PHYSICAL MODELLING OF THE EFFECT OF SLAG AND TOP- BLOWING ON MIXING IN THE AOD PROCESS
Tim Haas1,2, Ville-Valtteri Visuri1, Aki Kärnä1, Erik Isohookana1, Petri Sulasalmi1, Rauf Hürman Eriç3,Herbert Pfeifer2 and Timo Fabritius1
1Process Metallurgy Group, University of Oulu, PO Box 4300, FI–90014 University of Oulu, Finland.
2 Department of Industrial Furnaces and Heat Engineering, RWTH Aachen University, Kopernikusstraße 10, D–52074 Aachen, Germany.
3 Department of Materials Science and Engineering, Aalto University, PO Box 16200, FI–00076 Aalto, Finland.
Keywords: AOD, mixing time, slag, top-blowing Abstract
The argon-oxygen decarburization (AOD) process is the most common process for refining stainless steel. High blowing rates and the resulting efficient mixing of the steel bath are characteristic of the AOD process. In this work, a 1:9-scale physical model was used to study mixing in a 150 t AOD vessel. Water, air and rapeseed oil were used to represent steel, argon and slag, respectively, while the dynamic similarity with the actual converter was maintained using the modified Froude number and the momentum number. Employing sulfuric acid as a tracer, the mixing times were determined on the basis of pH measurements according to the 97.5%
criterion. The gas blowing rate and slag-steel volume ratio were varied in order to study their effect on the mixing time. The effect of top-blowing was also investigated. The results suggest that mixing time decreases as the modified Froude number of the tuyères increases and that the presence of a slag layer increases the mixing time. Furthermore, top-blowing was found to increase the mixing time both with and without the slag layer.
Introduction
The argon-oxygen decarburization (AOD) process is nowadays the most common process for refining stainless steel. Tuyères, which are mounted along the sidewall of the vessel, are used to inject O2-N2 and O2-Ar gas mixtures. Modern vessels are often equipped with a top lance in order to expedite decarburization in the high-carbon region. The advantage of the process is that the high gas injection rates provide violent stirring and thus good preconditions for a high decarburization rate, while the gradual dilution of the gas mixture prevents the excessive oxidation of valuable alloying elements.
Mixing in the AOD process has been the subject of numerous physical modelling studies, which have focused mostly on the effect of tuyères and their arrangement on the mixing time [1, 2, 3].
During the process, the viscosity of the slag varies as its composition changes due to chemical reactions. However, the effect of slag on the mixing has received relatively little attention, especially in connection with combined top- and side-blowing. Therefore, the aim of this work was to study the effect of a slag layer, top-blowing and the modified Froude number on the
Advances in Molten Slags, Fluxes, and Salts: Proceedings of The 10th International Conference on Molten Slags, Fluxes and Salts (MOLTEN16) Edited by: Ramana G. Reddy, Pinakin Chaubal, P. Chris Pistorius, and Uday Pal TMS (The Minerals, Metals & Materials Society), 2016
mixing time for both the reduction stage and the combined top- and side-blowing decarburization stage.
Materials and Methods
Experimental Setup
The experiments were performed in an acrylic glass model of a 150 t AOD converter. The employed physical model was built to geometric similarity with the actual converter in a 1:9 ratio and features seven tuyères along the sidewall of the vessel. Owing to its similar kinematic viscosity, water was used to represent the liquid steel bath, while rapeseed oil was used to simulate the slag phase. The water was deionized to avoid reactions between impurities and added acid. Pressurized air was used to represent the argon-oxygen mixtures used in the AOD process. Table I shows the employed physical properties and parameters.
Table I. Employed properties and parameters.
Parameter Unit Converter Model 1:9
Vessel properties
Vessel diameter mm 4,122 473
Bath height mm 2,400 295
Tuyère properties
Number of tuyères – 7 7
Exit diameter mm 10–14 2
Gas flow rate Nm3/min 60–140 0.14–0.56
Modified Froude number – 750–11,750 750–11,750
Top lance properties
Number of nozzles – 3 3
Exit diameter mm 33.9 3
Nozzle inclination angle ° 11 11
Lance height m 1.8 0.28
Gas flow rate Nm3/min 110 0.25
Modified Froude number – 4.26 4.20
Momentum number – 0.153 0.157
Blowing number – – 0.059
Steel/water phase properties
Density kg/m3 7,000 988
Kinematic viscosity m2/s 6.8×10-7 (1700 °C) 10.4×10-7 (20 °C) Slag/oil phase properties
Density kg/m3 2,990 899
Kinematic viscosity (reduction) m2/s 3.3×10-5 5.67×10-5 Kinematic viscosity (decarburization) m2/s 3.3×10-4 5.67×10-5 Dynamic Similarity Criteria
The gas flow rate through the tuyères was calculated by using the modified Froude numberNFr,t, which represents the ratio of the aerodynamic force to the gravitational force [1, 2, 3]:
, = , (1) where is the gas velocity at the tuyère exit (m/s), is the density of the gas (kg/m3), is the gravitational acceleration (m/s2), is the exit diameter of the tuyère (m) and is the density of the liquid phase (kg/m3). In order to obtain dynamic similarity of top lance blowing, four kinds of forces need to be considered: the gravitational force, the aerodynamic force, the inertial force and the viscous force [5]. In the literature, different criteria have been used to determine the lance gas flow rate and the lance height in the model. Often, the modified Froude number is also used [5, 6, 7, 8]. In the literature, lance height [5, 8], cavity depth [7] and lance nozzle diameter [6] have been employed as the characteristic length. In this work, the nozzle exit diameter was used as the characteristic length. The gas velocity at the bath surface ux, which is used to calculate the modified Froude number, is usually calculated by the following approximation formula [9, 10]:
= 0.97
0.07 + 0.29 , (2)
where is the gas velocity at lance exit (m/s), is the lance height (m) and is the radius of the lance nozzle (m). In this work, initial experiments were conducted to study whether this formula is appropriate to predict the gas velocity. Another similarity criteria for top lance blowing is the momentum number Mo, which is the ratio between jet momentum and displaced bath inertia [5]:
=0.7854 × 10 1.27 −1
, (3)
where is the diameter of the lance nozzle (m), is the number of nozzles in the lance, is the atmospheric pressure (bar) and is the supply pressure (bar). Subagyo et al. [11] identified the blowing number as the similarity criteria. The blowing number NB represents the ratio of inertia force to surface tension force:
=2 where = , (4)
where is the critical gas velocity (m/s), is the surface tension (N/m) and is the axial velocity of the gas jet at the bath surface. Originally, was treated as a constant [11]. However, Alam et al. have shown that is a function of the lance height [12] and the nozzle inclination angle [13]. The value ofη differs approximately 300% between lance heights of 250 mm and 900 mm [12]. However, no values or equations are available for a lance height over 900 mm, and for this reason the blowing number could not be calculated accurately for the actual AOD converter. With the these considerations in mind, the gas flow rate through the top lance as well
as the lance height were calculated by using both the modified Froude number and the momentum number, but not the blowing number. In the calculations, the employed three-hole lance was treated as a single-hole lance. This is reasonable as the effects of three-hole top- blowing become negligible at sufficient lance heights. For sake of simplicity, the influence of the slag layer was also neglected. Additional experiments were conducted in order to determine the gas velocityu at the surface of the bath. Therefore, a pitot tube was placed underneath the lance and connected to a manometer. The lance height varied between 2 cm and 30 cm, while the gas flow rate varied from 0.03 Nm/min to 0.275 Nm/min. For each lance height and volume flow rate, ten gas velocity measurements were conducted and the average of these values was calculated. Based on the experiments it was found that Eq. 2 over-predicts the gas velocity considerably. However, Eq. 7 showed reasonably good accordance with the experimental data:
= . , (7)
where is the gas exit velocity,c is the lance factor and is the lance nozzle diameter. The exponent –0.92 was found to be the best fitting one. However, it is should be noted thatc is not constant; it is a function of the volume flow rate through the lance. Based on linear regression, the lance factorc was found to vary between 2.84 and 3.18. The predicted and measured gas jet velocities are shown in Figure 1.
Figure 1. Predicted vs. measured dimensionless gas jet velocity.
Mixing Time Measurements
The experimental set-up of the mixing time measurements is illustrated in Figure 2.The drop in pH was measured with a pH meter using a sampling rate of 20 measurements per second. Ten milliliters of concentrated sulfuric acid (98%) was injected in one go with a pipette 60 seconds after the measurements started. Afterwards, the pH was measured for a further 240 seconds. Due to the decreasing temperature, the pH decreased slightly after the pH drop in some experiments.
Therefore, a temperature-compensating criterion was developed to identify the mixing time. The pH values of the last 180 seconds were taken to calculate a line of best fit. This line was
extrapolated backwards, and the function value of t = 60 s was used as a reference level for the employed 97.5% criterion. As a result, the yield of the experiments could be increased significantly and the variation in the results could be decreased.
Figure 2. Experimental setup.
For every data point, the number of required experiments was calculated using the following equation [14]:
≥ ⁄ , (6)
where is the standard deviation of the mixing times (s), ⁄ is the confidence interval factor and is the mixing time error (s). The errorE was set to one-third of the mixing time. The confidence level was chosen to be 95% and, therefore, zα/2was 1.96.
Results
Mixing Time Measurements (Reduction Stage)
In the first set of experiments, the effect of top slag on the mixing time in the reduction stage was studied without oil, with 2.9 l of oil and with 5.8 l; these values correspond to water-to-oil volume ratios of 0.1 and 0.2, respectively. The flow rates through a single tuyère were 20 l/min, 50 l/min and 80 l/min, which correspond to modified Froude numbers of 750, 4,750 and 11,750, respectively. Table II shows the results for average mixing time as well as the absolute and relative standard deviations. The standard deviations of the measured mixing times were found to be higher when the oil phase was used to represent the top slag.
Table II: Results for mixing time (reduction stage).
Parameter Series A Series B Series C
Oil volume [l] 0 2.9 5.8
Volume ratio 0 0.1 0.2
NFr,t 750 4750 11750 750 4750 11750 750 4750 11750
Average mix. time [s] 16.60 14.94 11.40 29.10 17.88 10.55 29.57 19.00 11.20 Standard deviation [s] 2.29 5.89 2.44 15.19 9.76 4.68 6.04 6.79 4.14 Relative std. dev. 0.14 0.39 0.21 0.52 0.54 0.44 0.20 0.36 0.37
Number of samples 5 9 5 10 12 10 7 6 5
Required samples 1 6 2 10 11 7 2 5 5
As shown in Figure 3, the mixing time decreased as the modified Froude number increased. The mixing time was found to increase along with the slag-steel ratio. However, the effect is different for different modified Froude numbers. While the effect of the steel-slag ratio is relatively high at the smallest modified Froude numbers, it is almost negligible at the highest modified Froude number. Due to the violent stirring, the employed oil layer emulsified into the water bath to a large extent.
Figure 3. Mixing time vs. modified Froude number (left) and vs. slag-steel ratio (right).
Different types of equations were tested to identify a mixing time equation that depends only on dimensionless process parameters. At first, a simple linear equation was tested:
= ⏟ − , + ,
(9)
where , and are fitting parameters, while is the slag-steel volume ratio. The experimental results showed that the effect of slag is smaller at higher modified Froude numbers.
Thus, slightly better accuracy was obtained by adding a cross-correlation term with fitting parameter k:
= ⏟ − , + + , .
(10)
The fitting parameters were determined by using the sum of the least squares method, and their values are shown in Table III.
Table III: Parameters of the mixing time equation for reduction stage.
Parameter y k R2
Value with linear equation 22.1 0.00123 28.0 – 0.8
Value with cross-correlation 18.8 0.00067 60.6 –0.00565 0.89 Mixing Time Measurements (Decarburization Stage)
In the second set of experiments, the physical model was fitted with a three-hole top lance in order to study the mixing behavior during the combined top- and side-blowing decarburization stage. The modified Froude numbers of the sidewall tuyères were again chosen to be 750, 4,750 and 11,750. The effect of a slag layer was studied using 2.9 l of rapeseed oil, which corresponds to a steel-slag volume ratio of 0.1. The experiments were conducted without lance or slag (Series A), with lance but without slag (Series B), and with both lance and slag (Series C). The results of these experiments are shown in Table IV .
Table IV: Results for mixing time (decarburization stage).
Parameter Series A Series B Series C
NFr,l 0 4.2 4.2
Slag volume [l] 0 0 2.9
Volume ratio 0 0 0.1
NFr,t 750 4750 11750 750 4750 11750 750 4750 11750
Average mix time [s] 16.60 14.94 11.40 24.17 15.10 14.70 39.50 21.40 12.00 Standard deviation [s] 2.29 5.89 2.44 8.76 3.53 1.72 6.19 3.84 2.35 Relative std. dev. 0.14 0.39 0.21 0.36 0.23 0.12 0.16 0.18 0.20
Number of samples 5 9 5 6 5 5 5 5 5
Required samples 1 6 2 5 2 1 1 2 2
Figure 4 shows the mixing time during the decarburization stage for the experiments conducted without slag as well as the experiments conducted with a slag phase. The results indicate that the presence of top slag increases the mixing time, although this effect decreases as the modified Froude number of the tuyères increases. In addition, it was found that top-blowing increases the mixing time. This result is in accordance with the experiments conducted by Wei et al. [3], who suggested that mixing is caused primarily by side-blowing, while top-blowing disturbs the flow pattern of the main vortex. The effect decreases slightly with increasing gas flow rates through the tuyères.
Figure 4. Mixing time vs. modified Froude number without (left) and with slag phase (right).
Again, a mixing time equation was developed by using the least squares method:
= ⏟ − , + + , + , + , , ,
(11) where and are fitting parameters. The values of the fitting parameters are given in Table V. It should be noted that the number of data points might not be sufficient to deduct the coherence between the investigated process values and the mixing time correctly.
Table V. Parameters of the mixing time equation for decarburization stage.
Parameter y k t w R2
Value 17.5 0.00057 104.3 –0.0095 2.48 –0.00023 0.92
Discussion
Wuppermann et al. [1] studied the mixing time in a 1:4 water model of a 120 t AOD converter and found that the mixing time decreases as the modified Froude number increases and that there is an ideal bath height-to-diameter ratio of between 0.7 and 0.75. Ternstedt et al. [2] investigated the effect of volume flow rate through the tuyères, bath height and vessel diameter during side- blowing in two models using potassium chloride as a tracer. The results of their experiments suggest that the influence of the bath height is almost negligible and that the mixing time decreases as the volume flow rate increases and with decreasing bath diameter. Wei et al. [3]
examined the influence of the angle between the side-blowing tuyères, the number of tuyères and the gas flow rate for both side- and top-blowing in a 1:3 model of a 120 t AOD converter. Their experiments showed that the mixing time decreased when the tuyère angle was increased from 18° to 22° and when the number of tuyères was decreased from seven to six. Moreover, the mixing time was found to increase with increasing top-blowing rate. In an earlier study, Wei et al. [4] employed a 1:3 model of an 18 t AOD converter and found that rotating jets decreased the mixing time. Table VI concludes the effect of all the parameters investigated so far on the mixing time within an AOD converter.
Table VI: Overview of the effect of certain parameters on the mixing time.
Parameter Effect on mixing time Reference
Modified Froude number (tuyères) Decreasing [1, 2, 3], this work Height-to-diameter ratio of bath Increasing* [1, 2]
Angle between tuyères Decreasing [3]
Number of tuyères Increasing [3]
Top-blowing Increasing [3], this work
Rotating jets Decreasing [4]
Slag/oil layer Increasing this work
Top-blowing with slag/oil layer Increasing this work
* The ideal bath height-to-diameter ratio was between 0.7 and 0.75.
There are certain points that need to be considered when interpreting the results. The viscosity of the AOD slag changes significantly during the process as solid phases might occur due to changing composition. Therefore, the viscosity values given in Table I should be understood as a snapshot in time and the identified equation describes merely one period during the process each time. Based on numerical simulations of an AOD vessel, Odenthal et al. [15] found that the damping effect of the top slag increases with increasing viscosity, which suggests that the mixing time would also increase. In addition, it should be noted that the steel–slag interfacial tension is considerably higher than that of oil and water and, therefore, the model exhibits more aggressive emulsification behavior than was expected to occur in an actual AOD vessel. In fact, it might be possible that the increasing mixing time is the result of an increasing effective viscosity and not of damping effects. In future work, it might be sensible to employ food coloring instead of sulfuric acid as a tracer, as suggested by Wuppermann et al. [1], as the standard deviation of their measurements was much lower than in this work. Nevertheless, it remains uncertain whether the color of the rapeseed oil would affect the reliability of this measurement method.
Conclusions
The aim of this study was to investigate the effect of different process parameters on the mixing time in an AOD converter. The mixing time measurements were conducted with a 1:9-scale water model of a 150 t AOD converter using concentrated sulfuric acid as a tracer. Rapeseed oil was used to represent the top slag phase. The experiments permit the following conclusions:
(1) In the reduction stage as well as in the decarburization stage, the mixing time decreases as the modified Froude number of the tuyères increases.
(2) Top-blowing increases the mixing time, which is believed to be due to the disturbance of the main vortex caused by side-blowing.
(3) The presence of a slag layer increases the mixing time, but the increment is not linear with the amount of slag used. The effect decreases as the modified Froude number increases. In the case of combined blowing, the presence of a slag layer increases the mixing time further.
Finally, mixing time equations were proposed for the reduction stage and the decarburization stage; the correlation factors for these equations were 0.89 and 0.92, respectively.
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