METHOD FOR OPERATING A SMELTING FURNACE INSTALLATION

Abstract

A method for operating a smelting furnace installation, in particular a blast furnace installation, the method including the following steps: feeding coke, iron oxide containing material and if required fluxing agents to the top of the smelting furnace; injecting a first reducing gas containing hydrogen at a tuyere level of the smelting furnace at a temperature above 1600 C.; and injecting a second reducing gas at a lower shaft level of the smelting furnace.

The coke is fed at a lump coke rate below 220 kg/t HM, and the density of the first reducing gas is below 0.80 kg/Nm.sup.3.

Claims

1. A method for operating a smelting furnace installation, in particular a blast furnace installation, the method including the following steps: feeding coke, iron oxide containing and other iron bearing material and, if required, fluxing agents to the top of the smelting furnace, injecting a first reducing gas containing hydrogen at a tuyere level of the smelting furnace, the first reducing gas being heated at a temperature above 1600 C., and injecting a second reducing gas at a lower shaft level of the smelting furnace within the gas-solid reduction zone of ferrous oxide above the cohesive zone, wherein coke is fed at a lump coke rate below 220 kg/t HM, and wherein the density of the first reducing gas is below 0.80 kg/Nm.sup.3.

2. The method as claimed in claim 1, further comprising injecting oxygen at the tuyere level of the smelting furnace at a rate below 120 Nm.sup.3/t HM, wherein the temperature of the injected oxygen is below 600 C.

3. The method as claimed in claim 1, wherein the first reducing gas is injected at the tuyere level at a total mass flow below 800 kg/t HM.

4. The method as claimed in claim 1, wherein the first reducing gas and/or the second reducing gas comprise(s) a gas produced by a reforming process, by reforming coke oven gas, natural gas, biogas and/or other hydrocarbon containing gases, with H.sub.2O, CO.sub.2, or a CO.sub.2 and/or H.sub.2O containing gas and/or more preferably with a steel plant offgas such as smelting furnace top gas, basic oxygen furnace gas, and/or open bath furnace gas.

5. The method as claimed in claim 1, wherein the first reducing gas and/or the second reducing gas comprise(s) a gas produced by applying a CO.sub.2 separation technique, such as absorption with monoethanolamine (MEA), membrane separation, Pressure Swing Adsorption (PSA) or Vacuum Pressure Swing Adsorption (VPSA) to a hydrogen and/or CO rich gas, to a steel plant gas such as smelting furnace top gas, basic oxygen furnace gas, and/or oxygen blast furnace gas.

6. The method as claimed in claim 1, wherein the first reducing gas has a hydrogen content above 30 vol.-%.

7. The method as claimed in claim 1, wherein the first reducing gas is injected at a temperature from 1600 C. to 2600 C.

8. The method as claimed in claim 1, wherein the first reducing gas is heated with one or more electric heaters before injection to the smelting furnace, within the tuyere stock(s) and/or the tuyere(s), by one or more plasma torches.

9. The method as claimed in claim 8, wherein the first reducing gas is heated with one or more plasma torches arranged within a blowpipe of the tuyere stock(s), the plasma torches being electrode-based plasma torches or electrodeless plasma torches, such as selected from inductively ignited plasma torches, microwave plasma torches, radiofrequency plasma torches or a combination thereof.

10. The method as claimed in claim 9, wherein the one or more plasma torches are direct current plasma torches and/or alternating current plasma torches and/or 3-phase alternating current plasma torches, wherein said plasma torches have an electric power rating of 1 to 10 MW.

11. The method as claimed in claim 1, wherein the first and/or the second reducing gas has a molar ratio (H.sub.2+CO)/(H.sub.2O+CO.sub.2) above 6.

12. The method as claimed in claim 1, wherein the second reducing gas has a hydrogen content above 25 vol.-%.

13. The method as claimed in claim 1, wherein the second reducing gas is injected at a temperature from 800 C. to 1200 C.

14. The method as claimed in claim 1, wherein a pressure level of the smelting furnace at the tuyere level is controlled to values above 2 barg.

15. The method as claimed in claim 1, wherein the first reducing gas and the second reducing gas have a nitrogen content below 35 vol.-%.

16. The method as claimed in claim 1, wherein the first and/or second reducing gas comprises a gas resulting from cracking ammonia.

17. The method as claimed in claim 1, wherein the coke is fed in layers and wherein the height of each coke layer is at least 10 cm.

18. The method as claimed in claim 1, further comprising a step of adjusting the average reduction degree of the iron oxide containing material reaching the cohesive zone to a value of above 85% by controlling the amount and/or composition of the second reducing gas injected at the shaft level as a function of the amount and/or composition of the first reducing gas injected at tuyere level and/or the amount of oxygen injected at tuyere level.

19. The method as claimed in claim 1, further comprising the step of reducing the channeling effect and flooding effect by controlling the top pressure of the smelting furnace in the range 1 to 10 barg.

20. The method as claimed in claim 1, further comprising the step of reducing the wall channeling effect of the gas coming from the cohesive zone by controlling the injection conditions of the second reducing gas, such as the injection speed and/or rate of the second reducing gas injected in the shaft of the smelting furnace.

21. The method as claimed in claim 1, further comprising the step of reducing the carbon dioxide content of any one or more carbon dioxide containing offgas and/or process gas produced during operation by carbon capture and utilization (CCU) and/or carbon capture and storage (CCS).

22. The method as claimed in claim 1, further comprising the step of converting carbon dioxide of any one or more carbon dioxide containing offgases produced during operation into a synthetic fuel, such as into synthetic natural gas by methanation, or into methanol and/or ethanol by methanol and/or ethanol production.

23. A smelting furnace installation, in particular a blast furnace installation, comprising a charging apparatus configured for feeding coke, iron oxide containing and other iron bearing material and if required fluxing agents to the top of the smelting furnace; a first injector arrangement positioned at a tuyere level of the smelting furnace and configured for injecting a first reducing gas containing hydrogen at said tuyere level of the smelting furnace at a temperature above 1600 C.; and a second injector arrangement positioned at a shaft level of the smelting furnace and configured for injecting a second reducing gas at a lower shaft level of the smelting furnace within the gas-solid reduction zone of ferrous oxide above the cohesive zone, wherein the charging apparatus is configured for feeding coke at a lump coke rate below 220 kg/t HM, wherein the first injector arrangement is configured for injecting the first reducing gas at a density below 0.80 kg/Nm.sup.3, at the tuyere level and wherein said first injector arrangement comprises an electric heating device configured for heating the first reducing gas at said temperature above 1600 C.

24. The smelting furnace installation as claimed in claim 23, wherein the first injector arrangement further comprises an oxygen injection port configured for injecting oxygen at the tuyere level of the smelting furnace at a rate below 120 Nm.sup.3/t HM, wherein the temperature of the injected oxygen is preferably below 600 C.

25. The smelting furnace installation as claimed in claim 23, wherein the first injector arrangement is configured for injecting the first reducing gas at the tuyere level at a total mass flow below 800 kg/t HM.

26. The smelting furnace installation as claimed in claim 23, further comprising one or more reformer configured for producing a gas as a first and/or second reducing gas by a reforming process, in particular by reforming coke oven gas, biogas, natural gas and/or other hydrocarbon containing gases, with H.sub.2O and/or CO.sub.2 or a CO.sub.2 and/or H.sub.2O containing gas and/or more preferably with a steel making offgas such as smelting furnace top gas, basic oxygen furnace gas, open bath furnace gas.

27. The smelting furnace installation as claimed in claim 23, further comprising one or more apparatuses configured for separation of CO.sub.2 by a CO.sub.2 separation process such as absorption with monoethanolamine (MEA), membrane separation, Pressure Swing Adsorption (PSA) or Vacuum Pressure Swing Adsorption (VPSA) for treating smelting furnace top gas, basic oxygen furnace gas and/or open bath furnace gas.

28. The smelting furnace installation as claimed in claim 23, further comprising a first source of hydrogen and a first hydrogen content controller configured to adjust a hydrogen content of the first reducing gas to values above 30 vol.-%.

29. The smelting furnace installation as claimed in claim 23, wherein the electric heating device of said first injector arrangement is configured for heating the first reducing gas to temperatures from 1600 C. to 2600 C.

30. The smelting furnace installation as claimed in claim 23, wherein said electric heating device comprises one or more electric resistance heaters and/or one or more plasma torches.

31. The smelting furnace installation as claimed in claim 30, wherein said electric heating device comprises one or more plasma torches arranged within a blowpipe of the tuyere stock(s), the plasma torches being electrode-based plasma torches or electrodeless plasma torches, such as selected from inductively ignited plasma torches, microwave plasma torches, radiofrequency plasma torches, or a combination thereof.

32. The smelting furnace installation as claimed in claim 31, wherein the one or more plasma torches are direct current plasma torches and/or alternating current plasma torches and/or 3-phase alternating current plasma torches, wherein said plasma torches have an electric power rating of 1 to 10 MW.

33. The smelting furnace installation as claimed in claim 23, further comprising a second source of hydrogen and a second hydrogen content controller configured to adjust a hydrogen content of the second reducing gas above 25 vol.-%.

34. The smelting furnace installation as claimed in claim 23, wherein the first injector arrangement comprises an upstream regulation device which is configured for controlling the mass flow rate of the first reducing gas injected in the smelting furnace at the tuyere level to 800 kg/t HM at the pressure level above 2 barg.

35. The smelting furnace installation as claimed in claim 23, further comprising regulating unit configured for adjusting the average reduction degree of the iron oxide containing material reaching the cohesive zone to a value of above 85% by controlling the amount and/or composition of the second reducing gas injected through the second injector arrangement at the shaft level as a function of the amount and/or composition of the first reducing gas injected at tuyere level and/or the amount of oxygen injected through the oxygen injection port at tuyere level.

36. The smelting furnace installation as claimed in claim 23, further comprising a carbon capture and utilization (CCU) unit and/or a carbon capture and storage (CCS) unit downstream of at least one carbon dioxide containing offgas producing element for reducing the carbon dioxide content of said offgas.

37. The smelting furnace installation as claimed in claim 23, further comprising a methanation unit configured for converting carbon dioxide to synthetic natural gas and/or process gas or any other device for producing synthetic hydrocarbons such as methanol and ethanol.

Description

BRIEF DESCRIPTION OF THE DRAWINGS

[0079] Preferred embodiments of the disclosure will now be described, by way of example, with reference to the accompanying drawings in which:

[0080] FIG. 1 is a schematic view of an advantageous embodiment of the present disclosure.

[0081] FIG. 2 is a schematic partially cut-out view of an advantageous embodiment of (part of) a first injector arrangement useful in the present disclosure.

[0082] FIG. 3 illustrates the zones within a blast furnace and effect of lower shaft injection of reducing gas B on the rising gas passing through the coke layers within the cohesive zone.

[0083] Further details and advantages of the present disclosure will be apparent from the following detailed description.

DETAILED DESCRIPTION OF THE DRAWINGS

General Considerations

[0084] The main issue of the blast furnace with regard to CO.sub.2 emissions is that its process is based on the reduction of iron ore based on coke and also on the injection of carbon rich auxiliary fuel/reductant at tuyere level.

[0085] The injection of a syngas at shaft or even at tuyere level has been proposed to help to decrease the coke consumption in the blast furnace and therefore to reduce the CO.sub.2 emissions. Nevertheless, even under these conditions, there is still a requirement for a non-negligible lump coke rate, generally still above 300 kg/t HM. This requirement of a still substantial coke rate comes mainly from process and to a minor extend from mechanical/fluid dynamic reasons.

[0086] Today, the blast furnace seems more and more overtaken by alternative process routes which need only small amounts of coke. In fact, these processes divide the blast furnace in two: an iron ore reduction part, performed in a shaft furnace, and a melting part, done in an electric melting furnace, the latter using a kind of electric arc/plasma or immersed electrodes. This concept however results in very big and complex installations which are very costly CAPEX-wise and also difficult to maintain. Also, the melting furnace still requires a significant amount of coke, about 60 to 110 kg/t HM in case of utilization of hydrogen-reduced direct reduced iron (DRI). This carbon is required to finalize the reduction of the DRI and to carburize the hot metal that will then be processed in a basic oxygen furnace.

[0087] The present method aims at being able to show that the widely used blast furnace technology, if converted appropriately as described herein, is able to provide solutions with low CO.sub.2 emissions and thereby allowing to use the (modified) blast furnace with a very low coke rate by improving the process conditions in the blast furnace.

[0088] In fact, the conventional blast furnace today uses coke for different requirements: [0089] Indirect reduction of the iron ore in the shaft:

##STR00001## [0090] Direct reduction of the iron oxide in the lower part of the blast furnace, reducing the molten iron oxide directly as: FeO+C.fwdarw.Fe+CO [0091] Solution loss: at a certain level of temperature the coke reacts with carbon dioxide and H.sub.2O to generate carbon monoxide and H.sub.2:

##STR00002## [0092] Carburization of the molten iron. The iron saturated with carbon at typical operating temperatures of the blast furnace at around 1500 C. has about 4.5% dissolved carbon. This is also the prerequisite for the blast furnace, since it allows the hot metal to be liquid at this temperature level. [0093] Burning of coke at the tuyere to generate the hot reducing gas required to melt and to reduce the iron ore. The hot reducing gas should reach the minimum flame temperature required for the operation of the blast furnace (1800-2600 C.). In fact, the flame temperature assures the heat transfer from the gas phase to the ore and its subsequent melting. If that temperature is too low, the blast furnace cannot achieve a high production. [0094] Due to chemical equilibria, approximately half of the reducing gas generated in the blast furnace is exiting at the top (top gas) without reacting. The remaining calorific value of the top gas is generally used to for heating the hot blast and are exported out of the blast furnace plant.

[0095] In a standard blast furnace application today, with a coke rate of 310 kg/t HM and 180 kg/t HM of injected pulverized coal (PCI), the coke consumption for the different factors is distributes as follows, see Table 1.

TABLE-US-00002 TABLE 1 In kg/t HM Case 0 PCI gasified/burned at tuyere 180 Coke gasified/burned at tuyere 120 Coke solution loss shaft 28 Coke for HM solution 47 Coke for direct reduction 114 Total coke 310 Note: approx. values only

[0096] The question is now how to diminish that coke rate?

[0097] The coke dissolved in the hot metal cannot really be reduced since the coke will dissolve in the hot metal until its saturation. The saturation of hot metal by carbon ensures also a long life for the carbon blocs in the hearth. It is therefore in principle not possible to diminish these 47 kg/t HM (see Table 1).

[0098] The necessary coke for the direct reduction in the lower part of the blast furnace could be heavily reduced if the reduction degree of the ore coming from the upper part of the blast furnace is increased. Increasing the reduction degree could be done by increasing the reductant to oxidant ratio (CO+H.sub.2)/(CO.sub.2+H.sub.2O) in the gas phase in the blast furnace shaft. This can be done by increasing the reducing gas flow, for example by injecting reducing gas at the bottom of the shaft. Another option would be by increasing the accumulated concentration of H.sub.2+CO mainly by decreasing the N.sub.2 content in the injected gas at tuyere level and at shaft level.

[0099] The necessary coke for the solution loss can be reduced as well by increasing the reductants to oxidants ratio in the gas passing through the shaft of the blast furnace. This can be done by increasing the reducing gas flow, for example by injecting reducing gas at the bottom of the shaft. Again, it could also be done by increasing the accumulated concentration of H.sub.2+CO such as by decreasing the N.sub.2 content in the injected gas at tuyere level and at shaft level.

[0100] The necessary energy for the lower part of the blast furnace is generated at the tuyere level where coke and auxiliary fuel like PCI or natural gas are burned and gasified with hot blast (between 90 and 1300 C.) and produce the tuyere gas at temperature of between 180 and 2500 C. A high amount of fuel needs to be burned for that purpose since it oxidizes only to CO and not to CO.sub.2. This results in a much lower release of reaction heat, thus a much lower flame temperature as when burning the fuel completely to CO.sub.2 and H.sub.2O as it can be done in an oxidizing atmosphere.

[0101] What the present disclosure proposes to do:

[0102] In the present disclosure, it is proposed to combine both types of injection, the utilization of a reducing gas in the blast furnace injected both at the bottom of the shaft and at the tuyere level, in which the reducing gas injected at the tuyere level is further heated up to a temperature similar or as close as possible to the usual blast furnace flame temperature. Thus, the present disclosure considers heating the tuyere reducing gas, e.g. syngas, to at least 1600 C., preferably above 1800 C., more preferably above 2000 C. and possibly even over 2300 C.

[0103] When injecting the hot reducing gas at shaft and tuyere level, the present method completely avoids a priori the requirement of oxygen containing hot blast. However, it is still possible to maintain a small amount of injected cold oxygen, such as below 120 Nm.sup.3/t HM, preferably below 112 Nm.sup.3/t HM, together or not with some auxiliary fuel as natural gas, coke oven gas, pulverized coal, etc., at the tuyere if desirable or deemed helpful. Usually, minimum injection rates for oxygen at tuyere level are above about 4 Nm.sup.3/t HM, such as above about 12 Nm.sup.3/t HM, above about 25 Nm.sup.3/t HM, e.g. above about 35 Nm.sup.3/t HM, or even above about 50 Nm.sup.3/t HM.

[0104] For heating the (first) reducing gas, different methods can be used to efficiently heat the reducing gas, e.g. from syngas production temperature (about 950 C. in case of catalytic reforming processes and partial oxidation/autothermal reforming, around 1200 C. in case of non-catalytic reforming processes, preheated or non-preheated temperature in case of (V)PSA (0 to 1300 C.) to the desired/required temperature. Technologies useful to that end are known. Advantageously, electric energy can be used for that purpose, especially green or renewable electric energy, e.g. through resistance heating and/or plasma technology.

[0105] Furthermore, the reducing gas can be produced with any appropriate method, such as in a CO.sub.2 removal plant, e.g. an amine type adsorber, PSA or VPSA starting from blast furnace gas, basic oxygen furnace gas or others in which the CO.sub.2 separation might be increased by using sorption enhanced water gas shift reactor. The reducing gas could also be the hydrogen-rich stream of a hydrogen removal plant.

[0106] Furthermore, it also explicitly foreseen to use pure hydrogen gas as a reducing gas or a mixture of N.sub.2 and H.sub.2 resulting from the cracking of ammonia (or even the cracking of methanol and/or ethanol) as mentioned above.

[0107] In a method as disclosed herein, it is in principle possible to reach coke rates of about 100 kg/t HM by almost completely eliminating the nitrogen in the gas passing through the blast furnace and by injecting the gas in the blast furnace at about the flame temperature.

[0108] Alternatively to the heating of the syngas injected at the tuyere, it is also possible to supply the electric power directly to the blast furnace, for example, by using a plasma torch, electrodes directly immerged in the blast furnace, by inductive heating, etc.

[0109] It is of particular benefit to reprocess part of the blast furnace top gas in a reformer, PSA, VPSA, CO.sub.2-removal plant, etc. for producing a reducing gas having an adequate reductant to oxidant ratio. The present method considers various processes for producing reducing gas and injecting it, also together with hydrogen, and/or cracked ammonia in the blast furnace at shaft and also tuyere level.

[0110] It is also considered to use part of the blast furnace top gas as heating fuel for the syngas production. Only a very small part of the blast furnace top gas could leave the blast furnace plant, which is an indication of high efficiency.

[0111] For permeability reasons, coke rates below 90 kg/t HM might not be reachable, but the present method allows reaching coke rates significantly below 200 kg/t HM.

[0112] As explained above, the permeability issue is the main reason why part of the reducing gas is not injected at tuyere level, but rather at the bottom of the shaft of the furnace. Indeed, decreasing the coke rate to minimum levels will decrease the passing area for the ascending gas in the cohesive zone. This raises the gas velocity and pressure drop in the cohesive zone and the upper dripping zone and which might also lead to abnormality in the distribution of gas radially and circumferentially. It is therefore very important to reduce the gas volume in the tuyere and bosh area, which is achieved by partially injecting the reducing gas at shaft level. It is also advantageous to decrease the density and viscosity of the reducing gas. This can be reached by the use of the hydrogen rich reducing gas, additionally enriched or not with hydrogen, which has a very low density and viscosity, which in turn results in less pressure drop and lower gas velocity.

[0113] The present method will still use some coke and may require e.g. methane for reforming of the recycled top gas. Additional CO.sub.2 savings can be achieved when producing the natural gas with help of the methanation of CO and/or CO.sub.2 containing gases such as the top gas of the blast furnace and/or the gas of the basic oxygen furnace. As an alternative to natural gas, methane produced by methanation can also be employed to further reduce the CO.sub.2 footprint.

[0114] Finally, the present method may also comprise the capture of the CO.sub.2 from waste gas of the reformers/heaters or any other steel plant process gases for usage or storage to further reduce the overall CO.sub.2 footprint.

Case Study: Reducing Gas Utilization in a Blast Furnace

[0115] To better illustrate the disclosure, several cases have been investigated in more depth. The details of the cases are presented in Table 2 below.

[0116] Case 1 and 2 are the blast furnaces running today with coarse coke rate of 255 kg/t HM and 205 kg/t HM respectively.

[0117] Case 3 and 4 are dedicated for the blast furnace with hot H.sub.2 injected to tuyere without and with hot hydrogen injection to the shaft respectively.

[0118] Cases 5 and 6 are about the new furnaces with about 180 kg/t HM and about 110 kg/t HM using a superheated syngas injection at tuyere level.

[0119] Cases 7 and 8 are about the new furnaces with about 180 kg/t HM and about 100 kg/t HM using a superheated syngas injection at tuyere level, and addition to that, an injection of a hot syngas into the shaft.

[0120] Case 1 represents a typical well operated blast furnace with high pulverized coal injection (PCI) as one can find several all over the world and specifically in Europe. The pulverized coal injection allows to decrease the lump or coarse coke rate to values of about 250 kg/t HM.

[0121] As can be seen, this furnace requires about 235 Nm.sup.3 of oxygen per metric t of produced hot metal (t HM). Part of this oxygen comes with the heated air, part of it comes from oxygen enrichment with pure oxygen coming from an air separation plant.

[0122] This oxygen enrichment is required since the blast furnace requires a certain flame temperature in order to operate correctly. In fact, the flame temperature assures that the reduced ore can be molten and that the injected coal can be burnt within the raceway. With high PCI, the required flame temperature is about 2200 C.

[0123] In this case, it is not possible to eliminate the natural blast completely, since a certain gas volume is needed in order to have enough latent heat required to supply the energy for the heating and reduction of the ore in the shaft of the furnace.

[0124] Case 2 represents a typical blast furnace reaching exceptionally high PCI injection. Such operations have already been sustained for long production periods of several months. This operation is however quite challenging and requires high operational skills and very good raw materials, in particular, high quality of costly cokes. In this case it is important to note that a coarse coke rate of 205 kg/t HM has already been reached.

[0125] If one compares the injection conditions at the tuyere, one can see that the oxygen enrichment had to be further increased, but altogether the conditions at the tuyere, hot blast temperature, gas volume flow rate, have not drastically changed.

[0126] Both these operations are of course not desirable if one wants to reduce the CO.sub.2 emissions from the hot metal production since the complete energy and reductant input is based on coal.

[0127] Cases 3 and 4: In these cases, the inventors have analyzed the operation of a blast furnace using high hydrogen injection rates in order to reduce part of the energy and reductant input from coal and use CO.sub.2 free hydrogen instead.

[0128] It is known that hydrogen cannot be injected in big quantities if injected cold at the tuyere together with oxygen enriched hot blast. The typical maximum amount of hydrogen utilization is restricted to below 30 kg/t HM.

[0129] Thus, the addition of hydrogen will require a change. It needs to be injected hot, only at tuyere level or at both tuyere and shaft level.

[0130] As injection temperature the inventors have assumed 950 C. at shaft level and 1200 C. at tuyere level. These are the temperature levels one can typically reach when using traditional heat exchangers and regenerative type heaters.

[0131] At tuyere level cold oxygen now needs to be added in order to burn some coke for reaching the required flame temperature for the melting of the ore. From furnaces which are not using PCI injection, it is known that in this case the flame temperature can be reduced to a minimum of about 1800 C., always requiring a sufficient flow rate to supply enough energy for melting the reduced ore and supply the energy for the heating and reduction of the ore in the shaft of the blast furnace.

[0132] As can be seen, it is possible to reduce the oxygen injection at the tuyere from about 240 to 150 and 180 Nm.sup.3/t HM, respectively. Less carbon will thus need to be burned at the tuyere. However, since there is no carbon available from PCI injection, the coke that needs to be burned at the tuyere is higher as in the Cases 1 and 2, leading to an increased coke rate.

[0133] This contradiction of additional coal/coke requirement when using hydrogen in a blast furnace, thus required CO.sub.2 emission reduction, cannot be overcome with conventional methods.

[0134] Additionally, it can be seen that pure H.sub.2 injection would in both cases, with and without shaft injection, require huge amounts of hydrogen, 1070 and 1280 Nm.sup.3 of hydrogen per t of hot metal with and without shaft injection respectively. This is by far exceeding the hydrogen requirement of other production routes such as the direct reduction process requiring about 660 Nm.sup.3/t HM hydrogen. Moreover, the required coke rate in these cases is quite high, making them uninteresting.

[0135] To overcome this problem the inventors are proposing the injection of a syngas with or without pure hydrogen addition, at superheated temperatures to the blast furnace.

[0136] The superheating can be done by plasma torches for example.

[0137] Cases 5 and 6: In these cases, superheating the gas that is injected at tuyere level has the advantage that we now do not need to burn coke with oxygen to have the temperature level at the tuyere required for melting of the reduced ore.

[0138] Cases 7 and 8: These cases have additional syngas injection into the shaft compared with the Cases 5 and 6. The inventors propose to operate the furnace by partly injecting the gas in the shaft of the furnace to be able to supply the required temperature level for melting with a lower electric energy requirement of the plasma torch. Moreover, with the shaft injection, less gas needs to pass through the high-resistance cohesive zone area.

[0139] Contrary to Cases 7 and 8, in Cases 5 and 6 all the reducing gas required for operating needs to go through the tuyeres and thus through the dripping zone and the cohesive zone, which will lead to the problems of flooding and hanging, respectively. Furthermore, in Cases 5 and 6, there is no way to avoid the wall channeling phenomenon described above.

[0140] By operating a blast furnace as described herein, it is now possible to actually reach the very low coke rates presented in Cases 7 and 8, i.e. of about 180 and about 100 kg/t HM, respectively.

[0141] It can also be seen, that the use of pure hydrogen is also very low in case of 100 kg/t HM with 440 Nm.sup.3/t HM.

[0142] This can be reached e.g. by recycling the top gas of the blast furnace back to the blast furnace. For this the content of H.sub.2O and CO.sub.2 needs to be reduced. The H.sub.2O can easily be reduced by cooling and condensation. The CO.sub.2 elimination is somehow more complicated and different methods have been proposed such as PSA, VPSA and also reforming of CO.sub.2 with hydrocarbons to form CO and H.sub.2.

[0143] In the example, a reforming with natural gas has been used.

[0144] Such a reforming can be done catalytically at temperatures of about 950 C. or without catalyst in a regenerative reformer type at elevated temperatures>1100 C. Since specifically at the tuyere level very high temperatures are desirable, the latter type is very well suited for preparing the gas at the tuyere level.

[0145] Also when reforming the gas at high temperatures, it is possible to reach a very high reduction degree of the gas (CO+H.sub.2)/(CO.sub.2+H.sub.2O)>7, preferably >8, more preferably >9. High reduction degrees are very important to reach a low coke rate since every CO.sub.2 and H.sub.2O in the gas will consume coke in the raceway.

TABLE-US-00003 TABLE 2 Case 5 Case 6 Case 4 New furnace New furnace Case 7 Case 8 Case 3 BF with with 180 with 110 New furnace New furnace BF with hot H2 kg/t HM, kg/t HM, with 180 with 100 hot H2 injected without without kg/t HM, kg/t HM, Case1 Case2 injected to tuyere shaft shaft with shaft with shaft BF1 BF2 to tuyere and shaft injection injection injection injection Total coke rate (kg/t HM) 301 256 274 324 180 114 177 102 Lump coke rate (kg/t HM) 255 205 274 324 180 114 177 102 Nut coke rate (kg/t HM) 46 51 0 0 0 0 0 0 PCl (kg/t HM) 192 232 0 0 0 0 0 0 Cold O.sub.2 (Nm.sup.3/t HM) 63 89 148 183 71 1.9 75 6 Natural dry blast volume 821 740 0 0 0 0 0 0 (Nm.sup.3/t HM) Total O.sub.2 (Nm.sup.3/t HM) 235 244 148 183 71 1.9 75 6 Flame temperature ( C.) 2 230 2 175 1 796 2 225 1 876 1 848 2 032 1 997 Top gas volume, dry (Nm.sup.3/t HM) 1392 1337 1285 1209 1005 1011 1084 1205 LHV of top gas, kJ/Nm.sup.3 3574 3673 9781 9496 7809 8 914 7 846 9 065 Reducing gas injected to the 0 0 0 310 0 0 300 565 shaft, (Nm.sup.3/t HM) Reducing gas injected to 0 0 1280 760 947 1200 760 889 tuyere (Nm.sup.3/t HM) Temperature of hot reducing 1250 1150 1200 1200 1750 2020 1820 2150 gas or natural hot blast injected to tuyere, ( C.) Electricity consumption for 0 0 0 0 273 516 242 435 plasma, kWh/t HM

[0146] The volume flow rate of these gases is summarized in the following

TABLE-US-00004 TABLE 3 Case 5 Case 6 Case 4 New furnace New furnace Case 7 Case 8 Case 3 BF with with 180 with 110 New furnace New furnace BF with hot H.sub.2 kg/t HM, kg/t HM, with 180 with 100 hot H.sub.2 injected without without kg/t HM, kg/t HM, Case1 Case2 injected to tuyere shaft shaft with shaft with shaft BF1 BF2 to tuyere and shaft injection injection injection injection Reducing gas into the 0 0 0 310 0 0 300 565 shaft, (Nm.sup.3/t HM) H.sub.2 used for reducing 0 0 0 0 0 440 0 440 gas production (Nm.sup.3/t HM) NG used for reducing gas 0 0 0 0 180 95 191 106 production (Nm.sup.3/t HM) Hot blast injected to 821 740 0 0 0 0 0 0 tuyere (Nm.sup.3/t HM) O.sub.2 injected to tuyeres 63 89 148 183 71 2 75 6 (Nm.sup.3/t HM) Reducing gas injected to 0 0 1280 760 947 1200 760 889 tuyere (Nm.sup.3/t HM) Gas volume at tuyeres, 1 247 1 217 1 538 1 076 1 135 1 240 920 939 i.e. in the raceway (Nm.sup.3/t HM) Gas volume in the bosh 1398 1332 1577 1147 1154 1294 966 969 (Nm.sup.3/t HM)

Coke Consumption

[0147] Coke in the BF is consumed by solution loss reaction in the shaft, carburization, direct reduction in liquid state (upper part of the dripping zone), and coke gasification/combustion in tuyere. Moreover, a little amount of coke fines is discharged into the dust during the charging of the furnace.

[0148] In Table 4 below, the amount of coke consumed by the mentioned means for all cases are presented (coke repartition).

TABLE-US-00005 Case 5 Case 6 Case 4 New furnace New furnace Case 7 Case 8 Case 3 BF with with 180 with 110 New furnace New furnace BF with hot H.sub.2 kg/t HM, kg/t HM, with 180 with 100 hot H.sub.2 injected without without kg/t HM, kg/t HM, Case1 Case2 injected to tuyere shaft shaft with shaft with shaft BF1 BF2 to tuyere and shaft injection injection injection injection Coke burnt/gasified 127 93 171 210 92 24 92 24 at tuyere (kg/t HM) Coke solution loss 22 29 23 11 17 17 13 10 shaft (kg/t HM) Coke for carburization 51 55 51 51 51 51 51 51 (kg/t HM) Coke for direct 91 70 19 40 13 18 14 12 reduction (kg/t HM) Coke in dust and 11 10 11 12 7 4 7 4 sludge (kg/t HM) Total coke with dust 301 257 274 324 180 114 177 102 and sludge (kg/t HM)

[0149] It can be seen that, the coke consumption by direct reduction in liquid state is significantly lower for the new furnace condition. The reason is that ferrous burden is reduced to a very high degree in the shaft, thanks to very high reducing power (H.sub.2+CO) in the furnace. Therefore, less FeO remains to be reduced to Fe by direct reduction and solution loss reaction. For the cases 5 to 8, coke consumption by direct reduction is below 20 kg/t HM whereas for the conventional furnace it is greater than 70 kg/t HM.

[0150] It can also be seen that in the Cases 7 and 8, the injected reducing gas into the shaft limits the coke consumption by the solution loss, consequently less chemical attack on the coke. It is therefore considered that the injection of reducing gas is not resulting in an increase in coke quality demand for the new furnace cases. Furthermore, as superheated reducing gas brings significant amount of thermal energy into the blast furnace, the need for the quantity of coke to be combusted in tuyere decreases. Nevertheless, our reducing gas contains small amount of H.sub.2O or CO.sub.2. These gases are converted to CO and H.sub.2 through the reaction with carbon in the raceway area. Therefore, some coke will be consumed in this area. This can be seen by comparing Cases 7 and 8. A non-negligible part of the coke can be gasified (instead of being burnt) by the injection of the very hot reducing gas, thereby contributing to creating and maintaining the raceway even with very low rates of injected oxygen.

[0151] What has been described here shows how the coke rate can be decreased to a very low level using a superheated reducing gas injected into the tuyere and a hot reducing gas into the shaft level. Only carburization cannot be reduced by the reducing gas.

[0152] Detailed explanation to flooding and pressure drop in cohesive (melting) zone:

[0153] The objective of this description is to show the relative and not absolute change of the flooding phenomena and the pressure drop in the dripping zone as well as the entrance of cohesive zone when changing the operation of a blast furnace with higher coke rate to the operation with lower coke rate:

[0154] For the flooding the following dimensionless parameters are used:

[0155] The dimensionless parameters of the Matsu-ura and Ohno diagram, as well as their variables and units used in the calculations, are presented below as Parameters (1) and (2):

Dimensionless Pressure Drop:

[00001]$\begin{array}{cc}{\left(\frac{P}{L}\right)}_d\left(\frac{1}{g{}_{\text{?}}}\right)& \left(1\right)\end{array}$$\text{?}\text{indicates text missing or illegible when filed}$

[0156] P, gas pressure drop between two points (Pa); L, bed length measuring P.sub.d(m); .sub.l, liquid density (kg/m.sup.3); g, acceleration due to gravity (m/s.sup.2).

Dimensionless Irrigation Density:

[00002]$\begin{array}{cc}{\left(\frac{{}^4}{{}_{\text{?}}^{\text{?}}{g}^{\text{?}}}\right)}^{1/8}\left(\frac{u_{\text{?}\left(1-\text{?}\right)}}{d_{\text{?}}}\right){\left(\cos \frac{}{2}\right)}^2& \left(2\right)\end{array}$$\text{?}\text{indicates text missing or illegible when filed}$

, liquid surface tension (N/m); .sub.l, liquid density (kg/m.sup.3), , liquid viscosity (Pa s); , void fraction (); dp, particle harmonic diameter (m); u, superficial liquid velocity based on empty column (m/s); , contact angle of liquid on solid ().
For all variables the liquid considered was the slag. Since the slag has a much higher viscosity and a much lower density than hot metal, flooding will occur first with slag.

[0157] In the first term of Parameter (1), the dimensionless pressure drop, only the pressure drop P depends on the gas characteristics.

[0158] The second term, the dimensionless irrigation density of Parameter (2) has no parameter that directly depends on the gas characteristics, but only indirectly through the coke particle voidage and the coke particle diameter.

[0159] In fact, increasing the flow rate and concentration of reducing agent (H.sub.2 and CO) in the gas will lead to a reduced coke consumption of the coke for the indirect reduction in the shaft (through the Boudouard reaction) and the direct reduction in the dripping zone. The coke particles in the cohesive zone and more importantly in the dripping zone will thus have a bigger diameter (since being less consumed by the direct reduction reaction) in these areas. Please refer to Table 4 showing the coke repartition for different cases. The bigger diameter will also result in a higher voidage. As a consequence, the dimensionless irrigation density will therefore decrease.

[0160] It is known that at lower values for the dimensionless irrigation density the dimensionless pressure drop can be higher.

[0161] Coming back to the pressure drop and thus the p/L directly depending on the gas phase.

[0162] In fact, that parameter is usually calculated with help of the Ergun equation:

[00003]$\begin{array}{cc}p=\frac{150L}{D_p^2}\frac{{\left(1-\right)}^2}{{}^3}{v}_{\text{?}}+\frac{1.75L}{D_p}\frac{\left(1-\right)}{{}^3}{v}_{\text{?}}.Math.v_{\text{?}}.Math..& \left(3\right)\end{array}$$\text{?}\text{indicates text missing or illegible when filed}$

[0163] The first term of this equation is the laminar term which is in the conditions of the cohesive zone and dripping zone negligible compared to the second term, the turbulent term.

[0164] As a good approximation, it could therefore be written:

[00004]$\begin{array}{cc}\frac{p}{L}\frac{1.75\left(1-\right)v^2}{D_p{}^3}& \left(4\right)\end{array}$

[0165] As already discussed, the particle diameter (Dp) and the void fraction () will be positively influenced by the high flux and concentrations of H.sub.2 and CO.

[0166] One can thus further simplify, always being on the safe side, the relation to:

[00005]$\begin{array}{cc}\frac{p}{L}{v}^2& \left(5\right)\end{array}$

[0167] The worst conditions in case of reduced coke rate will be in the entry from the dripping into the cohesive zone since the free passage of the gas is the lowest here. The velocity of gas going through the cohesive zone can be estimated as the gas volume flow rate divided by the free section in the coke layers in the cohesive zone, thus:

[00006]$\begin{array}{cc}\frac{p}{L}{\left(\frac{\mathrm{volume}\mathrm{flow}\mathrm{rate}}{\mathrm{Free}\mathrm{section}}\right)}^2& \left(6\right)\end{array}$

[0168] One can with good approximation say, especially in case of unchanged coke layer thickness, that the free section is proportional to the coke rate, thus

[00007]$\begin{array}{cc}\frac{p}{L}{\left(\frac{\mathrm{volume}\mathrm{flow}\mathrm{rate}}{\mathrm{coke}\mathrm{rate}}\right)}^2& \left(7\right)\end{array}$

[0169] Embodiments of the present disclosure:

[0170] It is known that blast furnaces have already successfully been operated with a coarse coke rate of below 210 kg/t HM using high pulverized coal injection rates, fed with highly oxygen enriched hot blast (hot air).

[0171] The oxygen consumes/burns the injected coal and coke, thus providing the energy to heat the tuyere gas to the high temperature level in the raceway required to melt the reduced ore.

[0172] In order to reduce the CO.sub.2 emissions coming from the blast furnace we want to eliminate the injected coal and reduce as far as possible the coke rate.

[0173] For this reason, the amount of total oxygen injected to the tuyeres should be decreased to a low level, preferably below 120 Nm.sup.3/t HM, preferably below 112 Nm.sup.3/t HM.

[0174] In order to provide the energy at high temperature level to melt the reduced ore, the tuyere gas volume flow rate can only be slightly decreased, by approximately 20%, if one wants to keep the typical raceway temperatures of 1800 to 2600 C.

[0175] Since the reducing gas is in the present process proposal no longer produced inside the raceway of the blast furnace, or at least only to a small extend, the reducing gas has to be produced outside and supply it to the furnace.

[0176] Table 4 shows that total mass flow injected to the furnace for all studied cases. It is clear that this value is significantly lower for the new furnace compared to the conventional ones, thanks to the quantity of H.sub.2 in the gas. The composition of gas injected to tuyere for all cases is presented in Table 6.

[0177] In order not to exceed the acceptable fluid dynamic conditions in the cohesive zone we claim therefore that this can only be reached when maintaining the mass flow of the reducing gas injected at the tuyere below 800 kg/t HM, such as below 700 kg/t HM (refer to Table 5).

[0178] In order to do so, the hydrogen content of the gas injected at the tuyere level must be higher as 30%. It needs to be mentioned that although the total mass flow in case of pure H.sub.2 (100%) is lower than our cases, the coke is as high as 274 kg/t HM and 324 kg/t HM.

TABLE-US-00006 TABLE 5 Total mass of gas injected to tuyere (N/A: not applicable) Case5 Case6 Case 4 New furnace New furnace Case7 Case8 Case3 BF with with 180 with 110 New furnace New furnace BF with hot H.sub.2 kg/tHM, kg/tHM, with 180 with 100 hot H.sub.2 injected without without kg/tHM, kg/tHM, Case1 Case2 injected to tuyere shaft shaft with shaft with shaft BF1 BF2 to tuyere and shaft injection injection injection injection Density of hot blast injected 1.29 1.29 N/A N/A N/A N/A N/A N/A to tuyere, kg/Nm.sup.3 Volume flow rate of hot blast 820.8 740.2 0 0 0 0 0 0 injected to tuyere, Nm.sup.3/tHM Mass flow rate of hot blast, 1056.8 953.0 0 0 0 0 0 0 kg/tHM Density of reducing gas N/A N/A 0.09 0.09 0.69 0.37 0.69 0.38 injected to tuyere, kg/Nm.sup.3 Volume flow rate of reducing gas 0 0 1280 760 947 1200 760 889 injected to tuyere, Nm.sup.3/tHM Mass flow rate of reducing gas 0 0 114 68 650 443 525 337 injected to tuyere, kg/tHM Density of O.sub.2, kg/Nm.sup.3 1.42 Volume flow rate of cold O.sub.2 62.7 89.2 147.6 182.5 71.2 1.9 75.1 6 injected to tuyere, Nm.sup.3/tHM Mass flow rate of cold O.sub.2 90 127 210.9 260.7 101.7 2.7 107 9 injected to tuyere, kg/tHM Total mass flow of injected 1146 1080 325 329 751 446 632 346 gas to the tuyere, kg/tHM

TABLE-US-00007 TABLE 6 Gas composition of reducing gas injected to the tuyere. Case4 Case5 Case6 Case7 Case 8 BF with New New New New Case3 hot H.sub.2 furnace furnace furnace furnace BF with injected with 180 with 110 with 180 with 100 hot H.sub.2 to kg/tHM, kg/tHM, kg/tHM, kg/tHM, injected tuyere without without with with Case1 Case2 to and shaft shaft shaft shaft BF1 BF2 tuyere shaft injection injection injection injection N.sub.2 vol.-% 0.0 0.0 0.6 0.6 0.5 0.5 CO.sub.2 + H.sub.2O vol.-% 0.0 0.0 4.0 4.1 4.0 4.2 CO vol.-% 0.0 0.0 46.9 20.4 47.4 21.3 H.sub.2 vol.-% 100.0 100.0 45.8 73.8 45.4 72.9 CH.sub.4 vol.-% 0.0 0.0 2.7 1.1 2.7 1.1 O.sub.2 vol.-% 0.0 0.0 0.0 0.0 0.0 0.0

[0179] To estimate the effect of such a lower coke rate on the cohesive zone in terms of pressure drop, the parameters of gas entering the cohesive zone are required. This gas is the results of injected gas into the tuyere, coke combustion (if O.sub.2 injected), coke gasification at tuyere (in case of presence of H.sub.2O and CO.sub.2), and produced CO by the final reduction reaction in the liquid state (FeO(I)+C(s).fwdarw.Fe(I)+CO).

[0180] Iron ores are reduced to nearly 100% in the case of new furnace (cases 7 and 8) by mainly high reducing power of gas injected to the shaft and tuyere, and to a limited amount by the final reduction and coke combustion. Therefore, the volume of gas travels through the cohesive zone is significantly lower for the new furnace.

TABLE-US-00008 TABLE 7 Conditions of bosh gas travels through the cohesive zone (CZ) and estimated pressure drop for all cases. Case 5 Case6 Case 4 New furnace New furnace Case7 Case8 Case3 BF with with 180 with 110 New furnace New furnace BF with hot H.sub.2 kg/tHM, kg/tHM, with 180 with 100 hot H.sub.2 injected without without kg/tHM, kg/tHM, Case1 Case2 injected to tuyere shaft shaft with shaft with shaft BF1 BF2 to tuyere and shaft injection injection injection injection N.sub.2 vol.-% 46.55 44.22 0.0 0.0 0.0 0.6 0.0 0.4 CO.sub.2 vol.-% 0 0 0.0 0.0 0.0 0.0 0.0 0.0 CO vol.-% 46.13 47.08 18 35.0 55.0 24.6 57 25.7 H.sub.2 vol.-% 7.22 8.7 82 65.0 45.0 74.8 43 73.8 H.sub.2O vol.-% 0 0 0.0 0.0 0.0 0.0 0.0 0.0 Gas volume flow rate NM.sup.3/tHM 1398 1332 1577 1147 1154 1294 966 969 of bosh gas Velocity in cohesive m/tHM 181.8 214.9 190.6 117.6 212.3 377.8 181.1 316.0 zone Density kg/Nm.sup.3 1.16 1.15 0.3 0.50 0.73 0.38 0.75 0.39 Mass flow rate of kg/tHM 1628 1530 486 569 839 495 731 381 bosh gas Passing area of gas m.sup.2 14.7 11.9 15.8 18.7 10.4 6.5 10.2 5.9 Pressure drop in CZ kPa/m 25.5 34.7 7.4 4.5 21.7 36.1 16.4 26.0

[0181] To estimate the pressure drop over the CZ, the inventors have assumed that the ore layers are impenetrable. Moreover, it is believed that some part of molten slag penetrates into the coke layer in the CZ and thereby clogs part of its height. Hence, this coke-ore interface layer is also considered as an impenetrable thickness. Thus, all ascending gas can only go through the remaining thickness of CZ coke layers (total coke layer heightthe height of the coke-ore interface layer)

[0182] Hence, the thickness of coke layer excluding the interface layers is generally about 15 cm for running BFs (case 1 and 2). The thickness of coke-ore interface layers is assumed to be about 4 cm. As the (effective) coke layer thickness is an important factor for stable furnace operation, this value is preferably kept unchanged for all cases.

[0183] As mentioned earlier the pressure drop is a function of density and velocity (volume flow rate and composition) of the ascending gas.

[0184] As can be seen from the Table 4, the pressure drop in the CZ for Case 2 is quite higher as in Case 1 due to the lower coke rate.

[0185] New furnace for coke rate of 180 kg/t HM can be well operated as the bosh gas can travel through the CZ much smoother. Nevertheless, to decrease the coke rate to the ultimate value of 100 kg/t HM without shaft injection seems to be challenging as the pressure drop in CZ is quite high. This means that the risk of flooding is high for the Case 6.

[0186] This issue can be overcome by increasing the pressure of the blast. Higher blast pressure increases the density and lowers the volume of gas coming to the furnace. The influence of velocity on the pressure drop is much higher than density since the pressure drop is related to the square of the velocity p/Lv.sup.2 (see Eq. (5)). Nevertheless, higher pressure requires more expensive mechanical devices as well as higher electricity energy demand.

[0187] However, this challenge can be mitigated by shaft injection which leads to a lower volume flow rate of injected gas into the tuyere resulting in a lower pressure drop. It can be clearly seen that the pressure drop in Case 8 (new furnace) is even lower than Case 2.

[0188] Furthermore, as the pressure drop in the CZ and dripping zone is lower for the new furnace, this allows an increase of the production rate. To achieve the level of pressure drop of the conventional BF (Case 2), the production rate could be increased by 15%. Thus, once again the advantage of the shaft injection is proved in the Case 8.

[0189] The following Table 8 shows the pressure drop for Case 2 compared to Case 8 with lower and higher production rates.

TABLE-US-00009 Case8 New New furnace furnace with 100 with 100 kg/tHM, with kg/tHM, shaft injection, Case 2 with shaft more BF2 injection production Production rate tHM/h 299 300 344 Gas volume flow NM.sup.3/tHM 1332 969 967 rate of bosh gas Velocity in m/t HM 214.9 316.0 312.7 cohesive zone Density kg/Nm.sup.3 1.15 0.39 0.41 Mass flow rate kg/tHM 1530 381 392 of bosh gas Passing area of gas m.sup.2 14.7 5.9 5.9 Pressure drop in CZ kPa/m 34.7 26.0 34.3

Electric Heating

[0190] Electrically driven heaters can be employed for highly efficient heating of the (first) reducing gas and attain the temperatures required into the smelting furnace. Advantageously, gas in plasma state can be used as a heating means.

[0191] Electrode-based plasma torches can serve as electrically driven heaters. The plasma is ignited on the surface of at least two electrodes. The electrodes can be made of graphite. At least one of the electrodes is in high potential and at least one of the electrodes is in lower potential. Direct current and 3-phase AC plasma torches are electrode-based plasma torches.

[0192] Besides electrode-based plasma torches, there are also electrodeless plasma torches. In the latter, the plasma is inductively ignited, thus no electrodes are needed. Classical electrodeless plasma torches are microwave (MW) plasmas and radiofrequency (RF) plasma. RF plasmas are usually referred as inductively coupled plasmas (ICP).

[0193] Direct current (DC) plasma torches have low volumetric footprint.

[0194] Alternating current (AC) plasma torches, particularly 3-phase alternating current plasma torches have higher volumetric footprint, but feature other advantages: the plasma is confined to the ignition area facilitating a better control of the position of the plasma. Due to the alternating current, thereby, periodic operating mode, the electrodes are naturally cooled which limits the electrode erosion and prolongs their lifetime. Alternating current plasma torches do not require either swirl gas or magnetic coils for plasma stabilization, like direct current plasma torches, thus, the design of the torch is less complicated.

[0195] Direct current plasma torch can comprise 2 or a multiplicity of 2 electrodes (e.g. 2, 4, 6, 8, 12, 14, 16, 18, 20). A higher number of electrodes can promote higher controllability of the plasma and increase the plasma torch power.

[0196] Alternating current plasma torch, particularly the 3-phase alternating current plasma torch can comprise 3 or a multiplicity of 3 electrodes (e.g. 3, 6, 9, 12, 15, 18, 21 electrodes or more). A higher number of electrodes can promote higher controllability of the plasma and increase the plasma torch power.

[0197] Plasma torches, especially direct current plasma torches or/and alternating current plasma and/or 3-phase alternating current plasma torches of 1-10 MW, preferably 2-6 MW, most preferably 4 to 5 MW power, may be employed in the furnace as defined above.

[0198] A plasma torch can be integrated with one or more gas injectors.

[0199] Preferably, one power supply powers 1-10, 1-5, 1-3 or 1 of the at least one plasma torches.

[0200] To allow for a stable plasma operation, the power supply might be designed with a reserve capacity. The reserve capacity may be as high as 3 times, preferably 2 times or even 1.5 times the plasma power needed for operating the plasma torches. Use of multiple power supplies of <30, <20 and/or <10 MW to power multiple plasma torches (one power supply drives one or more plasma torches) is more preferable than using one unique power supply of total capacity equal to or higher than the summation of the capacities of all low-capacity power supplies. The flickering imposed to the grid by plasma torch operation fluctuation in the former case is way less severe than in the latter. Moreover, using multiple power supplies of <30, 20 or 10 MW gives better controllability and flexibility: when a power supply fluctuates or crashes, there is no disturbance propagation to the whole system and the furnace may keep running without severe issues since the other power supplies can collectively provide the power input of the crashed power supply. In addition, one or more spare power supplies can be installed online and in case of failure of a main power supply, a switching over to the spare one assures steady-state operation at full power. During the next planned maintenance shift, the repair can be done, thereby, a production loss never occurs. Yet, high number of power supplies leads to high cost, high power losses and high space requirements, thus, an optimum number of power units should be selected.

[0201] In alternating current plasma torch, particularly in 3-phase alternating current plasma torch, the electrodes can be rod-shaped. When the plasma is on, the side of the electrode placed in plasma ignition region gets eroded and the length of the electrode is shortened. The plasma torch is equipped with a mechanism that ensures that the inter-electrode gap at plasma ignition region is maintained constant by moving the electrodes inwards, towards the plasma ignition region. After a predefined minimum length, the mechanism can exchange or amend the electrodes. In embodiments, the inter-electrode gap may be adjusted between 0 and 100 mm, such as e.g. between 5 and 50 mm or about 40 mm.

[0202] Amending the electrodes is attained by connecting, for example by screwing a new electrode at the backside of the old/used electrode (opposite the side of the electrode placed in plasma ignition region). The attachment of the new electrode to the old electrode is done by an electrode gripping device while the torch is in operation.

[0203] Another possibility for amending the electrodes may be using a carbon-based paste, when plasma operation caused the graphite electrodes to gradually erode and get consumed. The carbon paste, typically a mixture of graphite and binders, may be applied to the consumed or worn-out portions of the electrodes. This paste replenishes the carbon content, extends the electrode's life, and helps maintain efficient electrical conductivity and heat transfer during the plasma operation. The addition of carbon paste to amend graphite electrodes may typically be done using a device designed to amend electrodes without interrupting the plasma process, by bringing the electrode paste column onto the consumed portion of the graphite electrode. The device containing the carbon-based paste gradually pushes the paste into the electrode column, to fill the voids left by the consumed electrode material and restores the carbon content, allowing for a controlled and precise addition of carbon paste, ensuring that the electrode's performance and efficiency are maintained throughout the plasma operation, thereby helping to extend the electrode's operational life and contributing to the overall effectiveness of the plasma operation.

[0204] In direct current plasma torch, the electrodes may have a tubular shape. During operation, the side of the tube at the plasma ignition region erode and the thickness of the electrode is shortened. A plasma torch electrode exchanging device replaces the torn electrodes to new electrodes into the plasma torch. The replacement of the used electrode with the new electrode is achieved by a further electrode gripping device.

[0205] The mechanism of electrode replacing system can be equipped with a magazine for new electrodes and a magazine for used electrodes, so the plasma torch can operate without manual interaction for a long period.

[0206] In DC plasma torch, the plasma is controlled by tuning the electrical operating parameters. In 3-phase AC plasma torch, the plasma control is based on a camera that continuously captures pictures of the plasma. The plasma characteristics depicted in those pictures (i.e. luminosity, shape, diffusivity etc.) are benchmarked against plasma pictures that depict characteristics of plasma at steady state, using a relevant software. Then, changes are imposed (i.e. inter-electrode gap, voltage amplitude etc.) in order to retain the plasma in stable regime.

[0207] In some embodiments, the gas velocity across a plasma heater comprising a plasma torch may be between 10 and 120 m/s, preferable between 20 and 50 m/s. The incoming gas flow arriving to the plasma heater may be split into at least two streams, a first stream flowing centrally through the arc created by the plasma torch and a second stream flowing peripherally around the arc. Both streams may substantially identical velocities and/or flow dynamics. It is however preferred that the two streams have different velocities and flow dynamics, such as e.g. the central flow crossing the arc being preferably a flow of low velocity (5-60 m/s) with turbulences reduced as much as possible whereas the peripheral flow is preferably a flow of high velocity (30-200 m/s) and may be arranged as a vortex flow

[0208] In some preferred embodiments, the central stream flows through the plasma arc and is heated to very high temperature. The second stream is injected shortly downstream of the plasma arc and is orientated in such way that it acts as a protection of the refractory lined walls of the plasma burner chamber of the plasma torch from the heat of the central, high temperature, stream, thereby advantageously reducing the thermal load on the wall lining.

DETAILED DESCRIPTION OF EMBODIMENTS

[0209] FIG. 1 present an advantageous embodiment of a smelting furnace installation according to some aspects of the disclosure, such as a blast furnace installation modified to operate according to the disclosure.

[0210] The (modified) blast furnace installation comprises a blast furnace 10 with its conventional tuyeres and tuyere stocks 21 fed by a bustle pipe 20. However, unlike conventional blast furnaces, the bustle pipe 20 provides a first reducing gas A through the tuyere stocks 21 and the tuyeres to the blast furnace 10. Furthermore, a second reducing gas B is fed through a number of injectors (not shown) at shaft level, said injectors being fed by ring pipe 25.

[0211] The first reducing gas A and the second reducing gas B preferably comprise a reforming gas obtained by reforming the blast furnace's top gas D or other gases from the metallurgical plant. Optionally, H.sub.2 can be added to the second reducing gas B either to specifically increase the second reducing gas' hydrogen content or for adapting its temperature, such as for cooling it to the desired temperature. Reforming may be done in any appropriate reformer 40 or device operated as such, e.g. in regenerative heat exchangers as illustrated in FIG. 1, in the presence of H.sub.2O and/or natural gas (NG) or any other appropriate source. Hydrogen can be added to reformer 40 if desired.

[0212] Prior to reforming, top gas D is preferably cleaned in top gas cleaner 50 to eliminate any solid particles and thereby provide cleaned top gas C. If desired, top gas D or cleaned top gas C can be submitted to further treatment steps before entering the reformer 40 as for example (partial) removal of specific components such as H.sub.2O, chlorines, heavy metals, sulphur components such as COS and the like. Cleaned top gas C can also be directly (i.e. unreformed) added to the second reducing gas B.

[0213] The first reducing gas A is heated to temperatures above 1600 C. prior to injection through the tuyeres via tuyere stocks 21 into the blast furnace 10. This (additional) heating (or part of this heating) can be done through (resistance) heating device 30 before entering the bustle pipe 20. It is however preferred to heat the first reducing gas A to its desired temperature only shortly before its injection in the blast furnace 10, e.g. by plasma torches (not shown) installed on the tuyeres stocks 21, such as on the blow pipes.

[0214] If required or desired, a small amount of oxygen E can be added at tuyere level, such as for creating and maintaining the raceway voids at the tuyeres 21 within the blast furnace 10. The (optional) oxygen E is preferably added at rather low temperatures (compared to the first reducing gas) in order to allow for cooling the tuyeres 122 (FIG. 2).

[0215] FIG. 2 generally presents a possible embodiment of an electric heater useful in the context of the present disclosure. In this embodiment, three electrode plasma torches 1213 (only two are shown due to the cut-out view) are arranged in the blowpipe 1211 of the tuyere stock 121 such that their electrodes 1214 reach within the central duct of the blowpipe 1211, which blowpipe 1211 preferably comprises a refractory lining 1212. In FIG. 2, the tuyere stock 121 is fluidly connected to the tuyere 122 positioned within the wall (not shown) of the smelting furnace, the nose 1221 of the tuyere 122 reaching within the smelting furnace. An oxygen port 1222 is advantageously arranged within the tuyere 122 setup for injecting (preferably cold) oxygen as further disclosed above.

[0216] FIG. 3 shows a schematic (half-)section of a blast furnace in operation comprising stacked coke and iron ore layers fed from the top of the furnace. FIG. 3 also shows detailed views of the layers within the cohesive zone. The first reducing gas A is injected at tuyere level thereby forming a raceway due to the velocity of the gas and direct gasification of coke. The resulting gas rises within the furnace and must pass through the coke layers within the cohesive zone, where the now softened and molten iron (ore) layers in the cohesive zone have become essentially impermeable to gas. The injection of the second reducing gas B pushes the rising gas towards the center of the furnace, thereby largely preventing the wall channeling effect mentioned above.