optimization of bath mixing and steel cleanliness during

Optimization of bath mixing and steel cleanliness during steel refiningthrough physical and mathematical modeling

PRANAV KUMAR TRIPATHI* , D SATISH KUMAR, AMIT SARKAR and S C VISHWANATH

JSW Steel Ltd, Vijayanagar Works, Toranagallu, Bellary, Karnataka, India

e-mail: [email protected]

MS received 27 January 2021; revised 13 May 2021; accepted 16 June 2021

Abstract. During ladle refining process, argon gas is purged into the ladle for stirring the molten steel bath to

eliminate thermal and composition gradients and to achieve inclusion flotation. Operating parameters like

purging location, porous plug configuration and argon flow rate primarily affect liquid steel refining. The

efficiency of ladle processing is often quantified through mixing time. To optimize the mixing time and the

associated process parameters for improved bath homogenization and inclusion flotation under different oper-

ating conditions, water modeling studies using 0.2 scale perspex model and computational fluid dynamics

studies using ANSYS CFX v14.5 have been carried out through fluid profile assessment and mixing time

comparison. Comparative study was made between single plug, dual plug and top lance purging configurations.

The studies helped in identifying the optimum argon purging rates and configurations under normal operational

practices. Under abnormal operating conditions involving purging failure from either of the two porous plugs,

usage of a top lance along with the single working porous plug has been investigated and found to improve

mixing and inclusion flotation in the ladle equivalent to dual plug operation. The lab scale studies have been

validated on plant scale through inclusion mapping and found to be in close agreement.

Keywords. Ladle treatment; water modeling; argon purging; mixing time; steel cleanliness.

1. Introduction

During secondary stage of steelmaking, argon purging is

carried out from ladle bottom to enhance reaction rates,

eliminate thermal and composition gradients and remove

non-metallic inclusions from the steel. In such a process,

argon gas bubbles formed in the molten steel move up to

the slag-metal interface under the action of buoyant forces

and finally reach the top layer of slag phase. The rising gas

bubbles push the liquid steel up at local area, thereby

inducing a turbulent re-circulatory flow enhancing the rate

of chemical and thermal homogenization as well as accel-

erating the absorption of harmful non-metallic inclusions

into an overlying slag phase. Under industrial conditions,

typically moderate gas flow rates are applied to achieve

thermal and chemical homogenization, although intense

stirring conditions are also practiced for accelerating slag-

metal reactions. Consequently, depending on the specific

objectives of a ladle refining operation, a wide range of gas

flow rates are applied to maximize alloy recovery through

proper dissolution and refining operations during ladle

treatment along with removal of non-metallic inclusions to

produce ‘clean’ steel.

The physical characteristics of the gas-liquid or air-water

plumes of relevance to ladle processing have been inves-

tigated extensively by a number of researchers. Mazumdar

and Guthrie have extensively reviewed the physical and

mathematical modeling systems useful for gas purging

systems [1]. The effect of different ladle designs on bath

mixing have been reported by Mazumdar et al. They have

reported that a flat ladle bottom accounts for fastest mixing

in the ladle [2]. Castello-Branco and Schwerdtfeger have

studied the characteristics of bubble plumes and measured

the profiles of gas concentration, bubble frequency, and

liquid and gas velocities through water model studies and

concluded that the bubble plume is not at a fixed position

but wanders away from the vertical vessel axis [3]. Sahai

and Guthrie have studied the hydrodynamics of gas plume

systems and reported that the rising plume generates strong

recirculatory mixing zones in the ladle. They have postu-

lated the plume geometry as a function of gas flowrate,

vessel size and aspect ratio [4]. Geng et al have explained

that there are two types of recirculatory zones in the ladle

for dual plug purging – between two plumes and between

plume and ladle sidewall [5]. Schwarz has demonstrated a

two-fluid model to predict bath recirculation and bubble

plume structure in gas stirred vessels [6]. Haiyan et al havecarried out water model studies with different purging rates

for both the plugs. They found that due to different gas flow*For correspondence

Sådhanå (2021) 46:146 � Indian Academy of Sciences

https://doi.org/10.1007/s12046-021-01663-8Sadhana(0123456789().,-volV)FT3](0123456789().,-volV)

rates, a stronger and a weaker plume is formed. In such a

case, the interference and collision from the two plumes are

weakened and the dissipation of stirring energy is decreased

leading to a decrease in the mixing time [7].

It has been well accepted that during ladle refining of

steel, in the immediate vicinity of the nozzle, the input gas

kinetic energy as well as the mode of gas injection are

important variables governing the overall mixing rates in

the vessel. The refining mechanisms during a ladle treat-

ment are controlled by transport mechanisms and flow

patterns in the liquid phase which determine the reaction

rate. Bath stirring provided by argon injection helps in

uniform mixing of reagents in the ladle and better reaction

kinetics is achieved. The flow rates of injected argon have

been found to be the major factor controlling the stirring

rate in the ladle. It has been reported that higher gas flow

rates provide good stirring in the ladle whereas mild

purging increases the rate of inclusion flotation in the ladle.

Xie et al have studied the dynamic behavior of bubbles

emerging from various orifice sizes. They have reported

that bubble evolution and dispersion behavior depends

upon orifice sizes and gas flow rates [8]. Cao and Nastac

have studied the effects of the free surface set-up, injected

bubble size, gas flow rate, and slag layer thickness on the

slag-steel interaction and mass transfer behavior and

reported that as the gas stirring rate increases, the mass

transfer coefficient increases. By injecting finer bubbles,

increasing gas flow rate and slag layer thickness, the ladle

refining efficiency can be enhanced significantly [9]. Patil

et al have reported that along with gas flow rate, liquid

depth and vessel radius, thickness of the upper phase liquid

slag has the most significant effect on mixing times. They

have concluded that 95% bulk mixing time in slag covered,

axisymmetrical ladles appropriately represents the mixing

behavior in the ladle [10]. Amaro-Villeda et al have

reported that a thicker slag layer increases the mixing time

as it reduces the average recirculation velocity. An increase

in slag viscosity and a lower gas purging rate reduces the

top slag ‘eye’ opening. Mixing times are also increased by

decreasing the gas flow rate and increasing the number of

nozzles [11]. Gonzalez-Bernal et al have conducted water

model studies on a 1/7th scale model using modified Froude

number criterion and demonstrated that the number, size

and locations of recirculatory zones within the ladle have a

significant effect on the mixing time in the ladle. Increasing

the gas flow rate decreases the mixing time, however an

excess of gas flow results in adverse effects like re-oxida-

tion due to slag eye formation and reduction in the

refractory life of the ladle [12]. Chen et al have reported theeffects of nozzle arrangements, separation angle, radial

position and asymmetry as well as tracer adding position on

the mixing time for a ladle with dual plug purging system.

The most favorable mixing has been reported when the two

nozzles are symmetrically arranged at half radii in the ladle,

with 45� separation angle [13]. Liu et al have found that

eccentric gas injection in the ladle improves the mixing

efficiency. The mixing time decreases with increasing the

gas flow rates and porous plug angles. Shorter mixing times

are achieved in case of dual plug purging when plugs are

located diametrically opposite at mid-bath radius position.

Higher gas flow rates cause slag eye formation at the top

which decreases the mixing time due to higher turbulence

[14]. Ganguly and Chakraborty have examined the effects

of gas flow rate, bottom nozzle configurations and tracer

addition locations on mixing time. They have reported that

the arrangement of bottom nozzles has a great effect on the

mixing behaviour in a gas stirred ladle, with off centric gas

injection producing shorter mixing time. Mixing time is

also sensitive to the tracer addition position, particularly for

the axisymmetric bottom gas injection system [15]. Fan and

Hwang have concluded that the optimal addition location of

a Ca-Si injection wire in a gas stirred ladle is opposite to the

plume as it helps in avoiding the loss of additions through

vaporization [16].

The efficiency of ladle processing is often quantified

through mixing time. The stirring action imparted by the

rising gas in gas stirred ladles causes the movement of melt.

An optimized purging location tends to impart proper and

uniform mixing in the ladle with minimum percentage of

regions with no or ‘‘dead’’ flow. The choice of single or

multi-location purging is highly critical for final steel pro-

cessing. Apart from purging configuration, the purging flow

rates also need to be properly maintained in the ladle

depending upon the treatment being carried out. Gomez

et al have studied the effect of nozzle radial position,

nozzle separation angle, gas flow rate and slag thickness on

mixing time using water model studies. They have reported

shortest mixing time with a nozzle radial position of 0.67R

with 608 separation angle and without top slag layer. For a

radial position of 0.5R, both single and dual purging plugs

were found to be nearly equivalent. They have also reported

that tracer concentration also affects mixing times in the

ladle at low gas flow rates. However at high gas flow rates,

mixing time is independent of tracer concentration [17].

Chattopadhyay et al have measured mixing time on a 0.2

scale aqueous model of a single tapered ladle with different

bottom purging locations. They have reported that two plug

purging gives faster mixing than one plug purging. Also,

faster bath mixing is obtained when the plugs are located at

�R distance from the ladle centre. A higher gas injection

rate yields faster mixing in the ladle [18].

Even though a number of physical modeling and

numerical simulation studies have been carried out by a

number of researchers to study gas-metal dual-phase

interactions in ladle, the effects of different purging con-

figurations for argon injection ranging from multi-point

injection to the usage of a top lance have been scarcely

investigated. Additionally, the published research result

does not suggest any corrective action to be taken when

non-optimal purging configurations prevail during ladle

treatment in practical industrial conditions. In the present

work, different ladle purging configurations have been

146 of 14 Sådhanå (2021) 46:146

studied by means of physical and mathematical (CFD)

modeling techniques. The aim of the present work is to

identify the best purging configuration for faster bath

homogenization and inclusion flotation. Additionally, a

new purging configuration has been suggested as a pre-

ventive and corrective action when optimal purging con-

ditions are not achieved due to operational issues. The

suggested configuration was subsequently validated

through actual plant trials.

1.1 Ladle argon purging configurations

Steel melting shop II at JSW Steel is equipped with 180 ton

capacity ladles with dual porous plugs at ladle bottom for

argon purging. The normal operational practice at ladle

heating furnace (LHF) stations for ladle treatment involves

usage of two porous plugs for argon purging. These porous

plugs are located asymmetrically in the ladle for argon

purging. All the additions in the ladle are carried out from

the top. Majority of these additions are of lower density

than molten steel. Therefore, for their proper dissolution in

the steel bath vigorous mixing is carried out.

The different simulation conditions are listed in table 1.

Normally, both porous plugs are used for argon purging

(A1). However, due to some operational abnormality, if

argon purging fails from one of the plugs, the treatment is

continued using the other plug essentially reducing it to a

scenario of single plug purging (A3). When purging from

both the plugs fails then as a fail-safe option, a top lance is

introduced from the top into the ladle for argon purging. In

the existing setup, the top lance is used in a slant condition

at an inclination of 6-78 with vertical axis (B1). Unlike the

very fine openings of the plugs, the top lance has a quite big

9 mm opening for argon passage into the ladle. Apart from

these conditions, some additional situations have also been

studied to identify the effect of purging location on mixing

times in the ladle, namely - symmetrically placed dual

plugs (A2), single plug at ladle centre (A4) and top lance in

straight condition at ladle off-centre (B2) and centre (B3).

One additional condition was simulated when only one of

the two porous plugs remains operational viz. usage of top

lance along with the lone operative plug (C).

During ladle treatment, bulk additions (lime, aluminum,

ferro-manganese etc.) are carried into the ladle through

automated addition via hopper or through wire feeding

system. The addition location is fixed in such cases.

However, a majority of the additions like calcined petro-

leum coke (CPC), ferro-alloys and other micro-alloys are

carried out manually into the ladle. In such cases, some

variations in the addition location in the ladle are quite

likely which is expected to affect the mixing time in the

ladle. Thus, tracer addition locations were varied for certain

configurations (A1, A3, B1 and C) as shown by locations a,

b, c and d in figure 1.

2. Experimental

2.1 Physical model

To investigate the effect of different purging configurations

on the mixing time, a reduced scale plexi glass water model

was fabricated having 0.2 scale factor with respect to the

180 ton industrial scale ladle. The water model was

equipped with flow meters, multiple porous plugs at the

base and a top lance system. Water was used to simulate

molten steel whereas compressed air was injected through

porous plugs and top lance to simulate argon purging in the

ladle. Flow meters capable of controlling flow rates

between 0-10 litres/minute (l pm) were used to regulate air

flow into the ladle. Conductivity measurements were car-

ried out using the electrical conductivity measurement

technique through stimulus response of injected salt solu-

tion of known concentration. Flow visualization studies

were carried out using colored KMnO4 solution. The details

of actual ladle and model vessel are given in table 2.

The major forces influencing the flow in an argon purged

industrial ladle include gravitational, buoyancy and drag.

To establish similarity between plant and water model

Table 1. Cases studied through water modeling.

Purging configuration Description Tracer addition location

Dual plugs Same side

Opposite side

(A1)(A2)

1. Over plume

2. Opposite to plumes

3. Between plumes

(a)(b)(c)

Single plug Off-centre

Centre

(A3)(A4)

1. Over plume

2. Opposite to plume

(a)(b)

Top lance Slant & off-centre

Straight & off-centre

Straight and centre

(B1)(B2)(B3)

1. Over plume


(a)(b)

Top lance with single plug Slant lance with off-centre plug (C) 1. Over plume


3. Over non-working

plug

(a)(b)(d)

Sådhanå (2021) 46:146 of 14 146

system, modified Froude number criterion was used as it

can characterize the kinetic similarity between plant ladle

and water model for argon purging conditions

[3, 7, 11–14, 17, 19, 27]. Consequently,

ðF 0rÞmðF

0rÞp ð1Þ

qair:u2water

ðqwater � qairÞ:g:Hm¼ qAr:u

2steel

ðqsteel � qArÞ:g:Hpð2Þ

The characteristic velocity u. is expressed as:

u ¼ 4:Q

p:d2ð3Þ

Equating eg. (1) to (3) based on the data listed in table 3

gives :

Qm ¼ 0:0077195:Qp ð4Þ

Corresponding to the range of argon purging used in

plant (10-50 Nm3/h), the flow rates for compressed air

during water model studies is listed in table 1.

The water model vessel was filled with water up to the

required bath height corresponding to the level of the metal

bath in the plant ladle. The position of the lance and water

level were fixed for the entire set of experiments. The

extent of mixing was determined by conductivity mea-

surements using NaCl salt solution of known density (1.1

g/lt) as tracer and a standard conductivity meter (Eutech

CyberScan COND 600) for measuring the conductance of

the solution. Data generated was captured using a computer

equipped with data acquisition software. The experimental

setup is shown in figure 2. Compressed air was introduced

through the top of the lance or porous plugs at predeter-

mined flowrates using the flow meter and was monitored

continuously during the experiments. A fast speed video

camera was used to record the generation and dissipation

phenomenon of the bubbles in the ladle during the course of

the experiments. When steady state was attained, the tracer

was gently added from the ladle top using a beaker wall just

below the water surface shown as position ‘A’ in figure 3 to

simulate alloying additions in the ladle. The location A was

Figure 1. Experimental Cases.

Table 2. Details of water model and plant scale steel ladle.

Parameter Plant Model

Scale factor 1 0.2

Ladle height (mm) 3500 700

Ladle top diameter (mm) 3460 692

Ladle base diameter (mm) 3205 641

Ladle freeboard (mm) 575 115

Top lance tip diameter (mm) 9 1.8

Blowing Gas Argon Air

Density of Blowing Gas (kg/m3) 1.78 1.29

Density of liquid (kg/m3) 7020 1000

Table 3. Argon flowrates for industrial scale and water model

system.

Argon flowrate Values

Industrial scale ladle (Nm3/h) 10 20 30 40 50

Water model ladle (lpm) 1.3 2.6 3.9 5.1 6.4

146 of 14 Sådhanå (2021) 46:146

varied as per different addition locations mentioned in

table 3. The steel bath in the ladle was considered ‘ho-

mogenized’ when 95% mixing of tracer was achieved as

reported by Madan et al [10] and Patil et al [20]. There is asignificant difference between the mixing times in the

central volume of the ladle and near walls and hence the

measurement in the slowest flow region is critical for

estimating the bulk mixing time. Through a series of

experiments it was found that the conductivity probe dipped

650 mm into the ladle at 50 mm away from the wall of the

vessel shown as position ‘B’ in figure 3 was the most

suitable location as it took relatively higher time for com-

plete homogenization compared to other locations. 50 ml

tracer was used in each experiment and conductivity of the

solution was recorded continuously thereafter till the

reading becomes constant. The same process was repeated

for all the cases. The internal mixing time (T) was calcu-

lated from the conductivity vs time plots and compared for

all the design cases.

2.2 Mathematical model

Commercially available computational fluid dynamics

(CFD) code ANSYS-CFX 14.5 was used to perform com-

putations on a Dell Workstation with Intel Xeon 12 core

64-bit processor and 72 GB of RAM on Windows platform.

The fluid flow profile during ladle treatment was simulated

under different argon purging configurations and compared

with the water modeling findings. A full scale 3D model

was developed using momentum balance considering

molten steel as a Newtonian and incompressible fluid.

2.3 Governing equations

The following governing equations were used for devel-

opment of CFD model.

2.3a The continuity equation:

oqot

þr � qUð Þ ¼ 0 ð5Þ

where, q is fluid density, t is time and U is the flow velocity

vector field.

In the case of incompressible flows q is a constant,

hence, the above mass continuity equation simplifies to a

volume continuity equation.

r � u ¼ 0 ð6Þ2.3b The momentum equation:

o qUð Þot

þr � qU � Uð Þ ¼ �r pþ r � sþ SM ð7Þ

where, p is the static pressure, s is the shear stress and SM is

the momentum source term. The stress tensor, s, is

described as

s ¼ l rU þ rUð ÞT� 2

3dr � U

� �ð8Þ

where, l is the dynamic viscosity.

2.3c Buoyancy: For simulating the effect of buoyancy

due to difference in densities of the interacting fluids, a

source term is added to the momentum equations as in

Eq. (7).

SM;buoy ¼ q� qref� �

g ð9Þ

Figure 2. Experimental set-up.

Figure 3. Water model measurements

Sådhanå (2021) 46:146 of 14 146

The density difference q� qref� �

was evaluated using

full buoyancy model for modeling multiphase flow.

2.3d Turbulence model: Due to presence of re-circulatoryzones with high Reynolds number in the continuous fluid

domain, k � e model was used to model turbulence for the

continuous molten steel phase as it offers a good compro-

mise in terms of accuracy and robustness. k is the turbu-

lence kinetic energy and is defined as the variance of the

fluctuations in velocity.

o qkð Þot

þr � qUkð Þ ¼ r � lþ ltrk

� �r k

� �þ Pk þ Pkb

� qe

ð10Þe denotes the rate at which the velocity fluctuations

dissipate.

o qeð Þot

þr � qUeð Þ ¼ r � lþ ltre

� �re

� �

þ ek

Ce1 Pk þ Pebð Þ � Ce2qeð Þ ð11Þ

where, Ce1;Ce2; rk and re are constants.

For dispersed argon gas phase, dispersed phase zero equa-

tionmodel was used. It uses an algebraic equation to calculate

the viscous contribution from turbulent eddies to derive a

constant turbulent eddy viscosity is calculated for the entire

flow domain. The turbulence viscosity (lt) is modeled as the

product of a turbulent velocity scale (Ut) and a turbulence

length scale (lt) as proposed by Prandtl and Kolmogorov.

lt ¼ qflUtlt ð12Þwhere fl is a proportionality constant.

2.4 Modeling and boundary conditions

The model was developed in full scale. No symmetry

planes were considered so as to allow both symmetrical as

well as non-symmetrical flow conditions depending on the

selected turbulent conditions. The present scope of study

includes comparison of fluid flow profile, hence isothermal

conditions were assumed for all the simulations. It allows

for the assumptions that the density of molten steel doesn’t

change appreciably during steady state operating condi-

tions. The simulation involved a multi-phase study with

following two phases, phase 1 – molten steel, and phase 2 –

argon (purging) gas. Molten steel was taken as the con-

tinuous liquid phase with argon gas as the dispersed gas-

eous phase. As both these phases have different material

properties and a distinct flow field, hence, the inhomoge-

neous model was used for simulating multi-phase flow in

the ladle. This allows each fluid to possess its own flow

field with different fluids interacting via interphase transfer

terms. Free surface model was used for simulating

interfacial mass transfer between the continuous and dis-

persed phases. This allows a distinct interface between the

fluids to be modeled in multiphase flows. A drag coefficient

of 0.44 was used to model inter-phase drag on dispersed

fluid phase. A surface tension coefficient of 1.6 N/m was

used for steel-argon interface considering molten steel as

the primary fluid. k � e turbulence model with medium

turbulence intensity was used for continuous phase and

dispersed phase zero equation model was used for dispersed

phase.

The mass flow rate inlet boundary condition was used for

the inlets, and opening boundary condition was used at the top

of the ladle as it is open to atmosphere. As the viscosity of

continuous phase (molten steel) is very high as compared to

that of dispersed phase (argon gas), all the walls were assumed

as no slip walls for the continuous phase and as free slip walls

for the dispersed phase. Argon gas mean bubble size was

considered as 30 mm using Eq. 13 provided by Johansen and

Boysan for the simulations corresponding to the average

purging rates of 20-25 Nm3/hr under plant conditions [21].

db ¼ 0:35Q2

g

� �0:2

ð13Þ

Buoyant conditions were simulated based on density dif-

ference. The density of argon gas was used as buoyant ref-

erence density to avoid round-off errors during calculations.

2.5 Simulation process

During numerical simulations, the whole fluid domain is

divided into multiple small regions or grids, often called as

control volumes wherein the Navier–Stokes equation is

solved iteratively to obtain the approximate solution of the

fluid flow field. The accuracy of the solution and the

computational load depends on the mesh size and quality.

To establish mesh size independency of the results, initial

steady state simulations were carried out with different

mesh sizes of 15, 25, 30, 50 and 100 mm for purging

configuration A1 and A3. The resultant velocities of molten

steel phase were monitored at locations A and B (figure 3)

keeping argon purging rate constant (30 Nm3/h).

It was found that the steel flow velocities under different

conditions slightly increase with increase in mesh size and

become practically constant for mesh size B30 mm as

shown in figure 4. Consequently, a maximum mesh size of

25 mm was finalized to achieve good accuracy with opti-

mum computational load. The minimum mesh size allowed

was 1 mm for ladle domain and 0.5 mm for lance domain.

Three dimensional unstructured mesh was generated for the

fluid domain using a combination of tetrahedral, prism and

pyramid elements. Individual part meshing strategy was

used along with mesh refinement to achieve finer mesh

elements at intricate geometry regions. A minimum mesh

quality of 0.5 was maintained for[95% mesh elements to

ensure good accuracy of results.

146 of 14 Sådhanå (2021) 46:146

Initially the CFD model was set up for a transient run.

The molten steel in the ladle was considered stationary as

the initial condition. Argon gas was injected through the

purging locations at specified flow rates as per the actual

plant practice. The momentum from the impinging jet stirs

the molten steel bath in the ladle and a fully developed flow

profile gets established after some time (300 secs). The

results so obtained were used as an initial guess and then

solved again iteratively under steady state conditions to

obtain final flow profile in the ladle. The obtained flow

profiles for four different purging configurations (A1, A3,

B1 and C) were compared at a constant argon purging rate

(30 Nm3/h) (table 4).

3. Results and discussion

The rate of mixing in the ladle is affected by the intensity of

stirring which is a function of purging configuration and

argon flowrate. For the same argon flowrate, the purging

configuration which imparts higher stirring over a larger

area in the ladle will give better mixing. Stirring in a ladle

system with purging mechanism is a strong function of the

plume dynamics which causes re-circulation in molten steel

bath. Water modeling experiments were conducted to

measure the mixing time in the ladle for different purging

configurations. Mixing time is an indicator of fluid flow

profile and diffusion characteristics inside the vessel and

should be lower for faster process kinetics. It indicates that

the additions carried out in the ladle during treatment can

be quickly dissolved and thermal and compositional

homogenization is achieved at a faster rate. Figure 5 shows

the time taken for uniform mixing in the ladle under dif-

ferent purging configurations. Dual plugs (A1) yield faster

mixing than single plug (A3). Moreover, in case of dual

plug purging, same-side located porous plugs (A1) give

better mixing as compared to symmetrically located plugs

(A2). Similarly, a single plug located off-centre (A3) gives

better mixing in the ladle as compared to a centrally located

plug (A4). In the case of top lance usage, an inclined lance

(B1) gives a lower mixing time than a vertical one (B2) at

the same location. Similar to the scenario of single plug at

centre (A4), a top lance being used at ladle centre (B3) will

result in highest mixing time among all cases. Krishnaku-

mar et al have also reported that the bath mixing is lower

when purging is carried out from ladle center and gradually

increases as the nozzle is moved to a mid-radius position

[22]. Moreover, the mixing was found to be most sluggish

in case of a top jet

When a top lance is used with a single plug (C), the

mixing time is similar to dual plug purging scenario (A1). It

is evident that purging from off-centre (A1, A3, B1 & B2)

and asymmetric locations (A4) provides faster mixing in

the ladle as compared to central (A4 & B3) or symmetric

(A2) locations. This can be attributed to faster mass transfer

and recirculation rates in these configurations due to the

presence of stronger compositional gradients.

The mixing time studies imply that fastest mixing will be

achieved in the ladle when purging is carried out from two

locations in an asymmetric fashion. The reason for better

performance of these configurations (A1 & C) is demon-

strated in flow visualization studies. As shown in figure 6,

for both these configurations, the tracer gets uniformly

Figure 4. Effect of grid size on velocity measurements.

Table 4. Simulation/solver settings.

Total simulation time (transient run)300 secs

No. of iterations (steady state run) 10000

Advection scheme High resolution

Turbulence numerics High resolution

Transient scheme Second order backward Euler

Convergence criteria Residual type : RMS (target :

0.00001)

Solver scheme Double precision

Figure 5. Effect of different purging configurations on mixing

time.

Sådhanå (2021) 46:146 of 14 146

mixed at a faster rate as compared to single purging con-

figurations (A3 & B1). Even for a constant argon flow rate,

the mixing is considerably faster in dual purging configu-

rations primarily because of higher momentum being

imparted by the purging gas to the steel bath in the ladle.

This leads to vigorous mixing in the ladle which is reflected

in mixing times for respective configurations.

Flow visualization studies were validated through

mathematical model results by comparing the velocity

vectors for different cases. As shown in figure 7, during

purging from a single plug (A3) or a top lance (B1), the

melt velocities are higher only in the vicinity of the porous

plugs, providing good mixing only near such regions. For

locations away from the purging, the velocities are much

lower making such regions virtually ‘‘dead’’ towards fluid

flow. On the other hand, dual plugs (A1) provide good

recirculation in the ladle both near and away from purging

location. Similar effect is observed in the case when top

Figure 6. Comparison of ladle flow profile.

146 of 14 Sådhanå (2021) 46:146

lance is used along with a single plug (C). In such a sce-

nario, strong recirculation occurs in the ladle which yields

faster mixing in the ladle.

Even though temperature studies are outside the scope of

present studies, an indication of thermal stratification ten-

dency can be ascertained on the basis of velocities of the

steel melt in the ladle at ‘‘dead regions’’ where very low

mixing takes place. Such regions are located at extreme

ends of the flow streamlines originating from the porous

plugs near ladle walls. Two such locations have been pre-

viously shown in figure 3 as points P1 and P2 located near

ladle walls at half and quarter bath height from the ladle

base. It is imperative to assume that if the downward

velocities at these two locations are higher, faster mixing

will take place in such ‘‘dead regions’’ and hence, the rate

of compositional and thermal homogenization will be

enhanced. Figure 8 shows that velocities at these locations

are much lesser in case of single purging cases (A3 & B1)

but substantially higher in case of dual plug purging (A1).

Top lance with single plug (C) yields highest velocities at

such locations. Hence, fastest bath mixing and

Figure 7. Velocity vectors for different purging configurations (argon purging rate: 30 Nm3/h).

Figure 8. Molten steel downward velocities in dead flow regions

(argon purging rate: 30 Nm3/h).

Sådhanå (2021) 46:146 of 14 146

homogenization is expected in this case along with lower

thermal stratification.

3.1 Effect of addition location

Figure 9 shows the effect of different locations of the

alloying addition on mixing time for each purging config-

uration. Among all the purging cases, mixing time is lower

when addition is carried out over the plume. It implies that

the alloying additions tend to get ‘carried’ into the steel

melt with a greater momentum thus enabling faster mixing

and homogenization in the ladle. The fastest mixing is

obtained when both plugs operate and addition is done over

the plume over either plug (A1-a). The slowest mixing is

obtained in case of top lance usage with addition opposite

to plume (B1-b). In the case of a top lance being used with

single plug, the mixing times show least variations with

respect to different tracer addition locations (C-a, b, d).

This implies that the rate of mixing is much higher in the

entire volume of the ladle. This is expected to result in

higher alloy dissolution rates irrespective of their addition

location.

3.2 Effect of argon flowrate on bubble size

Argon flow rate plays an important role in governing the

fluid flow dynamics inside the steel ladle as it controls the

size of the generated argon bubbles. Hence, it is an

important criterion for governing the mixing and inclusion

flotation phenomena in the steel ladles during treatment.

There is immense difference in argon bubble formation

pattern between a porous plug (made of refractory block

with fine openings of*0.25 mm) and a top lance due to the

inherent design differences. Compared to the very fine

openings in a porous plug, top lance has a bigger opening

Figure 10. Argon bubble generation for porous plug and top lance.

Figure 9. Effect of tracer injection location on mixing time.

146 of 14 Sådhanå (2021) 46:146

(*9 mm) for argon injection leading to formation of big

argon bubbles.

Figure 10 compares argon bubble sizes for different

purging rates for a porous plug and a top lance. At lower

argon flow rates (1.3-2.6 l pm), fine sized argon bubbles

are generated and rise upwards in a non-turbulent fashion.

These bubbles provide mild purging to the steel melt and

tend to adhere themselves to the non-metallic inclusions

during ladle treatment and aid in their flotation as reported

by Li et al and Zhang et al [23, 24]. At intermediate

argon flow rates (3.9-5.1 l pm), argon bubbles form a

jacket over the porous plug generating bigger bubbles

which disintegrate into smaller bubbles as they move up.

This phenomenon of disintegration of large gas bubbles

imparts turbulent mixing in the ladle through distribution

of the inherent kinetic energy of rising gas bubbles. It

causes stirring of the steel melt and aids refining reactions

and composition and temperature homogenization. This

range is conducive for providing intermediate stirring in

the bath during later stages of ladle treatment. At higher

argon flow rates ([6 l pm) big bubbles are formed which

impart high momentum to the metal bath resulting in

turbulent mixing and faster homogenization of the steel

melt. Therefore, when high additions are carried out into

the bath and faster mixing is required for quick dissolu-

tion, argon must be purged at high flow rates. However,

too high flow rates don’t provide any additional benefit for

mixing in the ladle as specific consumption of argon is

heavily increased which offsets any additional benefits

[12]. This is because, beyond the critical flow rate, an

invariant flow pattern is established. Mandal et al have

reported that the critical flow rate corresponds to the onset

of the inertial and gravitational force dominated flow

regime [20]. When argon purging is carried out using a

top lance system, only big argon bubbles are formed at

both low and high flow rates. Thus, even though bath

mixing can be achieved, inclusion flotation is expected to

be highly diminished.

Figure 11 compares the argon bubble size distribution for

a porous plug and a top lance highlighting the minimum

and maximum size of the generated bubbles. Effective

bubble size for each case was calculated using the weighted

average from extreme size ranges and their relative con-

centration. It can be noted that the difference between

minimum and maximum bubble sizes is smaller at lower

purging rates. This is because at lower purging rates only

smaller bubbles are formed. At higher purging rates, bigger

bubbles are formed which get dissociated into smaller

bubbles due to the higher turbulent energy of the bath as

also been reported by Zhang et al [24]. This effect is more

pronounced for a top lance where only big bubbles are

observed at lower argon flow rates and both small and big

bubbles at higher purging rates. In such a scenario, even

though small bubbles are generated, their effect on inclu-

sion floatation is minimal due to the highly turbulent con-

ditions in the bath.

During initial stages of treatment, high argon flow rate is

preferable as it allows for faster mixing in the ladle. Con-

sequently, the alloying additions get homogenously mixed

at a faster rate thereby increasing their recovery. It results in

faster reaction kinetics and supports the refining processes.

On the other hand, too high flow rates during later part of

treatment might cause slag emulsification in the ladle and

open eye formation at the top surface of the ladle exposing

the molten steel to atmosphere leading to oxygen and

nitrogen pickup as also reported by Hoang et al and

Krishnapisharody et al [25, 26]. Therefore, lower flow rates

are desirable during later stages of the treatment so as to

generate small argon bubbles which help in inclusion

flotation to obtain ‘‘cleaner’’ steel. Due to the formation of

small bubbles at lower purging rates, porous plugs are

Figure 11. Argon bubble sizes for (a) porous plug and (b) top

lance.

Sådhanå (2021) 46:146 of 14 146

expected to be more effective for inclusion floatation as

compared to a top lance.

4. Plant scale validation

The industrial process is more complex and has other

additional variables which affect directly or indirectly the

mixing efficiency such as bath depth, temperature, top slag

layer etc. It is therefore important to validate the results at

the plant scale under the operating conditions which helps

in analyzing the limitations regarding the direct application

in practice.

For validation of lab scale studies, plant trials at ladle

heating furnace (LHF) in steel making shop were per-

formed for different purging configurations keeping other

operating parameters unchanged. After ladle treatment was

finished, lollypop samples were collected from the ladles.

Post treatment, the ladles were transferred to continuous

casting station. During continuous casting, at 50% ladle

weight, one additional sample was taken from the tundish

near shroud (metal entry) region. These samples were cut,

fine polished and observed under optical microscope to find

out inclusion area fraction through optical image analysis.

For each sample, a number of frames were captured to

cover the overall sample area. Out of these frames, 10 worst

frames showing the maximum number of inclusions were

selected for further analysis. Figure 12 shows the repre-

sentative frames from LHF lollypop samples showing the

inclusion distribution for different purging configurations at

100 x magnification.

The average inclusion area fraction comparison for each

of the purging configurations shown in figure 13 clearly

suggest that double plug purging has maximum effect on

inclusion flotation in the ladle evident from the lowest

values of inclusion area percentage in LHF as well as caster

samples. The samples with top lance purging show maxi-

mum inclusions indicating a lower degree of inclusion

Figure 12. Micrographs showing steel cleanliness for LHF lollypop samples for different purging configurations (argon purging rate =

20 Nm3/h).

Figure 13. Comparison of inclusion area fraction for different

purging configurations (argon purging rate = 20 Nm3/h).

146 of 14 Sådhanå (2021) 46:146

flotation in case of top lance purging. In line with water

model experiments, the plant trials also suggested that

usage of top lance along with single plug for argon purging

yields nearly similar effect as obtained from dual plug

purging. Thus faster bath mixing along with good inclusion

floatation can be achieved.

On the basis of plant trials, it was evident that in the case

of purging failure from one of the plugs, the mixing time as

well as inclusion flotation in the ladle can be improved by

using top lance for argon purging along with the single

operative plug.

5. Conclusion

To investigate the role of different configurations of porous

plus and top lance on fluid flow and mixing time in the ladle,

water modeling and CFD studies were carried out. The effect

of argon purging flow rates and relative location of porous

plugs on mixing kinetics in the ladle were studied. The opti-

mum location and conditions for usage of a top lancewere also

investigated.On the basis of the results obtained throughwater

modeling and CFD studies, it was concluded that:-

1. Argon purging through dual porous plugs gives the least

mixing time in the ladle followed by single plug purging.

Usage of a top lance results in highest mixing time

among all cases.

2. Asymmetric location of argon purging gives faster

mixing in the ladle as compared to central location.

3. Additions in the ladle must be carried out over the plume

to achieve lower mixing time and faster dissolution of

alloying additions.

4. Whenpurging fromoneof the porousplugs fails, a top lance

can be used in combination with the lone working plug to

achieve mixing time equivalent to dual plug purging.

5. Purging from two locations is expected to result in faster

compositional and thermal homogenization in the ladle

owing to strong steel melt velocities in relatively dead

flow regions in the ladle.

It is quite evident that the current asymmetric location of

dual plugs yields the best mixing time. The mixing time is

greatly reduced in case purging from one of the plugs fails.

In such a case, for non-critical steel grades, single plug

purging can be sufficient to achieve the desired cleanliness

levels. However, for critical grades, a top lance should be

used along with the lone working plug to achieve faster

mixing in the ladle along with higher levels of steel

cleanliness. Usage of a top lance alone is not recommended

and should be avoided as much as possible.

Symbols

CPC Calcined petroleum coke

LHF Ladle heating furnace

g Gravitational acceleration

t Time

T Temperature

CD Drag coefficient

Re Reynolds number

d Distance or length

db Mean particle diameter of phase bAab Interfacial area density

p Gauge pressure

Pk Turbulence production due to viscous forces

r Location vector

rb Volume fraction of phase bSM Momentum source

SM;buoy Momentum source due to buoyancy

Pkb, Peb Buoyancy production & dissipation terms

(represent influence of the buoyancy forces)

Ce1, Ce2 k � e Turbulence model constants

u Fluctuating velocity component in turbulent

flow

U Vector of velocity Ux;y;z

Ut Turbulent velocity scale

lt Turbulence length scale

a Subscript indicating phase ab Subscript indicating phase bq Density

qref Reference density

rk Turbulence model constant for the k equation

re k � e Turbulence model constant

s Shear stress

d Identity matrix or Kronecker Delta function

k Turbulence kinetic energy per unit mass

e Turbulence dissipation rate

l Molecular (dynamic) viscosity

lt Turbulent viscosity

r Gradient

� Scalar product

� Tensor product

db Bubble diameter

Q Flowrate (Nm3/s)

Acknowledgement

The authors would like to thank secondary steelmaking

operations team at JSW Steel Ltd. Vijaynagar Works for

the help and support extended during this study.

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