High Concentration, Coarse Particle, Hydraulic Conveying

*Lionel Pullum*

### **Abstract**

Conventional coarse particle hydraulic conveying is performed under turbulent flow conditions, usually at concentrations of less than 40% v/v. The last three or four decades have seen the development of much higher concentration conveying, with the successful transport of suspensions of 70% v/v or more. These suspensions can be conveyed at very low velocities and generally exhibit very benign characteristics, having the capability of being stopped and restarted at will. There are generally two methods of pumping coarse materials safely at low velocities. The first method, using a Newtonian such as water, can be applied when the particle size distribution is sufficiently broad to minimize percolation and the concentration sufficiently high to prevent particle restructuring during transport. A second method uses a non-Newtonian, visco-plastic carrier fluid, normally fine particle slurry, to convey the coarser particles. This second method removes the constraints of the first method, allowing a greater range of coarse solid distributions and concentrations to be pumped. In both cases, the conveying characteristics appear similar to laminar flow. Both methods are described and analyzed in this chapter.

**Keywords:** hydraulic transport, coarse particle, high concentration, no-Newtonian, hybrid suspension

### **1. Introduction**

Except for the transport of pastes and non-Newtonian slurries, most hydraulic transport applications operate under turbulent conditions at low to moderate volumetric concentrations. This contrasts with pneumatic transport systems which operate in both dilute, or lean, phase conveying and dense phase conveying. Dilute phase is analogous to this form of hydraulic conveying, while dense phase conveys solids at much higher solids' concentrations and at lower velocities. Over the last three or four decades, advances have been made in the development of higher concentration hydraulic transport systems, which share many characteristics in common with dense phase pneumatic systems.

Conventional hydraulic transport systems are turbulent, relying on the action of turbulent eddies to suspend and convey the solids in the pipe. At high velocities, the solids'suspension is almost uniform across the pipe, while at lower velocities, a pronounced concentration profile appears, forming sliding and even stationary beds of solids on the bottom of the pipe. Such flows are unstable and solids' transport will stop when the mean velocity falls below a minimum conveying velocity, and may even block the pipe. To minimize this risk of blockage and assist in restarting, a rule of thumb is used whereby the in-line solids' concentration, *Cvi*, is generally limited to less than 40% v/v. The minimum conveying velocity is a function of the solids' concentration, size and density, and can typically range from 1 to 2 m/s for fine particles (d < 1 mm say) to 7–10 m/s for coarse particles (d > 5 mm say).

Despite these unstable characteristics, such conveying has many advantages in the main areas where it is practiced, namely inter-process transfer, dredging, tailings disposal and long distant transport of ore or concentrates. It is simple to operate, requiring only that those solids are mixed with water in a suitably agitated vessel of some form, and pumped, normally using centrifugal pumps, through the pipeline to their destination. The low concentration also affords easy separation of the solids from the conveying fluid.

By contrast, outside of the food and some process industries, high concentration conveying is rarely seen, concrete pumping being the most obvious exception. However, high concentration systems have been developed and are finding increasing use in tailings disposal, hoisting and other niche applications. Their attractiveness lies in their benign operation, as they generally do not require a minimum conveying velocity, so lines can be stopped and restarted at will. Their low velocity, typically of order 1–2 m/s, several meters per second lower than their coarse particle low concentration counterparts, results in low particle attrition and pipe wear. This mode of conveying also allows very coarse particles to be conveyed at low particle pipe diameter ratios, e.g., *D*/*d* ≤ 5. The high concentration, 40<*Cv* ≤75% v/v, also produces low specific energy consumption (SEC), producing low operating costs (OPEX) and low greenhouse gas emissions. These advantages, however, are offset against a more complex operating system and the need for more specialized equipment (CAPEX).

Two forms of high concentration conveying have been developed, one using Newtonian fluids, like water, as the carrier fluid to convey the solids, and a second using non-Newtonian carrier fluids. Although sharing many characteristics, these methods will be described separately starting with the Newtonian system.

## **2. Newtonian high concentration conveying**

#### **2.1 Concentration considerations**

While weight concentration, *Cw* w/w, or worse weight of solids per volume of liquid, *Cwv* w/v, are common in industry, it is the volumetric concentration, *Cv* v/v, that characterizes and controls the way a suspension of solids behaves. At high concentrations, another more useful way of describing the solids' concentration is the ratio of the volumetric concentration to the maximum packing concentration of the solids, or more usually the maximum dilated or loose poured concentration, i.e., *Cv*/*Cb* known as the reduced concentration. The term *Cb* is used here to denote this packing concentration, as this value corresponds to the in-situ concentration of a settled or sliding bed of solids within the pipe. For non-Newtonian systems, described in §3, the bed concentration will be lower than for Newtonian systems, and will be described in greater detail there. This value may be considered to indicate the pipelines obscuration, or particle's mobility.

*Cb* is a function of the particle's size distribution (PSD) and shape distribution (SSD), from which this value can be calculated (for example [1, 2]), or measured

using a procedure outlined by Leuenberger [3], as follows. A tank of initially dry solids is continuously mixed with a pitch-bladed agitator, and the power drawn by the agitator monitored while water is progressively added. An example of such a test is shown in **Figure 1**.

The typical response to the addition of water (**Figure 1b**) is to change the solids from a purely granular flow in air, to one where an increasing number of water bridges are formed. Shearing these as well as the solids requires increasingly more energy. Once all the interstices are filled with water, further addition of water simply separates the particles, reducing the mixer energy requirements. Thus, the value of 1- *Cvwater* corresponding to the peaks in the characteristic curves approximates *Cb*.

The calculated value of *Cb*, based on the coal's PSD and SSD, is approximately 89%, whereas the empirical value is around 86%. The discrepancy, although small, is because the empirical value is for a sheared, and hence a dilated bed. It is recommended that the empirical value is used, or that calculated values should be reduced by a couple of percentage points.

#### **2.2 Conveying characteristics**

As mentioned in the introduction, both dilute and high concentration conveying is commonplace in pneumatic conveying systems, and the operational or transport characteristics are displayed in what is known as a diagram of state (**Figure 2**).

At moderate to high velocities, the transport characteristics exhibit the familiar hook curves of conventional hydraulic conveying systems, moving from a fully suspended state at high velocities, to sliding beds as the velocity is decreased. After going through a minimum, the pressure gradient rises due to partial obscuration of the pipe and increased particle wall contact. For hydraulic systems, which normally use centrifugal pumps, this is an unstable region, but for pneumatic systems, which use a variety of different feeding systems, e.g., blow vessels, transport can still be obtained down to very low velocities. At these low velocities, the line is filled with solids that

#### **Figure 1.**

*(a) Broad PSD of typical run of mine (ROM) coal, (b) experimental value for two different agitator speeds and calculated (dashed line) determination of Cb for the same ROM.*

**Figure 2.** *Diagram of state for coarse particle pneumatic conveying.*

move along as a continuous plug, although this plug is often deliberately broken into smaller plugs to reduce the pressure requirements. At intermediary velocities, however, the flow is unstable, exhibiting dune flow that can block the pipe.

Wilson et al. [4] examined the behavior of dense phase hydraulic hoisting, subsequently publishing an analysis of the frictional component of such plug flow [5], and such techniques were employed in the Hansa underground mine to hoist -60 mm coal, 850 m to the surface [6]. To the author's knowledge, though, it was not until the invention of a feeding system called a Rotary Ram Slurry Pump (RRSP), described in §2.7, that dense phase horizontal coarse particle hydraulic conveying was considered at an industrial scale.

The RRSP enabled very high concentrations of coarse particles to be loaded into a pipe using a valving system that opened to the full bore of the conveying pipe. These particles could be conveyed at very low velocities with little, if any, inter-particle movement. This feature was ideal for friable materials such as coal. Transport characteristics for ROM coal conveyed in a 300 mm pipe using this technique are shown in **Figure 3**.

The transport characteristics do not exhibit the typical hooked curves of lower density conveying, although transport below the lowest velocities shown were problematic. Suspensions with concentrations below 65% were also unstable for this particular material. Readings from an isolated electrode probe (see e.g. [7]), inserted into the top of the pipe, registered periodic high activity at these lower concentrations, indicating possible dune formation. A two-layer model, at concentrations close to plug flow, was used to model this flow with satisfactory agreement as shown. However, the adoption of a normal two-layer model is found to substantially overpredict the transport pressure gradients shown, and modification to such a model is needed before the agreement shown here is obtained.

### **2.3 Two-layer modeling modifications for coarse particles**

Only a simple two-layer model (e.g. [8]) will be used in this section for illustrative purposes, and the reader is referred to the chapter on Pipeline Modeling for a more complete description and a more nuanced approach. The early layered models were developed to model the behavior of small sand particle slurries, which become fluidized during transport. This is evident by observing slurry issuing from holes in the upper invert of the conveying pipe, where the slurry behaves as a fluidized jet [9].

*High Concentration, Coarse Particle, Hydraulic Conveying DOI: http://dx.doi.org/10.5772/intechopen.107230*

#### **Figure 3.**

*Typical high concentration, coarse particle, transport characteristics. Computed curves of constant delivered concentrations and partial deposition locus using a two-layer model. The bed depth* Ybed *is also shown.*

#### **Figure 4.**

*(a) Schematic of pipe cross section giving the normal force distribution for fine and coarse particles, (b) wear pattern in a 300 mm pipe used for conveying -32 mm solids, showing unworn original oxide coating in pipe upper invert.*

Such behavior is not observed in coarse particle flow. This difference has ramifications for the modeling of the resisting force of the bed in a two-layer model. Consider the schematic of the pipe cross section shown in **Figure 4a**.

If the solids are fine and fluidized, the bed presents a hydrostatic pressure at the pipe wall (**Figure 4a**, left hand side), which is normal to the pipe wall. If the density of the fluid and solids are *ρf* and *ρs*, respectively, then the normal force of the bed on the pipe wall is given by

$$N = 2\text{Cb}\left(\rho\_s - \rho\_f\right) \text{g}R^2 \int\_0^\beta \cos\left(\beta\right) - \cos\left(\theta\right)d\theta \tag{1}$$

Where *β* is the half bed angle shown in **Figure 4a** and which on integrating and rearranging becomes

$$N = 2\text{Cb}\left(\rho\_s - \rho\_f\right) \text{g}\frac{D^2}{4} \left(\sin\left(\beta\right) - \beta\cos\left(\beta\right)\right) \tag{2}$$

Conversely, for unfluidized coarse particles the normal force is simply

$$N = \text{Cb}\left(\rho\_s - \rho\_f\right) \text{g}\left(\frac{\pi D^2}{4}\right) (\mathbf{1} - A) \tag{3}$$

where *A* is the fractional cross-sectional area above the bed and (1-*A*) is the area of the bed 1ð Þ¼ � *A* ð Þ *β* � sin ð Þ *β* cosð Þ *β =π* ¼ *Cv=Cb*.

The difference between Eqs. (2) and (3) is then a multiplier which accounts for the effect on the resisting force of particle fluidization. Thus, combining Eq. (2), Eq. (3), the bed fractional area and a degree of fluidization 0 < *f* < 1, where a value of zero = the unfluidized state, 1 = the fully fluidized state, this multiplier can be written as

$$F = \mathbf{1} + f \left[ 2 \frac{\sin \left( \beta \right) - \beta \cos \left( \beta \right)}{\beta - \sin \left( \beta \right) \cos \left( \beta \right)} - \mathbf{1} \right] \tag{4}$$

and used to describe the bed's normal force for an arbitrarily fluidized bed by

$$N = \text{Cb}\left(\rho\_s - \rho\_f\right) \text{g}\left(\frac{\pi D^2}{4}\right) (\mathbf{1} - A)F \tag{5}$$

#### **2.4 Comparison with conventional conveying**

Data reported by the Saskatchewan Research Council (SRC) for coarse coal in water [10] show typical results in a similarly sized pipe for -50 mm solids, *d*50 = 3 mm, *Cv* = 0.457 v/v, *D* = 234 mm, to be *V* = 2.5 m/s, d*P*/d*x* = 0.64 kPa/m. The conveying velocity is higher than that shown in **Figure 3**, and this, combined with the lower concentration means that particle attrition would be considerable. The pressure gradient, however, is comparable to that of the 70% curve in **Figure 3**. A comparison of the operating costs can be achieved by comparing the *SEC* given by

$$\text{LECC} = \frac{\text{suspension pumping power}}{\text{solds mass flow x distance}} = \frac{dp/\_{\text{dx}}}{3.6 \text{C} \nu \nu\_{\text{/} \text{f}}} \text{ kWh/} \text{(tome km)}\tag{6}$$

where the pressure gradient, d*p*/d*x*, is in kPa/m.

Comparing the SRC data with the higher pressure gradient 75% v/v data, the *SEC* for the conventionally conveyed coal is 0.27 kWhr/(tonne km), compared with 0.23 kWhr/(tonne km) for the high concentration conveying. Furthermore, the high concentration system consumes 45% less water. A more detailed comparison of high concentration and non-Newtonian systems, described in §3, with conventional conveying is given in [11].

#### **2.5 Restart capabilities and practical considerations**

#### *2.5.1 Restart*

As mentioned above, transport at lower concentrations proved problematic through dune formation partially blocking the pipe, resulting in unsteady flow. Experience gained with the RRSP suggested that an operational concentration range should be limited to 0.75 < *Cv*/*Cb* < 0.9, and that the materials required a broad size distribution such that *Cb* > 0.75 to avoid excessive bed percolation of the low viscosity water carrier.

When operating within these constraints, restart of fully loaded pipelines several kilometers long proved to be immediate once the pumps were activated. The explanation for this is not entirely clear but is expected to relate to the low deposition velocities. The deposition locus for the material in **Figure 5** expressed as a function of *Cv*/*Cb* is shown here (**Figure 6**).

At the operational limits for this material the corresponding deposition velocities are approximately 0.6 and 0.2 m/s. Such suspensions started with ease, and it is expected that once the pumps were started, increasing the flow to the normal conveying velocity of a 1 m/s was sufficiently rapid so that reorganization of any particles

#### **Figure 5.**

*Normal force fluidization modifier for* f *= 0, unfluidized coarse particles,* f *= 0.5, partially fluidized intermediary particles,* f *= 1 for fully fluidized fine particles.*

#### **Figure 6.**

*Deposition locus as a function of* Cv/Cb *for the data shown in Figure 3. The concentration limits for successful operation are shown shaded.*

above the bed to form a plugging dune did not occur. This is despite the high velocity fluid flowing over the beds, which occupy 75 to 90% of the pipes CSA.

### *2.5.2 Practical considerations*

Production of such suspensions is comparatively easy as the mobility of these suspensions is very high, requiring only gentle agitation in suitable mixing tanks to ensure homogenization. More difficult is maintaining the high concentrations within the required limits, requiring conveyors etc. to be sheltered from rain, and sufficient surge capacity to be installed for constant operation. This last point is less of a problem, as the ability to reduce the transport velocity to very low values or stopping, provides a very large turn down ratio. Maintaining the correct feed to the pump also requires some care, as fluid can be preferentially drawn from the tank diluting the flow into the pump's inlet ducting, producing unstable operation, details of which are described in [12].


*Article* **Multi-Modal, Non-Invasive Investigation of Modern Colorants on Three Early Modern Prints by Maria Sibylla Merian**

**Olivia Dill, Marc Vermeulen, Alicia McGeachy and Marc Walton \***

The Center for Scientific Studies in the Arts (NU-ACCESS), Northwestern University, Evanston, IL 60208, USA<EMAIL_ADDRESS>(O.D<EMAIL_ADDRESS>(M.V<EMAIL_ADDRESS>(A.M.)

**\*** Correspondence<EMAIL_ADDRESS>
**Abstract:** Northwestern University's Charles Deering McCormick Library of Special Collections owns three hand-colored copperplate engravings that once belonged to an edition of *Matamorphosis Insectorum Surinamensium* by artist-naturalist Maria Sibylla Merian (1647–1717). Because early modern prints are often colored by early modern readers, or modern collectors, it was initially unclear whether the coloring on these prints should be attributed to the print maker, to subsequent owners or collectors, or to an art dealer. Such ambiguities posed challenges for the interpretation of these prints by art historians. Therefore, the prints underwent multi-modal, non-invasive technical analysis to assess the date and material composition of the prints' coloring. The work combined several different non-invasive analytical techniques: hyperspectral imaging (HSI), macro X-ray fluorescence (MA-XRF) mapping, surface normal mapping with photometric stereo, visible light photography, and visual comparative art historical analysis. As a result, the prints and paper were attributed to a late eighteenth-century posthumous edition of Merian's work while the colorants were dated to the early twentieth century. This information enables more thorough contextualization of these prints in their use as teaching and research tools in the University collection.

**Keywords:** Maria Sibylla Merian; colored prints; hyperspectral imaging; X-ray fluorescence spectroscopy; photometric stereo; Prussian blue; non-invasive pigment characterization


**2. Short history of forensic genetics**

In 1880, British anthropologist Sir Francis Galton published the first studies on digital finger‐ printing as an identification method of a certain person.

Another important step in forensic identification was discovering the protein polymorphism of the AB0 blood groups by the Austrian doctor Landsteiner at the beginning of the twentieth century.

After year 1950, in the forensic serology, laboratories were tested a number of blood and tissue antigens, culminating with researching the major histocompatibility complex (HLA).

Sir Alec Jeffreys introduced for the first time the DNA fingerprinting in the field of forensic genetics, proving that some regions from the DNA contain repetitive sequences which are variable among individuals [1]. He was the first to prove the importance of using genetic fingerprinting in the case of forensic personal identification (crimes, filiation, consanguinity, sexual abuse, immigration).

Due to this discovery, the first case of forensic genetics could be solved using the DNA analysis [2]. After murdering of two girls in 1983 and 1986, the police organized the blood sample collection from 5000 men living in the area where the murder took place, and finally found the killer by his DNA profile [3].

In 1983, Kary Mullis developed the polymerase chain reaction (PCR) technique which opened new ways in DNA analysis in the forensic genetics. Thus, from each biological human trace or micro‐trace containing nucleated cells, the DNA is extracted to be subjected to amplification reactions. The method is an enzymatic process by which regions of the DNA are replicated (multiplied) 28–34 times, generating about one billion (109 ) copies. This technique highlights the number of repetitions of the base unit and turns them into alpha‐numeric values, known as genetic profiles.

Genetic analysis of a very small number of nucleated cells, namely a very small amount of biological material, is made by a different approach from the usual situations and aims to generate, through the reaction of amplification, the sufficient quantity of copies of DNA fragments to obtain an exploitable genetic profile. This method is applicable to biological samples with a DNA matrix containing <50 pg., to biological micro‐traces, to biological samples that are in an advanced stage of decay, with single source or multiple sources.

For the last two decades, the results of the DNA analysis have been accepted as evidence in the court in many countries [4]. Since February 1992 when The European Council issued the **Recommendation No. 92**, regarding the use of the DNA analysis in the criminal justice the DNA test is accepted in the Court [5].


**4. Case study of evidence-based flood contingency planning**

The proposed method was applied to Calumpit Municipality in Bulacan Province.

**43**

**Figure 4.**

*Evidence-Based Contingency Planning to Enhance Local Resilience to Flood Disasters*

**Figure 4** illustrates overall activities related to the six steps to be proposed. The first and second steps were applied to the whole area of the municipality in April 2014. At the third and fourth steps, two flood-prone communities were selected as model sites, and their flood contingency plans were jointly developed through workshops with community leaders and members. In the workshops, the participants discussed problems they may face during the flood of each scale and proposed necessary response actions in order to make response strategies. The final workshop was held in March 2016, inviting all the community leaders in the municipality, and the experience in the workshops was shared among the participants. The following

In areas where the availability of natural and socio-economic data is limited, administering interviews and questionnaires to local government officials, community leaders and local people is useful to understand the current conditions of their localities. In this study, interviews were conducted at MDRRMO and selected communities [5]. The surveys found that population census data at the barangay level was available, while the spatial distribution data of buildings was not. Then, a questionnaire survey was administered to all 29 barangay leaders to understand the building conditions in each barangay. The houses in the 29 barangays were classified and tallied according to construction types and storeys. Of those, 62.5% were a

Interviews at the selected individual households were also conducted to understand recent flood damage, including the damage status of house structures, property and family members and their behaviours during the recent floods. During the interviews, the survey staff measured the heights of the first floor, ceiling and flood marks from the past floods with the permission of family members. From the household survey, it was found that the average first-floor height of the one-storey houses was 0.54 m from the ground, while that of the two-storey houses was 0.17 m, as shown in **Figure 5**. **Table 1** summarises the conditions of the houses at five inundation levels using the thresholds from the household interviews. Different damages to the livelihood of the residents are listed according to inundation levels. Inundation level 1 with a water depth of 0.17 m or lower did not inundate inside the house. At

*DOI: http://dx.doi.org/10.5772/intechopen.82312*

**4.1 Step 1: understanding current conditions**

one-storey structure, while the rest were two-storey.

*Activities at six steps of evidence-based flood contingency planning.*

subchapters detail each step:

*Evidence-Based Contingency Planning to Enhance Local Resilience to Flood Disasters DOI: http://dx.doi.org/10.5772/intechopen.82312*

**Figure 4** illustrates overall activities related to the six steps to be proposed. The first and second steps were applied to the whole area of the municipality in April 2014. At the third and fourth steps, two flood-prone communities were selected as model sites, and their flood contingency plans were jointly developed through workshops with community leaders and members. In the workshops, the participants discussed problems they may face during the flood of each scale and proposed necessary response actions in order to make response strategies. The final workshop was held in March 2016, inviting all the community leaders in the municipality, and the experience in the workshops was shared among the participants. The following subchapters detail each step:

### **4.1 Step 1: understanding current conditions**

*Recent Advances in Flood Risk Management*

barangay, municipality, city and province.

Pedring and Quiel in September 2011, and a large area of the municipality suffered massive flood damage. Due to an inundation of 1.2–1.5 m deep, Calumpit lost its government functions, which consequently impeded emergency response.

The Philippine Disaster Risk Reduction and Management Act (Republic Act No. 10121) [3], enacted in 2010, provides for the development of policies and plans and the implementation of actions and measures related to all aspects of disaster risk reduction and management. It defines the National Disaster Risk Reduction Management Council (NDRRMC) as the national organisation to coordinate, integrate, supervise, monitor and evaluate disaster policymaking. It also mandates the establishment of the Disaster Risk Reduction and Management Office in every

Calumpit has the Municipal Disaster Risk Reduction Management Council (MDRRMC) and the Barangay Disaster Risk Reduction Management Council (BDRRMC) in its 29 barangays. The act defines MDRRMCs and BDRRMCs to perform functions such as designing and coordinating disaster risk reduction and management activities, supporting risk assessments and contingency planning activities at the local level [3]. Following this act, the MDRRMC of Calumpit published a contingency plan [4], which describes governmental emergency response in case of a flood. It assumes casualties, structural damage and impacts on livelihood, infrastructure and facilities in the worst-case scenario, based on the experience during Typhoons Pedring and Quiel in 2011. However, the scenario based on the past flood experiences makes it difficult to assume future floods of different scales which have never occurred before. It is therefore recommended to make contingency plans by quantifying a spatial distribution of expected damage and necessary needs in consideration of dynamic changes in inundation depth provided from

flood inundation simulations performed for each community.

**4. Case study of evidence-based flood contingency planning**

The proposed method was applied to Calumpit Municipality in Bulacan

**42**

Province.

**Figure 3.**

*Land use map of Calumpit [4].*

In areas where the availability of natural and socio-economic data is limited, administering interviews and questionnaires to local government officials, community leaders and local people is useful to understand the current conditions of their localities. In this study, interviews were conducted at MDRRMO and selected communities [5]. The surveys found that population census data at the barangay level was available, while the spatial distribution data of buildings was not. Then, a questionnaire survey was administered to all 29 barangay leaders to understand the building conditions in each barangay. The houses in the 29 barangays were classified and tallied according to construction types and storeys. Of those, 62.5% were a one-storey structure, while the rest were two-storey.

Interviews at the selected individual households were also conducted to understand recent flood damage, including the damage status of house structures, property and family members and their behaviours during the recent floods. During the interviews, the survey staff measured the heights of the first floor, ceiling and flood marks from the past floods with the permission of family members. From the household survey, it was found that the average first-floor height of the one-storey houses was 0.54 m from the ground, while that of the two-storey houses was 0.17 m, as shown in **Figure 5**. **Table 1** summarises the conditions of the houses at five inundation levels using the thresholds from the household interviews. Different damages to the livelihood of the residents are listed according to inundation levels. Inundation level 1 with a water depth of 0.17 m or lower did not inundate inside the house. At

#### **Figure 4.**

*Activities at six steps of evidence-based flood contingency planning.*

**Figure 5.**

*Threshold of inundation based on measurement results.*

inundation level 2, the two-storey houses started being inundated. At inundation level 3, at which the water depth exceeds 0.54 m, both one- and two-storey houses suffered from inundation above the first floor, which suggests that the residents had to stay somewhere above the water level or evacuate to safer places near their houses.

The household interviews also found that the inundation above electric plugs caused severe damage to daily life. The height of electric plugs averaged 1.27 m and that of LP gas tanks 0.60 m. The residents usually move LP gas tanks to the second floor or to the rooftop to use them for cooking during an inundation. Inundation level 4 was set, based on the observed average height of electric plugs, as the condition cutting local people off from power. At inundation level 5, the inundation depth exceeds 2.83 m, the height of the second floor of a house. Under this situation, they could not find an evacuation space due to the rarity of buildings having three storeys or more, which means an inundation of this scale is likely to be a potentially life-threatening crisis for the residents.

Calumpit Municipality has its own community flood warning system called "colours of safety". This system uses power poles painted in three colours (yellow, orange and red) by every 2 ft to visualise the level of danger and help residents make decisions on evacuation. At present, 193 electric poles in the municipality are tricoloured for this purpose. The residents are advised to evacuate before the water reached the red colour.


**1. Introduction**

Reactive oxygen species (ROS) are directly associated with a variety of degenerative and age-related diseases, and other pathologies, including di fferent types of cancers. ROS are generated as by-products of cellular metabolism and their excessive production can damage lipids, proteins, and DNA, different cells, and tissues [1].

Resveratrol (RSV) (trans-3,4,5-trihydroxystilbene) is a natural antioxidant commonly found in grapes, berries, and nuts that has shown promising results in the treatment of a variety of degenerative and age-related diseases. It blocks the activation of nuclear factor κB (NF-κB), reducing the generation of ROS and pro-inflammatory cytokines (interleukin IL-1 β and IL-6), which results in the inhibition of chondrocyte apoptosis, inflammation, and the progression of several diseases [2,3]. It is also a direct inhibitor of cyclooxygenase 2 (COX-2) which produces pro-inflammatory lipid mediators (leukotrienes and prostaglandins) responsible for pain sensation [4].

In recent years, RSV has also shown beneficial results in modulation of tissue regeneration, microcirculation, the function of peripheral nerves, production of anti-inflammatory cytokines, and insulin [5–9]. However, RSV shows low long-term stability, rapid metabolism, and release; low aqueous solubility (0.05 mg/mL) and bioavailability. RSV is also unstable under the influence of light, certain pH levels, and temperature, which causes isomerization or degradation of RSV, making it difficult to apply in the medical–pharmaceutical area [10–13]. To overcome these limitations, RSV has been employed in pharmaceutical formulations in different drug delivery systems (DDS), including microparticulate system [14,15], micro/nanocapsules [16], cyclodextrin complexes [17,18], solid lipid nanoparticles (SLN) [19–22], nanosuspensions [23], vesicular systems liposomes [24], niosomes [25], nanosponges [18,26], microspheres [27], transfersomes and ethosomes [14,15,28], and nanostructured lipid carriers (NLC) [19,20]. The DDS are employed to improve the physicochemical stability of loaded drugs, provide a sustained-release profile, increase plasma half-life, decrease the risk of immunogenicity, improve the drug solubility and thereby its bioavailability and therapeutic activity, enhance antioxidant activity, and improve the permeation and targeted delivery [14,15].

NLC consists of a mixture of solid and liquid lipids, which creates an imperfect crystalline structure, providing more space between the lipid chains and the matrix [29,30]. The main advantages of the use of NLC are high encapsulation efficiency and storage stability, the possibility of controlling the release of several drugs, low toxicity due to the absence of solvents in the production process, and the possibility of production in an industrial scale [31,32]. Moreover, Gokce et al. [14] observed that RSV-loaded NLC penetrated deeper into the skin [19]. Jose et al. showed a negligible release of resveratrol over several hours, corroborating the high stability of RSV-loaded NLC [20].

Experimental designs have been the most used tools to simultaneously analyze the influence of different variables on the properties of NLC, aiming to ensure the high product quality, the economy of production, and reduction of production time, allowing the scale-up of the process [33].

To evaluate the optimum experimental conditions for NLC production, several authors have assessed the influence of different factors which can affect the final properties of formulations, including type, ratio, and concentration of lipids and stabilizers, cycle numbers, time and intensity of homogenization, and pressure [34–36]. Thus, this work reports the effects of the production process parameters, shear intensity, and homogenization time of RSV-loaded NLC by means of a 22 factorial design with triplicate of the central point, measuring the mean particle size (PS) and polydispersity index (PDI) as the dependent variables.

#### **2. Materials and Methods**


**Wettability of Carbonaceous Materials with Molten Iron at 1550°C**

### R. Khanna, I. Mansuri and V. Sahajwalla

Additional information is available at the end of the chapter

http://dx.doi.org/10.5772/61198

#### **Abstract**

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[116] Bačáková L., Filová E., Pařízek M., Ruml T., Švorčík V. Modulation of cell adhesion, proliferation and differentiation on materials designed for body implants. Biotech‐

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nology Advances 2011; 29 739–767.

356 Wetting and Wettability

In the direct iron smelting process, interfacial reactions of carbonaceous materials with molten iron are among some of the key factors that dictate the rate of carbon transfer into molten iron and establish a carbon concentrated melt to reduce iron oxide in the slag phase. Detailed wettability investigations on a range of carbonaceous materials, e.g. synthetic graphite, natural graphite, coke, coal-chars, coke-polymer blends, waste plastics in contact with molten iron at 1550°C were carried out using the sessile drop method. Experimental results on dynamic contact angles are presented in this chapter and are discussed in terms of the basic characteristics of carbons and the changing composition of the interfacial region as a function of time.

The influence of carbon and sulfur content in Fe-C-S melts on the wettability of synthetic graphite was investigated. The initial melt carbon content was in the range of 0.13 to 2.24 wt %, and the melt sulfur content was in the range of 0.05 to 0.37 wt %. It was found that the contact of solid graphite with Fe-C-S melts resulted in a nonequilibrium reactive wetting. It involved the transfer of carbon from the solid to the liquid and iron transfer from the liquid to the solid. The Fe-C-S melts exhibited relatively poor wetting in the absence of such a material exchange.

The wetting of natural graphite, which contained 8.8 pct ash, by iron was investigated to establish a fundamental understanding of the influence of ash on interactions between graphite and iron. It was found that the formation of an ash interfacial layer between the carbonaceous material and liquid iron had a strong influence on the mass transfer and interfacial reaction.

Wetting investigations of pure liquid iron with three types of metallurgical cokes at 1550°C showed a non-wetting behaviour for all cokes with contact angles ranging from 123° – 126° initially, which marginally decreased to 109°-114° after 60 minutes

© 2015 The Author(s). Licensee InTech. This chapter is distributed under the terms of the Creative Commons Attribution License (http://creativecommons.org/licenses/by/3.0), which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.

of contact. The slight decrease in contact angles was attributed to the transfer of carbon from coke to liquid iron, which then led to a lower the interfacial tension.

The wettability of four non-graphitic coal-chars with electrolytic pure iron was investigated at 1550°C. The ash concentration in chars ranged between 9.04 and 12.61 wt %, with alumina and silica as predominant ash components. The initial contact angles ranged between 106° and 137°, the contact angles for all coal-chars were quite similar after 60 minutes of contact (105 - 110°). While no well-defined correlations could be observed between the initial char structure and ash concentration/composi‐ tion and contact angles in the initial stages of contact, the contact angles over extended periods were significantly affected by the presence of reaction products and impurity deposits in the interfacial region.

**Keywords:** Carbonaceous materials, molten iron, wettability, contact angles

#### **1. Introduction**

Ironmaking is one of the largest consumers of carbon fuels in the world, with an annual energy consumption of 24 EJ (x 1018J) and 69% of the world's total energy consumption in iron/ steelmaking industry [1]. Conventional ironmaking involves temperatures of more than 1500°C, leading to high-energy consumption. It is also dependent upon high quality coking coal, and cannot use low-grade waste materials such as coal fines. To lower both the pollution and carbon emissions, the energy consumption levels need to be lowered and this can be achieved by lowering the temperature of ironmaking. Previous attempts to lower the temper‐ atures have focused on macro-scale reactions. There have been attempts to develop a lowertemperature ironmaking process to reduce energy requirements and emissions by replacing coke with coal fines; but what have been predominantly iron oxide reduction and industryoriented experiments, have encountered crucial issues of operational stability and lowtemperature carburization.

'Mechanisms for liquid iron carburization were extensively investigated in a previous study from our group [2] and the results showed a close relationship with composition of inorganic materials, carbon structure, and temperature. However, solid state carburization is rather complex due to the combination of gas phase carburization reaction and solid state diffusion of carbon [3, 4]. Due to difficulties associated with measuring contact angles with fine powders, very limited information is currently available on the wettability of non-graphitic carbons, such as coal-chars, and the role of carbon type and ash composition is not well determined.

In the direct iron smelting process, interfacial reactions of carbonaceous materials with molten iron are among some of the key factors that dictate the rate of carbon transfer into molten iron and establish a carbon concentrated melt to reduce iron oxide in the slag phase. Basic param‐ eters such as wettability, fixed carbon, ash content/composition, and crystalline order of the carburizing material influence the dissolution kinetics of carbon [5, 6]. Various metal oxides present as ash impurities can also form a viscous layer in the interfacial region affecting wettability and may act as a physical barrier for the transfer of carbon. An understanding of wettability and associated interfacial phenomena is required to establish carburizing funda‐ mentals of various carbonaceous materials. The interfacial phenomena between carbonaceous materials and molten iron is very important due to the extensive use of these materials in iron and steelmaking.

Wu and Sahajwalla provided an alternative explanation for the initial decrease in the contact angle for Fe-C-S melts on graphite followed by a recovery to the equilibrium value [7]. They postulated that the initial decrease was due to the diffusion of carbon and sulfur atoms to the interfacial layer resulting in a reduction in the interfacial tension between the solid and the liquid. Once the diffusing species reduced the chemical potential difference between the solid and the liquid to negligible levels, the Van der Waals forces begin to dominate the wettability. This results in the contact angle approaching its equilibrium value.

The presence of surface-active elements at the interface can affect the interfacial energy. Surface activity is defined as the magnitude of change in surface tension brought about by the addition of a unit quantity of the species. The ultimate surface activity of a surface-active species in a binary system generally decreases with increasing temperature. Additions of more than one solute to a system complicate the determination of surface activity. The compounded results cannot be predicted from the individual contributions. For example carbon, a non-surface active element, increases the activity coefficient of the dissolved sulfur in liquid iron, lowers the surface tension of Fe-C-S solutions thereby enhancing the surface concentrations of sulfur. In an atomistic computer simulation study on the Fe-C-S system, S and C were found to mutually displace each other. Carbon-rich regions were depleted of sulfur and vice versa [8].

Detailed wettability investigations on a range of carbonaceous materials, e.g., synthetic graphite, natural graphite, coke, coal-chars, coke-polymer blends, and waste plastics, in contact with molten iron at 1550°C were carried out using the sessile drop method. Experimental results on dynamic contact angles are presented in this chapter and are discussed in terms of the basic characteristics of carbons and the changing composition of the interfacial region as a function of time. The systems of interest in this study is rather complex: along with a hetero‐ geneous composition, the carbonaceous materials contain non-graphitic carbon and ash minerals; the liquid metal typically contains carbon, sulfur, and quite a few trace elements. In combination, these can result in a number of reactions occurring in the interfacial region, which could a have significant influence on the wettability of the system.

#### **2. Experimental**

of contact. The slight decrease in contact angles was attributed to the transfer of carbon

The wettability of four non-graphitic coal-chars with electrolytic pure iron was investigated at 1550°C. The ash concentration in chars ranged between 9.04 and 12.61 wt %, with alumina and silica as predominant ash components. The initial contact angles ranged between 106° and 137°, the contact angles for all coal-chars were quite similar after 60 minutes of contact (105 - 110°). While no well-defined correlations could be observed between the initial char structure and ash concentration/composi‐ tion and contact angles in the initial stages of contact, the contact angles over extended periods were significantly affected by the presence of reaction products and impurity

from coke to liquid iron, which then led to a lower the interfacial tension.

**Keywords:** Carbonaceous materials, molten iron, wettability, contact angles

Ironmaking is one of the largest consumers of carbon fuels in the world, with an annual energy consumption of 24 EJ (x 1018J) and 69% of the world's total energy consumption in iron/ steelmaking industry [1]. Conventional ironmaking involves temperatures of more than 1500°C, leading to high-energy consumption. It is also dependent upon high quality coking coal, and cannot use low-grade waste materials such as coal fines. To lower both the pollution and carbon emissions, the energy consumption levels need to be lowered and this can be achieved by lowering the temperature of ironmaking. Previous attempts to lower the temper‐ atures have focused on macro-scale reactions. There have been attempts to develop a lowertemperature ironmaking process to reduce energy requirements and emissions by replacing coke with coal fines; but what have been predominantly iron oxide reduction and industryoriented experiments, have encountered crucial issues of operational stability and low-

'Mechanisms for liquid iron carburization were extensively investigated in a previous study from our group [2] and the results showed a close relationship with composition of inorganic materials, carbon structure, and temperature. However, solid state carburization is rather complex due to the combination of gas phase carburization reaction and solid state diffusion of carbon [3, 4]. Due to difficulties associated with measuring contact angles with fine powders, very limited information is currently available on the wettability of non-graphitic carbons, such as coal-chars, and the role of carbon type and ash composition is not well determined. In the direct iron smelting process, interfacial reactions of carbonaceous materials with molten iron are among some of the key factors that dictate the rate of carbon transfer into molten iron and establish a carbon concentrated melt to reduce iron oxide in the slag phase. Basic param‐ eters such as wettability, fixed carbon, ash content/composition, and crystalline order of the carburizing material influence the dissolution kinetics of carbon [5, 6]. Various metal oxides

deposits in the interfacial region.

**1. Introduction**

358 Wetting and Wettability

temperature carburization.

The wettability of carbonaceous materials and Fe-C-S system was investigated in a laboratoryscale, horizontal tube resistance furnace using the sessile drop approach. The carbonaceous materials formed into cylindrical substrates (25 mm diameter, 3–4 mm height) and compressed under 7 tonne pressure (for approximately 1 minute) using a hydraulic press (ENERPAC, Australia®). The weight of the metal used was ~0.70 g of electrolytic-grade iron and Fe-C-S alloys, respectively. The static nature of the droplet in the sessile drop arrangement helps to preserve the interfacial layer that forms between the iron and carbonaceous materials, allowing a detailed investigation of the interface. Apart from Marangoni flow, there is minimal flow within the liquid iron droplet in the sessile drop method.

A schematic diagram of the experimental arrangement is shown in Fig. 1. The furnace tube had an inside diameter of 50 mm. Initially, the metal/carbon assembly was held on a specimen holder, which could be pushed to the center of the hot zone in the furnace with the help of a graphite rod. The metal/carbon assembly was held in the cold zone of the furnace until the desired temperature (1550°C) was attained. The assembly was then inserted into the hot zone; this eliminated any reaction that could occur at lower temperatures and possibly influence the phenomena to be studied at the temperature of interest. The melting of iron marked the beginning of contact time. The furnace tube was purged with argon throughout the duration of the experiment with a flow rate of 0.5 L/min.

**Figure 1.** A schematic representation of experimental arrangement.

The wettability behavior of the metal/carbon system was investigated using a closely control‐ led and visually monitored sessile drop technique. A high-quality, high-resolution chargecoupled device (CCD) camera fitted with an IRIS lens was used to capture the live in-situ phenomena in the furnace. The output from the camera was channeled to a video cassette recorder (VCR) and a television (TV) monitor to record the entire process as a function of time. This allows specific images, displaying the contact between the metal and carbonaceous material, to be captured from the videotape, as a function of time, into a computer using a frame grabber. A time-date generator is used in the system to display the duration of the process. For a better understanding of reaction dynamics, contact angles were recorded up to 2 hours in most cases. Fig. 2 shows a representative image of molten iron droplet during melting in the horizontal tube furnace.

Specially designed computer software was used to determine the contact angle from the captured images, on the basis of a curve-fitting exercise. The contact angle between the iron

**Figure 2.** Representative image of molten iron droplet during melting in the horizontal tube furnace.

droplet and the substrate was measured by using computer image analysis software based on a curve-fitting algorithm. The working principle of this algorithm is detailed in the work conducted by [7] and is shown in Fig. 2. Typical accuracy of contact angle measurements was within ± 5%.

#### **3. Wettability and interfacial phenomena**

Wettability investigations were carried out on a number of carbonaceous materials in contact with molten iron at 1550°C. Detailed results are presented in this section.

#### **3.1. Synthetic and natural graphite**

alloys, respectively. The static nature of the droplet in the sessile drop arrangement helps to preserve the interfacial layer that forms between the iron and carbonaceous materials, allowing a detailed investigation of the interface. Apart from Marangoni flow, there is minimal flow

A schematic diagram of the experimental arrangement is shown in Fig. 1. The furnace tube had an inside diameter of 50 mm. Initially, the metal/carbon assembly was held on a specimen holder, which could be pushed to the center of the hot zone in the furnace with the help of a graphite rod. The metal/carbon assembly was held in the cold zone of the furnace until the desired temperature (1550°C) was attained. The assembly was then inserted into the hot zone; this eliminated any reaction that could occur at lower temperatures and possibly influence the phenomena to be studied at the temperature of interest. The melting of iron marked the beginning of contact time. The furnace tube was purged with argon throughout the duration

The wettability behavior of the metal/carbon system was investigated using a closely control‐ led and visually monitored sessile drop technique. A high-quality, high-resolution chargecoupled device (CCD) camera fitted with an IRIS lens was used to capture the live in-situ phenomena in the furnace. The output from the camera was channeled to a video cassette recorder (VCR) and a television (TV) monitor to record the entire process as a function of time. This allows specific images, displaying the contact between the metal and carbonaceous material, to be captured from the videotape, as a function of time, into a computer using a frame grabber. A time-date generator is used in the system to display the duration of the process. For a better understanding of reaction dynamics, contact angles were recorded up to 2 hours in most cases. Fig. 2 shows a representative image of molten iron droplet during

Specially designed computer software was used to determine the contact angle from the captured images, on the basis of a curve-fitting exercise. The contact angle between the iron

within the liquid iron droplet in the sessile drop method.

360 Wetting and Wettability

of the experiment with a flow rate of 0.5 L/min.

**Figure 1.** A schematic representation of experimental arrangement.

melting in the horizontal tube furnace.

A number of researchers have investigated the wettability between graphite and pure iron, graphite and Fe-C alloys at temperatures of 1300–1550°C under atmospheres of helium, argon and hydrogen gases and in vacuum. These studies have shown that pure iron showed good wetting with graphite at high temperatures under these conditions. During sessile drop experiments of iron with graphite, it was observed that in the initial stages, when the iron had melted completely, the carbon content had reached 4.23%; and in the final stages, it had reached the saturation carbon level of 5.62%. The contact angle changed from 64° at the initial state to 38° in final stages when the carbon content had reached saturation.

Lower contact angles in C/Fe interaction have been explained through the lowering of interfacial tension between the two phases due to the mass transport across the interface. The low value of the contact angle was maintained even when the carbon transport rate slowed down at higher carbon content in the liquid droplet during the later period of carbon had dissolution reaction [9, 10]. Thus for the sessile drop experiment, the reaction area was

**Figure 3.** The flow sheet of "interface software" for analysis of the contact angle of a sessile drop image [7].

estimated reasonably well by the lower value of the contact angle. Some experimental results are listed in Table 1.

Sun et al. [13] found that the contact angle of iron on carbon with 0% initial carbon was 59° at 1500 °C, while that of a sample with 4.8 wt% carbon was 118°. It was considered that the contact angle for a low carbon sample is very small, because a lower interfacial tension exists during the rapid initial carbon pickup reaction occurring at the metal-graphite interface. This initial rapid carbon pickup is considered to occur during and immediately after melting. Once the interfacial area extends during the initial carbon pickup reaction, it cannot recover to the shape where a higher carbon content sample was used, even though carbon content reaches the saturation level. These studies showed that pure iron had good wetting with graphite at high temperatures under vacuum, helium and hydrogen and argon gases [7, 11].

**Table 1.** Some experimental results on the wettability of synthetic graphite.

estimated reasonably well by the lower value of the contact angle. Some experimental results

Select image area

Binary Image

Edge detection

Inclination ?

n

Setting top of substrate

Blanking extraneous information

> Parameters selection

Profile points initialization

Curve fitting calculation

Overall sessile drop drawing graph display

Best fitting polynomial selection

> Data ,image storage

> > End

**Figure 3.** The flow sheet of "interface software" for analysis of the contact angle of a sessile drop image [7].

y

Correction

Sun et al. [13] found that the contact angle of iron on carbon with 0% initial carbon was 59° at 1500 °C, while that of a sample with 4.8 wt% carbon was 118°. It was considered that the contact angle for a low carbon sample is very small, because a lower interfacial tension exists during

are listed in Table 1.

362 Wetting and Wettability

Wu and Sahajwalla [7] investigated the dynamic wetting behavior of Fe-C-S melts with graphite substrate using a sessile drop technique by focusing on the influence of melt carbon and sulfur concentration on the wettability at 1600° C. The wettability between the Fe-C-S melt and solid graphite has been presented in terms of the changes in contact angles with time (Fig. 4).

**Figure 4.** The variation of contact angles as a function of time for molten Fe-S-C and graphite system.

From Fig. 4, in the initial state (within ~8 seconds of contact), the system is observed to be nonwetting (~100°). Thereafter, this system becomes wetting (~60°) with a considerable decrease in the contact angles (in the time period ~8 to 90 seconds). The contact angles were observed to stabilize at ~60° for the rest of the experimental time. The changes in contact angles in this system were attributed to carbon and iron inter-diffusing into the solid graphite and the liquid iron.

Wu et al. [14] investigated the wetting behavior of natural graphite and liquid iron system using the sessile drop technique and the results are shown in Fig. 5. The contact angles of this system decreased from ~100° to become wetting within the first 10 seconds (~60°). Thereafter, it became non-wetting with the contact angles sharply increasing to stabilize at ~100° within 100 seconds. This dynamic wetting behavior of iron and natural graphite is very different from the behavior of molten Fe-C-S and graphite systems; the deposition of ash oxides at the interface played a strong role in this phenomena.

**Figure 5.** The variation in the contact angles as a function of time for molten iron on natural graphite substrate.

#### **3.2. Coal chars**

Raw coals were crushed in a jaw crusher and vibrating grinder, and wet sieved and the resulting size graded samples were dried in an oven prior to char making. The chars were produced from raw coals in a drop-tube furnace in an atmosphere of 23 % O2 in N2 at a temperature of 1200 °C.

The wettability for four chars was determined using the sessile drop approach and the contact angles were measured. Chars 1 and 2 had a very high initial contact angle of 145° and 140° respectively, which decreased to ~110° after 60 minutes of contact. While the contact angle for Char 1 showed a gradual decline with time, the contact angles for Char 2 showed a sharp reduction to 110° after 2 minutes of contact and then fluctuated around that value over long times of contact. Chars 3 and 4 had somewhat lower initial contact angles, 115° and 120° respectively, but these showed no improvement with time. Following a trend contrary to Chars 1 and 2, these chars did not show a sharp initial decline; there were only marginal fluctuations in contact angles as a function of time. After 60 minutes of contact, the contact angles for all four chars were quite similar ranging between 105–110°.

Only a marginal improvement in wettability (decrease in contact angles) was observed as a function of time. The authors identified the factors influencing the wettability as: transfer of carbon and sulfur across the metal/carbon interface, the formation of Al2O3 and CaS complexes at the interface and the reduction of reducible oxides, such as SiO2 and Fe2O3, and the transfer of Si into the liquid metal.

#### **3.3. Metallurgical coke**

in the contact angles (in the time period ~8 to 90 seconds). The contact angles were observed to stabilize at ~60° for the rest of the experimental time. The changes in contact angles in this system were attributed to carbon and iron inter-diffusing into the solid graphite and the liquid

Wu et al. [14] investigated the wetting behavior of natural graphite and liquid iron system using the sessile drop technique and the results are shown in Fig. 5. The contact angles of this system decreased from ~100° to become wetting within the first 10 seconds (~60°). Thereafter, it became non-wetting with the contact angles sharply increasing to stabilize at ~100° within 100 seconds. This dynamic wetting behavior of iron and natural graphite is very different from the behavior of molten Fe-C-S and graphite systems; the deposition of ash oxides at the

**Figure 5.** The variation in the contact angles as a function of time for molten iron on natural graphite substrate.

Raw coals were crushed in a jaw crusher and vibrating grinder, and wet sieved and the resulting size graded samples were dried in an oven prior to char making. The chars were produced from raw coals in a drop-tube furnace in an atmosphere of 23 % O2 in N2 at a

The wettability for four chars was determined using the sessile drop approach and the contact angles were measured. Chars 1 and 2 had a very high initial contact angle of 145° and 140° respectively, which decreased to ~110° after 60 minutes of contact. While the contact angle for Char 1 showed a gradual decline with time, the contact angles for Char 2 showed a sharp reduction to 110° after 2 minutes of contact and then fluctuated around that value over long times of contact. Chars 3 and 4 had somewhat lower initial contact angles, 115° and 120° respectively, but these showed no improvement with time. Following a trend contrary to Chars 1 and 2, these chars did not show a sharp initial decline; there were only marginal fluctuations in contact angles as a function of time. After 60 minutes of contact, the contact angles for all

interface played a strong role in this phenomena.

iron.

364 Wetting and Wettability

**3.2. Coal chars**

temperature of 1200 °C.

four chars were quite similar ranging between 105–110°.

An experimental investigation of the wetting behavior of pure liquid iron with three types of metallurgical coke at 1550° C was conducted by Cham et al [15]. It was observed that all cokes showed a non-wetting behavior with contact angles ranging from 123°–126° initially, which marginally decreased to 109°–114° after 60 minutes of contact. It was concluded that the slight decrease in contact angles was due to the transfer of carbon from coke to liquid iron, which then decreased the interfacial tension. On the other hand, the effects of sulfur, ash content and composition and carbon structure on the wettability could not be observed in this study.

The wetting images of molten steel droplet on coke after 60 minutes of reactions are shown in Fig. 6. It can be seen that the liquid steel droplet showed a non-wetting behavior with the coke (contact angle 129°) initially and the value increased slightly within the first few minutes. A very small change was observed throughout the experimental run with the contact angle reaching 128° after 60 mins of contact.

**Figure 6.** Wetting images of steel droplet after reaction with coke for 60 minutes at 1550°C.

#### **3.4. Metallurgical coke/Bakelite blends**

A number of polymers and metallurgical coke were blended in three different ratios: Blend 1, Blend 2, and Blend 3. These were labelled as BK1, BK2, BK3; H1, H2, H3; and P1, P2, P3 for Blends 1, 2 and 3 of Bakelite/Coke, HDPE/Coke, and PET/Coke, respectively. Fig. 7 represents the blending ratios used in these blends.

**Figure 7.** Relative proportions between the plastics and metallurgical coke in the polymer/coke blends.

**Figure 8.** Wetting images of steel droplet after reaction with blend BK1 for 60 minutes at 1550°C.

The raw polymer/coke blends were combusted in a drop tube furnace (DTF) at 1200°C under an atmosphere of 80% nitrogen and 20% oxygen with a gas flow rate of 1.0 L/min, and the chars obtained were used for steel/carbon interaction investigations. The devolatilization was carried out to simulate the real process in Electric Arc Furnace (EAF) steelmaking. In EAF steelmaking process, the carbonaceous materials are injected into the furnace and these undergo partial combustion before coming into contact the molten steel.

**Figure 10.** Wetting images of steel droplet after reaction with blend BK2 for 60 minutes at 1550°C.

Blends 1, 2 and 3 of Bakelite/Coke, HDPE/Coke, and PET/Coke, respectively. Fig. 7 represents

**Figure 7.** Relative proportions between the plastics and metallurgical coke in the polymer/coke blends.

**Figure 8.** Wetting images of steel droplet after reaction with blend BK1 for 60 minutes at 1550°C.

undergo partial combustion before coming into contact the molten steel.

The raw polymer/coke blends were combusted in a drop tube furnace (DTF) at 1200°C under an atmosphere of 80% nitrogen and 20% oxygen with a gas flow rate of 1.0 L/min, and the chars obtained were used for steel/carbon interaction investigations. The devolatilization was carried out to simulate the real process in Electric Arc Furnace (EAF) steelmaking. In EAF steelmaking process, the carbonaceous materials are injected into the furnace and these

the blending ratios used in these blends.

366 Wetting and Wettability

Wetting results showed variations in contact angle between the Bakelite/coke blends. A general improvement in the wettability of the liquid steel with Bakelite/coke blends was observed as compared to wettability with coke alone.

The wetting images of molten steel droplet on blends BK1, BK2, and BK3 and corresponding contact angles after 60 minutes of reactions are shown in Figs. 8–13, respectively.

**Figure 9.** Variation of contact angles of steel droplet with blend BK1 and coke substrates at 1550° C.

**Figure 11.** Variation of contact angles of steel droplet with blend BK2 and coke substrates at 1550°C.

**Figure 12.** Wetting images of steel droplet after reaction with blend BK3 for 60 minutes at 1550°C.

From these figures, it can be seen that the liquid steel droplet showed a non-wetting behavior with the coke (contact angle 129°) initially and the value increased slightly within the first few minutes. Bakelite containing blends were also found to exhibit non-wetting behavior with liquid steel. The liquid steel droplet exhibited better wetting with blend BK1 substrates with the measured contact angles being approximately 10° lower than that seen with coke alone. The initial contact angle was 118° that increased to 121° after 60 minutes of reaction. In the case of blend BK2, the initial contact angle was 123° that decreased to 120° after 15 minutes and

**Figure 13.** Variation of contact angles of steel droplet with blend BK3 and coke substrates at 1550°C.

then increased to almost 139° after 60 minutes. In the case of BK3, the initial contact angle was 131° that initially decreased to 120° after 25 minutes and then increased to 129° after 60 minutes.

The wetting behavior observed in the case of blends BK2 and BK3 was similar to that observed in natural graphite and liquid iron system by Wu et al. [14]. These authors investigated the wetting behavior of natural graphite with liquid iron at 1600°C using the sessile drop techni‐ que. They observed a decrease in contact angles (from ~100° to ~60°) along with a rapid increase in the carbon content in the metal droplet within the first 10 seconds of reaction. Thereafter, the melt carbon content was observed to slightly decrease and the contact angles were observed to increase from ~60° to ~100° within approximately 2 minutes of contact and then stabilize for the rest of the reaction times. Wu et al. [14] explained that the decrease in contact angles within the first 10 seconds was due to a decrease in the interfacial tension due to the dissolution of carbon into the melt. Thereafter, the contact angles increased because of a slowing down in carbon dissolution since the the formation of ash layer at the metal/carbon interface became significant (interfacial blockage and consumption of solute carbon by reducible oxides). The carbon dissolution into the metal droplet and the ash deposits in the interfacial region were also expected to influence the wetting behavior between molten steel and Bakelite/coke blends.

#### **3.5. Bakelite char**

From these figures, it can be seen that the liquid steel droplet showed a non-wetting behavior with the coke (contact angle 129°) initially and the value increased slightly within the first few minutes. Bakelite containing blends were also found to exhibit non-wetting behavior with liquid steel. The liquid steel droplet exhibited better wetting with blend BK1 substrates with the measured contact angles being approximately 10° lower than that seen with coke alone. The initial contact angle was 118° that increased to 121° after 60 minutes of reaction. In the case of blend BK2, the initial contact angle was 123° that decreased to 120° after 15 minutes and

**Figure 12.** Wetting images of steel droplet after reaction with blend BK3 for 60 minutes at 1550°C.

**Figure 11.** Variation of contact angles of steel droplet with blend BK2 and coke substrates at 1550°C.

368 Wetting and Wettability

Wettability investigation was carried out using Bakelite char substrate and molten iron. Video images and the measured contact angles between the substrate and molten iron is shown in Figs. 14 and 15. A non-wetting behavior (contact angle > 90o ) was observed for Bakelite char with liquid iron. Initial contact angle (at 0 minutes) between Bakelite char substrate was found to be 125.7o and decreased to 122.4o after 1 minute of reaction time. Not much change in contact angle was observed at longer times and angles remained more or less constant later on.

**Figure 14.** The wetting behavior of Bakelite char in contact with molten iron as function of time.

**Figure 15.** Contact angle as a function of time between iron and Bakelite char substrate

#### **3.6.  The contact angle between the substrate and molten iron was measured and is shown in Figs. 16 and 17.

**Figure 17.** Contact angle as a function of time between iron and waste CD char substrate.

**Figure 14.** The wetting behavior of Bakelite char in contact with molten iron as function of time.

**Figure 15.** Contact angle as a function of time between iron and Bakelite char substrate

A wettability study was carried out using waste CD char substrate and molten iron. The contact angle between the substrate and molten iron was measured and is shown in Figs. 16 and 17.

**3.6. Waste CD char**

370 Wetting and Wettability

During initial contact from 0 to 10 seconds, a small decrease in contact angle was observed from 79° to 70o , possibly due to the transfer of carbon and also due to a change in interfacial energy due to interaction between char and molten iron. Wu et al. had observed a similar behavior while using graphite as a carbon source. During later periods, an increase was observed in contact angles that reached 100o after 3 minutes and stabilized at 105° after 15 minutes thereby indicating a much poorer wetting with the carbon level in molten iron approaching saturation. This result points to a much lower driving force for carbon dissolution.

**Figure 16.** The wetting behavior of waste CD char substrate in contact with molten iron as function of time.

### **4. Concluding remarks**

Detailed results have been reported on the wettability of a wide range of carbonaceous materials with liquid iron at 1550°C and important fundamental data on contact angles have been provided. This study has identified important drivers in the form of interfacial reactions; the evidence suggests that the deposition of reaction products in the interfacial region strongly affected the wettability and therefore the contact area between carbons and liquid iron. Main findings of this investigation are:

The effect of structure on the initial wettability of carbons were found to be marginal, with no well- defined trends observed with Lc values. Wide variations were observed in contact angles for chars possessing very similar Lc values. Due to the likelihood of structure evolution and changes in particle morphologies, no comments however can be made regarding the influence of structure on contact angles over prolonged periods of contact.

While no correlations were observed between initial contact angles and ash concentration/ composition, final contact angles were significantly affected by the presence of reaction products and impurity deposits in the interfacial region. Aluminum, calcium, sulfur, oxygen, and iron were detected in the interfacial layer. Small concentrations of magnesium and silicon were also detected indicating the partial presence of the parent ash. The presence of ash impurities and the nature of the interfacial layer play a very important role in affecting the wetting behavior, as it is this material that the liquid iron is in quasi-equilibrium with.

It is quite likely that combinations of several reactions occur simultaneously to dictate the wetting behavior. Reactions that could occur and subsequently influence the contact angles are: the formation of an enriched interfacial layer containing calcium, sulfur, and alumina; transfer of carbon and sulfur by mass transport across the interface; and reduction of reducible oxides such as silica and iron oxides, and possible transfer of these elements into liquid iron.

The mass transfer between solid carbons and iron melts was observed to strongly enhance the wetting phenomena and lowered contact angles. The reduction of system free energy corre‐ sponding to the mass transfer reactions was found to influence the formation of the interface region and resulted in the progressive spreading of the wetting line. The composition and thickness of the carbon/iron interfacial layer was dependent on the intensity of mass transfer across the interface. The resulting change in the interfacial energy *γ*ls was a strong function of mass transfer, and was found to vary with the time of contact. The influence of carbon content on the wetting phenomena could only be seen in the initial stages, whereas the influence of sulfur on the wettability was found when the system approached equilibrium.

### **Acknowledgements**

This work was supported by the Australian Research Council (ARC) funds. We gratefully acknowledge their financial contribution. We thank Ikram-ul-Haq and Somyote Kongkarat for some of the experimental work reported in this article.

#### **Author details**

**4. Concluding remarks**

372 Wetting and Wettability

findings of this investigation are:

**Acknowledgements**

Detailed results have been reported on the wettability of a wide range of carbonaceous materials with liquid iron at 1550°C and important fundamental data on contact angles have been provided. This study has identified important drivers in the form of interfacial reactions; the evidence suggests that the deposition of reaction products in the interfacial region strongly affected the wettability and therefore the contact area between carbons and liquid iron. Main

The effect of structure on the initial wettability of carbons were found to be marginal, with no well- defined trends observed with Lc values. Wide variations were observed in contact angles for chars possessing very similar Lc values. Due to the likelihood of structure evolution and changes in particle morphologies, no comments however can be made regarding the influence

While no correlations were observed between initial contact angles and ash concentration/ composition, final contact angles were significantly affected by the presence of reaction products and impurity deposits in the interfacial region. Aluminum, calcium, sulfur, oxygen, and iron were detected in the interfacial layer. Small concentrations of magnesium and silicon were also detected indicating the partial presence of the parent ash. The presence of ash impurities and the nature of the interfacial layer play a very important role in affecting the wetting behavior, as it is this material that the liquid iron is in quasi-equilibrium with.

It is quite likely that combinations of several reactions occur simultaneously to dictate the wetting behavior. Reactions that could occur and subsequently influence the contact angles are: the formation of an enriched interfacial layer containing calcium, sulfur, and alumina; transfer of carbon and sulfur by mass transport across the interface; and reduction of reducible oxides such as silica and iron oxides, and possible transfer of these elements into liquid iron. The mass transfer between solid carbons and iron melts was observed to strongly enhance the wetting phenomena and lowered contact angles. The reduction of system free energy corre‐ sponding to the mass transfer reactions was found to influence the formation of the interface region and resulted in the progressive spreading of the wetting line. The composition and thickness of the carbon/iron interfacial layer was dependent on the intensity of mass transfer across the interface. The resulting change in the interfacial energy *γ*ls was a strong function of mass transfer, and was found to vary with the time of contact. The influence of carbon content on the wetting phenomena could only be seen in the initial stages, whereas the influence of

sulfur on the wettability was found when the system approached equilibrium.

some of the experimental work reported in this article.

This work was supported by the Australian Research Council (ARC) funds. We gratefully acknowledge their financial contribution. We thank Ikram-ul-Haq and Somyote Kongkarat for

of structure on contact angles over prolonged periods of contact.

R. Khanna\* , I. Mansuri and V. Sahajwalla

\*Address all correspondence to<EMAIL_ADDRESS>
Centre for Sustainable Materials Research and Technology, School of Materials Science and Engineering, The University of New South Wales, NSW Sydney, Australia

#### **References**

[11] Humenik M, Kingery WD. Metal-Ceramic Interactions: III, Surface Tension and Wettability of Metal-Ceramic Systems. Journal of the American Ceramic Society.

[12] Naidich YV, Kolesnichenko GA. Investigation of the wetting of diamond and graph‐ ite by molten metals and alloys. Soviet Powder Metallurgy and Metal Ceramics.

[13] Sun H, Mori K, Sahajwalla V, Pehlke R. Carbon solution in liquid iron and iron al‐

[14] Wu C, Sahajwalla V. Dissolution rates of coals and graphite in Fe-C-S melts in direct ironmaking: Influence of melt carbon and sulfur on carbon dissolution. Metallurgical

[15] Cham ST. Investigating factors that influence carbon dissolution from coke into mol‐

loys. High-temperature materials and processes. 1998;17(4):257-70.

and Materials Transactions B. 2000 2000/04/01;31(2):243-51.

ten iron. Sydney: The University of New South Wales; 2007.

1954;37(1):18-23.

374 Wetting and Wettability

1965;3(3):191-5.

### *Edited by Mahmood Aliofkhazraei*

On the liquid 's surface, the molecules have fewer neighbors in comparison with the bulk volume. As a result, the energy interaction shows itself in the surface tension. Traditionally, the surface tension can be assumed as a force in the unit of the length which can be counted by the unit of Newton on squared meter, or energy on the units of the surface. The surface tension, implies the interface between liquid and vapor, which is an example of the surface tensions. The equilibrium between these surface tensions, decides that a droplet on a solid surface, would have a droplet form or will change to layer form. This book collects new developments in wetting and wettability science.

Photo by sirirak / DollarPhotoClub

Wetting and Wettability

Wetting and Wettability

*Edited by Mahmood Aliofkhazraei*


*5.1. Range Measurement*

In Figure 8a, range measurement is implemented when a static target is placed 0.605 m from the receiver. The target is detected with an error of 0.005 m, as mentioned in Table 3. Based on Figure 8a, we can also observe multiple reflections originating from the target and wall; TS1 and TS2 indicate first and second reflections from the target, whereas W is the wall's reflection. Here, the second reflection refers to a signal that is reflected twice from the target and eventually detected by the receiver since the receiver reflects a portion of strong signal. Then, we perform the measurements for two targets placed at 0.6 m and 0.65 m, and Figure 8b shows that the closely placed targets are distinguishable from each other at their locations. Similarly, Figure 8c provides the measurement results for two targets standing at 1.313 m and 1.548 m, respectively. The change in the received power is due to a higher FSPL.

**Figure 8.** Measurement results for different arrangements of static targets. (**a**) Range plot of a target when placed at 0.6 m. (**b**) Range plot for two targets placed at 0.65 m and at 0.6 m. (**c**) Results of two targets placed at 1.313 m and at 1.548 m, respectively.


**5.1 Statement of the problem. Dispersion relation**

In this section we pay special attention to systems, the geometry of which is similar to the geometry of plasma microwave sources and possible development of the new types of DSI in such systems. The simplest theoretical model of plasma microwave generators assumes relativistic e-beam propagating along axis of a plasma filled waveguide of radius *R*. The beam and plasma are assumed to be completely charge and current neutralized. In the waveguide cross-section the plasma and beam are annular, with mean radii *rp* and *rb*. Their thicknesses Δ*p* and Δ*b* are much smaller, than the mean radii. Strong external longitudinal magnetic field is assumed to freeze transversal motion of beam and plasma electrons.

For theoretical study of the problem we use an approach [10], which gives result for arbitrary level of beam-plasma coupling. This condition is obligatory for obtaining comprehensive results. The DR, which follows from the approach, has a form, which clearly shows interaction of the beam and plasma waves. The approach proceeds from equation for polarization potential *ψ*

$$\frac{\partial}{\partial t} \left( \Delta\_{\perp} + \hat{L} \right) \varphi = -4\pi \left( J\_{bx} + J\_{px} \right), \hat{L} = \frac{\partial^2}{\partial x^2} - \frac{1}{c^2} \frac{\partial^2}{\partial t^2} \tag{42}$$

Here *Jbz*ð Þ¼ **r**⊥, *z*, *t pb*ð Þ **r**⊥ *j bz*ð Þ *z*, *t* and *Jpz*ð Þ¼ **r**⊥, *z*, *t pp*ð Þ **r**⊥ *j pz*ð Þ *z*, *t* are perturbations of the longitudinal current densities in the beam and plasma. Functions *pb*,*p*ð Þ **r**⊥ describe transverse density profiles of the perturbations of the longitudinal currents in the beam and the plasma. For homogeneous beam/plasma *pb*,*p* � 1 for infinitesimal thin *pb*,*p* � *δ r* � *rb*,*p* (*δ* is Dirac function), Δ⊥ is the Laplace operator over transverse coordinates, *z* is the longitudinal coordinate, *t* is the time, *c* is the speed of light. The longitudinal electric field expresses as *Ez* ¼ *L*^*ψ*. The equations for *j bz* and *j pz* are

$$\left(\frac{\partial}{\partial t} + V\_b \frac{\partial}{\partial x}\right)^2 j\_{bx} = \frac{\alpha\_b^2 \chi^{-3}}{4\pi} \frac{\partial}{\partial t} E\_x; \left(\frac{\partial}{\partial t} + \nu\right) j\_{px} = \frac{\alpha\_p^2}{4\pi} E\_x,\tag{43}$$

where *ωp*,*b* are the Langmuir frequencies for plasma and beam respectively, *ν* is the effective collision frequency in plasma, *γ* ¼ 1 � *V*2 *b=c*2 �1*=*2 , *Vb* is the velocity of beam electrons.

The DR, which follows from the statement, is still very cumbersome (of integral type). To reduce the DR to a simple algebraic form one should make following expedient for theoretical model assumption: the plasma and the beam are not just thin but infinitesimal thin. In this case the DR becomes

$$D\_p(\boldsymbol{\alpha}, \boldsymbol{k}) D\_b(\boldsymbol{\alpha}, \boldsymbol{k}) = G \, \kappa^4 \delta \varepsilon\_p \delta \varepsilon\_b,\tag{44}$$

$$\begin{aligned} \text{where } D\_{p,b}(\boldsymbol{\omega}, \boldsymbol{k}) = \boldsymbol{k}\_{\perp p,b}^2 - \kappa^2 \delta \varepsilon\_{p,b}, \,\delta \varepsilon\_p = \frac{a\_p^2}{\omega(\boldsymbol{\omega} + \boldsymbol{i}\boldsymbol{\omega})}, \,\delta \varepsilon\_b = \frac{a\_b^2}{r^3(\boldsymbol{\omega} - \boldsymbol{k} V\_b^2)}, \kappa^2 = \boldsymbol{k}^2 - a\boldsymbol{\alpha}^2/c^2, \,\varepsilon\_b = \frac{\kappa^2}{r^3} \end{aligned}$$

*k* is the wave vector along axis, *ω* is the frequency, *k*⊥*p* and *k*⊥*b* play role of the zero order transversal wave numbers for plasma and beam [11, 14].

*New Types of Dissipative Streaming Instabilities DOI: http://dx.doi.org/10.5772/intechopen.98901*

$$k\_{\perp p,b}^2 = \left\{ r\_{p,b} \Delta\_{p,b} I\_l(\kappa r\_{p,b}) \left[ \frac{K\_l(\kappa r\_{p,b})}{I\_l(\kappa r\_{p,b})} - \frac{I\_l(\kappa R)}{K\_l(\kappa R)} \right] \right\}^{-1} \tag{45}$$

(*Il* and *Kl* are modified Bessel and Mac-Donald functions, *l* ¼ 0, 1, 2 … is the azimuthal wave numbers). *G* is the coupling parameter. It depends on the overlap of the plasma and the beam fields and shows efficiency of their interaction

$$\mathbf{G} = \begin{cases} \frac{I\_l(\kappa r\_b)K\_l(\kappa r\_p)I\_l(\kappa R) - K\_l(\kappa R)I\_l(\kappa r\_p)}{I\_l(\kappa r\_p)K\_l(\kappa r\_b)I\_l(\kappa R) - K\_l(\kappa R)I\_l(\kappa r\_b)} & r\_b \le r\_p\\ \frac{I\_l(\kappa r\_p)K\_l(\kappa r\_b)I\_l(\kappa R) - K\_l(\kappa R)I\_l(\kappa r\_b)}{I\_l(\kappa r\_b)K\_l(\kappa r\_p)I\_l(\kappa R) - K\_l(\kappa R)I\_l(\kappa r\_p)} & r\_p \le r\_b \end{cases} \tag{46}$$

An important property of *G* is: *G* =1 for *rp* ¼ *rb* and *G* < 1 in other cases. In long wavelength limit (for definiteness *l* ¼ 0 and *rb* ≤ *rp*) we have *G*≈ln *R=rp* � �*=* ln ð Þ *R=rb* , but in opposite limit *G*≈exp �2*κ rp* � *rb* � � � � � � (for arbitrary *l*).

### **5.2 Growth rates**

The DR (44) determines proper oscillations of transversally no uniform beamplasma waveguide. The changes of the physical character of beam-plasma interaction must reveal themselves on its solutions. *Dp*,*b*ð Þ¼ *ω*, *k* 0 are the DR for waveguide with thin annular plasma and e-beam respectively. The spectra of fast (+) and slow (�) waves are

$$a\_{\pm} = kV\_b(1 + \mathfrak{x}\_{\pm}); \mathfrak{x}\_{\pm} = \left(\sqrt{a}/\gamma\right)\left(\pm\sqrt{\beta^4 \gamma^2 a + 1} - \beta^2 \gamma\sqrt{a}\right);\tag{47}$$

where *β* ¼ *Vb=c*. The parameter *α* ¼ *ω*2 *b=k*2 ⊥*bV*2 *bγ*3 is familiar (see above). It determines the beam current value: *α* ¼ *Ib= γ*2 ð Þ *I*0 (*Ib* is the beam current, *I*0 is the limiting current in vacuum waveguide). In the limit of under-limiting beams *x*� ! � ffiffiffi *α* p *=γ*. In opposite limit of over-limiting beam *x*þ ¼ 1*=*2*β*2 *γ*2 and *x*� ¼ �2*β*2 *α*. If one looks for solutions of (44) in form *ω* ¼ *kVb*ð Þ 1 þ *x* , *x*< < 1 it becomes

$$\left(\varkappa + q + i\frac{\nu}{ku}\frac{1 - 2\beta^2 \gamma^2 \varkappa}{2\gamma^2}\right)(\varkappa - \varkappa\_+)(\varkappa - \varkappa\_-) = G\frac{a}{2\gamma^4}\left(1 - 2\beta^2 \gamma^2 \varkappa\right)^2,\tag{48}$$

where *q* ¼ *k*2 ⊥*pu*2*γ*2*=ω*2 *p* � 1 � �*=*2*γ*2. The Eq. (48) presents sound way to study

instabilities in given system. First of all, it is easily seen that in conditions of growing negative energy wave *x*≈*x*� and collective Cherenkov resonance *q*≈ � *x*� the role of dissipation increases. For under-limiting e-beams *α* ≤1*=γ*2 and in case of strong coupling *G* � 1 the DR (44) leads to the well-known conventional beam instabilities of no-dissipative and dissipative type. The growth rates of these instabilities have well-known dependencies on beam density � *n* 1*=*3 *b* and on dissipation (� 1*=* ffiffi *ν* p ). Both for these instabilities proper oscillations of the beam are neglected. Only for explanation of the physical meaning of the DSI the conception of NEW should be invoked. However, if *G* < <1 (weak coupling) the growing of the NEW plays dominant role. In this case the growth rate of no-dissipative instability reaches its maximum under Collective Cherenkov resonance *q* ¼ ffiffiffi *α* p *=γ* and is equal

$$\left(\mathrm{Im}\rho\right)\_{\mathrm{und}}^{\left(\nu=0\right)} = \left(kV\_b/2\gamma\right)\left(\mathrm{G}\sqrt{a}/\gamma\right)^{\mathrm{i}\cdot\mathrm{i}}.\tag{49}$$

This expression coincides to (27). Dissipation coming into interplay transforms this instability to DSI of new type with growth rate (coincides to (29))

$$(\text{Im}\,\rho)^{(\nu\to\infty)}\_{\text{und}} = \mathcal{G}\sqrt{a}(kV\_b)^2/2\eta\nu\tag{50}$$

As it should be, this is the instability discovered under consideration of the classical problem of the initial perturbation development in weakly coupled beamplasma systems.

Of particular interest are limit of high, over-limiting currents of e-beam *γ*�2 < <*α* < < 1. In this case the DR (44) takes the form

$$\left(\varkappa + q + i\frac{\nu}{ku}\frac{1 - 2\gamma^2\varkappa}{2\gamma^2}\right)(\varkappa + 2a) = -G\frac{a}{\gamma^2}(1 - 2\gamma^2\varkappa) \tag{51}$$

For *ν* ¼ 0 the analysis of (51) leads to following. Under single particle resonance we have either instability of negative mass type (under *G* � 1) with the growth rate Im*ω* ¼ *ku* ffiffiffi *α* p *=γ*, or stability (under *G* < < 1). But under collective Cherenkov effect *q* ¼ 2*α* the growth rate of developing instabilities is

$$(\text{Im}\,a)^{(\iota=0)}\_{\text{ovl}} = \begin{cases} \sqrt{3}kV\_b a & \text{for} \quad G \sim 1\\ 2kV\_b a \sqrt{G} & \text{for} \quad G << 1 \end{cases} \tag{52}$$

The instability (52) under *G* � 1 has mixed mechanism: it is caused simultaneously (i) by a-periodical modulation of the beam density in media with negative dielectric constant and (ii) by excitation of the NEW. But the lower expression is the growth rate of instability caused only by excitation of the NEW of overlimiting e-beam. The presence of dissipation intensifies the growing of the slow beam wave. Instability turns to be of dissipative type with growth rate that again is inverse proportional to dissipation.

$$(\text{Im}\,\rho)^{(\nu)}\_{\text{ovl}} = \mathcal{Z}(ku)^2 \text{G}\alpha/\nu \sim \alpha\_b^2/\nu \tag{53}$$

However, the dependence on the beam density is completely different. This is the same DSI, which develops in uniform cross-section beam-plasma waveguide under over-limiting currents. Instabilities of the same type may be substantiated for finite thicknesses of the beam and plasma layers in waveguide. In this case one must use perturbation theory based on smallness of coupling coefficient.

As follows from this section, in the geometry of microwave plasma sources, the development of both new DSI is possible. Basic parameters of the both new DSI, and the conditions of their development should be taken into account upon design of the high power, high frequency plasma microwave devices.

## **6. Conclusion**

Thus, based on very general initial assumptions, we have found out that the number of DSI in the beam-plasma interaction theory is not limited by the only previously known type. Two new, previously unknown types of DSI are presented. The new DSI reveal themselves in the analysis of solution of the problem of initial perturbation development. This problem is classical in the theory of instabilities.

## *New Types of Dissipative Streaming Instabilities DOI: http://dx.doi.org/10.5772/intechopen.98901*

The first new DSI is the dissipative instability under weak beam-plasma coupling. In absence of dissipation the instability in these systems is caused by the interaction of the beam NEW with the plasma. With an increase in the level of dissipation this instability gradually transforms to the new type of DSI. Its maximal growth rate depends on the beam Langmuir frequency *ωb* and the frequency of collisions in plasma *ν* as *ωb=ν*. This, more critical (as compared to conventional), inverse proportional dependence on *ν* is a result of superposition of two factors those lead to growth of the beam NEW: weak coupling and dissipation.

The second new type of DSI is dissipative instability of over-limiting e-beam in uniform cross section waveguide. With increase in the beam current, its space charge and inner degrees of freedom reveal themselves more efficiently. If the beam current becomes higher than the limiting current in vacuum waveguide then the instability mechanism changes. In uniform cross section beam-plasma waveguide the instability becomes due to a-periodical modulation of the beam density in medium with negative dielectric constant. In this case the increase in the level of dissipation leads to one more new type of DSI with the maximal growth rate � *ω*2 *b=ν*.

The same types of DSI develop in systems having geometry, similar to microwave sources: cylindrical waveguide with thin annular beam and thin annular plasma. If the coupling between the beam and the plasma hollow cylinders is weak and the beam current is under-limiting the first type of DSI develops, but under over-limiting currents – the second. However, if the coupling of the beam and the plasma cylinders is strong, conventional type of DSI develops with well-known growth rate � *ωb=* ffiffi *ν* p .

Both new DSI are confirmed by conventional analysis of the respective DR.

Some words about the approach used. It has many advantages. First of all, it is based on very general initial assumptions and gives results regardless on geometry and specific parameters. The same approach is used for solving the same problem for conventional beam-plasma instabilities of all types (Cherenkov type, cyclotron type etc) [13], for the Buneman instability [22] etc. Obtained expressions for the spatial–temporal distribution of growing fields clearly show that with increase in the level of dissipation in background plasma, all these SI transform into DSI of conventional type. In addition, the analysis of obtained expressions gives much more detailed information on SI than other methods give. Part of the information on SI is not available in any other ways. The coincidence of other information to the results of conventional analysis confirms the validity of the approach (initial assumptions, mathematics etc).

Also, the presented approach shows that the DR describing the SI of given type can serve not only for solving of the initial/boundary problems and obtaining the dispersion curves. This point of view is very simplified. The approach shows that much more additional information is available from the DR. It, in fact, provides results on the initial perturbation development.

Summarizing, one can state that the presented approach can serve as an independent and very effective method for studying of any SI. There is no need to solve the problem again. One should only substitute the parameters of given instability in general expression for the field's space–time distribution. The usage of this approach instead of traditional initial/boundary problems gives complete picture of the instability development. At first glance, it might seem that this method of analyzing instabilities is more complicated. However, this complexity is only apparent. In addition, this complexity, if any, is overlapped by the completeness of the information received.

*Plasma Science and Technology*


**11. Voids**

*Welding - Modern Topics*

**10. Dendrites**

**Figure 14.**

**Figure 15.**

*Photo of the tombstone effect.*

Dendrites grow due to electrochemical migration when metal ions go into the electrolytic solution at the anode, plating out at the cathode and creating needle- or treelike formations on the PCB substrate (see photos in **Figure 16**) [16]. Electrochemical migration can be defined as the movement of a metal ion in an electrolytic solution

The time required to create dendritic bridges, which cause short circuits between two conductive paths, is determined by several factors including relative humidity,

between two neighbouring conductors with different electrical potential.

*Schematic representation of the wetting force (F1); upward push force from solvent vapours (F2).*

**128**

**Figure 16.** *Photo of dendrite.*

Voids are non-conductive cavities within the soldered joint, and their excessive presence poses a significant reliability risk for the manufactured product, especially in power electronics, where higher currents are present [20]. Voids are formed during the soldering process, their presence in the soldered joint causes the displacement of electrical and thermal paths, the resistance is higher and the temperature stress is non-uniform. This subsequently causes cracks and lowers the mechanical shock tolerance of the whole PCB.

With the advent of lead-free soldering (new material base, different temperature profiles, different types of fluxes, higher surface tension of solder, etc.), a higher incidence of voids was detected, leading to the lower reliability of soldered joints. Increased attention is therefore paid to the voiding issue.

Voids can be classified into several categories [21]. There are macro- and microvoids, shrinkage voids, voids in microvias, Kirkendall voids and pinhole voids (see **Figure 17**).

#### **11.1 Macrovoids**

Macrovoids are the most commonly occurring type of voids. Macrovoids are formed by the evaporation of gases from fluxes and soldering pastes during the reflow process. Macrovoids may occur anywhere in the solder joint, and their diameter is around 100–300 μm.

The factors that affect the size of the macrovoids are the solder paste's properties (particle size, composition, melting temperature and oxide content), flux (viscosity,

#### **Figure 17.**

*Soldered joint and the location of different types of voids. Intermetallic layer contains Kirkendall voids; pinholes are located on the boundary between the intermetallic layer and the PCB.*

surface tension, activator, solvent, etc.), components (geometry, shape, terminal oxidation, etc.) and the process (thickness, the shape and the parameters of the solder paste's printed layer, temperature profile) [22–26]. An example of macrovoids in the solder joints of a soldered BGA package on the PCB is given in **Figure 18**.