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Ice Adhesion —Theory, Measurements and Countermeasures Article  in  Journal of Adhesion Science and Technology · March 2012 DOI: 10.1163/016942411X574583

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Journal of Adhesion Science and Technology 26 (2012) 413–445 brill.nl/jast

Ice Adhesion — Theory, Measurements and Countermeasures Lasse Makkonen ∗ VTT Technical Research Centre of Finland, Box 1000, 02044 VTT, Espoo, Finland Received on 18 February 2011

Abstract In this paper concepts and models for theoretically estimating ice adhesion are presented. The effects of temperature, ice salinity and properties of the substrate material on ice adhesion are explained by these theoretical concepts. Major problems caused by ice adhesion are outlined and the applications of the theory in combating ice adhesion are discussed. Measurement methods of ice adhesion are described and results of ice adhesion for various material surfaces are reviewed and interpreted in view of the theory. Prospects for reducing ice adhesion to solve practical problems are discussed and the methods of anti-icing and de-icing are summarized. © Koninklijke Brill NV, Leiden, 2012 Keywords Ice adhesion, interface mechanics, surface energy

1. Introduction Adhesion science and technology is usually aimed at improving adhesion. With ice, however, the problem is of an entirely different nature. Ice adhered to various surfaces, such as roads, aircraft wings, ship superstructure (Fig. 1), offshore structures, TV-towers, overhead power line cables, meteorological instruments and various machinery components, etc. is a severe problem which causes loss of lives and huge costs in terms of structural damage and accidents. Accordingly, the aim of this review is to increase the understanding of ice adhesion with the ultimate intent of combating the problems related to it. Theoretical aspects of ice adhesion are first discussed and some theoretical concepts are introduced. Ice adhesion experiments are then described at a general level and the factors that affect the comparison of experimental data are analyzed. Finally, the use of the *

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© Koninklijke Brill NV, Leiden, 2012

DOI:10.1163/016942411X574583

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Figure 1. An example of the many practical problems related to ice adhesion: Mechanical ice removal from ship’s superstructure.

Figure 2. Water drop on a solid surface and the definition of the contact angle θ .

theory and the experimental results in developing countermeasures for problems related to ice adhesion is discussed and future prospects are highlighted. 2. Theory 2.1. The Relation between Water Contact Angle and Ice Adhesion Consider a drop of water (w) on a solid (s) with an interface (w, s) and the corresponding surface energies γ and a droplet contact angle of θ . The situation is schematically shown in Fig. 2. The Young equation for the equilibrium of this situation reads γw,s + γw cos θ = γs .

(1)

Consider next ice (i) that is frozen on the solid (s). In Fig. 2 this would mean that the drop is frozen into ice. We are now interested in the work that is required to

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remove the ice, i.e., break the bond (i, s) and form two new surfaces (i and s) in the absence of deformations. We call this the thermodynamic work of adhesion Wa which then is defined as Wa = γs + γi − γi,s .

(2)

Inserting γs from equation (1) into equation (2) shows that Wa = γi + γw cos θ + (γw,s − γi,s ).

(3)

Considering now that the surface energies of water and ice are approximately the same [1] and assuming that their interfacial energies at the solid interface are also approximately the same, we obtain Wa ≈ γw (1 + cos θ ).

(4)

According to equation (4) the thermodynamic work of ice adhesion can be closely approximated by the surface tension of water and the contact angle of water on the material in question. This is presented graphically in Fig. 3. Equation (4) and Fig. 3 show that, ideally, we should in ice removal expect a deterministic dependence between the work of adhesion and the contact angle of water. Finding a deviation from the curve in Fig. 3 in macro-scale experiments would imply that work is spent in deformations of the materials or that the ice– solid contact is somehow complex or incomplete. These possibilities are discussed in the following sections.

Figure 3. Thermodynamic work of ice adhesion scaled by the surface tension of water as a function of water contact angle θ .

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2.2. The Effect of Material Deformations The work Ws spent in separating an interface into two surfaces by a perpendicular force F (x) is  δx Ws = F (x) dx, (5) 0

where δx is the distance at which the surfaces are considered to be ‘separated’. Here both Ws and F (x) are defined as per unit area of the original interface. If the surfaces were perfectly smooth and rigid, δx would be of the molecular scale, because the atomic forces decrease very rapidly with distance. For such ideal surfaces the mechanical work spent would include a reversible component only and equal the thermodynamic work of adhesion Wa . However, real material surfaces are deformed upon adhesional failure. The deformations may be either brittle or ductile. Ice itself may fail in both modes, and it is noteworthy that the elastic modulus of ice strongly depends on its temperature [2]. The distance δx in ice removal may thus be orders of magnitude larger than in the ideal situation discussed above. Work may also be done in forming non-planar failure plains and micro-cracks within the ice. Accordingly, the work spent in mechanically removing ice from a substrate is typically much more than the thermodynamic work of adhesion Wa even in the case of a relatively rigid substrate material. As an example, the fracture energy of an ice–steel interface has been experimentally determined as 1.1 J/m2 [3] while Wa from equation (4) gives 0.09 J/m2 for the same interface. This shows that the irreversible contribution to the work spent in removing ice from surfaces due to dissipative processes can be much larger than the reversible contribution due to intermolecular interactions across the interface. However, since the intermolecular interactions are the cause of the dissipative processes, we may still expect a strong correlation between Wa and the total work spent in ice removal [4], similarly to other materials [5]. In any case, the correlation between Ws and Wa is likely to be less significant when, not only the contact angle, but also the elastic modulus of the substrate material varies. The work of ice removal Ws , discussed above, is not the usual way to consider the adhesion strength of ice, however. Rather, one is interested in the adhesion strength which is defined as the maximum value Fmax of F (x) required to mechanically separate ice from the substrate. As discussed above and, e.g., in [5] the connection between Ws and Fmax is not straightforward. In fact, it is not evident why the material deformations during ice removal should affect the adhesion strength of ice at all. This can be explained by an indirect effect that arises due to the brittle-ductile nature of ice. When the substrate material is highly elastic and/or the temperature of ice is close to its melting point, the ice interface behaves as that of a ductile material with little tendency to fracture even at high strain rates. On the other hand, on a rigid substrate and/or cold ice, mechanical removal of ice occurs in a brittle manner. Thus

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the failure mechanism at the interface may vary with the flexibility of the interface. On rigid surfaces, this may theoretically cause a temperature dependence of Fmax , so that at low ice temperature, a lower ice adhesion strength would occur. This is indeed observed at very low temperatures, as discussed later in this paper. At temperatures close to 0°C there appear to be other mechanisms that override the effect of deformations. These mechanisms affect the ice–substrate interface already prior to an attempt to remove the ice and are discussed in the following sections. 2.3. The Effect of Interface Morphology and Crystal Structure of Ice Obviously, the nominal adhesional force per unit area between ice and a substrate is weaker when the true interface area is less than the apparent one. This is the case when the contact is incomplete, e.g., due to minute bubbles of air at the interface. Also, micropores of the substrate that are not filled with water due to hydrophobicity of the substrate surface reduce the effective contact area. On the other hand, the true contact area is larger than the apparent one when the interface includes perturbations, i.e., is rough on the microscopic scale. For example, if the interface includes micropores that are filled with water, then the true interface area is large. In such a case, sometimes called ‘locking’ a purely adhesional failure may become impossible, and the fracture occurs partly cohesionally within the ice material. Therefore, the adhesion strength of ice on many materials is higher on a microscopically rough surface than on a smooth surface [6, 7]. The true interface area of the contact of water with a textured surface can be estimated theoretically in detail. Apparently the fine details that affect the contact angle hysteresis, such as the sharpness and spacing of the roughness elements, are important [8, 9]. The theory can be used as guide for estimating ice adhesion when bulk water is frozen on textured surfaces, for example those that have a ‘lotus’ or fractal structure in order to make them super-hydrophobic [8–13]. However, these ideas are perhaps over-optimistic because they are based on the mechanical equilibrium between capillary forces only. In many cases of ice accretion in nature the ice is formed from water droplets that impact the surface at a very high speed. This involves high droplet inertia and often a high stress due to wind, so that water droplets may penetrate and freeze into the porous structure of the material, particularly when they are smaller in diameter than the roughness element spacing. Small droplets that are highly supercooled also freeze so quickly that they do not re-bounce from a superhydrophobic surface. Consequently, the adhesion strength of bulk-formed ice on textured surfaces may be different from that of ice formed by droplet accretion and the sizes of the surface roughness elements and the impinging droplets may affect the latter. This conclusion is supported by the fact that Teflon has been found to be quite poor in reducing adhesion of ice produced by droplet accretion [14, 15], while the adhesion strength of bulk-formed ice on Teflon has been shown to be very low in many other studies (Section 3).

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Figure 4. Crystal structure of ice photographed through a polarized plate on two substrate surfaces, concrete (above) and Teflon (below).

The size and orientation of crystals in bulk ice at an interface depend on the material on which the ice is frozen, see Fig. 4. When ice forms due to droplet accretion the growth conditions have an effect as well [16–19]. Generally, hydrophilic and rough surfaces cause smaller and more randomly oriented ice crystals to form. It has been proposed that the crystal structure at the interface would affect ice adhesion. However, this is difficult to demonstrate experimentally, because materials show different ice adhesion strengths for other reasons. Theoretically, it is not clear why such an effect should be significant, since the surface energies of the crystal faces of ice are not much different from each other [20].

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2.4. The Effect of Temperature 2.4.1. The Role of Pre-melting of Ice One would, of course, expect that a liquid film at the solid–ice interface reduces the adhesion strength. Anticipating that a very thin layer of water may not mechanically behave as a liquid, one would also expect that the thicker the liquid film, the lower the adhesion. In 1859 Faraday [21] showed by experiments on adhesion of small ice pieces to each other that there exists a liquid layer on ice at typical natural temperatures. After a long controversy, and more recent direct measurements, it has been agreed that this is indeed so. Generally, see, e.g., [22, 23], this has been explained by minimizing the total surface free energy when a liquid layer forms, i.e., by assuming that γi > γw + γw,i . In terms of Fig. 2 and equation (1) this requires assuming that the contact angle of water on ice is zero. It was shown, however, by Makkonen [1] that the contact angle of water on ice under relevant experimental conditions is almost 40°. This finding outlined that the commonly adopted theory on the surface of ice based on minimizing the surface free energy is without a basis [1, 2]. A more appropriate explanation for the liquid layer on ice [1, 24] and its consequences are briefly explained in the following. Consider first a surface just exposed by separation assuming that its interatomic distances are still the same as in the bulk material. Then there is a net inward force due to the imbalance of atomic forces across the interface. This force per unit area is a pressure which may be interpreted to cause the reconstruction of the surface into its final equilibrium state. It is reasonable to assume that the structure and interatomic distances of this reconstructed layer correspond to those formed when applying the same external pressure on a bulk material. For ice, contrary to most other materials, an excess pressure causes reduction in the equilibrium melting temperature. This causes the fact that a liquid water film exists on an ice surface even well below 0°C. Applying the Lennard–Jones potential of the interatomic forces and the Clausius–Clapeyron equation Makkonen [1] estimated the change in the surface equilibrium melting temperature δT . This result can be simplified into the following form: δT = 290|γi,w − γi,s |,

(6)

where δT is in °C. Here γi,s is the interface energy of the ice–solid interface and γi,w that of the ice–water interface, both in J/m2 . The value of γi,w = 29 × 10−3 J/m2 can be used here [25]. As an example of the results of equation (6), the surface pressure is sufficient to keep the outermost surface layer of an ice–vapour interface (γi = 75 × 10−3 J/m2 , [1]) in a liquid state at temperatures above −13°C. However, the use of equation (6) for ice in contact with solid materials is not straightforward. First, the interface energy γi,s is unknown. It can be estimated based on γs and γi but only theoretically. Using, e.g., the Berthelot rule suggests that, qualitatively, γi,s is high for the

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materials for which γs is high. Second, it is not clear if this theory applies to the cases where the surface energy of the solid is higher than that of ice, and if both the dispersion and polar components need to be taken into account. Hence, the predictive value of equation (6) is, perhaps, limited to the low energy surfaces, such as those polymers that may have γi,s close to the value of the ice– water interface, i.e., 29 × 10−3 J/m2 . At such interfaces, pre-melting should occur at temperatures close to 0°C only, whereas for the polymers that have an even smaller surface energy should involve a liquid film in a wider temperature range. This may introduce another mechanism, in addition to the one presented in Section 2.1, which produces a low ice adhesion at low energy surfaces. Based on the same theory, the thickness of the liquid water layer on ice can be estimated [1]. The theory suggests that the water film thickness at an interface of ice decreases from infinity at 0°C to zero at 0◦ C − δT , being proportional to δT −1/3 . Such relationship has been found also experimentally [26, 27]. More studies would be necessary to evaluate such a nanoscale film thickness at interfaces and its relation with ice adhesion. In any case, qualitatively, the theory predicts an increase in the adhesion strength of ice with decreasing temperature down to the temperature at which the liquid film disappears. 2.4.2. Interface Cracking Due to Thermal Expansion It is noteworthy that an ice/substrate interface always forms at the freezing temperature of water, i.e., 0°C for pure water. Thus, at the time of freezing there exists no direct effect of ambient temperature on the properties of the interface. However, since the initial ice/substrate interface is at the freezing temperature, any such interface at a sub-freezing temperature must have been cooled to that temperature in a solid state. Thermal contraction of both the ice and the substrate is always involved in such a cooling. Obviously, this may initiate failures of the joint at the interface and cracking of ice when the thermal expansion coefficients of the ice and the substrate material are different from each other. In most cases they are quite different as the linear thermal expansion coefficient of ice is 50 × 10−6 /°C [20], whereas that of many other materials it is much smaller. For steel, as an example the value is 11 × 10−6 /°C. This means that the stress that arises at the interface upon simultaneous cooling of the ice and the substrate is mainly caused by contraction of the ice itself. The ice and the substrate material may also cool at different rates due to their different heat capacity and coefficient of thermal conduction, causing an additional stress at the interface. The significance of the stress depends on the cooling rate because ice creeps. When the substrate material remains elastic down to the temperature in question, this effect is absent. That failure of an ice–substrate interface actually takes place by cooling only can be demonstrated by numerical analysis. Figure 5 shows a finite element method analysis of the shear stress caused by thermal contraction in a rapid temperature change of 10°C in the U.S. Army Cold Regions Research and Engineering Laboratory (CRREL) ice adhesion test arrangement (see Fig. 7). This analysis and the corresponding ones for the other stress components indicate that the thermally

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Figure 5. Finite element method analysis of the shear stress (in Pa) caused by thermal contraction in a rapid temperature change of 10°C in the CRREL ice adhesion test arrangement (see Fig. 7).

induced stresses at the ice–substrate interface already at −10°C are higher than the adhesional and cohesional strength of ice. Thus, in this test arrangement the ice–substrate interface has been partly broken and/or the ice itself is damaged by cracking already prior to the adhesion test. In experiments on ice adhesion this effect can be alleviated by letting the samples cool very slowly so that the creep of ice partly relaxes the thermal stresses that arise. One would expect that thermally induced failures of the ice and of the joint at the interface will reduce the measured value of the adhesion strength. In other words, the adhesion strength of ice should theoretically decrease with decreasing temperature due to the effect of differential thermal expansion, except for elastic materials. This is indeed observed (see Figs 11 and 12) but only at rather low temperatures. Apparently, the disappearance of the liquid-like layer with decreasing temperature (Section 2.4.1) provides a more important mechanism than the thermal expansion at temperatures close to the freezing point of water. Moreover, as long as the liquid-

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like layer exists the interface may be ductile so that it adjusts to the thermal stresses without fracturing. 2.5. The Effect of Water Salinity Impurities and chemicals affect the adhesion of all materials. This general subject will not be discussed here. However, in the case of ice adhesion, salt in sea water requires attention due to its major role in many problems related to accretion of ice at sea [28]. Salinity at the interface always is a factor when the ice is formed out of sea water. The physical processes that are involved in saline ice adhesion are discussed in the following and an attempt is made to, at least qualitatively, explain some of the observed phenomena. Saline ice consists of salt-free ice and the so-called brine pockets that contain saline water. The salt concentration and relative volume of these brine pockets depend on the bulk salinity and temperature of the ice [20]. The reduction in the adhesion strength of ice caused by salt is apparently due to reduced effective solid– solid contact area at the ice–substrate interface. It is reasonable to assume that the adhesion strength depends linearly on the effective contact area. The effective contact area can be related to the brine volume assuming some geometrical form of the brine pockets. Oksanen [29] used these ideas to model saline ice adhesion by adopting the geometry of brine pockets as vertically oriented cylinders. He assumed in his model that the ice–substrate interface structure is equal to that of a vertical cut of the ice. However, this model predicted much higher values of the adhesion strength than observed since the portion of the brine at a plane cut area is small. It, thus, became obvious that a liquid layer of salt solution exists at the ice–substrate interface. Indeed, a liquid layer has been observed in adhesion tests at high salinities [30] and its existence even in the case of very low concentrations of potassium chloride has been confirmed [31]. The formation of this saline water layer is discussed below. Ice freezes at its equilibrium freezing temperature, so that it must cool in the form of ice to the temperature of the environment. During cooling, water freezes on the walls of the brine pockets [20]. This raises the salt concentration of the brine pockets allowing the phase equilibrium between the ice and the brine to be maintained. The ice that forms within the brine pockets occupies 9% greater volume than the liquid brine. There is no evidence that this would result in significant microfracturing of ice. Therefore, upon cooling, some brine is evidently forced out of the brine pockets and must eventually be expelled out of the ice sample. As a result of this brine expulsion, a layer of high salt concentration forms on the ice surface. When ice is adhered to a structure and cooled, a concentrated salt layer forms at the ice/substrate interface too, and reduces the adhesion strength. Details of the brine expulsion process during cooling of ice are poorly known, but the existing data indicate that the brine moves along the grain boundaries [32]. It is, therefore, likely that brine expulsion from an ice sample is not always three-

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dimensional as the orientation of grain boundaries depends on the crystal structure of the ice. It is also possible that, in contrast to long-term brine drainage of sea ice, the preferred direction of brine expulsion from an ice sheet is upwards, since the upward brine permeability of sea ice is higher than the downward permeability [33]. Whatever the factors that affect the preferred direction of brine expulsion are, it is apparent that gravity plays no significant role in the process. This is because the forces related to expansion of the water in the brine pockets are much higher than those related to gravity. Following [34], let us consider a simple case of a cube of ice adhered to a plate, in order to estimate the amount of salt expelled from the ice to the ice/structure interface. We will assume that the ice cube consists of polycrystalline isotropic material with random crystal orientation and that the plate adhered to one wall of the cube does not affect the brine expulsion process. With these assumptions the amount of salt, ms , expelled to the surface of an ice sample during cooling from temperature T1 to temperature T2 is ms (T2 ) = Mi (Si (Ti ) − Si (T2 )),

(7)

where Si is the salinity and Mi the mass of the ice sample. Equation (7) can be presented as   Si (T2 ) ms (T2 ) = Mi Si (T1 ) − Si (T1 ) . (8) Si (T1 ) Cox and Weeks [35] derived an equation for the salinity ratio Si (T2 )/Si (T1 ). When using Zubov’s [36] relationship between brine salinity and brine density (ρb in g/cm3 ), ρb = 1 + 0.8Sb , the equation by Cox and Weeks [35] reads     Si (T2 ) Sb (T2 ) (1−1/ρpi ) 1 + 0.8Sb (T2 ) 0.8 = exp (Sb (T1 ) − Sb (T2 )) . Si (T1 ) Sb (T1 ) 1 + 0.8Sb (T1 ) ρpi

(9)

(10)

Here ρpi is the density of pure ice in g/cm3 . The salinity of brine, Sb , is defined as the ratio of the mass of salt, ms , to the mass of brine, mb , so that mb (T2 ) =

ms (T2 ) . Sb (T2 )

(11)

When this amount of brine mb (T2 ) forms a liquid layer on the ice surface, its thickness h will be h=

1 mb , ρb Ai

(12)

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where Ai is the surface area. For a cube of ice with density of ρi and wall length of L the surface area, Ai , is 6L2 and the mass, Mi , is ρi L3 . It, therefore, follows from equations (10), (11) and (12) that for a cube of ice with a density of 0.9 g/cm3   Si (T2 ) 0.9L3 Si (T1 ) . (13) 1− h(T2 ) = Si (T1 ) 6L2 (1 + 0.8Sb (T2 ))Sb (T2 ) Inserting equation (10) into equation (13) and using the value of 0.917 g/cm3 for ρpi gives an equation for the thickness of the liquid film on the ice surface at T2 : h(T2 ) =

0.9LSi (T1 ) 6(1 + 0.8Sb (T2 )Sb (T2 ))    Sb (T2 ) −0.0905 1 + 0.85Sb (T2 ) × 1− Sb (T1 ) 1 + 0.85Sb (T1 )    × exp 0.872(Sb (T1 ) − Sb (T2 )) .

(14)

The salinity Sb of brine in ice is an explicit function of temperature which can be calculated [37] by Sb (T ) = −3.9921 − 22.700T − 1.0015T 2 − 0.019956T 3 .

(15)

When the wall length L and salinity Si (T1 ) of the original ice are known, equations (14) and (15) give the thickness, h, of the liquid film expelled when the ice is cooled from T1 to T2 . When we consider h as it affects ice adhesion at a temperature T2 , we are interested in total brine expulsion that has taken place before the temperature T2 has been reached. Therefore, the temperature T1 that we need to know to calculate Sb (T1 ) is the temperature from which cooling and brine expulsion have started, i.e., the temperature at the ice/water interface during freezing, Tf . The temperature Tf is basically unknown. This is because salt rejection from ice during freezing increases the salinity at the ice/water interface. The salinity and temperature of the interface layer depend on the growth rate of ice and effectiveness of mixing in the water [19, 38]. These factors are generally unknown. However, the problem can be solved by determining Sb (Tf ) from the definition of the interfacial distribution coefficient k ∗ : Si k∗ = . (16) Sb The experimental data [35, 38, 39] show that at typical growth rates of natural sea ice k ∗ is approximately a constant at k ∗ = 0.26 regardless of the growth conditions and water salinity. It has been pointed out by Makkonen [19] that the value of k ∗ = 0.26 is a reasonable approximation also in the case of icing due to sea spray droplets. A later analytical solution shows a value of 0.30 in close agreement [40]. Consequently Sb (Tf ) can be approximated by Sb (Tf ) =

S1 (T1 ) 0.26

(17)

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and the final equation for the thickness of the liquid film on the ice surface at a temperature T is 0.15 (1 + 0.8Sb (T ))Sb (T )    0.26Sb (T ) −0.0905 1 + 0.8Sb (T ) × 1− Si 1 + 3.077Si  × exp(3.354Si − 0.872Sb (T )) .

h(T ) = LSi

(18)

Here Si is the initial salinity of the ice cube and Sb (T ) is solved from equation (15). An iterative solution for h(T ) as a function of the ultimate salinity could also be obtained but this would not be justified because equation (17) is only an approximation. For practical purposes Si can be replaced by the measured ice salinity in equation (18), as the difference between Si (T1 ) and Si (T2 ) is always small and equation (18) is insensitive to these small differences. Examples of the brine layer thickness h calculated at various temperatures and ice salinities are given in Fig. 6 for an ice cube with L = 10 cm. According to Fig. 6, the brine layer thickness starts to grow at a certain temperature (e.g., −4.2°C for

Figure 6. Theoretical brine layer thickness on the surface of a three-dimensionally cooled isotropic 10−3 m3 ice cube as a function of its eventual temperature. The curves are for different ice salinities.

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Si = 20h). This is because there is no ice formation at temperatures higher than this. Equation (17) determines the brine salinity, Sb (Tf ), and, together with equation (15), gives an equilibrium freezing temperature Tf , which is an explicit function of Si . At temperatures lower than the equilibrium freezing temperature the brine layer thickness grows rapidly with decreasing temperature, as liquid water within the ice matrix freezes and brine is expelled from the ice cube. On the other hand, when temperatures drops, the brine layer formed on the surface partly freezes, and this reduces its thickness. Therefore, h reaches a maximum at a certain temperature (e.g., −6°C for Si = 10h in Fig. 6) and then starts to decrease. It is interesting to note that according to the results in Fig. 6 the maximum brine layer thickness is almost independent of ice salinity. As discussed earlier in this paper, the adhesion strength is probably in some way related to the calculated liquid layer thickness, which in the case of saline ice would be the brine layer thickness h. What do the results of equation (18) in Fig. 6 mean in terms of ice adhesion? First one may note that the values of h calculated by equation (18) are typically of the order of 102 µm, whereas the liquid-like layer thickness of fresh water ice is estimated to be of the order of 10−2 µm only [41]. This shows theoretically that the adhesion strength of saline ice is much smaller than that of fresh water ice down to the nucleation temperature of the salt. What actually happens to the ice–substrate bond when brine is expelled to the interface is unknown. It is also uncertain whether a continuous liquid film forms at the interface. Instead, a layer of mixed ice and brine may exist at the interface, allowing some ice/surface contacts to remain. Brine absorption by the structure surface because of its pores and roughness elements may also affect the phenomenon, although one should note that the calculated values of h are much higher than the height of the roughness elements on typical plastic, metal and painted surfaces. Finally, it is possible that some brine is removed from the surface by liquid flow along the ice/substrate interface. The equal theoretical maximum values of h at different salinities in Fig. 6 indicate that, if there is weak adhesion at some ice salinity, an equally weak adhesion will be found for any ice salinity at some temperature. This finding could be useful in mechanical de-icing, for example. When, say, a ship’s superstructure has been covered by saline spray ice and the temperature is low, one may minimize the effort of ice removal by waiting for the ice to cool to the optimum temperature predicted by Fig. 6. The results in Fig. 6 suggest that, at a given temperature, ice adhesion should depend on the salinity. This has been observed in experimental studies at low salinity [30, 42–46] but not in the range of Sw = 5–60h [30, 46], i.e., the reduction of the adhesion of saline ice from the value of fresh water ice occurs already at a very small water salinity. This suggests that ice adhesion may be insensitive to variations in the brine layer thickness when h is large. However, there may be a critical h, at which the continuous liquid film disappears and the adhesion strength considerably

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increases. If such a critical value exists, it probably depends on the roughness of the substrate surface. When h is larger than such a critical value, one should expect saline ice adhesion to be almost independent of temperature. The theory presented above provides a framework for further studies on the adhesion of saline ice. The major problem which should be studied further is the basic relationship between the liquid layer thickness and the ice adhesion strength. Before this relationship is determined we can only say that the thicker the brine layer, the lower the ice adhesion strength. 3. Adhesion Testing and Experimental Relationships The theory of ice adhesion outlined in Section 2 above underlines that there are many complex factors involved. Thus, predicting ice adhesion for the design of countermeasures is also complex. Moreover, all aspects of the theory have not, so far, been quantified sufficiently to allow purely theoretical estimates of ice adhesion. Consequently, laboratory and field experiments are still essential in developing and testing countermeasures for ice adhesion [13, 47–53]. The quantitative values of the measured ice adhesion strength vary from almost zero to about 1 MPa for the different coatings and test conditions. It is noteworthy that the highest values in this range are approximately the same as the shear strength of ice. Indeed, in such tests, traces of ice remaining at the substrate surface after the test are observed. This shows that the separation occurred at least partly within the ice material as a cohesional failure, rather than at the ice–substrate interface. Tests in the cohesion mode are, of course, useless in evaluating the true adhesion strength of ice with a material as they only measure the strength of the ice material. There are considerable difficulties in making comparative adhesion strength measurements even in laboratory conditions because so many factors affect the results. In addition to such factors as thermal expansion of ice, the variables which affect adhesion tests on all materials must be considered. These include the loading arrangement, i.e., the stress distribution at the interface during a test, and the loading rate. Moreover, in the case of ice, the environmental conditions during and after ice accretion as well as the time of contact affect the results. The rate at which the samples are cooled to the test temperature is important in affecting the rigidity of the interface due to thermal expansion, as discussed in Section 2.4.2. Furthermore, there is a stochastic effect caused by the brittle nature of ice fracture at low temperatures. The ultimate failure may be related to randomly spaced dislocations at which the initial fracturing process at the interface may start. This may also cause a sample size effect in ice adhesion tests. In case of saline ice, the size effect is unavoidable, as shown in Section 2.5. To obtain comparative results it is, therefore, important to remove as many variables from the test procedure as possible. Several test apparatuses for ice adhesion tests have been developed with this in mind in various research institutes [30, 47,

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54–71]. The principles of different methods have been reviewed by Sayward [72]. Many techniques exist and the size of the ice samples varies from nano-size [66, 67] to microscopic [68] to those faced in large scale applications. Test methods that are specific to some geometry, such as the leading edge of an airfoil, have also been developed [73]. Nano-scale experiments on ice adhesion are appealing in that many of the variables that affect macro-scale ice failure are absent. However, the true contact area is very difficult to determine in such experiments, and they do not necessarily relate to the practical situations of ice removal. Therefore, macro-scale ice adhesion measurement methods will be discussed in the following. As an example, the U.S. Army Cold Regions Science and Engineering Laboratory (CRREL) test arrangement is shown in Fig. 7. As another example, the ice adhesion test method presently used at VTT Technical Research Centre of Finland is described in more detail below.

Figure 7. CRREL cylindrical ice adhesion test arrangement (www.crrel.usace.army.mil). The joint between the coated metal pile and ice is broken by torque.

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Figure 8. Ice adhesion test apparatus of VTT and coating samples with ice cylinders frozen on them. Removed ice pieces are also shown.

In the VTT test the test specimens are 100 mm × 100 mm × 10 mm aluminium plates on which the coatings have been applied. Ice cylinders with a diameter of 30 mm are frozen on the coating at −10°C by placing a small container filled with water on the test plate. During freezing the samples are thermally insulated on the side on which the ice sample is frozen on the plate. Consequently, the latent heat of freezing is mainly removed by conduction through the aluminium plate. The purpose of this is to make the freezing front advance from the coating interface towards the bulk water. This assists in obtaining ice with little or no air bubbles close to the coating interface. The ice cylinders are tested by shearing them from the plates after at least 24 h of storage in a cold room kept at the test temperature. The test device and some samples are shown in Fig. 8. In the tests, failure at the interface is obtained by a shear force applied by a belt moving at a constant nominal rate of 3.2 × 10−4 m/s, which is also measured by a laser device. The resulting force is measured by a load cell and the force–time curve is recorded. Examples of force–time curves are shown in Fig. 9. Due to stretching of the belt, used in delivering the force to the ice cylinder, the increase of the force is somewhat nonlinear with time. The adhesion strength is determined as the peak shear force divided by the interface area. For each coating several ice samples are tested and the adhesion strength is calculated as the mean of the results.

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Figure 9. Example of four measurements on the same coating sample in VTT experiments. The adhesion strength in terms of force is based on the mean of the four peak values. Av. force = 441 N, pressure = 0.49 MPa.

One may evaluate the work done in an ice adhesion test of Fig. 9 by integrating the force–time curve with respect to the displacement, as the latter is known from the constant deflection rate applied. Such a calculation gives about 40 J/m2 for the total work, which can be compared to typical values of about 1 J/m2 for the fracture energy and about 0.1 J/m2 for the thermodynamic work of adhesion (Section 2.2). Clearly, the major part of the work spent goes for the irreversible deformations in the measurement system. This outlines that ice adhesion measurements of this type provide the adhesion strength only and are inapplicable to determining a relevant value for the work of adhesion. A fundamental aspect in measuring ice adhesion is the stress distribution at the ice–substrate interface upon loading. Ideally, the shear stress should be evenly distributed on the interface. This situation can be approached by measuring tensile strength instead of shear strength [57, 61], by a centrifuge adhesion test [63] and by a test method utilizing laser-pulse induced spallation in which a compressive stress pulse travels through a substrate disk which has an ice layer grown on its front surface [69]. However, one needs to consider not only the easiness of the theoretical interpretation of the ice adhesion measurements, but also their applicability: In most applications ice removal involves an uneven distribution of stress at the interface. In some applications ice removal in practice is done from curved surfaces and, accordingly, such laboratory experiments have also been done [73]. In any case, it is essential to understand how the stress in distributed in different test arrangements. The stress distribution in the basic shear ice adhesion test by VTT, as calculated by the finite-element method, is shown in Fig. 10.

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Figure 10. Shear stress (in Pa) in the VTT test (Fig. 8) as calculated by FEM analysis. The substrate is at the bottom of the ice sample and the pull is from right to left. The indentation is the ice deformation by the belt (dimensions exaggerated by a factor of 103 ).

In the VTT test the relative standard deviation of the measured ice adhesion strength is on average 12% showing that the method provides rather repeatable results on ice adhesion measurements. A large set of ice adhesion tests on various materials at different temperatures were made at VTT in the 1980’s using a similar test procedure but on a larger scale [30, 49]. In those tests the relative standard deviation was more than in the new method suggesting that it is better to use small samples in the testing. Some results of the previous tests at VTT are shown in Figs 11 and 12. The results in Fig. 11 demonstrate the theoretical predictions (Section 2.4) that the adhesion strength has a maximum at some intermediate temperature. ‘Inerta 400’ which was manufactured at the time by Teknos Oy is an exception, because it was the only coating which remained elastic at all test temperatures, thus preventing the brittle fracturing of the ice due to thermal contraction when cooling down to low temperatures (see Section 2.4.2). The results in Fig. 11 and in many other studies referenced in this paper confirm the theoretical prediction that the materials with a large water contact angle and the usually related low surface energy tend to have low adhesion strength with ice. In particular, equation (4) has been verified in that it describes not only the adhesion energy, but also the adhesion strength on polymers [71]. Experiments on nanoscale [68] have also confirmed this prediction. There are some experimental results in which this relationship has been less clear [70], but this is probably due to other complicating factors causing scatter in the measurement results. It is noteworthy that polymer surfaces have a high linear thermal expansion coefficient: For example the value for poly(vinylchloride) is about the same as for ice. Following the ideas proposed in Section 2.4.2, this should produce only a small temperature dependence at low temperatures. This is indeed observed experimen-

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Figure 11. Effect of temperature on the ice adhesion strength in shear. Inerta 400 manufactured by Teknos Oy was a coating that remained elastic at cold temperatures. Vellox 140 was an anti-icing coating supplied by Clifford W. Estes Co. (USA).

tally (Fig. 11) suggesting that the ice adhesion strength of polymers is affected not only by the surface energy but also the thermal expansion coefficient. Figure 12 shows the temperature dependence of the ice adhesion strength for one substrate material in more detail and also in the case of ice frozen from saline water. It can be seen that adhesion strength in the case of saline water does not have a maximum at an intermediate temperature but increases indefinitely. This is in accordance with the theoretical predictions discussed in Section 2. First, the salt causes a liquid film that makes the adhesion strength fall to a small fraction of its value for pure water. Second, this film makes the interface ductile enough, so that thermal expansion when going to low temperatures does not cause a reduction in the adhesion strength. That the adhesion strength of saline ice exceeds that of pure ice at very low tem-

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Figure 12. Adhesion strength of ice as a function of temperature for ice formed from fresh water and water with a 2h NaCl concentration. The substrate material Inerta 160 is an epoxy-based paint supplied by Teknos Oy.

peratures is because salt eventually nucleates, so that the liquid film completely disappears at the lowest temperatures. Adhesion strength of saline ice is typically only 10–20 kPa, except at extremely low salt concentrations and low temperatures as shown in Fig. 12. However, sometimes an order of magnitude higher values have been reported [45, 74]. The reason for this discrepancy is probably related to the formation and disappearance of the liquid layer at the interface. Different methods of freezing the samples, different test arrangements and varying storage times may all affect the thickness of the liquid film. Preferred direction of brine expulsion may also depend on crystal orientation and brine drainage may remove the liquid film. Furthermore, brine drainage may affect the adhesion strength more on horizontal than on vertical surfaces [44]. The adhesion strength drops to only a fraction of its value for fresh water ice already at ice salinities below 1h. Also, it has been shown [54, 66, 75] that the adhesion strength suddenly drops to a very small value when the temperature rises above the eutectic point of the salt solution, and that this happens also with very low salt concentrations. This is consistent with Fig. 12 in that the adhesion strength of saline ice was observed to be considerably higher at −30 to −50°C than at higher temperatures (the eutectic temperature of NaCl is −23°C). It was found by Makkonen and Lehmus [30] that the ratio of saline ice adhesion to fresh water ice adhesion was much higher on a rough polymer surface and on

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Figure 13. Field testing of ice adhesion on a navigation lock wall. Samples are made by a bore and then sheared off by a belt seen in the right hand upper corner.

a concrete surface when compared with other surfaces. It is noteworthy that these surfaces were much more porous than the other tested surfaces. This suggests that the liquid film is absorbed by a porous surface, thereby considerably increasing the adhesion strength. Therefore, roughness of the surface plays a much more important role in adhesion of saline ice than of fresh water ice. Measurements of ice adhesion strength have been made not only in the laboratory, but also in the field for naturally formed ice. This has been necessary for investigating the feasibility of coatings and de-icing systems in practical applications. Measurements with different coating materials have been made, for example, onboard ships and offshore structures [76–78] and on navigation lock walls [79, 80] (Fig. 13).

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4. Combatting Ice Adhesion 4.1. Anti-icing Coatings When developing synthetic coatings having low adhesion to ice, researchers have looked for relationships between adhesion strength and other material properties that are easy to measure. In accordance with the theoretical predictions discussed in Section 2, it has been observed experimentally that low adhesion to ice is generally connected with low permeability and absorption capacity and high hydrophobicity, i.e., a high contact angle with water [58–60, 81–84]. Materials that meet these requirements, such as synthetic polymer surfaces, have proved to be effective in reducing ice adhesion. The adhesion strength with polymer coatings is an order of magnitude less than on uncoated surfaces. In full-scale tests on a ship the best coatings were Teflon-4, organosilicone epoxy ‘G’, and a vinyl polymer sheet with perfluorinated film [43]. Laboratory tests with ice accreted in a wind tunnel [13] showed that the durability and effectiveness of silicone rubber made it a promising surface coating even in the case of ice accreted by droplets. Reports of the effect of a poly-bisphenol-A-polycarbonate block copolymer have also been encouraging [39, 48]. Crouch and Hartley [51] found that a mixture of watersoluble vinyl pyrrolidone resin and poly(ethylene glycol) in a room temperature curing silicone gave lowest adhesion in impact tests. A new silicone-based coating R-2180 supplied by NuSil Technology (USA) has a significantly lower adhesion strength with ice than Teflon [85, 86]. There are many commercial coatings especially designed for low ice adhesion available today, see, e.g., www.crrel.usace.army.mil and [63, 85, 87–89]. The most effective commercial materials are grease-type coatings, which, of course, have limitations in applications due to their low durability. In regard to their usefulness, a relevant question is: How low should the ice adhesion strength of an effective coating be? The best hydrophobic rigid materials existing today provide ice adhesion strength of the order of 100 kPa [49, 63, 85]. Let us consider a vertical coated plate and an ice layer with a thickness L and surface area A frozen on it. On such an ice cover gravity will impose a force gρAL. This divided by the surface area gives the critical adhesion strength Pc for the ice cover to be released by gravity alone as Pc = gρL.

(19)

Supposing in equation (19) an ice density ρ of 0.9 g/cm3 and thickness L of ice, say, 5 cm, gives the critical adhesion strength Pc of 0.44 kPa, which is more than two orders of magnitude smaller than that provided by typical polymer surfaces. In other words, only an ice layer that is more than 10 m thick will be released spontaneously due to gravity for such a surface. In many applications other forces, such as centrifugal force, winds shear and vibrations contribute to the forces at the ice–substrate interface. Nevertheless, it is clear from this simple calculation that none of the hydrophobic solid materials existing today provides an adhesion

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strength low enough to solve the practical problems of ice adhesion. Furthermore, the effectiveness of superhydrophobic nano-rough surfaces is hampered by impact of small cloud droplets and formation of frost within the texture [90–92]. Therefore, new ideas for anti-icing coating are required and have indeed been recently introduced as commercial products. As it is difficult to remove ice when it has frozen on a substrate, an obvious solution would seem to be preventing the freezing of water on a substrate surface. An obvious approach is to make the surface so hydrophobic that water drops will not remain on the surface for a sufficient time for nucleation to occur, see [9]. The other approaches are more fundamental in that they attempt to reduce the adhesion strength of ice by surfaces that prevent nucleation or freezing even in long-term contact. It is well known that the structure of water is modified in contact with a surface and that lowering of the freezing temperature of water occurs at some surfaces, for example to the benefit of some biological species [93–95]. However, theoretically, the matter is less than clear [96, 97] and the development of such new surfaces has been done mainly by experimental searching. One idea worth developing is based on extremely low ice adhesion (less than 10 kPa) measured for an organopolysiloxane resin mixed with alkali metal compound manufactured in Japan in the 1980’ [32, 65, 98]. This coating material was manufactured by Kansai Paint Inc., but is no more on the production line. Its mechanical durability was rather low but it appeared not to be self-sacrificing. This hydrophobic organopolysiloxane resin contained a small amount of polar ingredients. Lithium ions act as hydrogen bond breakers in the material. It was assumed in [65] that the four oxygen atoms from water molecules and two oxygen atoms present in a carboxylic residue associate together and form an octahedron coordination structure. DSC analysis showed that this water does not have a definite freezing point [65]. It is assumed by the developers of this coating that “the structural and energetic differences thereof and synergetic effect from silicone matrix result in preventing ice adhesion”. A second idea pursued recently at VTT and elsewhere [95, 99] is to synthesize natural anti-freeze proteins, apply them in coatings, and rely on their ability to reduce the freezing point and thus affect ice adhesion. Preliminary experiments on prototype coatings have, however, shown that once freezing occurs on these materials, the adhesion strength is high. It appears, therefore, unlikely that anti-freeze proteins will provide solutions to practical problem of ice adhesion generally. However, they may be useful in systems that have a very small size and an undisturbed environment. They may also have an indirect effect in reducing icing by delaying the freezing of water. Yet another approach is the ePaint coating which reduces the adhesion strength of ice using several processes [86]. This coating is hydrophobic but also includes phase-change material that is thermally activated. As the coating cools below 0°C the epoxy-like material contracts, and the embedded solid-phase change material expands, causing little net change in the surface area of the coating. However, as ice

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accretes, liberated latent heat from the ice warms the coating surface. This causes the phase-change material to warm and to expand. The simultaneous contraction of the epoxy-like material and expansion of the phase-change material causes shear stress within the coating and failure of the ice–substrate adhesion bond (see Section 2.4.2). 4.2. Mechanical Methods As discussed above, the use of the present low-adhesion coatings is not very effective when used alone. However, combined with mechanical and thermal methods, coatings and paints can make deicing easier. For example, the time needed for deicing a surface by heating is reduced considerably when the surface is coated with a suitable material [100, 101]. Also, the ease of manual and automatic mechanical ice removal is very sensitive to the strength of adhesion of the ice to the material. As an example, polymer coatings have proved to be useful in routine service when deicing, e.g., navigation lock walls [79]. Manual deicing has historically been the only method in combating ice, for example at sea. The ice is removed by crew members using mallets, axes, baseball bats, etc. (see Fig. 1). This method is unsatisfactory because conditions on a slippery deck during severe icing are hazardous, and de-icing actions are almost impossible when most needed. The use of motorized cutters for removing ice on a ship [102] is usually possible only after an icing storm. It is seldom possible at any time to manually remove ice from the upper parts of the structures most critical to ship’s stability or from other types of tall structures on ground. Because of these problems in manual ice removal, automatic and semiautomatic deicing methods have been developed. The most effective of these devices is a pneumatic deicer, a series of tubes standing alone or built into a rubber mat called a boot. When the boot is inflated with air it expands and breaks the ice adhered to the surface. The principle is in use for protecting aircraft wings from icing [103, 104]. Pneumatic deicers have proven to be effective also on small cylindrical objects and large flat surfaces, such as radar dishes and ship superstructure [105–107]. Disadvantages of this method are the cost and likelihood of damage to the deicers if used in working areas. More recent developments are the electro-expulsive (EESS) and electro-impulsive (EEIS) separation systems tested in, e.g., shipboard applications and aircraft [81, 98, 99]. These methods utilize high repulsive forces to impart expulsive movements to the flexible outer layer. The forces are generated by overlapping conductive ribbons that receive a very high instantaneous current pulse. A lightweight retrofitable EESS system consists of a 0.5 mm thick polyurethane blanket with embedded flexible conducting copper ribbons which are paired and separated by a dielectric. The instantaneous power pulse is very high, but the pulse width is only about 20 µs. Thus the energy consumption of the EESS is quite low. Test results and shipboard applications of the EESS system are described by Embry et al. [109].

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Economical mechanical deicing can be achieved by using a flexible coating that moves due to wind drag [78, 110, 111]. The method is not quite satisfactory because accreting ice makes the coating partly inflexible. To avoid this Alexeiev [112] covered small-scale models of meteorological masts with flexible coatings, fixing the coatings with the aid of guy ropes. Vibration of the mast and the guy ropes then keeps the coating moving sufficiently for ice removal. Also, a large number of small, conical plastic shelters that move in the wind have been tested with some success in small-scale tests. Flexible surface-coating materials whose deicing effect is not based on low adhesion strength but on the ease of ice removal on impact have also been tested on ships, see, e.g., [102]. The surfaces tested were plastic foam mats, either alone or covered with a sheet of neoprene rubber. This type of coating is not very successful, although it makes the ice somewhat easier to remove manually. However, ice has spontaneously detached from a rubber mat on the outer side of a ship’s bulwark. High costs, relatively poor durability and problems in attaching the rubber mats firmly are the main reasons for the limited use of the flexible coatings. Mechanical deicer based on metal plates and wires moving by means of electromagnetic induction, activated by discharge from a condenser through a solenoid situated near the surface of the plates, has been used in protecting meteorological instruments [110–112] and various other structures [100]. Robotics has also been developed and demonstrated, e.g., in mechanical de-icing of overhead power line cables [101, 113]. A piezoelectric de-icing system has also been developed [86]. 4.3. Thermal Methods The most obvious method for preventing icing is heating. However, it is far from being the most practical because of the large amount of latent heat required to melt ice or to prevent its formation. For example, the power required for anti-icing, when the icing rate is 30 kg/(m2 h), is about 2 kW/m2 . For example, anti-icing a cup anemometer in the field externally typically requires 300–700 W [114]. De-icing by heating typically requires about 300 kJ/m2 of energy [112, 114]. Large energy requirements of thermal methods have restricted their use mostly to small objects, although infrared heaters have been used even in the case of an aircraft [81]. Internal heating is much more effective than external heating unless the ice layer is transparent and thin [111]. For this reason ultrasonic heating films have been developed [115, 116]. Another significant problem with heating as a deicing or anti-icing method is that the melted water will freeze at some other location or form icicles. Thus, water drainage must be designed carefully if this method is to be applied. One way of heating structures is to use a thermosiphon. This method can be applied to structures such as masts and handrails, the structures themselves acting as heat pipes. It is also possible to cover flat surfaces with loops of heat pipes. These techniques have been applied when constructing a practical thermostatically controlled thermal deicing system on a ship [117], which can use different heat

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sources. The heat consumption of the system has been found to be about 1 kW/m2 in field tests under typical icing conditions. The heat pipe method has also been used in deicing navigation buoys [118]. Hot water is rather effective in short-term ice prevention and is sometimes used as a deicing method on ships [119]. A disadvantage of this method, in addition to high heat consumption, is that water for protecting the upper parts of the structures falls and flows along the surfaces and may increase icing at lower levels of the vessel. Thermal and mechanical methods are combined in the so-called seawater lance, which consists of a high-pressure jet of seawater capable of removing ice by melting and dynamic pressure. This method has been used on navy vessels [120]. Current-conducting coatings are used for heating small surface areas, such as automobile and aircraft windows [86]. This method works for larger surfaces too, when combined with coatings that reduce ice adhesion. Such an approach reduces the power consumption to a level realistic for some practical applications, such as wind turbine blades. Detailed analysis of the heat consumption and optimized design of a wind turbine anti-icing system has been made by Makkonen et al. [116]. The high energy consumption of thermal de-icing is not only due to melting a sufficiently thick ice layer but due to high loss of heat as conduction to ice and the substrate while the system is on. This is especially so when heating is applied from the air and not from the substrate [116]. It is easy to show [117] that the conduction loss is smaller the quicker the process, i.e., the higher the power used in the de-icing. Based on this, methods have been developed in which a very high power is applied directly at the interface by a heated film [121, 122]. This approach is applicable to many problems, see http://engineering.dartmouth.edu/thayer/research/ice-engg. html. In some applications, even a very small amount of rough ice may cause serious problems. These are related to aircraft wings, helicopter rotors and wind turbines where aerodynamic penalties due to accreted ice are unacceptable. In these applications de-icing may not be enough to prevent problems; instead, anti-icing is required. Such anti-icing systems are presently based on heating, e.g., [116]. 4.4. Chemical and Electrical Methods The application of chemicals on an icing surface to reduce ice adhesion has been tested, mostly with limited success [102]. Moreover, these chemicals deteriorate easily by weathering and by the accreting ice [123]. Another kind of chemical measure for preventing ice adhesion is to apply freezing point depressants on surfaces. The major problems of this method are optimizing the amount of the chemicals and distributing them uniformly on the surface. When these problems can be overcome, it is possible to reduce icing considerably by using organic anti-icing fluids, e.g., ethylene glycol [103, 124] or urea [125]. Salts such as calcium nitrate [126] have also been tested and, of course, various anti-freeze liquids are widely used in removing ice from road surfaces and aircraft. A detailed review on anti-icing chemicals is available in [86].

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Chemical methods can be appropriate for protecting small objects (windshields, bridge windows, automatic meteorological instruments, etc.) and objects that require ice removal for only a short time (aircraft before take-off, for example). Chemicals are difficult to apply on large structures for long-term ice prevention since their rapid deterioration makes the method uncertain and expensive. Other disadvantages include making horizontal surfaces slippery and contaminating the environment. A coating that ejects the encapsulated anti-icing chemical onto the surface would be one possible solution [127]. Phan and Laforte [128] showed that the adhesion strength of ice accreted on wires from cloud droplets depends on the applied DC negative electric field. The effect of applying an electric field when freezing bulk water has also been studied [129], resulting in the conclusion that the observed small effect is due to the effects of electrolysis and corrosion on metal surfaces. In some applications electrical phenomena might be of benefit in anti-icing [130, 131], but no practical tests have been made. An electrostatic model has been proposed for the basic atomic attraction in ice adhesion [130] and this theory may provide ingredients for further development [132, 133]. 5. Summary The review and theoretical analysis presented above points out the complexity of ice adhesion. Many mechanisms related to ice adhesion are still poorly understood and most practical problems caused by ice adhesion are far from being solved, see, e.g., [134, 135]. The adhesion strength of ice is related not only to the chemical composition, surface morphology, stiffness and thermal expansion coefficient of the substrate material, but critically depends on the temperature and the test arrangement as well. Further studies on these effects are necessary in order to obtain comprehensive understanding of ice adhesion. When salt is involved, as in marine icing and road maintenance, the adhesion strength of ice depends on even more factors. When the liquid film that forms at the ice–substrate interface is retained, then the adhesion strength of saline ice is very small at any salinity. Therefore, the concepts presented in this paper regarding the formation of the liquid film due to brine expulsion should be tested and developed further. For the same reason, the mechanisms of brine drainage and movement of brine along the ice/structure interface should be studied. Generally, the higher the water contact angle of the material the lower the ice adhesion strength. However, it seems that the routes to find practical solutions for many problems caused by ice adhesion using hydrophobic surfaces have been fully explored and will not result in further significant development. There are notable exceptions from this general rule suggesting that it may be more fruitful to investigate the materials that would prevent freezing on a surface, or at least make the true contact area smaller. The organosilicone material with lithium ions and antifreeze

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proteins synthesized from insects may provide such solutions. Further fundamental research on the freezing and properties of water at interfaces is required for development in this field. The work done in mechanically removing ice is typically orders of magnitude larger than the fracture energy and the thermodynamic work of ice adhesion. This suggests that there may be room for developing better mechanical de-icing methods by applying shockwaves or other very high energy pulses with a short duration. New technologies for manufacturing heated films for thermal anti-icing and deicing provide solutions in many applications. That these films now allow heating at a very high power makes the costs of combating ice more reasonable. However, energy consuming ice combating methods are expensive to use unless their operation is optimized by detecting the ice on the surfaces. Therefore, attention should be paid also to developing reliable ice detection methods. Acknowledgements Thank are due to Eila Lehmus, Pieti Marjavaara, Matti Halonen and Erkki Järvinen for assistance in the VTT ice adhesion projects and Kari Kolari for providing FEM simulations. This work was funded by the Academy of Finland and by the Nordic Council Top-level Research Initiative project TopNano, as well as the TekesRescoat project. References 1. L. Makkonen, J. Phys. Chem. B 101, 6196–6200 (1997). 2. V. F. Petrenko and R. W. Whitworth, Physics of Ice. Oxford University Press, Oxford, U.K. (1999). 3. Y. Wei, R. M. Adamson and J. P. Dempsey, J. Mater. Sci. 31, 943–947 (1996). 4. K. R. Jiang and L. S. Penn, J. Adhesion 32, 217–226 (1990). 5. K. L. Mittal, Polym. Eng. Sci. 17, 467–473 (1977). 6. M. F. Hassan, H. P. Lee and S. P. Lim, Measurement Sci. Technol. 21, 075701 (2010). 7. M. Zou, S. Beckford, R. Wei, C. Ellis, G. Hatton and M. A. Miller, Appl. Surface Sci. 257, 3786–3792 (2011). 8. S. A. Kulinich and M. Farzaneh, Appl. Surface Sci. 255, 8153–8157 (2009). 9. L. Mishchenko, B. Hatton, V. Bahadur, A. Taylor, T. Krupenkin and J. Alzenberg, ACSNano 4, 7699–7707 (2011). 10. F. Wang, C. Li, Y. Lv, F. Lv and Y. Du, Cold Regions Sci. Technol. 62, 29–33 (2010). 11. J. Bico, U. Thiele and D. Quere, Colloids Surfaces 206, 41–46 (2002). 12. S. A. Kulinich and M. Farzaneh, Langmuir 25, 8854–8856 (2009). 13. D. K. Sarkar and M. Farzaneh, J. Adhesion Sci. Technol. 23, 1215–1237 (2009). 14. J. R. Stallabrass, Canad. Aeronaut. Space J. 9, 199–204 (1963). 15. C.-L. Phan, P. McComber and A. Mansiaux, Trans. Can. Soc. Mech. Eng. 44, 204–208 (1978). 16. O. B. Naselle, L. Levi and F. Prodi, J. Glaciol. 33, 120–122 (1987). 17. J.-L. Laforte, C. L. Phan, B. Felin and R. Martin, Special Report 83-17, pp. 83–92, U.S. Army Cold Regions Research and Engineering Laboratory, CRREL (1983).

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