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Simulation of Fluid Flow and Oscillation of the Argon OxygenDecarburization (AOD) Process

HANS-JUERGEN ODENTHAL, UWE THIEDEMANN, UDO FALKENRECK,

and JOCHEN SCHLUETER

The oscillation of argon oxygen decarburization (AOD) converters is flow related and dependson the process parameters (e.g., vessel geometry, melt fill height, process gas type and blowingrate, vessel tilting angle, as well as geometry, number, and arrangement of the side-wall nozzles).For a 120-ton AOD converter with seven submerged side-wall nozzles, plant tests, physicalsimulations on a 1:4 scale water model, and computational fluid dynamics simulations havebeen done. The investigations show that the penetration depth of an inert gas jet into the meltdoes not exceed approximately 0.4 m. The plumes are located close to the nozzle-side converterwall and induce a large-scale primary vortex as well as intensive surface movements; both areresponsible for the oscillation. Several process mechanisms were investigated. The oscillation ishighest in the last stage of the dynamic blow and is still high during the reduction stage. As theamount of inert gas increases, the vibration level also increases. Inert gas has a greater influenceon the oscillation than oxygen. Tilting the converter around 8 deg clearly leads to moreintensive oscillations. Increasing the blowing rate increases the forces and torques acting on thevessel, whereas the oscillation frequency remains nearly constant. A varying fill level does notinfluence the vibration level the same way as the blowing rate. The operational test shows, for

example, that the maximum torque does not depend on the heat size when the latter variesbetween –8 pct and +21 pct of the nominal heat size. The water model test shows decreasingforces and torques with a rising fill level.

DOI: 10.1007/s11663-009-9335-y The Minerals, Metals & Materials Society and ASM International 2010

I. INTRODUCTION

IN 2008, around 26 million tons of stainless steelwere produced around the world. Stainless steel containsa minimum of 10 pct chromium and various amounts of

Ni, Mo, Mn, and other elements to arrive at the desiredmaterial properties. The corrosion resistance is causedby a thin layer of passive, stable chromic oxide, whichreacts slowly. The layer adheres extremely well and,when combined with oxygen from the air, is self-healing.

To produce stainless steel, a two-stage process isemployed. Scrap and alloying elements, sometimestogether with hot metal, are molten down in the electricarc furnace. Then the molten metal is refined in theargon oxygen decarburization (AOD) converter. In theAOD process, a highly chromium-alloyed melt isdecarburized by injecting and blowing oxygen and inertgases (N2, Ar) through submerged side-wall nozzles anda top lance. Depending on the mix of charge materials,

the beginning carbon contents are between 1.5 pct and4.5 pct, and the chromium contents are between 10 pctand 25 pct, depending on the steel grade to be produced.

A specific feature in the decarburization of high-chromium melts is that chromium itself has a high

affinity to oxygen, which requires special process-relatedmeasures to limit its conversion into the slag. Thecompetitive situation of carbon and chromium incontact with oxygen is demonstrated by the followingreaction equations:

2½C þ fO2g ! 2fCOg ½1

4½Cr þ 3fO2g ! 2ðCr2O3Þ ½2

where [ ] means it is dissolved in the melt, ( ) means itis included in the slag, and { } means it is gaseous. Atambient pressure, for every temperature, an equilib-rium exists between the carbon and chromium content

of the melt and the oxygen supplied (Richardson– Ellingham diagram). As the temperature increases, thefree reaction enthalpy DG0 for Eq. [1] decreases,although it increases for the slagging reaction inEq. [2]. Equations [1] and [2] can be combined as follows:

2½Cr þ 3fCOg $ 3½C þ ðCr2O3Þ ½3

Equation [3] makes it clear that the equilibriumbetween carbon and chromium in the melt depends onthe CO partial pressure in the gas bubbles. To supplyas much oxygen as possible to the carbon while at the

HANS-JUERGEN ODENTHAL, Deputy General Manager,R&D Division—Fundamentals and Models Melting/SAF, UWETHIEDEMANN, Deputy General Manager, Steelmaking/ContinuousCasting Technology Division—Product Development Steelmaking,UDO FALKENRECK, General Manager, R&D Division, andJOCHEN SCHLUETER, Vice President Special Technologies, arewith the SMS Siemag AG, Eduard-Schloemann-Straße 4, 40237Duesseldorf, Germany. Contact e-mail: [email protected]

Manuscript submitted May 5, 2009.Article published online January 26, 2010.

396—VOLUME 41B, APRIL 2010 METALLURGICAL AND MATERIALS TRANSACTIONS B

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same time minimizing the oxidation of the chromium,the reaction must proceed toward the left. Thissequence takes place preferably at high-carbon andlow-chromium contents, at high temperatures, andabove all, at a low-CO partial pressure pCO. Increasingthe process temperature is only reasonable up to1700 C because of excessive refractory wear at highertemperatures. The critical carbon content that marksthe lower limit where chromium starts to scorify as afunction of pCO is calculated as follows:

cpctC

¼

ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiK pCOf g

3 f Cr½

2cpctCr 2

aCr2O3 f c½ 3

3

v uut ½4where c is the pct per weights of carbon, K is the equi-librium constant, a is the activity, and f is the coeffi-cient of activity. The CO partial pressure in the gasbubble is expressed as follows:

pCO ¼ nCO

nCO þ nAr

pz ½5

where n is the molar mass and pz is the static pressure inthe melt at depth z.

The AOD process uses the dependence of the Cr–Cequilibrium on the partial pressure by breaking theprocess down into various phases and progressivelyadding more inert gas to the process gas as the carboncontent decreases. This process reduces the CO partialpressure so that the decarburization toward low carboncontents is promoted, while at the same time theoxidation of chromium is limited. However, for kineticreasons, the scorification of chromium cannot be fullyavoided. To reduce the loss of chromium, the input of oxygen at low-carbon contents of less than, for example,0.4 pct is, depending on the chromium content of themelt, increasingly reduced because the achievable decar-

burization rate here is a function of the melt bathcarbon-content proper and, moreover, chromium scori-fication increases because of the thermodynamic equi-librium situation.

Figure 1 shows the typical blowing stages of a 120-tonAOD converter with seven side-wall nozzles. Down to acarbon content of approximately 0.4 pct, a total of _V O2 = 240 m

3/minstp is blown for refining, with thelarger proportion of oxygen blown through the single-hole top lance. The decarburization rate is proportionalto the oxygen supplied. The maximum amount of oxygen is limited by the steel grade, the startingconditions, the oxygen supply, the waste gas exhausted,and the slag slopping. At this stage, the low amount of inert gas through the side-wall nozzles serves forcooling. In conformity with the laws of kinetics andthermodynamics at decreasing carbon contents, theamount of oxygen is reduced in stages, and at the sametime, the process gas is enriched with inert gas (dynamicblow). The aspect ratio of oxygen and inert gas dropsfrom the original 8:1 via 1:1 and 1:2 down to 1:3.3, withthe total amount of process gas remaining almostconstant during the dynamic stages. At the end of decarburization period, the carbon content is roughlyaround 0.02 pct for a lot of stainless steel grades.

Decarburization is followed by a pure inert gasstirring for reduction. During the reduction stage, theoxidized metals—chromium, above all—are reducedmostly by the addition of silicon carriers. In thisprocess, the initial Cr2O3-containing refining slagchanges from a solid to a liquid state.

Reduction and subsequent deslagging are followed bya separate desulphurization using a two-slag method.The basicity of the slag and, thus, the sulphur capacity isincreased by adding lime and fluxing elements, with

which simultaneous inert gas stirring causes an almostcomplete transition of sulphur from the melt into theslag. The last inert gas stirring stage (alloying) providesan opportunity for analysis trimming of the melt beforetapping.

The AOD converter is a metallurgical reactor thatoffers excellent mixing conditions because of the highmelt turbulence caused by the injection of large amountsof process gas through the side-wall nozzles. However,the AOD process is accompanied by intense convertervibrations, which from the outside can be recognized bymore or less staggering movements of the vessel aroundthe axis of rotation. Vessel vibrations act as torques on

the structural components and have to be taken intoaccount during the design phase, especially with regardto the dimensioning of the torque support arm.

Figure 1 also shows the vibration intensity through-out the process schematically as a red curve. The carboncontent of the starting melt was around 3 pct. Althoughthe highest amount of process gas is injected during themain blow, the vibration intensity during this stage isrelatively low. It is typically already during the mainblowing period that a clear reduction of the vibrationamplitude is to be noticed, which can be attributed todamping effects caused by the increasing weight of theslag with high-slag viscosity. Another reduction of the amplitudes usually is observed toward the end of

the first blowing stage, which can be explained by areduced decarburization rate and the decreasing forma-tion of CO bubbles. Similar phenomena were observedin vibration measurements on basic oxygen furnace (BOF)

Fig. 1—Blowing pattern of the 120-ton AOD converter with charac-teristic vibration intensity levels.

METALLURGICAL AND MATERIALS TRANSACTIONS B VOLUME 41B, APRIL 2010—397

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converters as well.[1 – 3] The authors explained thisbehavior by reducing the decarburization rate in favorof an increased iron scorification while approaching thecritical point of decarburization. The general idea wasthat strong oscillations were accompanied by an inten-sive mixing of the melt.

The part the top lance that contributes to theexcitation of vibrations is negligible when comparedwith the part supplied by the side-wall nozzles. Thedecisive factor in the excitation of vibrations is the

volumetric flow rate of the process gases. However, inertgas has a much stronger influence on the vibration thanoxygen. The higher the amount of inert gas, the strongerthe vessel vibrations become. The different effect causedby oxygen and inert gas is explained by the differentbehavior of these gases immediately after they havepenetrated the melt domain.

When injected into the melt, the oxygen jet immedi-ately collapses, with most of the oxygen first oxidizingchromium and then, because of the reduction of Cr2O3in the presence of overcritical carbon contents, reactingto form CO bubbles (see Eq. [3]). In contrast, inert gasbubbles have more time to grow after penetrating the

melt. Consequently, a mechanism must be assumed toexist in the formation of inert gas bubbles, withcorresponding effects on the bubble size and theirspatial distribution, which is much different from themechanism existing when injecting pure oxygen. Duringthe dynamic decarburization stage, the conversion of oxygen to CO additionally increases the vibration levelinduced by inert gas, especially when the decarburiza-tion rate is high, and a large amount of oxygen isconverted to CO.

In addition to the amounts of process gas, therefractory lining degree of wear influences the vibration.As the degree of wear increases, the bath geometry andthe center of gravity of the converter vessel change.

Thus, the detected torque shows a maximum at the startof a converter campaign, which decreases typically byaround 25 pct toward the end of the campaign.

Vibrations are caused by geometry conditions, pro-cess gases, rising gas bubbles (N2, Ar, CO), and theresulting movements of the free surface (i.e., vibrationsare closely linked with fluid flow and heat phenomena).Process variables that influence the flow conditions arethe converter geometry—mainly characterized by theratio of the bath diameter and fill level—the type andflow rate of process gases, the geometry and alignmentof the side-wall nozzles, and the inclination of theconverter vessel. The interaction of all variables has, tothis date, not been understood in detail. This is becauseof both the multitude of variables and the varyingboundary conditions prevailing in a melt shop, whichfrequently make it difficult to carry out a systematicinvestigation and to achieve reproducible results.

For reliable operation, a long service life of the plantand its components as well as optimal process condi-tions and knowledge about the factors influencing thevibration behavior are indispensable. Against this back-ground, SMS Siemag, in addition to worldwide vibra-tion analyses on existing AOD converters, also performsphysical simulations on water models and numerical

simulations on computational fluid dynamics (CFD)and the finite element method (FEM). A variety of theseinvestigations is presented in the following sections.

II. LITERATURE

Most of the authors carry out investigations on AODwater model converters,[4 – 10] using the modified Froudenumber for similarity criterion. The model scales vary

between 1:10[9] and 1:4.[10] Investigations on the pene-tration length of gas jets into liquids were made byBjurstro ¨ m et al.,

[2] Fabritius et al.,[6] and Tillianderet al.[9]

Bjurstro ¨ m et al.[4] performed laser Doppler anemom-

etry measurements on a water model converter. How-ever, the gas bubbles considerably complicated themeasurements. When the gas flow rates were high, onlya few velocity vectors characterized the flow field.According to the authors, the penetration length of thegas jet depends much more on the blowing rate than onthe fill level. With a low blowing rate, a circulating flowresults in the symmetry plane. As the blowing rate

increases, the area of recirculation is pushed away in thedirection of the wall located opposite the nozzle.Fabritius et al.[6] investigated the oscillation of the

free-water surface. Regular oscillations occurred whenthe gas jet reached up to the center of the bath, and theheterogeneous buoyant plume fluctuated uniformlyback and forth (type A oscillation). In case of a highblowing rate, the gas jet in the water model reachedbeyond the center of the bath, and the main flow wasreversed — although this result is not likely to occur inreality. With the same blowing rate, the penetrationlength with a small nozzle diameter is higher than witha large diameter. The dominating frequency duringthe decarburization stage of the original converter

(O2/N2 = 1:3) was 1.45 Hz. The frequency was inde-pendent of the blowing rate, nozzle diameter, and theangle between the nozzles.

Fabritius et al.[11] conducted vibration analyses on areal AOD converter. An acceleration transducer wasinstalled on one of the side-wall nozzles. The frequencydetected depended only on the height-to-diameter (H/D)ratio of the melt bath but not on the blowing rate or gastype. With an increasing H/D ratio, the frequency roseslightly and was constant for H/D > 0.45. The wornconverter (shallow bath and/or small H/D) induced flatsurface waves with a low frequency and long periods,whereas the freshly relined converter (high bath and/orhigh H/D) caused higher frequencies. For a low fill level,the converter bottom dampened oscillations at the phaseboundary. Vibrations were determined mainly by theway in which the converter was operated and, to a lesserextent, by the decarburization rate. The higher theblowing rate and the proportion of inert gas, the higherthe vibration amplitude (with unchanged frequency).Likewise, the amplitude increased with a rising fill level.With a constant flow rate of the process gas, theinjection of inert gas produced more vibrations than theoxygen injection (i.e., a correlation was observedbetween the length of penetration and the vibrations).


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The argon jet (nAr = 39 kg/kmol) was longer than thatof oxygen (nO2 = 32 kg/kmol), and consequently, theamplitudes were higher. According to Fabritiuset al.,[6,11] the vibrations measured for nitrogen (nN2 =28 kg/kmol) were higher than for argon, which may bebecause the maximum amplitude without a root-mean-square (rms) value was considered. Thus, no statisticalevaluation of the measurement was provided. Theauthors found the maximum amplitude in the processstage at _V O2þN2 ¼ 146m

3= minstp and O2/N2 = 1:3, which

coincides with our own operational measurements.Tilliander et al.[8,9] performed numerical and physical

simulations on the converter and on the nozzle. Thepenetration length of the gas jet did not extend tothe center of the bath, and gas bubbles rose close to thenozzle wall. The mean penetration length of the gas jetas quoted by Fabritius et al.[6] and Tilliander et al.[9]

was higher than the one indicated by Hoefele andBrimacombe.[12]

Different flow rates, nozzle diameters, angles betweenthe nozzles, and their influence on the mixing behaviorwere investigated by Fabritius et al.[6] and by Weiet al.[10] The mixing time was determined by a conduc-

tivity measurement after adding salt—this method wasinaccurate because of the influence of a variable density.The mixing time decreased with rising blowing rate andlarger angles between the nozzles.

Many vibration analyses also have been made onBOF converters. Grenfell et al.[1] evaluated the signalsfrom load cells installed in the bearing pedestals of 130-ton BOF converters. The authors made a correla-tion between the decarburization rate (i.e., the CO gasbubble formation rate in the bath and the vibrationamplitudes of the measured signals). This correlationwas used to determine the end-of-blow point based onthe vessel vibration and the carbon content.

Onishi et al.[3] investigated the vibration of a com-

bined blowing 250-ton Kawasaki-basic oxygen process(K-BOP) converter. They found that a clear reduction of the vibration amplitudes took place toward the end of the blowing process because the vibration amplitudedepended on the gas-bubble formation rate in the meltbath. Because the decarburization rate decreased at theend of the blowing stage in favor of progressing metalscorification, fewer gas bubbles were generated, whichled to a reduced melt movement and a lower vibrationlevel.

Mucciardi et al.[2] seized on the results of Grenfellet al.[1] and Onishi et al.[3] An oxygen blowing testperformed on a hot-metal charge, and for which anacceleration transducer was used, confirmed the findingsof both teams of authors.

III. AOD CONVERTER

AOD converters today are designed mostly as change-vessel units. This design provides for the suspensionsbetween the vessel and the closed trunnion ring withoutwater cooling to be released, the process and inert gaslines to be disconnected, the gas stack and lance systemto be taken to the change position, and the converter

vessel to be lifted out of the trunnion ring by the shopcrane. In contrast, U-shaped, water-cooled trunnionrings combined with a converter change car are ideal forconverters installed between columns. The trunnion ringis equipped with a tilt-drive system and hydraulicequipment.

A. Geometry and Process Data

Table I and Figure 2 show the nomenclature and

simplified rotational symmetric geometry of the 120-tonAOD converter on which operational tests were per-formed in the melt shop and for which both physicalsimulations on a water model and CFD simulationswere carried out. The diameter-to-height ratio of themelt bath was D=H st ¼ 1:58 at D ¼

D1þD2ð Þ2

. The processgases were injected through seven side-wall nozzlesarranged at the level H n. Each nozzle consisted of acoaxial inner tube (Dn,in) for the process gas (O2, N2, Ar)and a shroud tube (Dn,out) for the cooling gas (N2, Ar).The nozzles were arranged at an angle of an = 18 deg,with the horizontal nozzle axes aligned relative to thevertical centerline of the AOD converter. The distancebetween the refractory bottom and the trunnion axiswas H t. The parameters investigated referred to thestandard operating point of the reduction phase with ablowing rate of _V Ar = 120 m

3/minstp (blowing rate100 pct) and a fill level of H st = 2.096 m (fill level100 pct). In melt shop practice, the torque was measuredwith the help of strain gauges (DMS) arranged on thetorque support arm.

Figure 3 shows the water model at the Institute forIndustrial Furnaces and Heat Engineering of theRWTH Aachen University.[13] The model consisted of 40-mm-thick acrylic glass components (Figure 3(a)).The vessel was installed in an aluminum frame andconnected to the latter with a tilting axis and a trunnion

ring. The torque support arm used was a square solid-material bar (Figure 3(d)). The lower part of the vesselcontained the nozzles, which were adjustable in allspatial directions (Figure 3(e)). The flow rate of everynozzle was controlled by a float-type flow meter. Thewhole vessel could be rotated around the trunnion axis.

Nozzle 4 additionally had a pressure sensor and athermocouple. With the help of two ultrasonic sensorsarranged above the water surface, the oscillation wasmeasured with a resolution of 10 Hz and 0.35 mm(Figure(b)). Four strain gauges were mounted on theshaft—two were located opposite each other—whichdetected the shaft load after the signal conversion of thevertical force F

y and the horizontal force F

z. Another

pair of strain gauges measured the force F ts acting on thetorque support arm (i.e., its deflection). To increase theresolution of the strain gauges, the trunnion axis wasdesigned as a hollow shaft. Capacitative sensors tomeasure the acceleration and an inductive positiontransmitter were arranged at the bottom of the vessel.The model was equipped with an ink-injection systemfor residence time distribution (RTD) investigations.A defined amount of ink was injected, and the variationin brightness was detected at a certain position in thewater domain. For flow visualization, an argon-ion


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continuous wave laser was used, and the tracers usedwere polyamide particles (q = 1000 kg/m3).

Following the AOD’s standard operating practiceduring the reduction phase (Figure 1) and taking intoaccount the laws of similarity, various blowing rates andfill levels of 50 pct to 150 pct each and different nozzleconfigurations were investigated. Because of the differ-ent positions of the strain gauges, no direct comparisoncould be made between the operational and the watermodel measurements. In addition, the focus in operationwas on the maximum values of the torque oscillationbecause they supplied important reference values for the

design torque support arm, whereas in the water modeltest, the standard deviations were used because of theirmore accurate statistical statement.

B. Side-Wall Nozzle and Gas Injection

The process gases were seen as the driving forcebehind the melt homogenization. In addition to thebubble size, the length L with which the gas jet (index g)

penetrated the liquid (index l) played an important part.The behavior of gas bubbles in the melt was character-ized by weight, buoyancy, and inertia forces, andtherefore, the following modified Froude number:

Fr ¼qg

ql

u2g

gDn; where ql qg ½6

The operating point is Fr = 3268. In the followingconsideration, the shroud tube of the side-wall nozzlewas neglected. The gas (O2, N2, Ar, and/or air) exitedthe nozzle as a continuous jet at almost sonic speed(jetting regime) and, in the fluid (melt or water),disintegrated into a multitude of individual bubbles(bubbling regime).[14] The bubbles rose as a heteroge-neous buoyant plume, which entrained the surroundingmelt. To prepare for the CFD model, the differentbehavior of the various gases throughout the processhad to be neglected. For example, an oxygen bubbledirectly reacts with the chromium contained in themelt. The exothermal reaction (2) results in localtemperatures of more than 2500 C result, which causethe gas bubble to expand rapidly. If the bubble enters acooler region afterward, then it abruptly will contractagain. This effect is superimposed by the increase in

Table I. Nomenclature of the 120-ton AOD Converter and the 1:4 Scale Water Model

Converter Model

Scale S — 1:1 1:4Inner volume V 0 m

3 63.1 0.846Liquid steel volume (100 pct) V m3 17.3 0.292Liquid steel mass M t 120 0.274Vessel tilting angle c deg 0–8 0–15Vessel height H m 7.430 1.858Vessel cone diameter D0 m 1.182 0.300

Vessel inner diameter D1 m 3.724 0.931Vessel bottom diameter D2 m 2.885 0.721Vessel outer diameter D3 m 5.100 1.011Liquid steel level (100 pct) H st,w m 2.096 0.524Distance between bottom and trunnion axis H t m 3.007 0.752Number of side-wall nozzles nn — 7 7Nozzle height above vessel bottom H n m 0.498 0.125Inner diameter of the side-wall nozzle Dn,in m 0.016 0.002Outer diameter of the side-wall nozzle Dn,out m 0.018 — Length of the side-wall nozzle Ln m 1.2 0.1Angle between the side-wall nozzles an deg 18 18Density of melt/water qst,w kg/m

3 7033 998Density of slag qsl kg/m

3 2990 — Density of argon/air at standard conditions qAr,air,stp kg/m

3 1.784 1.293Process gas flow rate (100 pct) _V Ar;air m

3/minstp 120 0.286

Cooling gas flow rate (100 pct) _

V Ar m3

/minstp 15.4 — Modified Froude number Fr — 3268 3268

Fig. 2—Simplified, rotational symmetric geometry of the 120-tonAOD converter.


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bubble size and the decrease in ferrostatic pressure. Incontrast to this finding, an argon bubble used tointensify mixing behaves like an air bubble rising in

water because it will not react with the melt. As ageneral rule, it can be said that the higher the blowingrate, the smaller the bubbles are. The diameter of gasbubbles in a melt is between approximately 10 and100 mm.[14]

The penetration length of the gas jet is importantbecause the flow structure depends on the spatialdistribution of the induced plume. The plume generatestwo asymmetric mixing zones—a smaller zone on thenozzle wall and a larger zone opposite the nozzle wall.[6]

In this article, the mean penetration length L of the gas jet into the liquid fluid was established according toHoefele and Brimacombe[12] because this approachagreed most with our own experience and is expressed

as follows:

L ¼ 10:7Fr0:46Dnqg

ql

0:35½7

As the Fr number, nozzle diameter Dn, and qg/qlratio increase, the length of penetration L rises. Toensure that the flow structure in the model (index m)and original (index o) converter were similar, the ratioLm/Lo must have corresponded to the model’s scale of 1:4. After rearranging Eq. [7], the nozzle diameter

Dn,m obtained in the water model is expressed asfollows:

Dn;m ¼ 0:25Dn;oq

Arqst

qw

qair

0:35½8

Based on the use of argon as a process gas and the rel-evant fluid properties, Dn,m = 2 mm. The nozzle lengthis selected at Ln,m = 0.1 m. These values ensured thatthe penetration lengths of the jet in the model weresimilar to the original. The standard volumetric flowrate to be set at the model nozzle is expressed asfollows:

_V n;air ¼

ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiqAr

qair

qw

qst

Dn;m

Dn;o

5_V 2

n;Ar

s ½9

From the equation of continuity, the mean outletspeed from the nozzle is expressed as follows:

un;air ¼4 _V n;air

pD2n;m

p0

p

T

T 0½10

with the standard values p0 and T 0.To review the Hoefele and Brimacombe approach,[12]

the penetration of an air jet into water was evaluatedstatistically.[13] The theoretical penetration length LArof an argon jet in the AOD converter, as ascertained by

Fig. 3—1:4 scale water model of the 120-ton AOD converter equipped with innovative measurement technique; standard conditions of the watermodel are _V air = 0.29 m

3/minstp (seven nozzles), H w = 0.524 m.


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Eq. [7], was compared with the length Lair of a fluiddynamic similar air jet in the water model (Figure 4).The operating point of the AOD converter and theequations for the penetration lengths of argon and airare plotted. Generally, the penetration length was low.At the operating point of the AOD process, it wasLAr = 0.39 m for the argon–melt system andLair = 0.086 m for the air–water system. The theoreticalvalues according to Hoefele and Brimacombe agreedwith the measured penetration lengths in the watermodel. When raising the nozzle pressure beyond thedesign condition, the measured data were leveling, andLair remained more or less constant. No penetration of the air jet beyond the center of the vessel and up to theopposite wall was observed. The air bubbles always roseclose to the nozzle wall and generated a large-space,

quasistationary vortex. This general behavior coincidedwith the real wear pattern of the refractory liningobserved in AOD practice.

In a preliminary study, the injection of argon througha single side-wall nozzle was calculated with the help of a transient, two-phase CFD simulation according to theEuler–Euler approach. Here, a complete system of fluid-dynamic equations was solved for each fluid-flow phase.The injection of a compressible, cold gas, which attainedroughly sonic speed at the nozzle outlet and then entereda highly viscous incompressible hot-melt domain atan approximately 7000 times higher density, requiredextreme demands on the CFD solver. Thus, a numericaltrick was employed to stabilize the solution and make itconvergent. The results obtained are variables p, T , q, u,and Ma at the nozzle outlet but, above all, thepenetration length LAr of an argon jet in melt. Theseresults serve as input data for a simpler and lesscompute-bound injection model of Lagrangian formal-ism in which the gas downstream from the nozzle wasspecified in the form of an empirical bubble distribution.This process is referred to as the discrete phase model(DPM) (see Section IV).

Figure 5 shows results of the Euler–Euler simulationshortly after the argon injection began into a rectangular

vessel filled with melt up to an elevation H st = 2.096 m.The heat transfer between gas flow and surroundingrefractory material has been considered. Argon gasexited the nozzle at pAr,ex = 7.6 bar, T Ar,ex = 25 C,qAr,ex = 4.2 kg/m

3, and uAr,ex = 447 m/s, the pressuredrop of the nozzle was D pn = 1.7 bar. The penetrationlength varied between LAr = 0.35 and 0.4 m, whichagrees with the values in Figure 4. The gas accelerationin the nozzle agreed with Fanno’s theory. Coolingcaused by expansion of the gas in the nozzle wascompensated partly by the wall heat flux from therefractory lining. Outside the nozzle, an underexpansionup to Ma = 1.4 can take place, which because of thehigh static differential pressure between the gas at thenozzle outlet ( pAr,ex = 7.6 bar) and the pressure inthe melt ( pst,ex = 1.45 bar, above ambience), seemed tobe realistic. This phenomenon also is supported by theabrupt heating of the argon gas.

In operational practice, a small portion of the meltsolidifies at the tip of the nozzle because of coolingeffects. The adhered steel develops into a tubular shapeand extends the nozzle length by a few centimeters.This finding causes the process gas to penetrate the melt

Fig. 4—Penetration length LAr of argon gas into melt for nozzlelength Ln = 1.2 m and diameter Dn = 16 mm. Penetration lengthLair of air into water for nozzle length Ln = 0.1 m and diameterDn = 2 mm.

Fig. 5—CFD simulation (Euler–Euler approach) of argon gas injected into the melt domain, H st = 2.096 m, Ln = 1.2 m,Dn = 0.016 m. The flow structure at t = 0.5 s after the start of gasinjection is shown. Argon boundary conditions at the nozzle inletare p Ar,in = 9.3 bar (above ambience), T Ar,in = 20 C.


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a little deeper and the exothermal reactions of theoxygen in front of the nozzle to take place further awayfrom the refractory lining.

IV. NUMERICAL MODELS

The CFD simulations are based on the unsteadyReynolds averaged Navier–Stokes (URANS) equationsin conjunction with a turbulence model. For the water

model, the realizable ke model was used, and for the realAOD converter, the shear stress transport—scale adap-tive simulation (SST-SAS) model was used.[15 – 18] Insubjective terms, the SST-SAS model shows a finerresolution of the stochastic phenomena than otherturbulence models, especially at the phase boundaries.For the multiphase flow (water–air and/or melt– slag– gas), the volume of fluid (VoF) model was used. [19] Thismodel is a special Euler–Euler approach for two or moreimmiscible fluids in which a surface-tracking techniqueis applied to a fixed Eulerian mesh to predict theinterface position between the fluids. A single set of momentum equations was shared by the fluids, and anadditional equation for the volume fraction of each fluidin each grid cell was tracked throughout the computa-tional domain.

The fluid movement, especially near the phase inter-faces, induced time-depended forces and torques on theconverter wall. These forces caused the converteroscillation. In every time step of the CFD simulation,the total pressure force ~F p;i and the friction force ~F s;i acting on the wall element and, thus, the resulting force~F i are calculated as follows:

~F i ¼ ~F p;i þ ~F s;i ½11

The tilting torque ~M i around point Pt on the trunnionaxis (Figure 2) was calculated as follows:

~M i ¼ ~r0;i ~rt

~F p;i þ ~r0;i ~rt

~F s;i ½12

Here ~r0;i is the position vector of the i th wall elementthat referred to the coordinate origin P0. With the helpof a user-defined function (UDF), the center of massfor all phases and the integral variables ~F p; ~F s, and ~M were established from the transient simulation. Theinfluence of the oscillating converter on the fluid wasneglected (i.e., one-way coupling of domains). Inevaluating the forces and torques that are variable interms of time, the standard deviation (rms) was used asa quantitative criterion because the mean value is, bydefinition, equal to zero. However, this finding alsomeans that peak forces that occurred in real operationwere filtered out.

Because of the computing time, the gas injectionprocess was not realized by a Euler–Euler technique butby a so-called DPM method. This method provided forthe specification of an empirical bubble distributionclose to the nozzle outlet. The DPM model performedLagrangian trajectory calculations for the dispersedphases (i.e., particles, droplets, or bubbles), whichincluded coupling with the continuous phase. To thisend, the inertia, mass, buoyancy, and drag forces acting

on the bubble were considered.[15] To include theinfluence of the turbulence on the bath movement, astochastic bubble tracking method was used. Interac-tions between the continuous and the dispersed phasealso were taken into account (two-way turbulencecoupling). Small bubbles generally behaved like rigidspheres, whereas larger bubbles became cup-shaped witha constant drag coefficient (cd,b). In this article, aBro ¨ der

[20] approach was used in which the drag coeffi-cient was a function of the bubble Reynolds number.

The expansion of the bubble while rising in the melt alsowas taken into account. To this end, the local bubblediameter Db as a function of the vertical y-coordinatewas described on the simplified assumption of anisothermal change of state and is expressed as follows:

Db ¼ Db;Hn 1 þql g y H nð Þ

p

0:333½13

Db,Hn is the local bubble diameter at the level of bubbleinjection H n (Figure 2). For the water model, theincrease in diameter was approximately 1.6 pct, andfor the melt, it was around 35 pct. The gas bubbles wereadded to the relevant fluid in the form of a Rosin– Rammler distribution. This method described the cumu-lative mass portion of the gas bubbles with a diameterlarger than Db, as a function of Db, and a parameter thatstated the width of the distribution. For the watermodel, bubble diameters of Db,air = 1 to 35 mm withDb;air = 20 mm were assumed, and for the AOD con-

verter Db,Ar = 5 to 70 mm with Db;Ar = 40 mm. Thebubbles were eliminated from the computing domain atthe phase boundary (i.e., water–air or melt–gas and/orslag–gas). The process data and fluid properties neededfor the DPM were obtained from the describedpreliminary simulation using the Euler–Euler approach.

The numerical grid of the water model consisted of

around 0.5 million hexahedron cells and, for the AODconverter, of around 1.5 million hexahedron cells.

V. SOME LAYOUT CRITERIA

The AOD process was affected largely by the con-verter geometry (Table II and Figure 6). Efficient reac-tion kinetics required a rapid, intensive mixing of themelt. The mixing time and mixing intensity depend, interalia, on the ratio of the bath diameter and the bath

Table II. Basic Process Conditions that Affect the Flow

Structure in an AOD Converter

D=H st1

Fill level of liquid steel high lowHomogenization process good badMixing time short longWear rate of the refractory lining high lowMovement of the free surface intensive weakDistance between melt center

of gravity and trunnion axislow high

Vibration intensity high low


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height D=H st: The results shown in Reference 21demonstrate that the mixing time increased with anincreasing D=H st ratio.

Unaffected by geometric restrictions, ideally ( D=H st = 1,Figure 6(a)) a circular primary vortex with a highangular momentum and a high kinetic energy results,which covers a large region of the bath. On the oppositeconverter wall, the melt flows downward. The meltvelocity adjacent to the wall is high, mixing is good,and the mixing time is low. When D=H st rises with aconstant melt capacity (Figure 6(b)), the vortex formsan elliptical shape, and the kinetic energy is dissipated.

The rotational energy decreases and the mixingtime increases. With an even higher ratio of D=H st(Figure 6(c); e.g., at the end of a converter campaign),the maximum rising velocity of the gas bubbles may nolonger be achieved and part of the kinetic energy of theplume becomes lost for the melt drive. Because theD=H st ratio then is unfavorable to form a circular

vortex, the low-momentum melt close to the freesurface will not reach the opposite converter wall butreturns to the interior and the primary vortex decaysinto several secondary vortices. Inactive dead-flowregions are induced, which are only slowly involved inthe chemical reactions. Although the melt velocity nearthe converter wall is low, it minimizes wear, but mixingis poor.

Closely associated with transient flow phenomena arerefractory wear and converter vibrations. Reactionstaking place at temperatures higher than 2500 C,process-related temperature variations, high-melt flowvelocities, and turbulence lead to erosion of the nozzle-side converter wall, especially at the melt–slag–gas phaseboundary. Typical wear rates are 3–5 mm/heat. Wear of the nozzle-side converter wall is counteracted by athicker refractory lining. High velocities near the wall,which occurs for a low D=H st ratio, increase the wear

rate and vice versa. Moreover, the dissolution capacityof the slag decreases with an increasing converterdiameter, which may require the excessive addition of a fluxing agent. Consequently, all requirements, a shortmixing time, low wear rates, and vibration cannot besatisfied at the same time. For a freshly relined AODconverter, D=H st 1.5 to 2 is a good compromise.

The side-wall nozzles have to be arranged at a lowlevel to ensure a high rising level (H st – H n) of gasbubbles (Figure 6(d)) and to avoid dead flow regions.The nozzles may not be located at a low level either, orelse the converter bottom will erode. If nozzles are

installed at a high level (Figure 6(e)), then a part of theferrostatic height to drive the melt will be lost, and aninactive flow region will be induced close to the bottom.Investigations are currently under way that deal with theconcentric arrangement of nozzles in the converterbottom (Figure 6(f)). A positive influence is expectedon the reaction kinetics as well as on the vibrationamplitude. The inclination of the refractory wall isrelevant for mixing because the wall is not allowed tocounteract a uniform, concentric vortex formation. Theusual number of degrees is 20 deg to 35 deg. In thecourse of the converter campaign, the refractory liningwill adapt itself to the existing flow structure and adopta dished contour.

VI. RESULTS

A. Plant Test

SMS Siemag carries out vibration analyses on con-verters across the world, with statistical evaluationbased on several converter campaigns. The subject of the analysis in the following discussion is the 120-tonAOD converter shown in Figure 2.

Fig. 6—Fundamental flow structure for different geometries of the AOD converter.


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To counteract the high wear rate, many plant ownersincline the converter vessel by 5 deg to 10 deg during theprocess. The intention is to put the plumes away fromthe refractory lining (i.e., to reduce the wear rate abovethe side-wall nozzles and to improve the conditions forthe species conversion, or both). Operational measure-ments made with both a vertical (c = 0 deg) and aninclined (c = 8 deg) converter clearly indicate the ten-dency toward a higher vibration intensity when thevessel is inclined.

Figure 7 shows the difference DM = Mc = 8 deg – Mc = 0 deg of the measured torque standard deviationM both with an inclined and a vertical vessel for theprocess stages. This result is the mean deviation fromthe zero position (i.e., the higher the standard deviation,the more intensive the vibration). To minimize theinfluence of the progressing wear of the refractorylining, the tests were made one immediately after theother after expiry of 40 pct and 60 pct of the converter

campaign. Although the results were not necessarilyconsistent at the various process stages, all cases indicatea significant increase in vibrations when the converter isinclined. For example, the standard deviation during thereduction phase after 60 pct of the converter campaignrose by DM = 8.4 pct.

The operational tests also prove that the measuredtorque (i.e., maximum value as well as standard devi-ation) at a variation range between 110 t and 145 t isindependent of the heat size. The variation between

–8 pct and 21 pct of the nominal melt capacity isaffiliated with a fill level variation of DH st = –0.13 mand DH st = 0.32 m, respectively. This unexpected resultis reduced to the fact that the movement of the freesurface and, thus, the vibration level is barely affected bythe variation of the fill level around the nominal height.Only at low melt heights will the surface waves bedamped by the converter bottom. This behavior isconfirmed largely by water model tests, which show onlya moderate dependence of the vibration amplitude onthe fill level so that the amplitude decreases slightly witha rising fill level (see Section VI-B).

Moreover, the operational tests indicate that the

torque oscillation was highest at the start of theconverter campaign, which decreased by around 25 pctby the end of the campaign. This statement is based onthe evaluation of approximately 450 converter heats andhas been confirmed recently by other plant tests.Generally, tendency is clear for N2 to induce slightlyhigher oscillation amplitudes than Ar. The measuredfrequency of the 120-ton AOD converter dominating inthe frequency spectrum is typically at around 2.5 Hz.

B. Water model

Figure 8 shows laser-light-sheet visualizations at avariable blowing rate (for 100 pct fill level) and a

variable fill level (for 100 pct blowing rate). The air jetspenetrated the water model from the left and quickly

Fig. 7—Difference of the torque standard deviation DM = Mc = 8 deg –

Mc = 0 deg measured for the upright and the inclined 120-ton AODconverter after 40 pct and 60 pct of the converter campaign.

Fig. 8—Laser-light-sheet visualization at the water model for (a) different blowing rates and (b) fill levels. Standard blowing rate with seven noz-zles _V air = 0.29 m

3/minstp (blowing rate 100 pct) and standard fill level H w = 0.524 m (fill level 100 pct).


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collapsed into a variety of individual bubbles ascendingon the vessel wall and caused a primary vortex thatrotated clockwise. The size and location of the vortexremain nearly unchanged. A rough, heavily bubblingwater surface was generated, which may be regarded asthe cause of the vessel vibration. The higher the blowingrate, the more intensive the vortex rotation and the moreturbulent the water became. Even at a high air-inletpressure into the side-wall nozzle, the air jet did notpenetrate far into the water. At a fill level of 75 pct

(

DH w = 2.10), several stochastic individual vortices wereformed, and the primary vortex arbitrarily changed itssense of rotation. From a fill level of around 100 pct(

DH w= 1.58) on, the primary vortex statistically behaved

almost at a steady state. The higher the fill level, thefaster the primary vortex rotated. Here again, themovement of the water surface increased but not tothe same extent as with increasing the blowing rate.

Several investigations have been performed on thewater model. For the following statements, the mea-sured data are always referred to the standard (index:stand) mode of operation at a blowing rate of 100 pct,fill level of 100 pct, and c = 0 deg. For example, the

relative deviation of the x-force component wasDF x = (F x – F x,stand)/F x,stand.

(a) When inclining the water model converter byc = 15 deg toward the nozzle wall, the rms value rel-ative to the standard mode of operation increased byaround 18 pct (DF y,rms = +6 pct, DF z,rms = +32 pct,DF ts,rms = +16 pct). This order of magnitude wasobtained in the operational test as well (seeSection VI-A). For all blowing rates and fill levelsinvestigated, the detected model forces and, thus, thevibration amplitudes were higher with an inclinedvessel rather than with an upright vessel.

(b) When simulating the influence of the dampeningslag layer, which is highly viscous during thedecarburization phase, in the water model by addingwooden spheres of a suitable density onto the watersurface, the rms value was reduced by around 10 pct(DF y,rms = –7 pct, DF z,rms = –9 pct, DF ts,rms = –15 pct). Similar results were obtained for otherblowing rates, fill levels, and c values.

(c) When the torque support arm was removed, thevertical force DF y.rms decreased by around 13 pct,whereas the horizontal force DFz,rms increased byaround 56 pct. This behavior applied to all blowingrates, fill levels, and c values.

(d) An additional weight of 50 kg fixed to the bottom of themodel suppliedDF y,rms = –3 pct,DF z,rms = +14 pct

and DF ts,rms = +6 pct.(e) Figure 9 illustrates that the higher the blowing rate

_V air becomes, the shorter the mixing time s95 is. Asthe fill level H w increases, the mixing time s95 firstincreases then remains constant for high fill levels.

(f) The natural frequency f E of the water model test rigwas measured for the neutral position at a fill level of 100 pct. The dominating peak in the frequencyspectrum was f E1 = 4.3 Hz. More sub maximawere found at f E2 = 5.6 Hz, f E3 = 25.3 Hz, and

f E4 = 29.5 Hz. The test rig was designed with the

help of Pro/E, and the theoretical natural frequencywas calculated from the total stiffness and the totalmoment of inertia by taking into account the tor-sional stiffness of the trunnion axis and torque sup-port arm. f E,theo = 4.5 Hz agreed with the firstmaximum of the measurement.

(g) For standard conditions, the maximum levels in thefrequency spectrum were found at f 1 = 1.0 Hz,

f 2 = 4.8 Hz, f 3 = 13.5 Hz and f 4 = 20.5 Hz, withdominating f 2 and f 3. An increase in the fill level didnot reveal a uniform frequency response. Althoughthe first maximum f 1 increased slightly here, thevalues for f 2-4 decreased. In the water model, avariable blowing rate had no influence on the fre-quency spectrum.[11]

The results of the vibration and frequency measure-ment with variable blowing rates and fill levels arepresented in the following sections, which include acomparison with the results of the CFD simulation.

C. CFD Simulation and Comparison with Plant Testand Water Model Investigation

The movement of the free surface and the deviation of the velocity across time for the standard case (blowingrate 100 pct, fill level 100 pct) are shown in Figure 10.The centerline of the converter and the fill level in therest position also are shown. The plumes—of which onlythe rear ones are shown—do not reveal any individualbubble but an iso-volume that enclosed the respectiveplume with a constant bubble concentration. The colorsof the iso-volume corresponded to the local meltvelocity. In the CFD simulation, the physical boundaryconditions for the Rosin–Rammler distribution of theinjected air bubbles have been adapted so that thecalculated length of penetration Lair agreed withLair = 0.086 m as measured on the water model. The

Fig. 9—Mean mixing time s95 and corresponding standard deviationmeasured at the water model for different blowing rates _V air and filllevels H w.


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bubbles rose near the nozzle wall and induced theclockwise rotating primary vortex described earlier aswell as a small elliptical secondary vortex close to thewall. The size and location of the vortices that affectedthe homogenization process were similar for all blowingrates and fill levels. When the plumes reached the freesurface, high wave amplitudes with a low-frequencyoscillation were induced. The wall-bounded water sur-face moved up and down. This behavior correspondedto the type-B oscillation according to Fabritius et al.[11]

(Figure 10(c)). The type-A oscillation, which means theswashing back and forth between the vessel walls,neither was calculated by CFD nor observed in anywater model test. However, the characteristic bubbling

known from the water model cannot be simulatednumerically because the bubbles defied a spatial andtime-dependent resolution. If the free surface above theplumes rose, then velocities in an upward direction weregenerated in the water and vice versa. The higher theblowing rate, the higher the amplitudes of the surfaceare. At a blowing rate of 150 pct, the maximumdeviation from the fill level at rest position on theside opposite the nozzle wall was, for example,D y ± 20 mm. Adjacent to the wall, high velocitiesled to high-wall shear stresses and, in reality, to a highdegree of wear. High shear stresses resulted around thenozzles as well and in the area of the free surface,especially on the side of the nozzle wall. The maximumshear stress was found when the water surface passedthe location of the rest position, because at this time, theinduced velocity reached a maximum. As a result of theperiodic excitation, the system was supplied continu-ously with energy (i.e., this system was a forced wide-band oscillation).

For the water model, Fig. 11 gives an example of thesimulated distribution of force F y according to Eq. [11]and the distribution of torque M y according to Eq. [12],

Fig. 10—CFD simulation for the water model including the phaseinterface movement for _V air = 0.29 m

3/minstp (blowing rate 100 pct),H w = 0.524 m (fill level 100 pct), and free surface level.

Fig. 11—CFD simulation for the water model with _V air =0.435 m3/minstp (blowing rate 150 pct) and H w = 0.524 m (fill level100 pct). (a) Distribution of the force F y according to Eq. [11] and(b) distribution of the torque My according to Eq. [12]. Minimumand maximum value of the standard deviation as well as thecalculated frequency f

1 = 1.2 Hz and f

2 = 2.4 Hz.


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based on a blowing rate of 150 pct and 100 pct fill level.In sum, the simulated forces and torques are at a lowlevel. The corresponding table shows that maxima resultthat are clearly higher than the rms value plotted as adash-dot line. This behavior is known from the real AODprocess as well. All rms values of the forces are of thesame magnitude. Oscillation has no preferred direction,but the wide-band excitation induced a stochastic oscil-lation with a tumbling movement in all spatial directions.This phenomenon was reflected by the rms values of the

torques. Torques M x,rms and M z,rms around the horizon-tal converter centerlines were generally higher thantorque M y,rms around the vertical centerline; the latterwas characterized by F x and F z. In most cases, M x,rmsyielded the highest values, presumably because somesurface waves proceeded preferably from the nozzle sidetoward the opposite vessel wall. The frequency spectrumcalculated from the forces and torques showed twodominating frequencies at f 1 = 1.2 Hz and f 2 = 2.4 Hz.

Figure 12 shows the comparison between the mea-sured and the calculated forces/torques for the watermodel at different blowing rates and fill levels. On thetrunnion axis of the model, the components F y and F z

have been measured (Figure 2). These values are plottedin the diagrams of the force distribution as coloredsymbols without a connecting line. The detected datasupplied by the strain gauge pair installed on the torquesupport arm were too small, and the torque support armshould be required to be designed as a hollow section infuture investigations. Despite the basically differentapproach in establishing the occurring forces (i.e., on

the one hand, the direct strain gauge measurement onthe trunnion axis of the physical model, and on the otherhand, the integration of the numerically calculatedpressure and shear stresses), one arrives to the sameorder of magnitude and a similar behavior for ~F rms. Thegeneral tendency (i.e., increasing ~F rms values with arising blowing rate _V air and slightly decreasing ~F rmsvalues with a rising fill level H w, and starting at a filllevel of around 75 pct) is confirmed by both simulationmethods.

As the blowing rate increases, the forces and torquesand, therefore, vibrations increase because the movementof the surface becomes more intensive; this behavior isconfirmed by the physical simulation at the water model.As the fill level increases, the amount of water rises,thereby reducing the distance between the model’s centerof gravity and the trunnion axis. Consequently, thewhole system receives more mass inertia. In the physicalmodel, the increasing movement of the water surfacewith a rising fill level was at least partially compensatedby the simultaneously increasing mass inertia so that theforces, torques, and vibrations decreased slightly. Inregard to the surface movement, the numerical model

revealed a contrary behavior. With small fill levels, anintensive surface movement could be recognized, whichat approximately 100-pct fill level reached its maximum,will cause it to drop again while the fill level continued torise. This behavior is reflected both by the ~F rms and ~M rmsvalues as well as by the peak forces.

Figure 13 shows the results of the transient, three-phase, nonisothermal CFD simulations for the 120-ton

Fig. 12—Comparison between experiment and CFD simulation for the water model. In all cases, the basis for the shown data is the standarddeviation (rms value) of the forces and the torques for different (a) blowing rates _V air and (b) fill levels H w.


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AOD converter filled with melt (yellow) and a 0.3 mthick slag layer (red). The simulations were based on theSST-SAS turbulence model in conjunction with the VoFand the DPM models. The blowing rate was constant in

all cases; _V Ar ¼ 120 m3/minstp (blowing rate 100 pct). At

the start of the reduction stage (Figure 1) the slagconsists of around 40 pct CaO, 5 pct MgO, 5 pct SiO 2,25 pct Cr2O3, 10 pct FeO, 5 pct MnO, and 10 pct

Fig. 13—CFD simulation for the original 120-ton AOD converter. The movement of the phase interface and the velocity distribution during thereduction phase is shown. All forces and torques are determined according to Eq. [14] and are referred to the standard process with_V Ar= 120 m

3/minstp (blowing rate 100 pct), H st = 2.096 m (fill level 100 pct), c = 0 deg, and lsl = 0.001 Pas; melt, slag.


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metallic granules and is highly viscous. The arbitrarilychosen viscosity of lsl = 50 Pas is that of thick honey,lsl = 0.05 Pas is like glycerin, and lsl = 0.001 Pas is likewater. For the analyses, no individual components wereused at this stage anymore but rather the integral forcesand torques. The latter were referred to standardoperation at _V Ar = 120 m

3/minstp (blowing rate100 pct), H st = 2.096 m (fill level 100 pct), c = 0 deg,and lsl = 0.001 Pas. The following equation thenapplies, for example, to the relative change in the force:

D ~F ¼ ~F

~F stand ~F stand

with ~F ¼ ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiF 2rms;x þ F 2rms; y þ F 2rms;zq

½14

Figure 13(a) shows CFD results of the fundamentalinfluence of the slag layer being molten during thereduction phase. As the viscosity of the slag decreased,the dampening effect of the slag layer declined. Move-ments at the melt–slag and slag–gas interface increased, asdid the melt turbulence in the fluid domain. Atlsl = 50 Pas the resulting force was around 60 pct, andthe resulting torque was 57 pct lower than the standardcase with lsl = 0.001 Pas. These values are high, yet whatmust be considered here is that the numerical modelinterpreted the sole fluid movement as oscillation withoutthe dampening influence of the additionalconverter mass.In terms of operating practice, this result means that theconverter vibration—despite a constant inert gas flowrate—can only increase because of slag liquefaction,which recently has been confirmed by operational tests.Yet, the reduction of the slag viscosity below a certainlevel, here lsl 0.01 Pas, was not accompanied byanother increase in vibrations. In sum, the large-spaceflow structure was similar to that of the water model.However, the structure of the large-space primary vortex,

which is characteristic for water, has been lost for melt infavor of many small individual vortices. The localturbulence inside the melt domain can be extremely high.

Figure 13(b) shows sequence-vs-time images of themelt flow with Dt = 0.5 seconds. At some points in time,the lighter weight slag layer was displaced by the risingmelt and intensive melt splashing may have occurred. Along-stretched secondary vortex rotating anticlockwiseformed at the nozzle wall. An irregular type-B move-ment was simulated on the surface.

Figure 13(c) shows the situation for a converter thathas been inclined by c = 8 deg. In this case, the risingargon bubbles moved away from the nozzle wall. Thismovement intensifies the bath one more time becausethe bubble plumes now have more space available formoving back and forth. The mean deviation of the free-melt surface around the rest position is about twice ashigh as the vertically positioned converter. Whencompared with the standard case, (right-hand picturein Figure 13(a)) the acting forces increase byD ~F ¼ þ113 pct and torques by D ~M ¼ þ122 pct. The

same applies to the vibration behavior because itcorrelates with the forces and torques. This phenome-non—predicted based on CFD simulations—was

confirmed by the operational test in Figure 7. Thetorque standard deviation DM , as measured at thetorque support arm, was a mere 8.4 pct. However,the following two aspects must be considered: On onehand the measurement was based on the evaluation of just a few heats and also may vary considerably becauseof permanently changing boundary conditions in themelt shop. On the other hand, the mass inertia of theconverter plant was not taken into account in the CFDsimulation.

Figure 13(d) shows the situation at H st = 2.620 m (filllevel 125 pct). Although the intensity of the surfacewaves subjectively agreed with the standard fill levelH st = 2.096 m (fill level 100 pct), the forces and torquesdecreased by D ~F

¼ 28 pct and D ~M ¼ 47 pct whencompared with standard operation. It is obviously notonly the pure movement near the free surface of the meltresponsible for the amount of D ~F

and D ~M , but alsothe local turbulence within the melt bath. For the100-pct fill level, the flow was characterized by amultitude of turbulent individual vortices, whereas withthe 125-pct fill level it was characterized by a relativelystable vortex rotating clockwise. Following the consid-

erations from Figure 6, the primary vortex at

D=H st= 1was stable, whereas for D=H st „ 1 it became moreunstable. For this reason, the flow structure in standardoperation (H st = 2.096 m, fill level 100 pct fi D=H st =1.58, right-hand picture in Fig. 13(a)) was also signifi-cantly more stochastic than the higher fill level(H st = 2.620 m, fill level 125 pct fi D=H st= 1.26, seeFigure 13(d)). But then again, higher static forcesresulted in the latter case because of the higher converterweight.

The melt movement correlates with the minimum andmaximum forces as well as with the rms forces.Figure 14(a) illustrates the influence of a variable tiltingangle c by referencing the horizontal component Fy in

the main vibration direction. A process time of 50 sec-onds was shown after the full transient response of theconverter. Both the mean amplitude and the rms valueclearly increased. This was particularly obvious whenlooking at torque M x, which formed from componentsF y and F z and increased from 15.9 kN to 59.9 kN(Figure 14(b)). The frequency established from CFD isshown in the diagrams as well. For c = 0 deg,

f 1 = 0.48 Hz, and f 2 = 1.30 Hz, for c = 8 deg, f 1 = 0.47 Hz, and f 2 = 0.93 Hz. In all cases, the fastFourier transform (FFT) supplied several peaks in thefrequency spectrum. However, it did not matter whetherthe FFT was applied to the force, the torque, or the localmelt speed; the result was always the same. Similarstatements applied when the fill level was changed fromH st = 2.096 m (fill level 100 pct) to H st = 2.620 m (filllevel 125 pct) in Figure 15. The fill level, however, had asmaller influence on F y,rms and M x,rms than the tiltingangle.

Table III summarizes the most important results andcompares the resulting minimum and maximum forcesand torques as well as the rms values for the casesinvestigated (i.e., the standard case (1), the inclinedconverter (2), and the higher fill level (3)).


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VII. CONCLUSIONS

By introducing process gases under the surface of the

melt bath, the AOD process causes considerable vibra-tions that may be intensive enough to make componentsof the structural steelwork, foundations, and plantcomponents suffer from them. For safe design, it isimportant to understand the causes and relationships of these vibrations. For this reason, the chemical-physicalprocesses in the AOD converter have been investigatedby a combination of operational measurements, physicalsimulation on a downscaled water model, and numericalsimulation using ANSYS FLUENT. In particular theCFD simulation provided an effective tool to make the

AOD process more transparent. Some inconsistenciesshould be noted between the results of the differentsimulation approaches. This concerns the height of themeasured and calculated torques for the 120-ton AODconverter as well as the respective frequencies. Futureinvestigations to clarify these discrepancies are currentlyon their way.

The present work allows the following importantconclusions for the AOD process:

(a) The converter oscillation is essentially caused bytransient movements of the phase-boundary sur-faces. The process gases injected through thesubmerged side-wall nozzles and the plumes,

Fig. 14—CFD simulation for the upright (c = 0 deg) and theinclined (c = 8 deg) 120-ton AOD converter with _V Ar=120 m3/minstp (blowing rate 100 pct) and H w = 2.096 m (fill level100 pct). Time-dependent distribution of the (a) forces and (b) tor-ques as well as simulated frequencies.

Fig. 15—CFD simulation for the fill level H st = 2.096 m (100 pct)and H st = 2.620 m (125 pct) of the 120-ton AOD converter at_V Ar= 120 m

3/minstp (blowing rate 100 pct) and c = 0 deg. Time-dependent distribution of the (a) forces and (b) torques as well assimulated frequencies.

Table III. Results of the CFD Simulation for the 120-ton AOD Converter with Different Process Conditions

Component F x [kN] F y [kN] F z [kN] M x [kNm] M y [kNm] M z [kNm] Case

Min./max. peak –76/81 –1460/–1379 –34/38 –38/51 –0.04/0.10 –96/97 1. Fill level 100 pctrms value 35 19 13 16 0.04 43 c = 0 degMin./max. peak –126/127 –1450/–1388 89/292 –359/–124 –0.14/0.18 –152/154 2. Fill level 100 pctrms value 68 24 50 60 0.07 82 c = 8 degMin./max. peak –38/28 –1860/–1744 –46/67 –52/64 –0.09/0.05 –38/27 3. Fill level 125 pctrms value 13 18 20 21 0.03 13 c = 0 deg


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respectively, act as a vibration stimulus with wide-band excitation. Characteristic torque oscillationsare highly determined by the defined process stan-dard (i.e., by the oxygen and inert-gas (N2, Ar) flowrate at the blowing stages and their allocation todefined bath-carbon contents). The vibrationamplitude was highest during the last stage of thedynamic blow.

(b) The length with which the inert gas jet penetrates themelt domain can be determined with good approx-

imation by using an empirical approach accordingto Hoefele and Brimacombe.[12] Even at a high inletpressure, the penetration length of a highly pres-surized inert gas is not more than approximately 0.4to 0.5 m.

(c) The vibration amplitude—characterized by themaximum value on one hand, and by the rms valueon the other hand—depends on the process gas typeand flow rate. Vibrations are low when the propor-tion of oxygen in the process gas is high. Vibrationsbecome more intensive as the amount of inert gas inthe process gas rises, and the inert gas flow rateincreases. The type of inert gas has an influence on

the oscillation. N2 has a tendency to induce slightlyhigher oscillation amplitudes than Ar. The detectedfrequency is independent of the blowing rate.

(d) The operational tests demonstrate that the maxi-mum level of the torque oscillation does not dependon the heat size when the latter varies between –8 pct and +21 pct of the nominal heat size. Boththe physical and the numerical simulation indicatethat the standard-deviation forces and torques dropslightly with a rising fill level. The relating frequencyspectrum, though, does not reveal a clear-cutbehavior. However, the frequency is droppingslightly with rising fill level.

(e) A small proportion of the vessel vibrations is caused

by the oxygen injected through the side-wall nozzlesand its reaction with carbon and a clearly largerproportion by the injection of inert gas.

(f) The slight inclination of the converter vessel duringthe ongoing process as practiced by many steelmakersto put plumes away from the refractory liningincreases converter vibrations. This general phenom-enon, whose occurrence is independent of the blowingrate and fill level, indeed may intensify homogeniza-tion caused by the higher melt turbulence.

(g) Torque oscillations are highest with a freshly relinedconverter, which decrease with increasing wear of the refractory lining. During the converter cam-paign, the oscillations decrease by around 25 pct.

(h) The CFD simulation shows, among others, thateven at a constant inert gas flow rate, the oscillationof the converter vessel can increase because of theslag liquefaction.

(i) The higher the blowing rate, the shorter the mixingtime. As the fill level rises, the mixing time firstincreases and remains almost constant for high filllevels.

TABLE OF SYMBOLS

~r 0 Position vector_V Volumetric flow ratea Angle between the side-wall nozzlesq Densityc Tilting angle of the converter vessels95 Mixing timea Activityc Mass concentration

cd Drag coefficientD Diameter f Coefficient of activity f Frequency f E Natural frequency (eigenfrequency)F ForceF p Pressure forceFr Modified Froude numberF s Friction force

g Acceleration due to gravityH HeightK Equilibrium constantL LengthM Torque

n Molar massn Number of Nozzles

p PressureP0 Origin of coordinates

pCO Partial pressureS ScaleT Temperatureu VelocityV Volumex,y,z Coordinates

INDICES

– Mean value0 Referred to standard fluid flow conditionsair AirAr Argonb BubbleC CarbonCO Carbon monoxideex Exitg Gasi Indexin Inletl Liquidm Model

n Nozzleo Originalst Liquid steelstand Referred to standard operationtheo According to theoryts Torque support armw Waterx,y,z Components in direction of coordinates


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Oden Thal 2010