Lubricant Development for Hard Steel Rolling: From Design to Validation by Quaker Houghton

• Project & Plant Management, Energy & Utilities and Environmental Technologies Pre-Show Issue

29 June–1 July 2021 / Nashville, Tenn., USA

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A Publication of the Association for Iron & Steel Technology

AIST Transactions

Vol. 18, No. 2

DOI 10.33313/TR/0521

Lubricant Development for Hard Steel Rolling: From Design to Validation

T Author J.B.A.F. Smeulders Quaker Houghton, Uithoorn, The Netherlands bas.smeulders@quakerhoughton.com

his paper discusses aspects of the development of lubricants suitable for rolling hard steel types. First, it is important to gain an understanding of the role of the lubricant, based on theories on cold rolling and tribology. Mathematical modeling can be useful to explore the demands on the lubricant in the process; for instance, for very hard steel types, but also to explore the limitations of lubrication. Second, it is important to assess lubrication performance in laboratory testing under realistic conditions. Benchtop laboratory testing can be done by using tribometers, where the performance of lubricants is investigated in well-defined tests in either boundary or elasto-hydrodynamic regimes. Pilot mill trials can quantify lubricant performance under realistic rolling conditions.

Lubrication Background When rolling very hard steel types, the achievable reductions and mill speeds are limited by allowed roll force and available mill power. Both are influenced by the pressure distribution in the roll bite, the friction hill. The pressures in the friction hill increase strongly with steel yield stress but are also influenced by frictional conditions in the roll bite. Roll force and mill power can therefore be reduced by improving lubrication. Cold rolling occurs under mixed lubrication conditions. This means that in the contact, areas of close surface contact coexist with areas of higher separation. The friction in the areas of close contact, where boundary lubrication (BL) prevails, is governed by surface chemistry, i.e., adsorbed and reacted molecules that facilitate sliding and prevent metalto-metal contact. The friction in the

areas of higher separation, where elasto-hydrodynamic lubrication (EHL) prevails, is governed by the viscosity of the lubricant in the contact under conditions of high pressures, and sometimes high temperatures and high shear rates. The proportion of BL and EHL regions is determined by the entrained lubricant film thickness. The film thickness is increased by higher rolling speeds and lower bite angles and is further influenced by processes in the emulsion application zone and the inlet zone immediately before the roll bite by, e.g., pre-existing oil layers (from a previous stand), pre-applied oil layers (by direct application), emulsion concentration, oil particle size and oil viscosity. Additionally, film formation can occur within the bite: microplasto-hydrodynamic lubrication (MPHL). This leads to the formation of a thin layer of lubricant, replacing areas of boundary lubrication. This in-bite film formation is mainly influenced by sliding speed in the contact and the viscosity of the lubricant. The friction in the MPHL regime is governed by the tribological properties of the lubricant under high pressures and shear rates. The total coefficient of friction in the roll bite is determined by the proportion of BL, EHL and MPHL regions and the intrinsic friction in these three lubrication regimes. Lubrication in the roll bite can thus be influenced by several lubricant parameters, which can be assessed in the laboratory. More detailed information can be obtained in earlier publications.1,2 As was mentioned, the frictional conditions in the roll bite influence the friction hill and therefore the roll force and mill power, but there are limits to what a lubricant can achieve. Depending on the mill type and setup (especially strip thickness

3 Figure 1

Schematic illustration of the subsequent zones accounted for in the tribological roll bite model (TRBM) (geometry not to scale).

and reductions), and under normal lubrication conditions, the proportion of friction in the roll force may be relatively low, and, as was shown in an earlier publication,3 only a limited improvement can be achieved by adapting the lubricant. This is especially true when rolling hard steel types, due to the typically low reductions. The potential of the lubricant is further limited by the requirement that no slip should occur in the process. This means that when improving lubrication, even before the maximum effect on roll force is achieved, slip may occur. This is the reason that in the calculations in this paper, on the effect of steel and process parameters and lubrication on rolling, focus is on forward slip as well as on roll force and power.

Mathematical Modeling

∆T =

εYav ρC p

(Eq. 1)

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where ε = the accrued true strain in the stand (ε = ln(hin/ hout) with hin and hout the strip entry and exit thickness, respectively), Yav = the average yield stress and ρ and Cp = the density and specific heat of the steel, 7,800 kg/m3 and 500 J/kgK, respectively.

The rolling model that was used in this work is the tribological roll bite model (TRBM), developed in a collaboration project between the University of Linz, Primetals Technologies, voestalpine and Quaker Houghton.4 The TRBM model successively calculates processes in the emulsion application, inlet and deformation zones, schematically shown in Fig. 1, using an extensive range of lubricant and emulsion properties. In the application zone, account is taken of pre-existing oil layers from a previous stand and/ or pre-applied by direct application. This is also the zone in which impingement of oil droplets applied from the emulsion spray leads to the growth of the oil layer on strip and work roll. The strong emulsion backflow in the inlet zone causes the oil layers to be partially washed off, leading to a local oil concentration increase in the emulsion. The combined effect of the medium’s viscosity, the strip and roll speed in the converging contact and the resulting flow field, and an increasing asperity contact leads to a pressure

buildup that ultimately reaches the strip yield stress (Y in Fig. 1). This is the start of the deformation zone, where the entrained lubricant film thickness (h0 in Fig. 1) and entrained oil concentration are calculated. The lubrication conditions in the deformation zone are defined by asperity contact evolution along the roll bite, the accompanying evolution of the lubrication regimes of BL, EHL and MPHL (noted in Fig. 1), and the friction in these regimes. The TRBM model calculates, among others, the complete pressure hill, the local and global coefficient of friction, the roll force, forward slip and power. The calculations in this paper are done for a rolling mill with four reduction stands (a fifth texturing stand is ignored). The work roll diameters are 0.5 m and their Ra roughness 0.4 µm, except for stand 1, where it is 0.8 µm. It is assumed that the strip enters stand 1 unoiled, but in stands 2–4 “inherits” 0.25 µm of oil from the previous stand. The strip width is assumed 1 m. One of the input parameters for the model is the incoming strip temperature. This temperature is influenced by the deformation heat in preceding stands and a cooling effect of the applied emulsion. The temperature increase due to deformation can be estimated by Eq. 1:

AIST Transactions The cooling effect by the applied emulsion can be estimated by using Newton’s law of cooling and assuming a very high heat conductivity within the steel (“lumped parameter” assumption). This approach results in Eq. 2, for the strip temperature after a cooling zone of length x:

T (x ) = TC + (Ti − TC ) e −βx with β =

concentration and oil particle size of the recirculating coolant emulsion, the oil viscosity, and the coefficient of friction in the boundary lubrication regime. Note that the combined effect of emulsion concentration and particle size is the growth of the oil film in the application zone on top of the preexisting and pre-applied oil layer, as schematically shown in Fig. 1.

2H hρC pv

Calculations — The Influence of Steel and Process Parameters on the Rolling Process — First some calculations were done to investigate the influence of steel properties (hardness, gauge) and process parameters (reduction, rolling speed) on rolling performance. These parameters may influence, for example, the forward slip and the proportion of friction in the roll force, which together define the potential for the lubricant, i.e., to which extent the roll force can be reduced before the mill runs into slip. Thus, the values of the steel and process parameters where the role of the lubricant becomes critical can be identified. The influence of four steel- and process-related parameters was explored with the model. Their range can be seen in Table 1a. In Table 1a, the steel hardness is represented by a hardness parameter K and a strain-hardening parameter n, which define the tensile stress σ as function of strain ε as by: σ = K(ε + 0.04) n MPa. The tensile yield stress s0 of the incoming material to stand 1 is also given. In Table 1b, the values of the five chosen lubricant parameters are shown, which are typical values for cold rolling lubrication. First, a representative example is given for the hardest steel type and medium values for incoming gauge, reduction and mill speed (see Table 1a), and for the lubricant shown in Table 1b. In Table 2, some data for this mill are given, showing the strip thickness, the plane-strain yield stress development

(Eq. 2)

where

The factor 2 accounts for the fact that the strip is cooled from both sides. A realistic value for the heat transfer coefficient of sprayed emulsions5 is 25,000 W/m2K. Another input parameter for the model is the strip roughness. The incoming strip roughness in stand 1 is assumed to be 2 µm. The strip roughness entering the subsequent stands, i.e., the exit roughness from the preceding stands is calculated in the model, using the fact that here is a relation between roughness, film thickness (known) and contact area (known).6 As lubrication becomes better, i.e., exit film thickness becomes larger, the strip roughness decreases less. As mentioned, an extensive range of lubricant parameters is included in the model. In this paper, only five are highlighted/investigated: the thickness of pre-applied oil layers (on top of any pre-existing oil inherited from the previous stand), the emulsion

Table 1 Overview of the Range of the Four Steel and Process-Related Parameters Explored in the Model (a) and the Chosen Lubricant Properties for the Steel and Process Calculations (b). The emulsion temperature is 50°C.

Tc = the temperature of the cooling emulsion, Ti = the strip temperature at the beginning of the cooling zone, H = the heat transfer coefficient, h = the strip thickness and v = the strip speed.

Low value Medium value High value

K (MPa)/n (-)/s0 (MPa)

Gauge (mm)

Reduction (%)

Exit speed (m/min)

800/0.230/382

250

950/0.195/507

375

1,100/0.160/657

500

(a)

Chosen lubricant value

Pre-applied oil on strip (µm)

Emulsion concentration (%)

Oil particle size (µm)

Oil viscosity (mPas)

CoF in boundary contact (-)

0.11

(b)

5 Table 2 Mill Data for the Hardest Steel Type and Medium Values of Gauge, Reduction and Exit Speed, for a Typical Lubricant Before stand 1 Gauge (mm)

3.00

Reduction (%) Reduction cumulative (%)

Stand 1

Before stand 2

Stand 2

2.40 20

Before stand 3

Stand 3

1.92 20

Before stand 4

Stand 4

1.54 20

After stand 4 1.23

48.8

756

1,041

1,151

1,223

1,278

Tensions (MPa)

100

Strip speed (m/min)

154

192

240

300

375

Plane-strain yield stress (MPa)

Temperature rise deformation (°C)

51.4

Temperature decrease cooling (°C)

62.7

67.9

71.6

15.0

42.0

56.7

Strip entry temperature (°C)

25.0

61.4

82.1

93.3

Strip roughness (µm)

2.00

1.31

0.84

0.63

Work roll roughness (µm)

0.8

0.4

Figure 2

Pressure hills for the four stands of the rolling mill shown in Table 2 for hard steel and a typical lubricant.

I I

fraction of MPHL (Fig. 3c) can be seen to increase from stand to stand, mainly due to the higher sliding speeds in the contact and the progressively lower strip roughness. It also strongly increases with steel hardness: the high contact pressures greatly increase the lubricant viscosity, which is another main driving force for MPHL. Forward slip (Fig. 3d) is relatively high in stand 1, which is due to the lower back tension in stand 1. The significant in-bite film formation in stand 4 for the hardest steel causes a strong decrease in forward slip. It was verified with extra calculations that the decrease in forward slip is indeed much more pronounced toward the hardest steel. The mill stand power (Fig. 3e) increases from stand to stand, due to the increasing speeds, but also increases significantly with steel hardness, due to the higher roll forces. The proportion of the roll force due to friction (Fig. 3f) is highest for the hard steel in stand 1 (due to the longer contact length by stronger

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(related to the tensile stress, i.e., higher by a factor 2/√3), speeds, the temperature development due to deformation and cooling, and the strip roughness. In Fig. 2, the pressure hills can be seen for stands 1–4. The contact pressures typically peak at about 1.5 GPa. With four steel and process parameters that assume three values each, a large number of calculations would have to be done for an analysis of all possible combinations. Therefore, for this work, the number of calculations was reduced significantly by adopting a quadratic design-of-experiments approach. An added advantage is that mill performance can be estimated/predicted for combinations of intermediate values for hardness, gauge, reduction and speed. For the study in this section, 20 calculations were done with carefully chosen combinations of the four steel and process parameters. Note that the reductions in all four stands are assumed equal. In Fig. 3, graphs of selected mill performance and friction related parameters for stands 1–4 can be seen, as predicted by the design-of-experiments software. The graphs show how these parameters depend on the steel hardness, going from relatively soft (blue color) to the hardest steel (red color). The results are given for the medium values for the other parameters, i.e., gauge 3 mm, reduction 20% and mill exit speed 375 m/minute. The first graph (Fig. 3a) shows that the maximum contact pressure for hard steel rolling, for the typical conditions specified in Tables 2 and 1b, is ca. 1.5 GPa (see also Fig. 2). The roll force (Fig. 3b) increases significantly with hardness, and slightly from stand to stand, due to strain hardening. For the hardest steel, in stand 4, roll forces tend to decrease slightly, which is due to a significant amount of in-bite film formation (MPHL) under these conditions. The

0.4

0.52

AIST Transactions Figure 3

(a) (b) (c) (d)

(e) (f) (g) (h)

Various mill and friction-related properties plotted for stands 1–4, for soft, medium and hard steel: maximum contact pressure (a), roll force (b), fraction of micro-plasto-hydrodynamic lubrication (c), forward slip (d), power (e), percentage of roll force due to friction (f), coefficient of friction (g) and fraction of asperity contact (g). The legend for all graphs is shown in the first graph.

elastic deformation of the work roll under the high roll forces), but toward stand 4 it decreases due to the MPHL effect. The global coefficient of friction (CoF) (Fig. 3g) decreases from stand to stand, which is due to the higher speeds and lower entry angles, leading to better film formation in the inlet zone, and progressive contributions from MPHL within the roll bite. The fraction of asperity contact (BL) averaged over the roll bite (Fig. 3h) decreases from stand to stand due to the better film formation at the higher speeds. Comparing Figs. 3g and 3h, it can be seen that the CoF is strongly influenced by this fraction. These calculations are just a cross-section; no calculation results were presented for the 2- or 4-mm gauge, the 15 or 25% reduction, and the mill exit speeds of 250 or 500 m/minute. Nevertheless, for the hardest steel type, the used values for gauge, reduction and speed are quite realistic. Therefore, it can be concluded that for hard steel rolling, in case a typical lubricant is used, the percentage friction in the roll force is of the order of 20%. This means that there is some scope for an improved lubricant to reduce roll force and mill power, but it has also been seen that this will be limited by slip in stand 4. From the design of experiments, it also follows that the proportion of friction in the roll force

decreases as the reduction decreases. In those cases, there is limited room for improvement of a typical lubricant. Similarly, it was shown that with increased mill speeds or reductions, due to the higher sliding speeds, the amount of MPHL increases rapidly, accompanied by lower forward slip. Calculations — The Influence of Lubricant Properties on the Rolling Process — For a lubricant supplier, it is of course most important to investigate how the rolling process can be influenced by the lubricant. There are several ways in which lubrication can be influenced, some of which were already introduced in an earlier publication: 3 pre-applied oil layers on the strip, emulsion concentration, oil particle size, viscosity and the coefficient of friction in the boundary lubrication regime. It was decided to explore the same parameters, but now in a more systematic way. A range of realistic values was investigated, which is shown in Table 3a. Just as in the previous section, a large number of calculations would have to be done for an analysis of all possible combinations of five lubricant parameters which assume three values each. Therefore, also in this section, the number of calculations was reduced by adopting a quadratic design-ofexperiments approach, in this case involving 26

7 Table 3 Overview of the Range of the Five Lubricant-Related Parameters Explored in the Model (a) and the Chosen Steel and Process Properties for the Lubricant Calculations (b) Pre-applied oil on strip (µm)

Emulsion concentration (%)

Oil particle size (µm)

Oil viscosity (mPas)

CoF in boundary contact (-)

1.0

0.090

Medium value

0.25

2.5

0.105

High value

0.50

4.0

0.120

Low value

(a)

Chosen steel and process value

K (MPa)/n (-)/s0 (MPa)

Gauge (mm)

Reduction (%)

Exit speed (m/min)

1,100/0.16/657

375

(b)

calculations. The calculations are based on the hardest steel type explored in the previous section, but with intermediate values of gauge, reduction and mill speed (see Table 3b). Again, a complete tandem mill was calculated with four reduction stands, and again with a higher work roll roughness in stand 1. The design-of-experiments software can be used to obtain predictions for any combination of lubrication parameters, as long as they are within the original range. In Figs. 4–8, the effect of each of the five lubricantrelated parameters on the mill will be illustrated. This is done by varying the value of each parameter as low-medium-high, while keeping all other parameters at their medium value, unless otherwise indicated. In the figures, the lubricant that is generally considered the worst is indicated in red, medium in green and best in blue. In Fig. 4, the influence of application of extra oil layers is shown. In Fig. 4a, it can be seen that the roll force is the highest when no extra oil is

applied. When 0.25 µm of oil is applied, the roll forces decrease by roughly 5%, but increasing the oil thickness even further, to 0.5 µm, has no additional effect; a saturation seems to have been reached. The reason for this can be explained as follows. Fig. 4b shows that as application conditions become worse, i.e., thinner pre-applied oil layers, the film thickness decreases. An important reason for this is starvation, which occurs when the oil pool in front of the roll bite is insufficiently replenished by the emulsion, leading to entrainment of water as well as oil into the contact and a decrease of the film thickness. This mechanism has played a role here. To illustrate this, Fig. 4c shows the effect of the application of extra oil layers on the entrained oil concentration. For the worse lubrication situation, with little pre-applied oil, there is significant starvation, i.e., entrained oil concentrations as low as 50% are found. Upon improving the lubrication condition, entrained oil concentrations increase, sometimes to values close to 100%.

Figure 4

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(a) (b) (c) (d)

The influence of the thickness of pre-applied oil layers (0, 0.25 and 0.50 µm) on roll force (a), film thickness (b), entrained concentration (c) and forward slip (d) for stands 1–4. The other lubricant properties are: emulsion concentration 2.5%, oil particle size 4 µm, viscosity 40 mPas and CoFBL 0.105. The legend for all graphs is shown in the first graph.

AIST Transactions Figure 5

(a) (b) (c) (d)

Comparing Figs. 4b and 4c, it can be concluded that oil starvation effects are a main trigger for the film thickness. In cases where, going from the medium to the best lubrication situation, the entrained oil concentration increases little further (Fig. 4c) film thickness increases no further (Fig. 4b), with ultimately no further reduction of roll force (Fig. 4a). In Fig. 4a, another effect can be seen: the decrease in roll force toward stand 4. This cannot be explained by film formation and starvation. Instead, it is caused by significant in-bite film formation (MPHL), mainly in stand 4 due to the high sliding speeds. This was also seen in Fig. 3c: for hard steel rolling, for typical rolling conditions (Table 2) and a typical lubricant (Table 1b), there is a significant amount of MPHL, especially toward the later stands. Fig. 4d plots the effect of the application of extra oil layers on the forward slip. The forward slip is relatively high for stand 1, which is due to the lower back tension. Further,

it can be seen that going from stand 2 toward stand 4, forward slip values tend to decrease, a result of the lower friction due to the higher entrained film thickness at progressively higher speeds (Fig. 4b) and more pronounced MPHL effects. In line with the effect on roll force, the forward slip also saturates at application of 0.25 µm of oil. In Fig. 5, the influence of the emulsion concentration is shown. A clear saturation effect is seen in roll force and forward slip, although the entrained concentration has not quite reached 100% and the film thickness still increases slightly going from the medium to the highest concentration. In Fig. 6, the influence of oil particle size is shown. Saturation has not yet been reached at the highest particle size, but considering the approach of the entrained concentration to 100% at the highest particle size, it is very close to the maximum film thickness, and thus the lowest roll force achievable

Figure 6

The influence of the concentration (1.0, 2.5, 4.0%) on roll force (a), film thickness (b), entrained concentration (c) and forward slip (d), for stands 1–4. The other lubricant properties are at their medium value: oil particle size 4 µm, viscosity 40 mPas and CoFBL 0.105. There is no extra oil applied. The legend for all graphs is shown in the first graph.

(a) (b) (c) (d) The influence of the oil particle size (2, 4, 6 µm) on roll force (a), film thickness (b), entrained concentration (c) and forward slip (d) for stands 1–4. The other lubricant properties are at their medium value: emulsion concentration 2.5%, viscosity 40 mPas and CoFBL 0.105. There is no extra oil applied. The legend for all graphs is shown in the first graph.

9 Figure 7

(a) (b) (c) (d) The influence of the viscosity (20, 40, 60 mPas) on roll force (a), film thickness (b), entrained concentration (c) and forward slip (d) for stands 1–4. The other lubricant properties are at their medium value: concentration 2.5%, oil particle size 4 µm and CoFBL 0.105. There is no extra oil applied. The legend for all graphs is shown in the first graph.

with this lubricant parameter. Note in Fig. 6c the significant starvation at the lowest particle size. In Fig. 7, the influence of the viscosity is shown. Increasing the viscosity continues to be effective up to the highest value. The saturation effect that was seen in the previous figures is absent: film thickness continues to increase with viscosity due to the hydrodynamic effect and the entrained oil concentration shows no strong dependence on viscosity. For the highest viscosity, the tendency for the roll force to decrease toward the later stands is stronger compared to what was seen in Figs. 4–6. This is caused by a strong MPHL film formation for the higher viscosities, which is as expected because viscosity is an important driving force for the MPHL mechanism, especially when viscosities are greatly increased under the high pressures when rolling hard steel types. In line with the roll force, forward slip values tend to decrease with increasing viscosity, and for the

highest viscosity even tends toward a slip condition in stand 4. Fig. 8 shows the influence of the coefficient of friction in the boundary lubrication regime (CoF BL) on the mill performance (roll force, forward slip, global (average) coefficient of friction and the mill stand power). There is a steady decrease of the roll force, forward slip and global CoF as the boundary friction decreases. This is expected to continue as CoF BL values would decrease further but, as can be seen in Fig. 8b, will be limited by slip, starting at the last stand. The effect of the CoF BL on the global CoF is highest in stand 1 and lowest in stand 4 because the proportion of boundary lubrication is highest in stand 1 and much lower in stand 4, analogous to Fig. 3h. Mill stand power increases from stand to stand, which is directly linked to the rolling speed, but the difference between the lubricants is limited.

Figure 8

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(a) (b)

The influence of the CoFBL (with values 0.090, 0.105, 0.120) on roll force (a), forward slip (b), global CoF (c) and mill stand power (d) for stands 1–4. The other lubricant properties are at their medium value: concentration 2.5%, oil particle size 4 µm and viscosity 40 mPas. There is no extra oil applied. The legend for all graphs is shown in the first graph.

AIST Transactions The above leads to several important conclusions. First, starvation effects play a very important role in the inlet zone. The extent of starvation is determined by the combination of pre-existing oil, pre-applied oil, emulsion concentration and oil particle size. This determines whether the inlet zone immediately before the deformation zone is fully flooded with oil, maximizing film thickness, or whether water is entrained in the contact, reducing the film thickness. Once a fully flooded situation occurs, increase of pre-applied oil thickness, concentration or particle size has no further effect, with no accompanying effect on roll force. In contrast, increased viscosity will continue to increase film thickness by its hydrodynamic effect. Second, in-bite film formation is an important factor involved in reducing friction and roll forces, and is significantly increased by viscosity, especially at the high pressures occurring when rolling hard steel, in some cases even leading to a tendency for slip. Third, the CoF in the boundary regime has a noticeable effect on roll force, especially under conditions of low speed, low reductions, and little inlet zone or in-bite film formation. Table 4 shows the achievable reductions in roll force going from the medium to the best lubrication condition for hard steel rolling and process parameters as in Table 3b. Note that this corresponds to improving an already adequate lubrication. The achievable reductions shown in Table 4 are relatively low, which corresponds to the conclusions in an earlier article.3 The conclusions can now be extended because it has been seen in Fig. 7d that, especially in stand 4, an improvement of an already adequate lubricant may result in low to negative forward slip values, i.e., a tendency for slipping. Note that this occurs despite the fact that the theoretical, maximum reduction in roll force, i.e., the ~20% friction in the roll force (Fig. 3f), has not yet been reached. Of course, this further limits the potential of the lubricant. On the other hand, there are several ways in which a mill may be optimized to reach a good compromise, for instance by adjusting the reduction schedule or front and back tensions, by optimizing lubricant and emulsion properties, or by strategic emulsion application (e.g., flexible lubrication7). It is important to realize that the calculations are done for one specific rolling condition, and trends may be different for other steel types, other

reductions or even other combinations of lubricant parameters. Nevertheless, as it has been seen, the calculations are illustrative and are useful to investigate the trends by which film formation, entrained oil concentration and, e.g., MPHL effects may influence mill performance, and how this may be influenced by the lubricant. The calculations are also useful because they allow for exploration of the limits of lubrication and to determine where there is room for improvement. It must also be realized that several lubricant properties incorporated in the TRBM were not varied in these studies, such as the parameters quantifying the lubricant’s viscosity dependence on pressure, temperature and shear rate. Furthermore, there are many more factors that determine satisfactory rolling, such as the way emulsions behave over prolonged durations, their effect on strip cleanliness and on downstream processes, such as annealing, and the way the lubricant influences work roll wear and strip surface quality. Therefore, the development of lubricants cannot rely on calculations alone. Experience on how emulsions perform on the mill, how chemical composition influences this, as well as further developments in laboratory testing, remain as important as ever.

Experimental As it has been seen, mathematical modeling can be useful to explore the effect of a range of physicochemical properties of the lubricant on mill performance. However, not all lubricant parameters can be addressed in the TRBM model. For instance, as mentioned, the surface protection provided by antiwear additives is not accounted for in the model. Also, the particular chemical structure of emulsifiers, potentially influencing emulsion properties and film formation, cannot be accounted for in the model. This is the reason that experimental evaluation of the lubricant remains necessary in lubricant development. For a full lubricant evaluation, several tests are required because a wide range of tribological mechanisms is involved in cold rolling lubrication, as was mentioned in the introduction: inlet and in-bite film formation, friction in the boundary lubrication regime, and high-pressure, high-temperature and

Table 4

Percentage Roll Force Reduction (% RFR) for Stands 1-2-3-4 (rounded off to the nearest integer), by the Indicated Change of Lubrication-Related Parameters, While Keeping the Other Values Constant, as Indicated in the Captions of Figs. 4–8

% RFR

Pre-applied oil on strip (0.25 → 0.5 µm)

Emulsion conc. (2.5 → 4%)

Oil particle size (4 → 6 µm)

Oil viscosity (40 → 60 mPas)

CoFBL (0.105 → 0.090)

0–1–2–0

0–0–0–2

3–4–4–6

2 – 3 – 7 – 12

3–4–4–3

11 high-shear rate tribology of the base lubricant. The following section gives an example of a laboratory test that is used to develop and optimize lubricant performance prior to offering the products to production mills. As production mills are increasingly faced with the challenges of rolling hard steel types, the laboratory tests should account for the mill conditions characteristic for rolling these steel types. Most prominently, hard steel rolling occurs at very high pressures (see Fig. 2 and Fig. 3a), so in this scope, laboratory testing is aimed mainly at choosing realistic pressures. Laboratory Testing — In an earlier paper,8 it was shown that the ability for an emulsion to protect the mating surfaces under severe conditions could be evaluated in so-called roll bite-mimicking tests. In these tests, the frictional conditions in the roll bite of a particular stand/pass (frictional power intensity, energy pulse, flash temperature) are mimicked in a tribometer, by specific combinations of ball speed, disc speed and load. The tests are able to distinguish between emulsions/lubricants by differences in coefficient of friction and the tendency for scuffing, and were demonstrated to correlate with field experience. In order to develop lubricants suitable for rolling hard steel types, the tests were modified to account for the very high pressures occurring when rolling these steel types: higher loads were applied and hard steel disc specimens used. Fig. 9a shows the result of the original roll bite mimicking test for three emulsions. The coefficient of friction is plotted for six tests; three tests

mimicking the frictional power intensity and energy pulse in stands 2 and 4 of a sheet tandem mill and stand 5 of a tinplate mill (FPI 2,4,5), and three tests mimicking the flash temperature in the same stands (TF 2,4,5). The performance of the emulsions is similar. In Fig. 9b, the results of the modified roll bite mimicking tests at higher pressures, for the same emulsions, can be seen. There are now significant differences in CoF and the number of passed tests (tests that resulted in scuffing are not plotted). The emulsions are clearly distinguished by their CoF and by their tendency for scuffing, with emulsion 2 giving the best overall performance. The results show that these modified roll bite–mimicking tests can distinguish emulsions based on test conditions relevant to rolling hard steel types. Note that a tendency for scuffing in the test does not necessarily mean there will be a serious problem on the mill; it merely means that on the microscopic scale some deficiency in surface protection may occur, leading to more metal-to-metal contact and an increase in friction, and thus roll force. Pilot Mill Trials I — Pilot mill trials can be used to assess cold rolling lubricants under real rolling conditions. Lubricants that were first evaluated in the laboratory can thus be tested under real rolling conditions prior to use on a production mill. Pilot mill trials were carried out to test the performance of eight emulsions. A hard steel type was rolled, with tensile stress σ as function of strain ε described by the equation σ = 1,060(ε + 0.04)0.17 MPa. The original strip thickness was 2.5 mm and

Figure 9

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(a) (b)

Original roll bite mimicking tests (a) and modified roll bite mimicking tests at high pressures (b). The bars show the CoF in the tests mimicking the frictional power intensity in three stands (FPI in the legend) and the three tests mimicking the flash temperature in three stands (TF in the legend). The bars are not plotted if the test fails (when scuffing occurred).

AIST Transactions Table 5 Incoming gauge (mm)

Reduction (%)

Back tension (MPa)

Front tension (MPa)

Exit speed (m/min)

Pass 1

2.50

7.6

Pass 2

2.31

11.8

110

100

Pass 3

2.04

12.2

118

150

Pass 4

1.79

14.0

140

200

Pass 5

1.54

14.9

150

250

Pass 6

1.31

14.9

150

300

Pass 7-a

1.11

16.3

150

100

Pass 7-b

1.11

16.3

150

350

its width was 0.15 m. Prior to the trials the rustprotective oil on the strip was removed by four times open-gap washing using a cleaning emulsion of the same composition as the test emulsion, but from a separate emulsion tank, so as to avoid contamination of the test emulsion with rust-protective oil. The first rolling pass, carried out at low reduction and low speed served to correct for deviations in strip thickness. The trials consisted of six subsequent passes at reductions of ca. 12–16%. Rolling speed increased from pass to pass, and in pass 7, a low and high rolling speed were chosen. See Table 5 for an overview of the rolling conditions. The tested emulsions varied mainly in type and amount of lubricity additives and extreme pressure/anti-wear agents. The emulsions were prepared on the day prior to the trial and were circulated overnight to allow equilibration. The roll forces in these trials are plotted in Fig. 10. Significant differences were observed between the

emulsions. Compared to the average roll force for these eight emulsions, roll forces vary by up to ±6%. Fig. 10b shows that for all emulsions roll forces decrease with rolling speed, on average by ca. 4% going from 100 to 350 m/minute. On the whole, the ranking of the eight emulsions is similar for all passes. Note that the emulsion properties of the eight emulsions (such as concentration and oil particle size) and the lubricant viscosities were similar, so that the differences can indeed be ascribed to the lubricity additives.

Pilot Mill Trials II — Further pilot mill trials were carried out with the same steel type, this time to investigate the influence of emulsion concentration on roll force. This was done in analogy with results presented in an earlier paper,3 in which the effect of oil particle size and viscosity was investigated. In view of this, the trials were carried out under the same conditions as in the earlier paper, i.e., slightly different as described in Table 5, and the pass numbers differing by 1. The mill setup and some information concerning the calculation of the CoF are described in an earlier paper.3 The trials were carried out in duplicate, with freshly prepared emulsions, equilibrated overnight. Table 6 gives an overview of selected properties of the emulsions. Analogous to how the results were presented in the earlier paper,3 Fig. 11a plots the coefficient of friction from pass to pass and in Fig. 11b at the two speeds in pass 8. The CoF decreases with pass number, which is due to the increasing speed and decreasing bite angle, both leading to better film formation, and in

Figure 10

Pass Schedule for the Pilot Mill Trials Evaluating the Eight Emulsions

(a) (b) Roll force for eight emulsions: from pass to pass (for speeds see Table 5) (a), and at two speeds in pass 7 (b).

13 Table 6 Selected Lubricant/Emulsion Properties Concentration (%)

Oil droplet size (µm)

Oil viscosity (at 40°C, mPas)

Emulsion (2.5%)

2.56

10.1

53.7

Emulsion (0.5%)

0.51

10.8

53.7

pass 8 decreases with speed, due to better film formation. Strikingly, for these emulsions, the CoF does not depend on the concentration. This seems to be in contrast to the experience that film formation, an important influencing factor for lubricity, normally depends on concentration. To illustrate this, Fig. 12a shows the film formation of the emulsion, at 0.5 and

2.5% concentration, measured with an ultrathin-film interferometer (PCS, London). The film thickness of the 0.5% emulsion is clearly CoFBL lower than that for the 2.5% emulsion, at all (-) speeds, suggesting that this emulsion suffers 0.104 from significant starvation. 0.104 The pilot mill trials were also modeled with the TRBM model, using measured lubricant properties, some of which are shown in Table 6. Fig. 12b shows that, also for the calculations, the film thickness is significantly lower with the lower emulsion concentration. Fig. 12c confirms that this is due to starvation of the 0.5% emulsion, which gives entrained oil concentrations significantly lower than 100%. Nevertheless, in the calculations, the increase in film thickness by the 2.5% emulsion

Figure 11

(a) (b) Coefficient of friction in the pilot mill trials for the emulsion at two concentrations, from pass to pass (a), and in pass 8, for the two chosen speeds (b). For the pass schedule refer to an earlier article.3

Figure 12

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(a) (b) (c) (d)

Film formation of the two emulsions, measured with the interferometer at 55°C (a), film thickness and oil concentration entrained in the roll bite as calculated with the TRBM model (b and c), and the effect of an increased concentration on the roll force, as seen on the pilot mill (line with large dots) and as calculated with the TRBM model (dotted line) (d).

AIST Transactions compared to the 0.5% emulsion is apparently insufficient to have a significant influence on the roll force: 0±1%, as shown in Fig. 12d; in fact just as in the pilot mill trials (shown in the same figure). In that sense, the model provides a good prediction for the potential for the lubricant to reduce the roll force, as it did in the previous paper.3 Contributing to the small effect of concentration on roll force is the comparatively low proportion of friction in the roll force in these trials: it was calculated that this is only ca. 10%, mainly connected to the low reductions. The pilot mill trials show that the effect of the lubricant on roll force and coefficient of friction can be measured well. However, the trials also indicate that the effects on roll force may be relatively low. This is demonstrated by this paper and an earlier article,3 which have shown that relatively large variations in concentration (2.5% vs. 0.5%), particle size (10 µm versus 2 µm) and viscosity (80 mPas versus 40 mPas) have a limited effect on roll force on the pilot mill (varying between +1 and –4%). Some of this may be due to the absence of strong starvation. In contrast, Fig. 10 has shown that relatively minor variations in chemical additives have a relatively large effect on roll force in the pilot mill (varying between –6% and +6%). In the calculations, it was found that starvation effects in the pilot mill may be less strong than in a tandem mill. This is due to the smaller work roll diameter and larger bite angles, leading to less powerful wash-off effects. This has the advantage that in a pilot mill the effect of chemistry on roll force can be investigated with less interference from starvation effects.

Conclusion Mathematical modeling has shown that when rolling hard steel types with a typical lubricant, an improved lubricant has some scope to reduce roll force and required mill power, but also that there is an increased risk for slippage, notably in stand 4. The effect on roll force and required mill power is caused by increased film formation, in turn caused by higher entrained oil concentrations or oil viscosity. It was also confirmed that in the case of a fully flooded contact no further improvement of, e.g., pre-applied oil or emulsion concentration can be achieved. At the later stands, due to the higher sliding speeds, and especially for high-viscous lubricants, the amount of in-bite film formation often becomes significant, which is accompanied by a further reduction of forward slip. Laboratory testing allows the overall lubrication of emulsions to be tested under conditions relevant to cold rolling hard steel types, notably at realistic high

pressures. Pilot mill trials provide an opportunity for the testing of emulsions under real rolling conditions. Because in some cases less starvation occurs in pilot mill trials, and the inlet zone immediately before the deformation zone is often fully flooded with oil, the effect of chemistry can be well investigated. Pilot mill trials suggest that the influence of chemical additives on roll force may be stronger than typical variations in emulsion properties. Future pilot mill trials will further investigate the influence of the lubricant on starvation. This is important in its own right, but will also define the conditions under which starvation effects are minimized. These will then provide optimum conditions to study the effects of lubrication additives and extreme pressure/antiwear agents on rolling performance.

Acknowledgment The author would like to thank Martin Bergmann, Primetals Technologies Austria, for his assistance during model calculations; the pilot mill team at Quaker Houghton in China: Deniel Zhang, Ze Feng, Kai Ye, Damon Zhu and Wenbing Jiang; and Pablo Bakermans, Maarten van Ham, Jan Melsen and Peter Schellingerhout, Quaker Houghton in the Netherlands, for their help in laboratory testing, model calculations and useful discussions.

References 1. J.B.A.F. Smeulders, “Roll Cooling and Lubrication in Cold Rolling,” The Making, Shaping and Treating of Steel ®, Flat Products Volume, Association for Iron & Steel Technology, Warrendale, Pa., USA, 2014, Chapter 13.2. 2. J.B.A.F. Smeulders, “Lubrication in the Cold Rolling Process Described by a 3D Stribeck Curve,” Iron & Steel Technology, Vol. 11, No. 2, 2014, pp. 37–45. 3. J.B.A.F. Smeulders, “Lubrication Strategies for Rolling Hard Steel Types,” Iron & Steel Technology, Vol. 16, No. 2, 2019, pp. 32–42. 4. M. Bergmann et al., “Enhanced Modeling of Friction and Lubrication in Cold Strip Rolling,” 9th International & 6th European Rolling Conference Proceedings, June 2013, Venice, Italy. 5. A. Horák et al., “Research on Cooling Efficiencies of Water, Emulsions and Oil,” Eng. Mech., Vol. 18-2, 2011, pp. 81–89. 6. W.R.D. Wilson and D.F. Chang, “Low Speed Mixed Lubrication of Bulk Metal Forming Processes,” J. Trib., Vol. 118/1, 1996, pp. 83–89. 7. M. Laugier et al., “Flexible Lubrication Concept: The Future of Cold Rolling Lubrication,” J. Eng. Trib., Pt. J, Vol. 225, 2011, pp. 949– 958. 8. J.B.A.F. Smeulders, “Novel Laboratory Lubrication Tests for Cold Rolling Emulsions,” Iron & Steel Technology, Vol. 12, No. 2, 2015, pp. 56–64. F

AIST Transactions seeks to publish original research articles focusing on scientific or technical areas relevant to the iron and steel community. All articles are peer-reviewed by Key Reviewers prior to publication. The scope of the Transactions spans all fields of iron and steel manufacturing, from extractive metallurgy to liquid metal processing, casting, thermomechanical processing as well as coating, joining/welding and machining. Articles related to structure and properties, i.e., characterization, phase transition, chemical analysis and testing of creep, corrosion or strength are accepted. In addition, papers on process control, testing and product performance are also encouraged. The Transactions will publish papers on basic scientific work as well as applied industrial research work. Papers discussing issues related to specific products or systems can be considered as long as the structure follows a scientific investigation/discussion, but not for the aim of advertising or for the purpose of introducing a new product. Papers should be no longer than 5,000 words, or 12 pages including figures. Papers will either be accepted unconditionally, conditionally based on reviewers’ comments, or rejected. The review process should be completed in less than two months. Upon acceptance, authors will be asked to complete a form to transfer copyright to AIST. A PDF proof of the accepted manuscript will be sent electronically to the author(s) approximately one month prior to publication for final approval. Reprints of published manuscripts can be ordered through AIST.

AIST Transactions Key Reviewers

Dr. Fred B. Fletcher ArcelorMittal Global R&D Research and Development

Dr. P. Kaushik ArcelorMittal Global R&D Steelmaking Process Research Prof. Patricio F. Mendez Department of Chemical and Materials Engineering, University of Alberta

Dr. Ronald J. O’Malley Missouri University of Science and Technology

Please contact me with any questions regarding AIST Transactions. P. Chris Pistorius POSCO Professor Carnegie Mellon University 5000 Forbes Ave. Pittsburgh, PA 15213 USA Phone: +1.412.268.7248 pistorius@cmu.edu

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Dr. Stefanie Sandlöbes Max-Planck-Institut für Eisenforschung GmbH Dr. Il Sohn Neo-Metallurgical Processing Lab, Department of Materials Science and Engineering, Yonsei University Prof. Lifeng Zhang School of Metallurgical and Ecological Engineering, University of Science and Technology Beijing

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Prof. Hatem S. Zurob McMaster University

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