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1 Feature Article Fundamental Evaluation of Natural Gas Co-Injection With Coke Oven Gas or Coal Blast furnace injection of hydrogen has been studied as a way to lower the carbon footprint of ironmaking. An application of these studies in North America is the co-injection of natural gas with H 2 -bearing fuels. Tests have been performed to compare the combustion of different CH 4 -H 2 - CO gases using a blast furnace blowpipe and tuyere arranged on a special test rig. The results of these tests are presented, and their implication on the design and operation of co-injection systems is discussed. Authors Louis W. Lherbier Jr. technical manager, Raw Materials and Ironmaking, U. S. Steel Research and Technology Center, Munhall, Pa., USA lwlherbier@uss.com Michael F. Riley senior development professional, Metals Processing Technology, Praxair Inc., Indianapolis, Ind., USA mike_riley@praxair.com United States Steel Corporation (U. S. Steel) and Praxair Inc. jointly studied the injection of hydrogen into the blast furnace. A unique feature of the studies has been the testing of CH 4 -H 2 -CO injection into a commercial blast furnace tuyere using a special test stand. The investigation anticipated the need to lower CO 2 emissions from ironmaking and the future availability of low-cost hydrogen production and carbon sequestration technologies. With the rapid increase in shale-derived natural gas (NG), however, a more immediate application of the work to the co-injection of NG with hydrogen-bearing injectants, such as coke oven gas (COG) and high-volatile coal, arose. Extensive operating experience at U. S. Steel has demonstrated that molecular H 2 in COG behaves differently as a blast furnace injectant compared to chemically bonded hydrogen in the methane found in NG. More recent experience has shown that NG co-injected with high-volatile coal behaves differently from either fuel injected separately. Observation of H 2, CH 4 and mixtures in the test tuyere, theoretical analysis of these results, and comparison with observations from operating furnaces suggest that the mixing rate of the injected fuel with the hot blast controls the rate of combustion. When this rate is too fast, it can cause excessive back pressure on the blast bustle pipe and restrict productivity. The mixing rate can similarly affect other performance factors, such as heat loading on the tuyeres and the amount of Boudouard and steam-coke reaction that must be supported in and above the raceway. Possible measures to control this mixing rate are examined for insight into the limits and best modes of injecting these fuels. Test Equipment The test equipment is shown in Figure 1, and a schematic is shown in Figure 2. A 110-mmdiameter blast furnace tuyere, blowpipe and tuyere stock are mounted on a test skid. The back of the tuyere stock is equipped with a specially designed, compact burner fired on NG and oxygen. Oxygen-enriched air is added to the burner combustion products to simulate a hot blast. Simple straight pipe lances are inserted through ports in the side of the blowpipe at an angle of 13 to the blowpipe axis. The discharge end of the lances was This article is available online at AIST.org for 30 days following publication. 292 Iron & Steel Technology A Publication of the Association for Iron & Steel Technology

2 Figure 1 Figure 2 Tuyere test stand. Test tuyere schematic. positioned at the axial centerline of the blowpipe at approximately the joint between the blowpipe and the tuyere, around 250 mm from the tuyere tip. Three different injectants were tested: NG, hydrogen, and a mix of simulated coke oven gas and natural gas (COG-NG). The composition of the NG and simulated COG used in the tests is shown in Table 1, along with the composition of the simulated COG-NG mix. The composition of the hot blast generated by the hot blast burner is shown in Table 2. For comparison, the composition of a similar actual hot blast humidified to 3.6% (12 gr/scf) is also shown. Table 3 summarizes the injection conditions in the tests. Conditions were selected to match the injection and blast velocities of a furnace at 1,065 C blast temperature and 200 kpa (gauge) blast pressure. Three different lances with outlet diameters of 21.2, 30.5 and 38.1 mm were used with each injectant. Results Experimental The most noticeable result of the tuyere testing was the formation of a stable flame at the lance tip at high H 2 levels in the injected gas, compared to an unstable flame formed with NG. Figure 3 shows a series of photos taken from a video recording during the injection of NG and of the COG-NG mixture through the 21.2-mm-diameter lance. Note that the COG-NG injection produces a stable flame (H 2 injection produced a very similar, stable flame). In contrast, the flame from NG injection blows off the lance. A periodic, short-lived flame appears outside the tuyere (indicated by the arrows), accompanied by an audible pop, but no stable flame is seen. When NG is injected at a lower velocity through a 38.1-mmdiameter lance, a stable flame is achieved, also shown in Figure 3. The highest stable velocity observed with Table 1 NG and Simulated COG-NG Compositions, in % NG COG COG-NG H CH C 2 H C 3 H CO N CO NG content of COG-NG 30.0 Table 2 Test Blast and Actual Blast Compositions, in % Test Actual N O H 2 O CO Table 3 Injection Test Conditions NG COG-NG H 2 Blast flow, Nm 3 /hour 1,500 1,500 1,500 Blast temperature, C Blast velocity, m/second Injectant flow, Nm 3 /hour Volume CO+H 2 injected, Nm 3 /hour NG content, % AIST.org May

3 Figure 3 Flames formed in tuyere tests. NG was 34 m/second (with the 38.1-mm lance), and the lowest unstable velocity observed with NG was 52 m/second (with the 30.5-mm lance). This gives a critical blowoff velocity of 43 ±9 m/second. Temperature measurements at the tuyere nose taken by an infrared (IR) thermometer varied strongly with injectant flame stability, as shown in Figure 4. The results with COG-NG and NG using the mm lance are qualitatively similar to those of Rabold, et al., discussed later. 1 Analysis of Blowoff Results Numerous authors describe the blowoff of flames as a competition between the rate of the combustion reaction and the residence time for the combustible mixture. 2 This competition is generally described using a Damkohler number, Da, defined as: Da = t t res chem (Eq. 1) where t res is a residence time and t chem is a chemical time related to reaction rate. Broadwell, et al. used a Lagrangian analysis to develop an expression for the Damkohler number for a fuel jet flowing into a quiescent atmosphere. 3 In their analysis, the residence time is based on the jet velocity and the distance required to fully mix the fuel and oxidant to the stoichiometric ratio. The reaction time is estimated from the thermal diffusivity and the laminar flame speed. Their analysis gives Equation 2 for the Damkohler number, i1 d Da = o ( 1+ f) ( r r ) n 2 1/ 2 o / b S k 2 (Eq. 2) i Broadwell, et al. give the effect of the stoichiometric mass fraction of oxidant as φ 2, but one of the co-authors later corrects that to (1 + φ) 2 in Reference Iron & Steel Technology A Publication of the Association for Iron & Steel Technology

4 where Figure 4 d o is the exit diameter of the fuel lance, φ is the stoichiometric mass ratio of oxidant to fuel, ρ o is the density of the injected fuel, ρ is the density of the surrounding oxidant, v b is the blowoff velocity, S is the laminar flame speed and κ is the thermal diffusivity of the fuel-air mixture. The first ratio on the righthand side of Equation 2 represents the residence time, and the second ratio is the inverse of the chemical time. Broadwell, et al. give the critical value for Da = 4.8. There are two issues in applying Equation 2 to the tuyere tests described here. First, the injectant flows into a rapidly flowing blast, not a quiescent atmosphere. However, the effects of the surrounding flow on jet mixing will affect the jet velocity in a similar fashion, so the impacts on Equation 2 should largely cancel out. The second issue is that while laminar flame speeds are well documented for combustion with room-temperature air, it is difficult to estimate the effects of blast preheat, oxygen enrichment and humidification. It is more convenient to replace Broadwell s chemical time with an ignition time calculated from kinetic models. Making this adjustment to Equation 2 and re-arranging it produces Equation 3: Infrared pyrometer temperatures from tuyere tip for conditions shown in Figure 3. Figure 5 n b d = 2 1/ 2 o ( 1+ f) ( ro / r ) 48. tign (Eq. 3) Calculated ignition time (ms) for NG-H 2 -CO mixtures reacting with test blast air at 830 C. where t ign is the calculated ignition time. Values for the ignition time can be calculated from published kinetic models. The GRI 3.0 mechanism is used here. Figure 5 shows the calculated ignition time for mixtures of NG-H 2 -CO reacting with the hot blast of the test system at 830 C. It is apparent from this figure that the ignition time varies considerably with the NG/H 2 ratio, but is relatively insensitive to CO content. For pure NG, the reaction time is 61 ms. The values used for the other parameters in Equation 3 AIST.org May

5 Table 4 Parameters for Equation 3 ρ o, kg/m ρ, kg/m φ Note: NG at 27 C are shown in Table 4. Figure 6 compares the blowoff velocity calculated from Equation 3 with the experimental data for NG. Agreement is very good, with a predicted blowoff velocity of 40 m/second, coinciding with the experimental value of 43 ±9 m/second. Table 5 Application to Co-Injection of Coke Oven Gas and Natural Gas The blast furnaces at U. S. Steel Mon Valley Works, Edgar Thomson Plant (ET) are equipped to inject NG, COG or mixtures thereof. ET operates two furnaces designated as No. 1 and No. 3. Key design parameters for each furnace are summarized in Table 5. COG injection rates depend on mainly availability, which varies with coke production and with seasonal changes in fuel demand for heating purposes across the Mon Valley Works complex. A combination of NG and COG, or all NG, is injected when COG supplies are limited. A total COG supply of up to 790 Nm 3 /minute has at times been injected. Noticeable increases in blast pressure are typically observed when even small amounts of COG are injected at ET. This effect is illustrated in Figure 7, which plots the average pressure differential across each furnace as a function of the composition of the injection fuel blend. The data in each case are taken from operating periods when each furnace was operating within a fixed range of bosh gas flow. As illustrated, the lowest pressure loss (top pressure less blast pressure) is observed when each furnace injects 100% NG. The pressure loss typically increases as COG is introduced and can increase even more as the proportion of COG in the mixture increases. Each furnace responds in a similar fashion, although the No. 3 blast furnace appears to tolerate Design Parameters for Mon Valley Works and Fairfield Works Blast Furnaces Variable Mon Valley Works Fairfield Works Furnace number Hearth diameter, m Working volume, m 3 1,541 1,380 3,326 Charging method Bell Bell Bell-less Tapholes Tuyeres Tuyere diameter, mm Injectants COG/NG COG/NG Coal/NG Figure 6 certain mixtures of COG and NG more favorably than the No. 1 blast furnace. Other factors can influence the pressure differential, including burden filling practices and raw material quality. Nonetheless, COG use typically yields a higher blast pressure, and this can constrain wind acceptance and productivity. Under the test conditions with flame blowoff discussed above, the combustion of injected fuel is controlled by both the chemical reaction rate and the bulk gas mixing rate. At higher blast temperatures and pressures, mixing alone will control the combustion of injected fuel. Repeating the ignition time calculations for NG-H 2 -CO mixtures reacting with actual blast (composition given in Table 3) at 1,090 C and Comparison of observed blowoff conditions with Equation Iron & Steel Technology A Publication of the Association for Iron & Steel Technology

6 Table 6 Ignition Times, milliseconds Blast composition Test Actual Blast temperature 830 C 1,090 C NG H CO Table 7 Fuel 1 Fuel 2 NG, scfm 7,930 2,882 COG, scfm 0 10,481 NG content, % Volume CO+H 2 injected, scfm 23,758 23, kpa (gauge) gives similar trends as in Figure 5, but the times are much faster. Table 6 compares the calculated ignition times of NG, H 2 and CO with the values from Figure 5. These much faster ignition times suggest that blowoff is unlikely to occur at the higher blast temperatures and pressure, and that the rate of mixing alone will control how quickly heat is released in the blowpipe and tuyere. This, in turn, affects the back pressure on the blast bustle pipe. Excessive back pressure interferes with the flow of blast and can limit furnace productivity. Rabold, et al. measured the heat release within the ET blowpipe and tuyere for the two injected fuels, as shown in Table 7. 1 If mixing controls the combustion of these fuels, then the extent of reaction of each fuel depends on the distance required to mix in the Figure 7 stoichiometric amount of blast. The stoichiometric mixing distance L is given by: ( )( ) L µbd o f + 1 r o / r 1/ 2 (Eq. 4) where β is a function of the relative density and velocity of the jet and the surrounding gas, or what will be termed the density adjusted velocity ratio. Chigier and Beér measured the mixing of a jet in a concentric surrounding flow, and their results show that β varies as shown in Figure 8. 5 The range of relative velocities given by Rabold, et al. is also shown in the figure, and for these conditions, β is essentially constant. Then, for a constant lance diameter, when mixing alone controls combustion, the ratio of the extent of reaction of two fuels is then just the inverse of the ratio of their stoichiometric mixing distances, or: f f 1 2 L2 = = L 1 ( f + 1) r f 12 / / 1+ 1 r 1 ( ) (Eq. 5) Pressure losses for Mon Valley Works blast furnaces as a function of injection fuel composition. The curves for each furnace were developed from operating periods with comparable bosh gas flowrates. where f is the fraction of the fuel reacted and the subscripts 1 and 2 refer to two fuels. Using the compositions in Table 1, the ratio of reaction extent for the fuels in Table 7 is 0.66, according to Equation 5. This compares with reported values in actual furnace operation of 35% reaction for fuel 1 within the blowpipe/tuyere zone and 50% reaction for fuel 2 a ratio of 0.7, in good agreement with the equation. 1 To prevent excessive back pressure on the blast, the extent of reaction between the injectant and the blast must be kept low. The most obvious options are to AIST.org May

7 Figure 8 Value of β as function of density corrected velocity ratio. Figure 9 Maximum value for product βd o. increase tuyere diameter to lower back pressure or to move the lance discharge closer to the tuyere tip to minimize the time for mixing. Both require redesign and replacement of all blowpipes on a furnace, with the associated costs and downtime. It may also be undesirable if flexibility in the COG/NG ratio is needed. Changing the lance dimensions may be an easier and more flexible option. Figure 9 shows how d o and the product β d o vary with the density adjusted velocity ratio. The maximum value of β d o occurs when the density adjusted velocity ratio equals 0.5. This point represents the maximum mixing length for a given set of conditions, and would be the optimum velocity ratio for minimizing reaction in the blowpipe and tuyere. For NG and the COG-NG mix, this occurs at an injection velocity of m/second for a blast velocity of 175 m/ second. Application to Co-Injection of Pulverized Coal and Natural Gas Similar issues with blast pressure have also occurred when coinjecting coal and NG on the No. 8 blast furnace at U. S. Steel Fairfield Works (FFW). Design parameters for No. 8 are summarized in Table 5. The furnace was recently equipped with new blowpipes that were designed to accommodate a second fuel lance for simultaneous injection of coal and NG. The prior single-lance design permitted only coal or NG injection. Since co-injection began, excessive blast pressure has been observed, particularly when hot metal demand was highest. As shown in Figure 10, the furnace operated with a pressure drop of about 200 kpa, while injecting coal and NG at rates of about 80 kg/thm and 55 kg/thm, respectively. High pressure during this period did not have a noticeably adverse effect on furnace stability, but did occasionally limit wind output from the turboblower and furnace productivity. In comparison, a mid-year transition to an all-ng practice resulted in an average pressure loss of 172 kpa, or 16% less. The natural gas rate for this period averaged about 93 kg/t. The average furnace productivity level during each period was comparable, as were filling practices, raw material quality, bosh gas flowrates and lump coke usage. As with COG 298 Iron & Steel Technology A Publication of the Association for Iron & Steel Technology

8 Table 8 Data for Coal-NG Mixing Calculation Coal injection, kg/thm NG injection, kg/thm NG density, kg/m φ, mass oxidant/mass NG Gaseous only volatile Char + volatile density, kg/m φ, mass oxidant/mass char + volatile CO + H 2 generated by NG + volatile, Nm 3 /thm Gaseous + tar volatile Char + volatile density, kg/m φ, mass oxidant/mass char + volatile CO + H 2 generated by NG + volatile, Nm 3 /thm Note: Coal assumed to be 80% C and 5% H 2 by weight Figure 10 injection at ET, the difference in pressure loss with co-injection at FFW is largely attributed to combustion of the injectants in the blowpipe and tuyeres. Simultaneously injecting solid and gaseous fuel streams is a far more complex situation to analyze. A few simplifying assumptions, however, can provide some insight on the potential role of mixing. A relatively wide stream of fine particles tends to act similarly to a gas jet of like density. 6 For a medium- to high-volatile coal, approximately 10% of the particle mass will quickly convert to a volatile gas similar in composition to COG, and another 25% forms volatile tar. 7 These volatiles surround the remaining char and are the first to see and combust with blast air. From this, the mixed coal-ng combustion can then be analyzed similarly to COG-NG injection, except that the mixing rate is controlled by the total mass of NG and coal; but combustion in the blowpipe and tuyere involves only NG and the volatile component of the coal. Lacking specifics on the combustion kinetics of the gaseous and tar components of the volatiles, the actual case can be bracketed by comparing results for the fast-burning gaseous volatiles only with results for the combined gas and tar volatiles. Pressure differential and apparent burden resistance index (BRI) across Fairfield Works No. 8 blast for periods of co-injection and NG-only injection. Burden materials and lump coke rates were similar in all periods. Consider the range of coal-ng injection rates shown in Table 8. In this table, the combination of injectants is selected to provide essentially the same volume of CO+H 2 in each case when the char is included, and lance sizes are adjusted to maintain injection velocities of 150 m/second for NG and 30 m/second for coal. If the composition of the tar is approximated by C 6 H 6, then the parameters for Equation 3 are as shown in the table. The stoichiometric mixing distance becomes shorter with increasing coal injection. AIST.org May

9 Figure 11 Relative gas generation for coal-ng injection. Figure 12 extent of combustion increases with increasing coal until all the NG + volatile is consumed within the tuyere. With further increase in coal injection, less NG + volatile is produced, so even though reaction is complete, less combustion occurs. Figure 11 summarizes this calculation and shows that the amount of reaction in the blowpipe/tuyere region may be higher for coal-ng co-injection than for either NG or coal alone. Thus, coinjection could present problems not seen with single-fuel injection. To control the extent of reaction, the methods described for NG injection can also be used for coal-ng injection. A further tool is to decrease the coal injection nozzle diameter. According to Field, et al., as the diameter of a particle-laden jet decreases, the eddy size decreases, making slippage between gas and particle more likely. 6 This slippage leads to a dampening of turbulence and lowers the mixing rate. The utility of this option may be limited, since it will inhibit dispersion of the coal in the blast. Schematic of relation between mixing rate/combustion rate and injectant velocity/lance diameter. However, the total amount of NG + volatile decreases. Take the Rabold, et al. NG injection case as a basis for comparison, where 30% of the NG combusts in the blowpipe/tuyere region. With co-injection, the Summary and Conclusions Experiments with injecting NG and simulated COG into a commercial tuyere demonstrated the importance of the mixing rate between injected fuels and the hot blast to furnace performance. The mixing rate appears to control the extent of combustion within the blowpipe and tuyere. The mixing rate and the associated combustion rate are affected by the lance diameter and injection velocity according to three regimes, shown schematically in Figure 12. The injection in commercial furnaces discussed here falls into regime 2. Here the injection velocity and blast velocity are similar, and the dominant factor is the lance diameter. A smaller diameter, with a larger velocity, produces a smaller jet with faster mixing and reaction. At some small diameter 300 Iron & Steel Technology A Publication of the Association for Iron & Steel Technology

10 and high velocity, however, the mixing is so rapid that the fuel-blast mixture passes through its combustible range before ignition can occur. This leads to blowoff of the fuel jet and the very low combustion rate of regime 3. At the other extreme, with increasing lance diameter and small injectant velocity, mixing slows, but again there is a lower limit. When the injectant velocity falls well below the blast velocity, the velocity mismatch dominates and increases mixing in regime 1. The minimum mixing for NG and COG in hot blast is about m/second. COG mixes more rapidly than NG does and appears to cause significant back pressure on the bustle pipe and limited wind delivery in industrial practice. Redesign of the blowpipe and tuyere to move injection points closer to the tuyere exit and to increase tuyere diameter is a possible solution, but can be expensive in material and downtime. Changes to the lance diameter could offer a shorter-term option. Jet dynamics suggest that there is an optimum injection velocity for gaseous fuels to minimize mixing. This technique could provide more flexibility in co-injection of NG and COG mixtures. Extending the analysis to coinjection of NG and coal suggests that coal volatiles play a role similar to COG and that the problem could be worse for co-injection than for either fuel injected separately. Modification of existing injection lances according to the ideas presented here is being evaluated for commercial trial. Acknowledgments The authors would like to thank the United States Steel Corporation and Praxair Inc. for permission to publish this manuscript. The work of Bryan Bielec, Robert Churpita and Robert L. Bell in assembling the tuyere test apparatus and conducting the tests is gratefully acknowledged. References 1. C.J. Rabold, et al., Experimental Probing of Temperatures in Blast Furnace Tuyeres, AISTech Conference Proceedings, 2007, Vol. I, pp Q. Zhang, Lean Blowoff Characteristics of Swirling H 2 / CO/CH 4 Flames, Ph.D. dissertation, Georgia Institute of Technology, May 2008, pp J.E. Broadwell, et al., Blowout of Turbulent Diffusion Flames, 20th Symposium (International) on Combustion, The Combustion Institute, 1984, pp W.J.A. Dahm and A.G. Mayman, Blowout Limits of Turbulent Jet Diffusion Flames for Arbitrary Source Conditions, AIAA Journal, Vol. 28, No. 7, July 1990, pp N.A. Chigier and J.M. Beér, The Flow Region Near the Nozzle in Double Concentric Jets, Journal of Basic Engineering, December 1964, pp M.A. Field, et al., Combustion of Pulverized Coal, Cheney & Sons, Banbury, U.K., 1967, pp J.B. Howard, Fundamentals of Coal Pyrolysis and Hydropyrolysis, Chemistry of Coal Utilization, Second Supplementary Volume, M.A. Elliott, ed., Wiley & Sons, New York, N.Y., 1981, pp Disclaimer The material in this paper is intended for general information only. Any use of this material in relation to any specific application should be based on independent examination and verification of its unrestricted availability for such use, and a determination of suitability for the application by professionally qualified personnel. No license under any United States Steel Corporation or Praxair Inc. patents or other proprietary interest is implied by the publication of this paper. Those making use of or relying upon the material assume all risks and liability arising from such use or reliance. F Nominate this paper Did you find this article to be of significant relevance to the advancement of steel technology? If so, please consider nominating it for the AIST Hunt-Kelly Outstanding Paper Award at AIST.org/huntkelly. This paper was presented at AISTech 2013 The Iron & Steel Technology Conference and Exposition, Pittsburgh, Pa., and published in the Conference Proceedings. AIST.org May