature to form an equilibrium mixture comprising an oxide or alloy of the selected metal.
20. The method of claim 19, wherein the method forms an alloy of titanium and a second metal, and the reactant stream comprises titanium chloride and a gaseous or volatizable compound of the second metal.
21. The method of claim 20, wherein the second metal is vanadium.
22. The method of claim 19, wherein the method forms a metal oxide of the selected metal, and the reactant stream further comprises oxygen.
23. The process of claim 22, wherein the process forms titanium oxide and the reactant stream comprises titanium tetrachloride and oxygen.
24. The method of claim 1, wherein:
the hot gas comprises a plasma generated from a gas comprising an inert gas, hydrogen, or a mixture thereof;
the reactant stream comprises gaseous or volatile hydrocarbons:
operating an axial reactor, comprising forming an equilibrium mixture comprising at least one desired product, the equilibrium mixture being thermodynamically stable at a reaction temperature; the method further comprises the following steps:
cooling the stream comprising at least one desired product as it exits the outlet end of the reactor by reducing the velocity of the stream while removing heat energy at a rate sufficient to prevent an increase in its motive temperature;
at least one desired end product is separated from the gas remaining in the cooled stream.
25. The method of claim 24, wherein the reactant stream comprises natural gas or methane.
26. The method of claim 24, wherein at least one desired product comprises acetylene.
27. An apparatus for thermally converting one or more reactants into at least one desired product, the apparatus comprising:
an axial reactor comprising an inlet end, an outlet end, and a reaction zone therebetween, the axial reactor being configured to maintain the reaction zone at a substantially uniform temperature throughout its length, wherein the axial reactor is operated at conditions sufficient to heat the reactant stream to a selected reaction temperature such that a stream comprising at least one desired product is produced at a location at least proximate to the outlet end of the axial reactor;
a flare portion configured to generate hot gas ahead of the axial reactor;
an injector section positioned and configured to introduce one or more reactants into the hot gas to produce a stream flowing toward the inlet end of the axial reactor.
28. The apparatus of claim 27, further comprising a converging-diverging nozzle coupled to the outlet end of the axial reactor and configured to rapidly cool a stream containing at least one desired product by producing adiabatic and isentropic expansion of the hot gas stream as it flows axially through the nozzle to convert thermal energy to kinetic energy.
29. The apparatus of claim 28, wherein the converging-diverging nozzle has a converging portion leading to the restricted opening throat and a diverging portion extending from the restricted opening throat, the diverging portion having a conical configuration centered along the axis of the axial reactor.
30. The apparatus of claim 29, wherein the conical configuration of the nozzle diverging portion has an included angle of less than about 35E.
31. The method of claim 28, wherein the converging-diverging nozzle has a converging portion with a high aspect ratio configured to rapidly accelerate a stream comprising at least one desired product into the throat of the nozzle while maintaining laminar flow therethrough.
32. The apparatus of claim 28 further comprising a cooling section after the converging-diverging nozzle.
33. The apparatus of claim 32, wherein the cooling section is configured to reduce the velocity of the stream comprising the at least one desired product while removing thermal energy at a rate sufficient to prevent a motive temperature thereof from rising and retain the at least one desired product therein.
34. The apparatus of claim 27, further comprising a cooling section connected to the outlet end of the axial reactor.
35. The apparatus of claim 34, wherein the cooling section is configured to reduce the velocity of the stream comprising the at least one desired product while removing heat energy at a rate sufficient to prevent a motive temperature thereof from rising and retain the desired product therein.
36. The apparatus of claim 34 further comprising a converging section interposed between the outlet end of the axial reactor and the cooling section.
37. The device of claim 27, further comprising a multi-channel syringe located in the syringe portion or the torch portion.
38. The apparatus of claim 27, wherein the torch portion comprises an anode injector.
39. The apparatus of claim 38, wherein the anode injector comprises a multi-port injector.
40. The apparatus of claim 27, wherein the hot gas comprises a plasma and the torch portion comprises a plasma torch having a plasma arc inlet for introducing a plasma arc gas stream into the plasma torch to generate the plasma in an axial reactor extending to an outlet end thereof.
41. The apparatus of claim 40, wherein the torch portion comprises a plasma torch comprising at least one pair of electrodes positioned in front of the inlet end of the reaction zone.
42. The apparatus of claim 41 wherein the plasma arc inlet is positioned in front of an electrode for introducing a plasma arc gas stream between the electrodes at a selected plasma gas flow rate while applying a selected plasma input power to the electrodes.
43. The apparatus of claim 27, further comprising at least one reactant inlet positioned before the inlet end for thoroughly mixing the reactant stream into the hot gas prior to entering the reaction zone.
44. The reactor of claim 43, wherein the reactant inlet comprises at least one of a multi-channel injector located in the injection portion and an anode injector located in the flare portion.
45. The reactor of claim 43, wherein the reactant inlet comprises a multi-channel injector proximate to a plasma torch in the torch portion.
46. The apparatus of claim 27, wherein the torch portion comprises one or more lasers or structures for generating an oxidizing flame from a suitable fuel.
47. The apparatus of claim 27, wherein the injector portion comprises an injection channel having a diameter smaller than the diameter of the reaction zone, the injection channel being located entirely within the flare portion, partially within the flare portion and partially between the flare portion and the inlet end of the axial reactor, or between the flare portion and the inlet end of the axial reactor.
48. The apparatus of claim 27, further comprising an insulating layer surrounding the reaction zone of the axial reactor.
49. The device of claim 48, wherein the insulating layer comprises a material selected from the group consisting of carbon, boron nitride, zirconia, silicon carbide, and combinations thereof.
50. The device of claim 48, further comprising a cooling layer surrounding the insulating layer.
51. The apparatus of claim 50, wherein the cooling layer comprises a layer of water.
52. The apparatus of claim 28, wherein the converging-diverging nozzle is configured to allow the reaction zone to operate at atmospheric pressure or higher while the cooling section is maintained in a vacuum state.
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