liquid steel flow in continuous casting machine: modelling and measurement

Liquid steel flow in continuous castingmachine: modelling and measurement

M. M. Yavuz*

Single phase (liquid steel) and two-phase (liquid steel and argon bubbles) three-dimensional

computational fluid dynamic and heat transfer models were developed for the continuous casting

machines of ArcelorMittal. The computational domains include tundishes, slide gates, submerged

entry nozzles and moulds. The effects of buoyancy, tundish design, tundish practices, nozzle

design and caster practices on flow structure were investigated. Mathematical modelling is

discussed in detail. In addition, submeniscus velocity measurements in the slab caster mould are

performed with the method of torque measurement. A consumable probe is inserted into the liquid

steel meniscus from the top of the mould through mould powder and slag layer. The liquid steel

flow applies a drag force to the probe, which then generates a torque. This torque value is

measured and then converted back to velocity. The concept and challenges of the technique are

discussed, and the effects of casting parameters on mould flow structure are investigated.

Product quality in relation to real time meniscus velocity measurements is also discussed.

Keywords: Continuous casting, Tundish flow, Mould flow, CFD, Mould flow measurement

IntroductionOver the years, due to stringent customer requirements,steel producers have had to continuously improvecastability and product quality in order to reduce rejectrates (internal and/or external). Extensive research hasbeen conducted to fundamentally understand theproblems and to propose effective solutions. It has beenshown that the castability and the product qualityprimarily depend on the flow in the continuous castingmachine, which has effects on many phenomena, such asheat transfer, transport of inclusions, meniscus freezing,shell thinning from the jet impinging upon the solidify-ing shell, transient waves and fluctuations of themeniscus, thermal stress and crack formation.1–9 Flowstructure ultimately determines the castability of aparticular steel product and the total quantity and finallocation of defects in the product. Therefore, under-standing and controlling liquid steel flow in continuouscasting machine, tundish and mould have been ofconsiderable interest in recent years.

The flow in tundish and mould is usually very complex,turbulent and therefore unsteady by its nature. Most ofthe time, it involves three different materials at threedifferent phases that are temperature and pressuredependent (liquid steel, solid inclusions and argon gas).In addition, it is quite challenging to perform any type ofreal time flow measurement in a casting environment.Therefore, comprehensive understanding is required to

model and/or measure the flow structure in a continuouscasting machine. Currently, only a few techniques arepossible to perform an actual flow measurement in themould, particularly in the meniscus region, while model-ling using both experimental and computational techni-ques is a common approach to characterise the flow.

This paper aims to provide an overview of liquid steelflow in continuous casting machines while discussing thechallenges and applicability of computational modellingand real time measurements. An attempt is made toaddress numerous casting issues via computational fluiddynamics (CFD) and real time flow measurements. Thecomputational modelling study addresses the effects oftemperature and/or buoyancy, unbalanced (uneven)throughput extraction from strands and tundish designon tundish flow. It also aims to investigate the effects ofsubmerged entry nozzle design, nozzle bottom and argoninjection on mould flow. The flow measurement studyprimarily demonstrates the effects of casting practices andnozzle designs on meniscus flow and product quality.

ModellingSingle phase (liquid steel) and two-phase (liquid steel andargon bubbles) CFD and heat transfer models weredeveloped for continuous casting machines using ANSYSCFX. The models were applied to tundishes and slab castersof various ArcelorMittal steel shops in order to solve specificproblems for each operation. Figure 1 demonstrates someof the simulation domains that were used in the presentstudy. Table 1 is constructed to list the specific informationregarding tundish designs, nozzle designs and simulationparameters used in the present study. For the mouldsimulations, Unsteady Reynolds averaged Navier–Stokes

Assistant Professor, Mechanical Engineering Department, Middle EastTechnical University, Ankara 06531, Turkey; formerly Senior ResearchEngineer, ArcelorMittal Global R&D, East Chicago, IN 46312, USA

*Corresponding author, email [email protected]

� 2011 Institute of Materials, Minerals and MiningPublished by Maney on behalf of the InstituteReceived 23 January 2011; accepted 11 March 2011DOI 10.1179/1743281211Y.0000000013 Ironmaking and Steelmaking 2011 VOL 38 NO 6 453

(URANS) k–e models were used, and time averagedpatterns were demonstrated for 200 s of liquid steel flow.For the tundish simulations, RANS k–e models were used.For the computational models, the governing equations(conservation of continuity, conservation of momentumand conservation of energy) along with the transportequations of turbulent kinetic energy k and turbulent eddydissipation e are solved. The equations are as follows10

Conservation of mass

Lr

Ltz+(rU)~0 (1)

Conservation of momentum

L(rU)

Ltz+(rU6U)~{+pz+tzSM (2)

Conservation of energy

L(rhtot)

Lt{

Lp

Ltz+(rUhtot)~

+(l+T)z+(Ut)zUSMzSE (3)

Transport equation of turbulent kinetic energy

L(rk)

Ltz+(rUk)~+ mz

mt

sk

� �+k

� �zpk{re (4)

Transport equation of turbulent eddy dissipation

L(re)

Ltz+(rUe)~

+ mzmt

se

� �+k

� �z

e

k(Ce1

pk{Ce2re) (5)

The corresponding constants in equations (4) and (5)Ce1

, Ce2, sk and se were 1?44, 1?92, 1 and 1?3

respectively. In addition, for turbulence viscosity, theconstant Cm was set to 0?09.

For the single phase flow, the relative importance ofbuoyancy due to temperature variations in a convectiveflow can be estimated using the ratio of Grashof numberand square of Reynolds number

Gr

Re2~

gbLDT

U2(6)

where b is the thermal expansion coefficient. For theabove relation, a value approaching or exceeding unityindicates that buoyancy effects are significant in theflow, while small values indicate that buoyancy effectscan be ignored. Calculations show that the buoyancyeffects can be important in the region of interest fortundishes but not for moulds. Therefore, non-isothermalmodels, which are described as ‘thermal’ in the text, arealso developed for tundish flow simulations. For thethermal model, the energy equation is applied with thespecified boundary conditions for heat flux values, asshown in Table 1.

For the buoyancy calculations, the Boussinesq modelis implemented. This uses a constant density fluid modelbut applies a local gravitational body force throughoutthe fluid, which is a linear function of fluid thermalexpansivity b and the local temperature difference withreference to a datum called the buoyancy referencetemperature.10

For the two-phase flow modelling, as also documen-ted in Ref. 11, an Eulerian–Eulerian approach is usedfor the coupling of two phases (liquid steel and argonbubbles). Conservation of continuity, momentum andenergy equations for the argon phase are also added tothe equations and interrelated with the interphasemomentum transfer rate. During the solution of theequations, since the volume of one phase cannot occupyanother, the sum of the volume fraction of the phases isequal to one. Spherical bubbles with a constant diameterare used for the argon phase.12 For the interphase drag,the Ishii–Zuber drag model is used, and the interphasecoefficient and the drag coefficient are set as follows

cab~Cd

8Aabra Ub{Ua

� �(7)

Cd~24

Re

1z0:15Re0:687

� �(8)

Table 1 List of tundish internal design, SEN (submerged entry nozzle) design and simulation parameters

T1 width max. 1300 mm T2 width max. 1200 mm Density of steel 7000 kg m23

Density of argon 0.326 kg m23

T1 width min. 800 mm T2 width min. 500 mm Viscosity 0.0056 N s m22

T1 length max. 7700 mm T2 length max. 6900 mm Surface tension(steel and argon)

1.6 N m21

T1 length min. 7200 mm T2 length min. 6450 mm Tundish topsurface heat flux

230 000 W m22

T1 height – liquidlevel max.

1400–1200 mm T2 height – liquidlevel max.

1200–850 mm Tundish side wallheat flux

25000 W m22

T1 height – liquidlevel min.

1150–950 mm T2 height – liquidlevel min.

1000–650 mm Port up SEN 1 7u up75675 mm

T1 bottom design Two Steps T2 bottom design Single Step Port down SEN 2 15u down75675 mm

T1 impact pad Deep T2 impact pad Shallow Circular roofbottom SEN

80 mm

T1 liquid weight 60 tonnes T2 liquid level 40 tonnes Cup bottom SEN 80 mm15 mm cup

T1 ladle shroudsubmergence

100 mm T2 ladle shroudsubmergence

150 mm Mould width(argon study)

1000 mm

T1 flowrate controltechnique

Stopper rod T2 flowrate controltechnique

Slide gate Mould width(SEN bottom study)

1900 mm

T1 throughput 7 tpm T2 throughput 6 tpm SEN A(flow measurement)

0u 80690 mm

T1 throughput %differences

15% T2 throughput %differences

12% SEN B(flow measurement)

25u down75675 mm

Yavuz Liquid steel flow in continuous casting machine

454 Ironmaking and Steelmaking 2011 VOL 38 NO 6

Re~ra Ub{Ua

�� dp

mm

(9)

Argon is introduced as an injection through the SEN inspherical bubble shape at 2 mm in diameter. The size andshape of a bubble are kept constant when it travels throughthe SEN and the mould domain. The effects of shellformation, mushy zone, slag layer, standing wave atmeniscus, mould level fluctuations, argon bubble coales-cence and argon bubble break-up on the flow structure areneglected. For the top surface of the mould domain(meniscus region), a degassing boundary condition is usedwhere argon bubbles leave the domain, while liquid steel isfree to move on the horizontal axis. The bubbles escape thedomain at the top surface and the mould bottom and reflectat the other domain walls. For the walls, including tundishand mould domain, and inside and outside of the SEN, noslip (Uw50) boundary conditions are used for the liquidphase. For the gas phase when present, the free slip (tw50)

boundary condition is used. At the mould domain exit, theopening boundary condition is imposed, where inward andoutward flows are allowed.

Extensive validation is performed for the CFD modelsthat were developed. Similar strategies for mesh genera-tion, case set-up and solver set-up are used for eachapplication. Reference 13 was primarily dedicated to theextensive validation of CFD models both qualitativelyand quantitatively. Because the focus of the presentreport is not the validation, it is not shown here. TheRANS and unsteady RANS k–e models are quitepowerful in predicting the average velocity field intundishes and mould within reasonable deviation, eventhough they have relatively simple formulation. Owingto this simplicity, the stability of calculations is superior.

Tundish modellingThis part of the study aims to investigate the effect oftundish design, thermal effects and unbalanced (uneven)

1 a simulation domains for tundishes used in CFD modelling and b typical simulation domains for slab casters used in

CFD modelling


Ironmaking and Steelmaking 2011 VOL 38 NO 6 455

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throughput extraction on the tundish flow. The tundishdomains that were used in the CFD models, including theboundary conditions, are shown in Fig. 1a. The specifica-tions of each design are shown in Table 1. Two differenttundishes have been analysed, both with two strand outlets,i.e. outlets 1 and 2. There are characteristic differencesbetween the two designs, as specified in Table 1. Thefollowing four different cases have been investigated forboth tundishes: A, isothermal, even (balanced) throughputextraction; B, non-isothermal (thermal), even (balanced)throughput extraction; C, isothermal, uneven (unbalanced)

throughput extraction; and D non-isothermal (thermal),uneven (unbalanced) throughput extraction.

Figures 2 and 3 represent the velocity vectors on thecentre/symmetry plane for the cases A, B, C and D fortundish 1 (T1) and tundish 2 (T2) designs respectively.The velocity vectors are grey scaled and normalisedwith their length. Thus, the length of a vector does notrepresent any quantitative information whereas the greyscale represents the velocity magnitude. The same stra-tegy to demonstrate the velocity field in the domain isused for Figures 5 and 6.

3 Velocity vectors at centre /symmetry plane for tundish 2 (A: balanced isothermal; B: balanced thermal; C: unbalanced

isothermal; D: unbalanced thermal)

2 Velocity vectors at centre/symmetry plane for tundish 1 (A: balanced isothermal; B: balanced thermal; C: unbalanced

isothermal; D: unbalanced thermal)





When considering the T1 design shown in Fig. 2, theoverall flow structure in the tundish does not show anoticeable variation when including thermal effects (caseB) and/or the difference in throughput extractions fromthe strands (case C). A large scale circulation dominatesthe flow in the tundish for all cases, although a slightdifference is detected on the flow structure near thesidewalls. However, for the T2 design shown in Fig. 3,thermal effects and unbalanced throughput extractionhave significant impact on the overall tundish flowstructure. There are two easily noticeable differences.The first is the large scale circulation region seen on bothsides of the tundish when thermal effects are included.Buoyancy effects cause this kind of circulation. Owingto this large scale circulation, there are regions where theflow is directed towards the centre of the tundish. Thescales of circulation regions are different on both sidesof the tundish when unbalanced throughput is applied,

as opposed to T1 shown in Fig. 2. The side that has alower throughput extraction has a larger circulationregion. The second difference is the overall flow activityat the top surface. For the thermal model, the flow at thetop surface is more active, and the velocity magnitudesare quite high compared to the isothermal model.

To clearly visualise the overall flow structure in thetundish, three-dimensional streamlines are also con-structed. Figure 4 represents three-dimensional stream-lines for the isothermal balanced throughput extractioncase for both tundishes 1 and 2. The streamlines are greyscaled with time. These images are more representativein demonstrating how liquid steel flows in a tundish. Asshown in Fig. 4, the general flow characteristics aredifferent in T1 and T2. In T1, the liquid steel jet comingfrom ladle shroud creates a very strong, large scalecirculation (almost covering the whole tundish), whereasin T2 design, most of the jet coming from ladle shroudbounces back from the impact pad, reaches the topsurface and does not create a large scale circulation;instead, follows its own path and curls from the wall bycreating a spiral-like flow. This is an indication of a highplug volume. It can be inferred from the field that theflow structure in T1 is more stable compared to T2. Thewell defined large scale circulation pattern supports thisidea. Having more stable flow should cause lesssensitivity to the disturbance in the inputs and/orboundary conditions and in turn would result in lessflow variability. This is also confirmed by the resultsshown in Figs. 2 and 3, which illustrate less change inthe overall flow pattern for T1 when thermal effects areincluded, and/or unbalanced throughput extraction isapplied in contrast to the T2 design.

Both tundishes include a single furniture, impact pad.Nevertheless, the overall flow characteristics are quitedifferent. The T1 design seems to be less sensitive tothermal effects and uneven throughput extraction.

Slab caster modellingThis part of the study aims to address the effects ofnozzle bottom design, nozzle port angle and argon

5 Velocity vectors in nozzle at different views for roof bottom circular port design

4 Three-dimensional streamlines scaled with time (s) for

isothermal balanced throughput extraction cases for

tundishes 1 and 2





injection on mould flow. The flow structures induced bydifferent nozzles, which have different port angles andnozzle bottom, were investigated. The specifications ofeach design are listed in Table 1. Figure 5 demonstratesthe grey scaled velocity vectors in the nozzle at differentviews for roof bottom design. This nozzle design hascircular ports. In the images, the effect of sliding gate onthe flow structure is clearly seen. The circulation at thebottom of the nozzle is quite strong. Circular jet exitsthe port. Port efficiency is quite high, which would po-tentially reduce port clogging.

Figure 6 shows the mould flow structure for thecorresponding nozzle and the same nozzle design withcup bottom. The velocity vectors on the centre planeparallel to the broad face and the vorticity contours atthe meniscus are shown. Both nozzles generate a doubleroll flow pattern in the mould. Detailed visualisationreveals that the flow characteristics are quite different.Jet expansion in the mould (shown with grey arrow) issignificantly wider with the roof bottom nozzle. This isprimarily due to the fact that strong jet circulation at thebottom of the nozzle expands into the mould domainand spreads three-dimensionally. In addition, it could bethe indication of high fluctuation of port jet in thecasting direction (up and down), which creates anexpanded jet in the mould in time averaged patterns.This is also supported in water model test results.14 Asexplained in Ref. 14 for this design, high mould flowfluctuations are evident, which in turn cause switching ofmould flow from single roll to double roll flow pattern.

When considering the vorticity levels at the meniscus,quite different patterns are witnessed. For the cupbottom design, traditional vorticity patterns having highlevels of vorticity near the nozzle are obtained. Thesepatterns are quite typical and potentially cause tradi-tionally known ‘vortex formation’ and slag entrainment.Intermediate levels of vorticity are also evident with roofbottom design at the same locations. However, thehighest levels of vorticity for the roof bottom design areobtained in a very unusual place at meniscus, which isclose to narrow faces. This is most likely due to the jet

behaviour coming out from the nozzle ports. The nozzlebottom (roof), the port shape (circular) and havingbiased flow inside the nozzle due to the sliding gate areall contributing factors to this behaviour.

Another study regarding the application of CFD onmould flow addresses the effects of port design andargon injection through the nozzle on mould flowstructure. Figure 7 shows the greyscaled velocity vectorson the centre plane parallel to the broad face of themould for two different nozzle designs, i.e. port up SEN1 and port down SEN 2, with and without argoninjection. The images are cropped to be shown in thesame figure and do not demonstrate the whole domainlength used in the simulations. This portion of the studywas also documented in Ref. 11. As shown at the toprow of Fig. 7, when no argon is injected, the standarddouble roll flow is evident in the mould. Although theport orientations are different, the liquid steel jet at theport seems to have a similar angle for each SEN design.The meniscus velocities with port down SEN 2 aresignificantly higher than the ones obtained with port upSEN 1. For port down SEN 2, the liquid jet is verypredominant at the bottom section of the port, which inturn creates a back flow region on the top portion of theport. However, for port up SEN 1, in addition to the jetat the bottom portion of the ports, jet expansion isevident. That expansion results in having no back flowregion on the top portion of the port. This causessignificantly low velocity magnitudes at the meniscus.

When argon is injected, the mould flow structurestransform to different patterns compared to the onesobtained under the absence of argon. Particularly, themeniscus region clearly indicates significant variationsand contains two counter rotating circulations for bothdesigns. Argon bubbles affect the liquid phase in similarways for both SENs. The argon dampens the upperloop. The upper circulation gets smaller in scale anddeteriorates according to its strength in single phaseflow. Thus, for port down SEN 2, since the meniscusvelocities in single phase flow are relatively highercompared to port up SEN 1, the deterioration of the

6 Velocity vectors on centre plane parallel to broad face and vorticity contours at meniscus for roof bottom and cup bot-

tom nozzle designs




upper loop is less when argon is introduced. Unlike theno argon cases, a significant difference is noticedregarding the liquid jet stream angle and the impinge-ment points on the narrow face when gas is injected.However, the velocity magnitudes at the meniscus donot show significant variations.

This part of the present study shows how critical it isto choose what nozzle to use in a slab caster. Smallvariations in both port angle and bottom design couldinduce significant changes on the mould flow pattern. Inaddition, argon flow is another parameter that has asignificant effect on mould flow; thus, the flowrate has tobe adjusted cautiously.

Mould flow measurement

Concept of measurementObtaining continuous, real time flow measurements atmeniscus could help achieve many fundamental objec-tives in the area of continuous casting. These include,but are not limited to, understanding the effects ofcasting parameters on meniscus velocity, evaluating andvalidating CFD and water models, obtaining correlationbetween meniscus velocity and defects and updating thetheory on how to optimise the mould flow. Directvelocity measurement in a mould is not possible due tothe extreme temperatures in the measurement domain.Therefore, indirect measurement techniques need to beimplemented. These techniques are usually simple innature but require a deep fundamental knowledge inorder to be properly applied. In this part of the paper,one of the indirect flow measurement techniques alongwith its challenges and implementation has beenintroduced.

Submeniscus velocity measurements were performedwith the torque measurement technique. The concept ofmeasurement is shown in Fig. 8. The idea is to measurethe torque generated by the drag force due to liquid steelflow. Special probes were dipped into the liquid steelthrough the mould slag layers. The liquid steel flowapplies a drag force to the probe with a well knownconcept, i.e. ‘force on a solid body due to drag’. Thefollowing equation shows the relationship between drag

force and velocity

Fdrag~1

2rCDAV2 (10)

where V is the average velocity of the liquid steel facedwith the submerged portion of the probe, A is theprojection area of the submerged probe that is themultiplication of probe submergence and probe dia-meter, CD is the drag coefficient that depends on theflow regime and the shape of the probe and r is theliquid steel density. Water model tests are used to obtainthe most accurate value of the drag coefficient. This ispreferable but not required, since the drag coefficientsfor simple geometries are well known and can beobtained from the literature.

The drag force, as described above, generates atorque, which is equal to the following

T~FdragL (11)

where L is the length between the position of the forceapplied and the axis of the measurement. By combiningboth equations and leaving the velocity on the left handside, the following relation for the velocity from themeasured torque values can be obtained as

V~2T

LrCDA

� �1=2

(12)

As shown in Fig. 8, two measurement devices (MD)were used in the plant trials. Each one was used on one

8 Sketch summarising measurement principle

7 Velocity vectors on centre plane parallel to broad face of mould for port up SEN 1 and port down SEN 2 with and

without gas injection11





side of the SEN. With this, the velocity informationfrom both sides of the SEN was obtained. It is importantto mention that this technique provides a single pointmeasurement, which corresponds to an average velocityof the flow in the area of the submerged length.

Since the slag/liquid steel interface creates a significantamount of wear on the probe, the submerged portion ofthe probe into the liquid steel could easily be detectedafter the trials when the probes are removed from themould. This helps to recalculate the velocity values withthe actual probe submergence values after the trial inorder to obtain more accurate results.

Measurement challengesThe measurement concept is fairly simple; however, thedifficulty comes with its application. Figure 9 shows themeasurement chain and the list of challenges inperforming the measurement. The measurement chainfollows the sequence shown in the figure. The signalcoming from the sensor is conditioned first before it issent to the data acquisition board. In addition, digitalfiltering is applied before the real time conversion oftorque to velocity is performed. The real time velocitymeasurements from both sensors are displayed duringthe measurements. The sensor and the signal condition-ing units were designed and developed by ArcelorMittalGlobal R&D Maizieres.15–17

There are four main challenges that are encounteredduring the measurements. These are probe life, verti-cality, safety and noise. Probe life is an importantparameter since the measurement time depends on it.The longer the probe life, the longer the measurementtime without interruption would be. There is a timerequired to replace the probes. In addition, each probechange causes disturbance in both process and meniscusflow. Thus, it is desired to keep the probe life as long aspossible. However, there is a cost associated with this.The life of the probe primarily depends on the slag wearrate at the slag/liquid steel interface. Selecting theappropriate material for the probe is one of the mostcrucial steps. In addition, the bigger the probe diameter,

the longer the probe life would be. However, since thethickness of a standard slab caster varies between 8 and10 in, there is a limit in selecting the diameter of theprobe. This limit depends on many things, but the mostimportant ones are the blockage effect and thedisturbance generation.

Another challenge of the measurement is the conceptof verticality. The measurement is performed to measurethe torque generation by the drag force due to liquidsteel flow. The force and torque values are small,particularly at low flow velocities. If the probe couldbe located perfectly straight (perpendicular to the mouldlevel), the only contributor to the torque generationwould be the liquid steel flow. Thus, there would not beany measurement error. However, practically, this is notpossible. When there is any angular misalignment,gravitational and buoyancy forces start generating‘artificial’ torque on the measurement axis that couldbias the measurement significantly. This generationdepends on the misalignment value, density of the probeand submerged volume. The percentage error dependson this generated torque relative to the torque generatedby the liquid steel. Sample calculations have beenperformed to quantify this error. As shown in Fig. 9,the per cent error at different liquid steel velocities iscalculated when 1u misalignment is present. At low flowvelocities, even a 1u misalignment contributes to asignificant error, which can be up to 65% in themeasurement. Thus, proper countermeasures need tobe taken to make sure that the alignment is as good aspossible. In addition, under any circumstances, velocities,0?05 m s21 are questionable. It is also important tomention that angular deflection of the torque sensor(due to the sensor’s working principle) is a contributorto the error associated with non-verticality. Thus, itneeds to be measured and accounted in the calibrationprocess as well.

The challenges regarding safety and noise are verytypical in any steel plant measurement and are notdiscussed here due to space limitation.

9 Measurement chain and measurement challenges




Results of measurementsDifferent casting conditions, including casting speed,argon flowrate, nozzle design and nozzle submergence,were tested. It was aimed to investigate the effects ofcasting parameters on mould flow structure and productquality. The measurements were performed at the quartermould width location for both sides of the nozzle.

Figures 10–13 were constructed with the followingapproach. The positive sign for measurement device 1(MD1) shows that the flow goes towards the SEN,whereas the negative sign for MD2 represents the samebehaviour. The velocity data shown here correspond tothe values of 25 s moving average. The x axis is specifiedas time. The primary y axis is for the meniscus velocityvalues for both measurement devices. The secondary yaxis represents the indexes of casting parameters, whichare submergence and cast speed in Fig. 10, gate positionin Fig. 11, argon flowrate in Fig. 12 and argon flowrateand cast speed in Fig. 13. The boxes (slabs), when shownon the top of the figures, represent the quality feedbackfor the corresponding slabs that are in the mould at thetime specified in the chart. A certain threshold decideswhether a slab is black or grey. Grey is considered ‘ok’,and black is considered ‘not ok’. The measurementswere performed for sliver sensitive grade. Thus, the greyand black boxes represent the sliver index coming fromthe coils of the corresponding slabs. Figures 10–12represent the results for nozzle design A, and Fig. 13 isfor nozzle design B. The design differences are describedin Table 1.

As seen in Fig. 10, there is a significant correlationbetween cast speed and meniscus velocity. At highcasting speed, the flow is double roll, thereby themeniscus moves towards the SEN (z for MD1 and 2

for MD2). The flow is usually symmetric. The reductionin casting speed causes a significant drop in meniscusvelocities and an increase in velocity fluctuations. Thissometimes generates bias or single roll flow patternhaving very low velocity magnitudes. Ramping up to thehigh casting speed shows recovery in the meniscusvelocity magnitudes and in the overall flow pattern. Theindication of the reductions and increases in cast speedcan be clearly seen on the meniscus velocity measure-ments. Since the mould width was kept roughly constantin each case, any remarks that are made for the castspeed is valid for the throughput as well. The flow seemsvery stable when the mean values are in the range of 20–40 cm s21. No significant correlation is noticed betweenthe quality data and the meniscus velocity.

A different characteristic was tried to be tracked in themeasurements shown in Fig. 11. Only the measurementshown in Fig. 11 was conducted later at the sequence,i.e. it did not start with a clean SEN. During this test, asignificant increase in the gate index is detected. It isimportant to mention that the cast speed and the tundishlevel are constant during the time span of thismeasurement shown in Fig. 11. Based on the measure-ment results, the flow at the meniscus is completelydisturbed and is biased all the time. The velocitymagnitudes are quite different from one side to theother side of the SEN. The velocities on one side of theSEN could sometimes reach a value of 60 cm s21, whichis considered extremely high for meniscus flow. Eventhough at the end of the sequence single roll flow isapparent, the velocity magnitudes still show significantvariation for each side of the nozzle, i.e. 10–30 cm s21.No symmetry is obtained at any time. These 50 min datashow one of the worst flow structures that could beachieved. During this measurement, it was also noticedthat there was a pressure loss in one of the argoninjection/sealing sources. Thus, there are two possibleexplanations for this kind of flow behaviour at themeniscus, which are air aspiration in the system andnozzle clogging. It is quite surprising to see that there isno correlation between the meniscus velocity and thequality feedback.

Figure 12 demonstrates how the meniscus velocityreacts to the argon flow change. The velocities on oneside of the nozzle seem to be insensitive to the changes in

12 Meniscus velocity measurements and casting para-

meters for casting sliver sensitive grade with SEN A

11 Meniscus velocity measurements, casting parameters

and quality feedback for casting sliver sensitive grade

with SEN A

10 Meniscus velocity measurements, casting parameters

and quality feedback for casting sliver sensitive grade

with SEN A






the argon flowrate. For that corresponding side, i.e.MD1, the meniscus velocity is within the range of 20–30 cm s21 and quite stable. However, the other sidereacts to the argon flowrate change dramatically. At lowargon flowrate, MD2 shows similar values with MD1.The overall flow pattern is seen as double roll andsymmetric. When the argon rate increases, the deteriora-tion of the meniscus flow on the MD2 side is evident.The higher the rate, the more it deteriorates. At themaximum rate, very low meniscus velocities are achievedon the corresponding side. For that period, it can beconcluded that the overall meniscus flow is biased.Reducing the argon flowrate causes a recovery of themeniscus velocity for that side of the nozzle. The overallflow pattern turns into a strong double roll once againwhen low argon injection is maintained. This SENdesign is more prone to biased flow.

Figure 13 shows the results of meniscus velocitymeasurements when different casting speeds and argonflowrates are tested for SEN B. In this case, the meniscusflow seems predominantly symmetric. The only excep-tion happens at very low casting speed. The flow isstrong, stable and double roll most of the time withmeniscus velocities in the range of 20–30 cm s21. Whenthe argon flow is suddenly increased to high values, themeniscus velocities significantly change, and the mea-surement values drop to 0–10 cm s21. Single roll flowbecomes evident at the highest argon flowrate. Loweringthe argon flow causes a recovery of the meniscus velocityon both sides of the SEN. Further reduction in the argonflowrate increases the meniscus velocities. The suddenchanges in the argon flowrate and therefore in themeniscus velocities are noticed as spikes in the mouldlevel signals, shown as red circled areas at the top ofFig. 13. A similar trend was obtained in many othermeasurements but not as dramatic as the case demon-strated. This means that sudden changes in the meniscusvelocities are also detected with the mould level sensors.Gradual but significant changes in the meniscusvelocities have also a footprint in the mould level signal,as shown in Fig. 13. This indication is not as significantas the ones obtained in sudden changes; however, it isstill apparent.

Based on all the measurements conducted, Fig. 14was constructed and shows the relationship betweenargon flowrate and throughput in order to reach a

favourable meniscus flow for different SEN designs. Afavourable meniscus velocity is considered when themeniscus flow is symmetric, double roll and havingrelatively less fluctuations. Thus, these three conditionshave to be met in order to consider the meniscus flowfavourable. The argon flowrate on the y axis shows themaximum allowable argon flow at the correspondingthroughput value. It shows the upper limit of the argonflow. That limit can shift according to the nozzlesubmergence. Basically, Fig. 14 suggests that there is amaximum argon flowrate beyond which the meniscusflow is not favourable. This rate is not only a function ofthroughput but also a function of nozzle design andnozzle submergence. It is expected to have a similartrend for the lower limit, which would probably besupported by the nozzle clogging index. Thus, thereshould be an area in the throughput versus argonflowrate plot where the favourable flow is obtained foreach nozzle design. The optimum meniscus flow shouldbe somewhere in that favourable area. This optimummeniscus flow is expected to be supported by the qualitydata. A well established quality feedback and morevelocity measurements should help to determine theoptimum operating conditions along with the optimummeniscus flow. A compressive study is needed tocomplete this part of the study.

It is important to mention that the correlationbetween the product quality and the meniscus flowcould not be found in the present study primarily due tothe fact that the product quality feedback was notsufficient both quantitatively and qualitatively.

ConclusionsIn this paper, an overview of the liquid steel flow incontinuous casting machines is provided, and thechallenges and applicability of computational modellingand real time flow measurements are discussed. Singlephase and two-phase CFD models for the tundishes andthe moulds were developed for various ArcelorMittalcontinuous casting machines. The effects of variousoperational parameters, including tundish and casterpractices on the flow structure, are investigated. Inaddition, real time submeniscus velocity measurementswere performed with the concept of torque measure-ment. The concept and challenges of the technique areintroduced, and the effects of casting parameters onmould flow structure were investigated. Some of themajor findings are the following.

14 Relationship between argon flowrate and throughput

to reach favourable meniscus flow for different SEN

designs

13 Meniscus velocity measurements and casting para-

meters for casting sliver sensitive grade with SEN B





Two different tundish designs, i.e. T1 and T2, wereconsidered for the CFD study of tundish flow. Theeffects of tundish design, uneven throughput extractionand temperature on tundish flow structure wereinvestigated. Even though both designs include singleinternal structure and impact pad, the general flowcharacteristics are quite different. In T1 design, theliquid steel jet coming from ladle shroud creates a verystrong, large scale circulation, whereas in T2 design,most of the jet coming from ladle shroud bounces backfrom the impact pad, reaches the top surface and doesnot create a large scale circulation; instead, follows itsown path and curls from the wall by creating a spiral-like flow. The flow structure of the T1 design isinsensitive to thermal effects and uneven throughputextraction from strands, whereas for the T2 design,thermal effects and uneven throughput extraction have asignificant impact on the overall tundish flow structure.

Four different SEN designs were considered for theCFD study of mould flow. The effects of circular port,roof and cup type bottom designs, port orientation, portangle and argon flow on the mould flow structure wereinvestigated. Circular ports along with a roof bottomdesign generate highly unsteady, expanded jet in themould, which in turn causes high levels of vorticity atmeniscus very close to the narrow faces. Argon injectioncauses significant deterioration of flow structure at themeniscus. This deterioration depends on the strength ofthe upper loop at the single phase flow condition whenno argon is injected. Owing to the fact that port downSEN provides a stronger upper loop in the absence ofargon compared to port up SEN, argon injection causesrelatively less deterioration of meniscus flow with portdown SEN when argon is introduced.

Two different SEN designs were considered for themould flow measurement study. Different castingconditions, including mould width, nozzle submergence,cast speed and argon flow, were tested. Significantcorrelations were found between throughput, nozzlesubmergence, nozzle design and meniscus velocities. It isevident that the throughput and the argon flow havemajor effects on the meniscus velocities. There is amaximum argon flowrate beyond which the meniscusflow is not favourable. This rate is not only a function ofthe throughput but also a function of the nozzle designand the nozzle submergence. In addition, any significantchange in the meniscus velocities causes significantchanges in the mould level.

Acknowledgements

The author would like to express his many thanks toB. Chukwulebe, D. White, M. Ozgu, J. Sengupta,J. Thacker, A. Elnenaey, V. Gueugnon, M. Atkinson,D. Kruse, S. Kinkel, J. Bradley, R. Kostyo, D. Sena,M. Alavanja, D. Idstein, S. Chakraborty, S. Kipp,C. Kennedy, S. Schreiner, J. F. Domgin, T. Tsai andG. Lawson from ArcelorMittal for their support.

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liquid steel flow in continuous casting machine: modelling and measurement

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