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\n
1. Introduction
\n
The cornea and the sclera are two conjugated quasi-spherical segments with unequal curvature radii; together they form corneoscleral (fibrous) tunic—the supporting structure of the eye capsule. Their mechanical properties play a crucial role of holding together the inner ocular structures. Despite them both being composed of connective tissue, they differ in physical (particularly, optical) and biomechanical properties [1].
\n
The cornea is the anterior part of the fibrous tunic of the eyeball, and it takes up 1/6 of its length. Despite it being relatively thin, its main function is protection—assured by its high durability. But the cornea also participates in light ray refraction, making up an important part of the visual apparatus; as such, it is characterized by high optical homogeneity and complete transparence.
\n
The cornea is an anisotropic, inhomogeneous structure; it mainly consists of highly specific connective tissue formed by parallel collagen fibrils that extend from one limb to another and act as load supporting elements [2].
\n
The sclera takes up the other 5/6 of the eye length and represents the posterior part of the fibrous tunic of the eyeball. Scleral tunic is the main supporting structure of the eyeball; it consists of dense collagen fibers. In contrast to cornea, the sclera has high dispersive power due to its chaotically distributed fibrils and fibers, which prevents light from entering the eye cavity from the side. In natural conditions, in the living eye, scleral elements are constantly in a strain-stress state determined by intraocular pressure and mechanical properties of the scleral tissue, as well as by anisotropy and inhomogeneity of these properties [3].
\n
Studying the biomechanical properties of the cornea is relevant for certain clinical needs associated with the appearance of new biomechanics examination methods, as well as the need to diagnose and monitor ectatic diseases of the cornea, to adequately select the parameters for keratorefractive surgeries, to correctly interpret the intraocular pressure (IOP) values, and, consequently, to appropriately assess IOP and monitor glaucomatous process.
\n
In addition, conducting studies on the biomechanical properties of the sclera is a necessary step in the research of pathogenic factors relevant for occurrence and progression of myopia, as well as finding effective means and methods of influencing the sclera in order to correct its biomechanical state.
\n
However, the lack of standardized terminology and uniform classification hurts the ability to compare research results and consequently hinders their introduction into the knowledge area of ophthalmology.
\n
\n
\n
2. Classification of approaches to study the biomechanics of the eye
\n
In accordance with different approaches, eye biomechanics can be divided into the following types:
theoretical;
physical (i.e., experimental); and
clinical.
\n
\n
2.1. Theoretical biomechanics of the eye
\n
Theoretical biomechanics is a science that employs mathematical methodology and mathematical analysis. As applied to ophthalmology, it handles with specific physical constants characterizing elasticity, strength, and other mechanical parameters of the tissues (usually measured in vitro).
\n
The main theoretical approach is mathematical modeling. The research may target separate structures of the eyeball and the tunic, or the eye in its entirety. It can also include modeling of physiological or pathological processes, changes induced by specific stimuli or effects of surgical treatment.
\n
The results obtained from modeling can be used in experimental and clinical studies. In turn, all models are based on the figures acquired in experiments or from clinical diagnostic.
\n
Disadvantages of the theoretical approach in studying eye biomechanics are associated with structural complexity of the eyeball, inhomogeneity, and variability of morphology of the ocular structures, and dependence on the technological advancement of experimental and clinical research methods.
\n
\n
\n
2.2 Physical (experimental) biomechanics of the eye
\n
Experimental biomechanics of the eye is based on studying individual tissues and the eyeball as a whole in vitro or by conducting animal experiments using physical methods. It is the most developed subdiscipline of biomechanics with many years of research history. Capabilities of the approach are limited by post-mortem changes in eye tissues, and anatomical and physiological differences between humans and animals. The main purpose of experimental studies is to find potentially useful methods of studying biomechanical properties of eye tissues in clinical environment and to acquire data for mathematical modeling. The main advantages of experimental researches are the absence of restrictions for employed methods and approaches, and the choice of which is only limited by technological and scientific advancement.
\n
Methods of experimental biomechanics allow measurement of a big number of physical parameters of the cornea:
Young’s modulus (E),
Poisson’s ratio (μ),
Durability (σ),
Deformation capacity (Σ), etc.
\n
However, they do not fully reflect the properties of fibrous tunic of the eye in vivo.
\n
\n
\n
2.3. Clinical biomechanics of the eye
\n
Clinical biomechanics of the eye studies the influence of biomechanical properties of the fibrous tunic on the results of diagnostics, development, and treatment of various eye diseases. Clinical biomechanics operates on data obtained with specialized examination methods used in ophthalmology (in vivo) that characterize biomechanical properties of the fibrous tunic. Its research subject is strictly the eyeball as a whole, only allowing arbitrary delineation of the internal structures. This complicates the interpretation of data. However, in order to improve diagnostics and treatment of eye diseases, clinical methods for eye biomechanics need to have higher priority in research and development.
\n
The following corneal parameters can be measured clinically:
Friedenwald’s rigidity coefficient;
Corneal hysteresis;
Corneal resistance factor;
Coefficient of elasticity;
Corneal deformation.
\n
The biggest number of already existed studies are dedicated to investigation of biomechanical properties of the cornea, which is probably related to the specifics of corneal structure, or to its accessibility for examination.
\n
\n
\n
\n
3. Theoretical (mathematical) biomechanics
\n
The originator of mathematical approach to study biomechanical properties of the cornea was F.A. Rachevsky. In 1930, in his theoretical study, he pointed “…for the first time at the paramount importance of the radius of the corneal curve and especially of its thickness for specific results of intraocular pressure tonometry.” Besides that, he proved mathematically that under effect of external and internal forces, tangentially directed stress occurs in the cornea, particularly during applanation tonometry [4].
\n
At present, the research of corneal biomechanics is conducted in two main directions. Mathematical modeling is generally used for the calculation of parameters and prediction of results of keratorefractive surgeries [5, 6, 7, 8], as well as for the determination of possible procedural errors of applanation tonometry methods when biomechanical properties changed as the result of a surgery or a disease [9].
\n
The main obstacle for proper mathematical modeling is anisotropy of the cornea. The majority of the proposed models does not consider it, which limits their application in practical ophthalmology [10, 11, 12].
\n
According to Pinsky et al., the anisotropy of the cornea primarily depends on its structural features, that is, specific architectural organization of collagen fibers [13]. X-ray structural analysis revealed that collagen fibrils of the central area have orthogonal orientation predominantly in vertical and horizontal directions, while fibrils of the periphery have tangential orientation [14]. Pinsky et al. developed a mathematical model for corneal anisotropy mechanics that accounts for these findings [13]. Based on the finite element method, the model allows predicting biomechanical response of the cornea to tunnel cutting, radial keratotomy, and LASIK [15, 16, 17].
\n
In order to determine the possible error margin of applanation tonometry methods, several mathematical models have been developed [18]. Liu et al. used mathematical modeling to study isolated effects of various biometric and biomechanical parameters of the cornea on Goldmann tonometry readings [19]. Kwon et al. developed a mathematical model demonstrating the need to take into account not only corneal thickness, but also its biomechanical properties when interpreting Goldmann tonometry readings [20].
\n
\n
\n
4. Physical (experimental) biomechanics
\n
\n
4.1. Normal (intact) cornea
\n
Experimental studies based on extensiometry revealed distinguishing biomechanical anisotropy and heterogeneity of the cornea. Corneal material acquired with a radial cut has the best durability and margin of deformation capacity. Those parameters decrease with distance from the radial direction. Corneal material stretched tangentially shows approximately the same elastic properties along the corneal disc. The samples stretched radially appeared to have the highest rigidity. In the course of the study, Poisson’s ratio was determined for various parts of the cornea. This ratio characterizes the transverse deformation lateral to stretch direction, for radial direction, it was in the range of 0.445–0.450, and for tangential direction, it was in the range from 0.290 to 0.310 (middle periphery) and from 0.340 to 0.350 (perilimbal) [21, 22].
\n
A variety of studies is dedicated to measuring the main elastic and strength properties of the cornea, but analysis of the data shows that isolated corneas exhibit big spread in the values—from 0.3 to 13.6 MPa. The phenomenon can be attributed to different experimental conditions and nonlinear nature of biomechanical properties of corneal material [23, 24, 25, 26]. Andreassen et al. studied the biomechanics of corneal discs with extensiometry; the discs were taken from patients with keratoconus after they underwent penetrating keratoplasty. The study revealed significant decrease of mechanical strength properties in pathologically altered corneas [27].
\n
Soergel et al. used dynamic mechanical spectroscopy to evaluate viscoelastic properties of the cornea in experimental environment. They found that elastic and shearing deformation depend on the hydration, time elapsed after death, and temperature of the tissue [28].
\n
Wang et al. calculated Young’s modulus by measuring the speed of ultrasound transmission through cadaver cornea and processing the data with Fourier analysis [29].
\n
Like ultrasound spectroscopy, Brillouin microscopy can determine intrinsic viscoelastic properties decoupled from the structural information and applied pressure. In contrast, it can measure the local acoustic properties with much higher spatial resolution and sensitivity, and the measurement is performed optically without the need for acoustic transducers or physical contact with the cornea [30].
\n
One of the techniques, holographic interferometry, is used to calculate Young’s modulus. The method is to some degree similar to videokeratography, that is, holographic technologies are used to examine the changes in corneal surface. A study conducted on an intact bull’s cornea showed Young’s modulus being two orders lower than when measured in an experiment with corneal tissue samples. The authors summarized that localization and hydration level plays the primary roles during measurement. This method, however, is limited in terms of practical use due to requiring maximum permissible laser emission in order for the resulting images to be high quality [31].
\n
\n
\n
4.2. Cornea after refractive surgery
\n
Some studies showed significant increase of tangential elasticity of the cornea after it was incised with radial cuts (up to 46.5% with an incision depth of 0.6 mm), that is, in the direction of the lesser material rigidity [32]. In certain cases, the changes led to severe complications in the long-term postoperative period. Particularly, it manifested as a significant decrease of eyeball’s resistance to trauma with potential disruption of corneal cicatrices and loss of membranes [33].
\n
Luminescent polariscopy revealed that after radial keratotomy, the main mechanical strain fell on the middle periphery of the cornea, particularly on the bottom of keratotomic incisions. An increase of intracameral pressure (analogue to intraocular pressure) raises the strain on peripheral part of the cornea and off-loads its central part, which can cause hypermetropic shift in refraction [34].
\n
However, with the appearance and widespread implementation of excimer laser technologies for correction of refraction errors, such risks have greatly decreased. It can be attributed to different mechanisms of corneal refraction change, that is, its thinning in the central area.
\n
Experimental studies on biomechanical properties of the cornea after excimer laser intervention indicate that thinning of the cornea in 6.0-mm optic zone for more than 15–20% results in significant changes of its mechanical properties. In terms of clinical relevance, the most meaningful change appears to be the significant (mean 20%) decrease of breaking load for experimental samples in comparison to the control samples. Additionally, changes in deformation properties of the cornea after laser ablation should also be taken into consideration, which manifested as lowered amount of movement the punch had to do before the cornea broke in experimental eyes in comparison to the control subjects in average by 10.72% [35].
\n
However, the mechanical properties of the data obtained using an isolated cornea cannot objectively reflect the parameters of the tissues in natural environment. Adequate information on the biomechanical state of the cornea can only be obtained from a living eye.
\n
\n
\n
\n
5. Clinical biomechanics
\n
\n
5.1. Normal (intact) cornea
\n
Clinical studies on the biomechanical properties of the relatively healthy cornea have been conducted since the middle of the twentieth century, but those methods remained widely unused due to various reasons.
\n
In 1937, Friedenwald suggested that rigidity coefficient could be calculated based on a logarithmic dependence between IOP changes and eye volume employing differential tonometry with Schiotz tonometer [36]. Friedenwald depicted the relation between pressure and volume in a coordinate system. As was shown by further clinical studies, the proposed coefficient strongly depends on the corneal curvature and thickness, as well as on the IOP level [37]. According to research results, the parameter suggested by Friedenwald—the rigidity coefficient—was inaccurate in eyes with deviations in biomechanics (thickness and curvature) from the norm. It was also strongly influenced by IOP.
\n
In 1936, S. F. Kalfa proposed a method of elastotonometry, that is, differential tonometry with four Maklakov tonometers weighing 5, 7.5, 10, and 15 g. Connecting the dots marked on a coordinate system forms an elastotonometric curve, which appears ascending line. The difference in mm Hg between the starting and ending points of the curve, that is, between IOP value obtained using 5.0 and 15.0 g tonometers, is called elasto-ascent. Essentially, Friedenwald’s rigidity coefficient and S. F. Kalfa’s elasto-ascent are different expressions of the same thing. In norm, the two figures are closely related, albeit not functionally [38].
\n
There are a number of techniques described by their authors as potential intravital methods for examination of biomechanical properties of the cornea, but they have not been adopted into clinical practice: electronic speckle interferometry [39], dynamic cornea visualization [40], corneal applanation and indentation [41], ultrasound elastometry [42], and photoelasticity method [43].
\n
As an alternative to holographic interferometry, a noncontact, nondisruptive method of electronic speckle interferometry was suggested; it is equally sensitive because it employs close wavelength for measurement. Advantages of the method include the absence of requirement of photographic hologram recording, which simplifies the procedure and enables real-time acquisition of corneal surface shift data using a television camera. The method is recommended for evaluation of changes in corneal biomechanics after excimer laser refractive surgery [39].
\n
Grabner et al. proposed a method of dynamic visualization of the cornea. It involves applying dosed pressure to the central area of the cornea during videokeratography by means of a special indenter and subsequent analysis of the topographic pattern. As the result, high correlation between the bending curve and depression depth was found. The form of the curve was noted to be affected by central corneal thickness, intraocular pressure, and patient age. Moreover, bending curves were different in keratoconus patients, as well as in patients who had underwent keratorefractive surgeries [40].
\n
Chang et al. studied biomechanical properties of the cornea in vivo using corneal applanation and indentation on rabbit and human eyes, regarding the cornea as a transversely isotropic material. The study showed normal Young’s modulus to vary from 1 to 5 MPa and transverse shift modulus from 10 to 30 KPa [41].
\n
Some authors used photoelasticity method to evaluate mechanical stress in the cornea involving the measurement of its polarization and optical properties [43].
\n
Scoping a large amount of clinical data, Edmund calculated Young’s modulus adhering to the hypothesis that the final form of cornea is the outcome of counteraction between tissue elasticity and intraocular pressure. The modulus values were significantly lower in keratoconus eyes when compared to norm. The study also showed significant difference between healthy and ectatic patients in relative distribution of stress in the central and peripheral areas of the cornea, which can help with the understanding of keratoconus pathogenesis. However, this method generally does not find much use in clinical practice [44].
\n
The one method most widely used in present day clinical practice involves ocular response analyzer (ORA)—a device that analyses corneal biomechanical properties based on bi-directional corneal applanation by an air pulse [45]. The method’s authors proposed to evaluate biomechanical response of the cornea by quantifying the differential inward and outward corneal response to an air pulse and thus obtaining two parameters—corneal hysteresis (CH) and corneal resistance factor (CRF). Corneal hysteresis characterizes the viscoelastic properties of the cornea responsible for the partial absorption of the air pulse energy. Corneal resistance factor is a derived parameter with high correlation to central corneal thickness that reflects the elastic properties of the cornea.
\n
Multiple studies have confirmed the usefulness of bi-directional corneal applanation for the evaluation of biomechanical properties of the cornea: they rise with the increase of the corneal thickness [46, 47]. Corneal hysteresis was in the average 10.8 ± 1.5 mm Hg and corneal resistance factor—11.0 ± 1.6 mm Hg. Statistically, a significant difference in the mean values of CH and CRF between groups of varying age was absent, with the exception of patients older than 60 years for whom the values were on lower. It is possible that the phenomenon reflects the changes in elastic properties of the cornea associated with age, but the authors also note the potential influence of other parameters (intraocular pressure and anterior-posterior axis length) that were disregarded in the study [48]. The comparison of CH and CRF in children and adults did not reveal any age-related differences [49].
\n
Studying the diurnal variations in CH and CRF parameters revealed their hourly stability, while minor changes observed between the morning and evening measurements can be explained by diurnal IOP fluctuations [50]. CH and CRF correlated strongly with corneal thickness and to a lesser degree with an amount of astigmatism. No correlation was found with keratometry, age, gender, spherical equivalent, or IOPcc [51]. Moreover, ORA shows good repeatability of biomechanical and tonometry measurements [52].
\n
Avetisov et al. studied the possibility of applying the dynamic pneumo-impression of the cornea approach to the existing corneal biomechanical properties analyzer (ORA). The fundamental principle was that at the curvature start point laying on the border of the impression area, the pneumatic jet is subject to the counter-force of IOP and corneal elasticity, in equal amounts. At the moment of maximum impression, the pneumatic jet is mainly countered by corneal elasticity—due to the maximum deformation of the cornea. As the result, a parameter named elasticity coefficient was calculated characterizing the elastic behavior of the cornea regardless of the IOP level [53].
\n
The same principle was used in CorVis device (Oculus, Germany), in which corneal deformation responding to a pulse of air is monitored with high-speed Scheimpflug camera. The device can help to measure a whole range of parameters that characterize the particularities of corneal deformation during the impression process. It records the process between the initial and the second applanations involving the cornea recovering its initial form, captures the maximum indentation point, and measures IOP [54, 55].
\n
\n
\n
5.2. Keratoconic cornea
\n
Intravital measurement of biomechanical properties of the cornea in keratoconus patients performed with dynamic bi-directional pneumo-applanation showed lower CH and CRF values than in healthy eyes. Apparent negative correlation between the CH and CRF parameters and the degree of keratoconus were also evident [56].
\n
Additionally, CH was significantly higher than CRF in the keratoconus group. The authors suggested the CH decrease of less than 8 mm Hg in conjunction with positive CH-CRF difference to be considered a stronger sign of keratoconus than isolated decrease of CH. Glaucoma patients showed reverse tendency: CRF value was higher than CH [57].
\n
Studying the parameters obtained with dynamic Scheimpflug analysis (Covis ST) showed the possibilities of the examination method for differential diagnostics of patients suspected of keratoconus or with early keratoconus from patients with normal cornea [58].
\n
\n
\n
5.3 Cornea after refractive surgery
\n
Intravital measurements of biomechanical properties of the cornea after excimer laser surgery performed using dynamic bi-directional pneumo-applanation also confirmed the loss of corneal strength. In patients who had underwent LASIK, examination showed decrease of IOP-related parameters such as corneal compensated IOP, as well as parameters reflecting the biomechanical properties. Along with that, significant correlation was observed between the amount of myopia correction and the deterioration of the biomechanical properties [59].
\n
Another study analyzed the results of dynamic bi-directional pneumo-applanation of the cornea and assessed the correlation between CH decrease and ablation depth in three patient groups: after photorefractive keratectomy, after LASIK with mechanical corneal flap creation, and after LASIK with femtosecond flap creation. The authors found that the strongest correlation was present in femto-LASIK group, while in the two other groups, it was significantly lower [60].
\n
Isolated creation of corneal flap was also found to cause minor changes in corneal refraction [61]. Roberts explained the phenomenon with a theory stating that after lamellar dissection, the corneal biomechanics change in such a way so that severed fibrils contract causing traction in the direction of limbus. With that, central corneal area deflates under the action of released fibrils inducing the so-called “hypermetropic” shift [62, 63].
\n
In parallel, a comparison of changes in biomechanical properties of the cornea after superficial and intrastromal keratectomy was done using OCULUS Corvis (Germany) tonometer. Both types of keratectomy were found to cause statistically significant decrease of biomechanical parameters [64].
\n
Despite the existing methods of measuring biomechanical properties of the cornea and the developed biomechanical models, the detection of ectasia after excimer laser vision correction varies from 0.04 to 0.6% of cases, but according to some researchers, the numbers may be an underestimation [65, 66].
\n
Iatrogenic keratectasia is known to develop due to two factors: an ectatic corneal disease that was undiagnosed in the preoperative stage and excessive thinning of the cornea [67]. In the first case, early detection of keratoconus poses objective challenges [68, 69].
\n
At the same time, even when keratoconus was timely diagnosed, the selection of candidates for keratorefractive surgeries is still difficult, and the evaluation of corneal biomechanics by means of dynamic bi-directional pneumo-applanation does not yield the necessary data.
\n
\n
\n
\n
6. Correction of corneal biomechanical properties
\n
Presently, the most common method of correcting (strengthening) biomechanical properties of the cornea is corneal cross-linking [70].
\n
The first specialists who in the 90s of the twentieth century created corneal cross-linking method for treating keratoconus were Wollensak, Spoerl, and Seiler [71]. They developed the protocol (“Dresden protocol”) for using this method of strengthening the cross-link bonds of collagen for treating progressive keratoconus involving riboflavin and ultraviolet A irradiation of the corneal stroma (UVA with a wavelength of 370 nm for peak absorption of riboflavin) [72].
\n
Careful preclinical experimental validation showed that the combination of riboflavin and UVA leads to a significant improvement of biomechanical stability of the cornea (increase of elastic modulus approximately by 300%) and the formation of large collagen molecular aggregates, including the appearance of cross links—predominantly between the fibril surface molecules and also between proteoglycans in the interfibrillar space [73, 74, 75].
\n
In the following decades, the corneal cross-linking technique has seen widespread clinical application with indications for its usage expanding significantly. Effectiveness of the method for strengthening biomechanical properties was confirmed for the treatment of not only progressive keratoconus, but also pellucid marginal degeneration and iatrogenic ectasia caused by excimer laser surgery [76].
\n
An important suggestion has been made recently for reinforcing the effect of corneal cross-linking—to combine the procedure with implantation of corneal segments [77, 78]. Comparative studies of different treatments—individually and in combination—showed the most pronounced effect to be from the combination therapy starting with the implantation of corneal segments and followed by cross-linking, and not in the reverse order. Such combination therapy also helps to achieve better results (weakening of manifest refraction and keratometric indicators) in cases with keratectasia after excimer laser surgery [79].
\n
There is another method described in the literature as directed laser ablation; it involves biomechanical approach to ablation calculations. Vaporization of the tissue thus happens on the middle periphery, which contains certain relatively flat spots, and not in the central (thin) area. The rationale is that thinning of the area leads to steepening of the cornea subsequently flattening the unablated area, which has more optical power [80]. It should be noted that in clinical practice, this method requires very careful consideration and cautiousness due to insufficient studies on its after effects.
\n
Furthermore, a multimodal approach involving implantation of intrastromal rings, CXL, and laser ablation in different configurations may provide not only stability of corneal topography, but also positive refraction result, thanks to the combination of all the methods’ advantages [81, 82, 83, 84, 85]. That said, the lack of established standards and clinical recommendations for combining different methods for the correction of corneal biomechanical properties may lead to various complications and unexpected aftermaths; it should be kept in mind when planning such treatment.
\n
\n
\n
7. Conclusion
\n
In summary, clinical relevance of studying biomechanics of the fibrous tunic is difficult to overestimate. The diversity of methods used for examination of biomechanical properties of the cornea means there is no single method that could fully satisfy the needs of practical ophthalmology. Further studies are necessary to develop simple, available, and sufficiently informative method for clinical assessment of ocular biomechanics. Moreover, the demand for techniques of correcting biomechanical properties keeps growing, and so this field of research has wide potential.
\n
\n
Conflict of interest
The author has no conflict of interest.
\n',keywords:"corneal biomechanics, refractive surgery, LASIK, keratokonus, IOP",chapterPDFUrl:"https://cdn.intechopen.com/pdfs/63146.pdf",chapterXML:"https://mts.intechopen.com/source/xml/63146.xml",downloadPdfUrl:"/chapter/pdf-download/63146",previewPdfUrl:"/chapter/pdf-preview/63146",totalDownloads:937,totalViews:169,totalCrossrefCites:0,totalDimensionsCites:0,totalAltmetricsMentions:0,impactScore:0,impactScorePercentile:44,impactScoreQuartile:2,hasAltmetrics:0,dateSubmitted:"April 29th 2018",dateReviewed:"July 14th 2018",datePrePublished:"November 5th 2018",datePublished:"January 30th 2019",dateFinished:"August 21st 2018",readingETA:"0",abstract:"Knowledge of biomechanical properties of eye globe is necessary both for correct selection of candidates for refractive surgery and right choice of operative intervention parameters. No less important, it is for corneal ectatic disease diagnostics and monitoring. Also it gives inestimable contribution for interpretation of intraocular pressure (IOP) indices especially in cases with irregular eye shape or after past corneal surgical procedures. Moreover, it allows studying injury mechanism by glaucoma process on optic nerve head fibers. Above it, scleral biomechanical properties research is necessary for the investigation of pathophysiologic factors of myopia manifestation and progression. This chapter is devoted to review of existed to date methods of study of eye fibrous tunic biomechanical properties. It describes mathematical, experimental, and clinical models provided evaluation of unsearchable by direct measurement parameters. It also observes effective technics of impact on both sclera and cornea with the aim of correction of its biomechanical condition.",reviewType:"peer-reviewed",bibtexUrl:"/chapter/bibtex/63146",risUrl:"/chapter/ris/63146",book:{id:"6843",slug:"biomechanics"},signatures:"Irina Bubnova",authors:[{id:"256707",title:"Ph.D.",name:"Irina",middleName:null,surname:"Bubnova",fullName:"Irina Bubnova",slug:"irina-bubnova",email:"bubnova.irina@gmail.com",position:null,profilePictureURL:"//cdnintech.com/web/frontend/www/assets/author.svg",institution:null}],sections:[{id:"sec_1",title:"1. Introduction",level:"1"},{id:"sec_2",title:"2. Classification of approaches to study the biomechanics of the eye",level:"1"},{id:"sec_2_2",title:"2.1. Theoretical biomechanics of the eye",level:"2"},{id:"sec_3_2",title:"2.2 Physical (experimental) biomechanics of the eye",level:"2"},{id:"sec_4_2",title:"2.3. Clinical biomechanics of the eye",level:"2"},{id:"sec_6",title:"3. Theoretical (mathematical) biomechanics",level:"1"},{id:"sec_7",title:"4. Physical (experimental) biomechanics",level:"1"},{id:"sec_7_2",title:"4.1. Normal (intact) cornea",level:"2"},{id:"sec_8_2",title:"4.2. Cornea after refractive surgery",level:"2"},{id:"sec_10",title:"5. Clinical biomechanics",level:"1"},{id:"sec_10_2",title:"5.1. Normal (intact) cornea",level:"2"},{id:"sec_11_2",title:"5.2. Keratoconic cornea",level:"2"},{id:"sec_12_2",title:"5.3 Cornea after refractive surgery",level:"2"},{id:"sec_14",title:"6. Correction of corneal biomechanical properties",level:"1"},{id:"sec_15",title:"7. Conclusion",level:"1"},{id:"sec_19",title:"Conflict of interest",level:"1"}],chapterReferences:[{id:"B1",body:'Iomdina EN, Bauer SM, Kotliar KE. Eye Biomechanics: Theoretical Aspects and Clinical Applications. Moscow: Real Time; 2015. p. 208\n'},{id:"B2",body:'DelMonte DW, Kim T. Anatomy and physiology of the cornea. 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DOI: 10.1016/j.jcrs.2009.05.033\n'},{id:"B78",body:'Kılıç A, Kamburoglu G, Akıncı A.Riboflavin injection into the corneal channel for combined collagen crosslinking and intrastromal corneal ring segment implantation. Journal of Cataract & Refractive Surgery. 2012;38(5):878-883\n'},{id:"B79",body:'Coskunseven E, Jankov MR II, Hafezi F, Atun S, Arslan E, Kymionis GD. Effect of treatment sequence in combined intrastromal corneal rings and corneal collagen crosslinking for keratoconus. Journal of Cataract & Refractive Surgery. 2009;35(12):2084-2091\n'},{id:"B80",body:'Cennamo G, Intravaja A, Boccuzzi D, Marotta G, Cennamo G. Treatment of keratoconus by topography-guided customized photorefractive keratectomy: Two-year follow-up study. Journal of Refractive Surgery. 2008;24(2):145-149\n'},{id:"B81",body:'Stojanovic A, Zhang J, Chen X, Nitter TA, Chen S, Wang Q. Topography-guided transepithelial surface ablation followed by corneal collagen cross-linking performed in a single combined procedure. Journal of Refractive Surgery. 2010;26(2):145-152. DOI: 10.3928/1081597X-20100121-10\n'},{id:"B82",body:'Krueger RR, Kanellopoulos AJ. Stability of simultaneous topography-guided photorefractive keratectomy and riboflavin/UVA cross-linking for progressive keratoconus. Journal of Refractive Surgery. 2010;26(10):S827-SS32. DOI: 10.3928/1081597X-20100921-11\n'},{id:"B83",body:'Kamburoglu G, Ertan A. Intacs implantation with sequential collagen cross-linking treatment in postoperative LASIK ectasia. Journal of Refractive Surgery. 2008;24(7):S726-S7S9\n'},{id:"B84",body:'Kanellopoulos AJ. Comparison of sequential vs same-day simultaneous collagen cross-linking and topography-guided PRK for treatment of keratoconus. Journal of Refractive Surgery. 2009;25(9):S812-S8S8. DOI: 10.3928/1081597X-20090813-10\n'},{id:"B85",body:'Kymionis GD, Kontadakis GA, Kounis GA, Portaliou DM, Karavitaki AE, Magarakis M, et al. Simultaneous topography-guided PRK followed by corneal collagen cross-linking for keratoconus. Journal of Refractive Surgery. 2009;25(9):S807-SS11. DOI: 10.3928/1081597X-20090813-09\n'}],footnotes:[],contributors:[{corresp:"yes",contributorFullName:"Irina Bubnova",address:"bubnova.irina@gmail.com",affiliation:'
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1. Introduction
Titanium aluminide (TiAl) is a member of group material referred to as intermetallics, consisting of various metals resulting in ordered crystallographic structures formed when the concentration of the alloy exceeds the solubility limit [1]. Properties as low density, high strength and elevated temperature properties make TiAl replacement candidates for nickel-based superalloys used in the aerospace and automotive industries [2, 3, 4]. One such alloy tried and tested by General Electric [5] for commercial turbofan engines is Ti-48Al-2Cr-2Nb. Despite the attractive high-temperature properties attained in research to date, the inherent poor ductility of TiAl at ambient temperatures remains a concern [6]. Over the past 20 years and recently, much work has been devoted to material tailoring through compositional variations and alloying aimed at improving room temperature ductility [7, 8, 9, 10, 11].
Phase evolution in TiAl alloys governs the mechanical and physical properties to be obtained. Primarily, two ordered structures exist, namely, γ-TiAl (L10) and hexagonal α2-Ti3Al (D019), resulting from different thermo-mechanical treatments. Furthermore, the mechanical properties to be obtained are dependent on the microstructure. Three microstructures exist, namely, equiaxed single γ phase, fully or near (γ/α2) lamellar and duplex (consisting of colonies of lamellar γ/α2 and pure γ phase grains). The achieved microstructure is significant for its mechanical properties, especially in structural applications. Duplex microstructures with enhanced ductility measures such as fracture strength, yield strength and strain have been reported [12, 13, 14]. Fully lamellar structures, in particular, have shown the best creep performance as contrasted to other microstructural modifications [15, 16, 17].
For the intended application, considering the inherent brittle nature of TiAl alloys, material tailoring through microstructural evolution is often necessary. Additionally, the low ductility and brittleness of TiAl alloys at ambient temperatures make their processing using conventional methods difficult. To overcome problems associated with conventional processing, such as microstructural inconsistencies inherited from solidification and phase evolutions resulting in the scattering of mechanical properties, heat treatment cycles are often designed [18, 19, 20, 21]. Traditional methods requiring post-treatment are time-consuming, labour and capital intensive, waste a lot of start-up material, and require unnecessary production costs. Therefore, there is a need to manufacture TiAl alloy components without the above-mentioned technical deficiencies and limitations and satisfy industrial needs for component fabrication [22].
For the last decades of the 20th century [23], the Additive Manufacturing (AM) method has been employed to obtain objects by the subsequent material supply. AM mainly aims to complete a collection of traditional subtractive manufacturing practices while avoiding and limiting the need for post mechanical processing such as machining. Laser powder bed fusion (L-PBF) is an AM technique, historically referred to as Selective Laser Melting (SLM) developed by F&S Stereolithographietechnik GmbH with Fraunhofer ILT [24], where a component is manufactured by melting a powder bed in a layer-by-layer sequence employing laser beam irritation [25]. The L-PBF process is initiated by creating a 3D digital part model (usually scan data or a CAD file), followed by slicing the model into thin layers using special software. The powder bed is achieved by spreading powder onto the substrate surface. The powder bed is selectively melted through cross-sectional scanning generated from the 3D part model by applying a laser beam. After cross-section scanning, powder bed layering is achieved by sequentially adding layers one after the other repeatedly until the part is complete. Recent studies [25, 26, 27, 28, 29] have shown that L-PBF is an innovative and efficient process employed to manufacture TiAl alloys compared to historically employed traditional manufacturing processes such as casting [30, 31, 32], ingot metallurgy [33, 34, 35], or even solid-state powder sintering [36, 37, 38]. The benefits of L-PBF include short production cycles and cheaper production costs. Also, parts produced are of high quality and have been found to exhibit desirable performance [39].
Exploring AM technologies to improve on properties of TiAl and its alloys is essential. As such, mechanical properties like compressive and tensile ductility measures [40, 41, 42], wear resistance [43, 44], elevated temperature creep and oxidation resistance [45, 46, 47, 48] superior to those processed by conventional means have been reported. Operation temperatures in new-generation gas turbines have fast-tracked progress in material development in the aerospace industry.
The dual combination of high temperatures and contaminant-containing aircraft environments shifts focus to hot corrosion and oxidation. Hot corrosion and oxidation can lead to catastrophic failures through material consumption at an unpredictably rapid rate. Much work has been devoted to understanding the hot corrosion and oxidation of TiAl alloys already [49, 50, 51, 52, 53]. As such, this research paper serves as a summary of the laser additive manufacturing of TiAl alloys. Particular attention is also given to the mechanisms, kinetics, prevention control and recent developments in hot corrosion and oxidation of TiAl alloys.
2. Titanium aluminide (TiAl) alloys: phase, microstructures and mechanical properties
2.1 Phase and microstructural evolutions
2.1.1 Phase evolutions
The three main phases of the Ti-Al system consist of various TiAl compounds, namely, γ-TiAl, α2-Ti3Al and TiAl3 [1]. Of the three phases, only γ-TiAl and α2-Ti3Al have shown to be of engineering significance [54] with outstanding properties. They are lightweight and can be implemented for structural parts, automotive and elevated temperature aerospace applications. The γ-TiAl phase is a face-centred tetragonal ordered phase with an L10 structure. It consists of atomic layers at 90° to the c-axis [55] with lattice parameters a = 0.4005 nm, c = 0.4070 nm and a tetragonality ratio (c/a) of 1.02 [56, 57]. The compositional range of the γ-TiAl phase is from 48.5 to 66.0 at.% of Al. The α2-Ti3Al phase has a hexagonal DO19 structure with a compositional range from 20 to 38.2 at.% of Al.
The α2-phase has high hydrogen and oxygen absorption rates and suffers from severe embrittlement, though it exhibits optimum high-temperature strength. The γ-phase has low gaseous absorption rates, outstanding oxidation resistance and poor room-temperature ductility. To maximise engineering benefits, dual-phase TiAl alloys consisting of γ + α2 phase are used. These alloys show excellent ductility [13, 58] at room temperatures due to the availability of refined lamellar colonies aiding γ-phase deformation [54, 59, 60]. The most known dual TiAl alloys with outstanding tensile properties are referred to as duplex alloys of the nominal (at.%) composition of Ti-(46–49) Al.
The four significant microstructures which may result in a Ti-Al system are namely, duplex (DP), near-gamma (NG), nearly lamellar (NL) and fully lamellar (FL). The obtained microstructures are greatly dependent on the processing route, Al compositional variations and thermo-mechanical treatments employed. Of the four, only fully lamellar and duplex have been considered necessary in engineering applications [54]. The evolutions (in Figure 1) of the microstructures mentioned above were be summarised in works by Cobbinah et al. [6] and Clemens et al. [61].
Figure 1.
The central portion of the binary Ti-Al phase diagram together with microscopic optical (left) and backscattered scanning electron (right) images showing NG, DP, NL/NLγ and FL microstructures achieved via heat-treating within α and (α + γ) phase-field. The phases obtained are identified using contrast, where a light contrast is representative of α2-Ti3Al and γ-TiAl of a darker contrast [61].
NG microstructures are obtained via thermal treatments slightly above the eutectoid temperature (Teu), while DP microstructures are achieved between Teu and α-transus temperatures. The thermal treatment implemented significantly affects the volume fraction of lamellar grains present. As a result, NL microstructures are obtained at (Teu) and Tα relative temperatures, slightly under Tα. NL microstructures exhibiting a specified globular γ-grain volume fraction are shown as NLγ. FL microstructures are achieved by thermal treatments above Tα. Generally, the obtained properties compensate for other properties [22] as represented in Figure 1 and should be considered when the material is designed for structural applications. Furthermore, the microstructure-property relationship in TiAl alloys makes it easier to modify the material for the anticipated application.
3. Additive manufacturing (AM) of TiAl
3.1 Process overview
Additive manufacturing (AM) presents an opportunity to manufacture TiAl alloys with minimal processing difficulties compared to those experienced during conventional processing, such as near-net-shape forging or investment casting [62]. For tailoring TiAl alloys with optimum properties, laser powder bed fusion (L-PBF) and electron beam melting have been considered suitable [63, 64, 65, 66]. Recently, the production of TiAl alloys using L-PBF has gained special attention [29, 67, 68, 69, 70, 71] owing to the benefits offered. Some of these benefits [6] complex geometry formation, ease of part dimension control, production of highly defined parts with orifices, mass customisation and material flexibility. Furthermore, during local melting of the powders, high solidification rates are obtained. These result in more refined microstructures.
The component is manufactured (in Figure 2) by melting a powder bed in a layer-by-layer sequence employing laser beam irritation [25]. The process is initiated by creating a 3D digital part model (usually scan data or a CAD file), followed by slicing the model into thin layers using special software. The powder bed is achieved by spreading powder onto the substrate surface. In preparation for part manufacturing, powders are preheated below their melting temperatures to promote bonding and minimise distortion [6]. L-PBF part manufacturing is executed in an inert gas (preferably argon) sealed environment to prevent reactive powder oxidation.
Figure 2.
Graphical representation of laser powder bed fusion method [72].
3.2 Research milestones
The need to replace previously used Ni-based superalloys in aerospace components has fast-tracked research and development of lightweight and cost-efficient TiAl alloys. To date, Ni-based alloys still outmatch TiAl alloys in fabrication costs and mechanical performance. This is mainly due to the poor room temperature ductility of TiAl alloys and the delay in engineering design practices for low ductility materials [54]. Additionally, the high part fabrication costs involved in producing TiAl alloys are related to the knowledge that low ductility fabrication processes, which also produce high melting point alloys, are unavailable. As such, there has been much investment in exploring complex part fabrication techniques, requiring minimal post-processing steps such as L-PBF.
The evidence of many research breakthroughs concerning the production of TiAl alloy parts using L-PBF does not make the processing technique immune to limitations. Efforts have been invested in overcoming processing limitations such as part cracking, micro-pore formation and uneven powder deposition through processing parameter optimisation [73]. Processing parameters can be varied to develop TiAl alloys with excellent mechanical properties in application. Some of these properties are beam size, laser power, scanning speed, scan hatch spacing and powder layer thickness [74].
Polozov et al. [75] confirmed that TiAl-based alloy crack-free samples could be built via L-PBF processing with a high-temperature platform preheating of 900°C. Fully densified samples (highest relative density of 99.9%) were attained at volume energy density 48 J/mm3. The refined microstructure consisted of equiaxed grains, lamellar α2/γ colonies and retained β-phase. As compared to conventionally produced TiAl alloys, high ultimate compressive strength and strain values were obtained.
Process parameters can be optimised to aid the fabrication of TiAl specimen, and unfortunately, the resultant part still shows pores, cracks and low densities. One needs to understand the crack and pore formation mechanisms and the defect-process parameter relationships in such a case. Shi and associates [70] investigated optimal L-PBF process window and the effect of substrate preheating. Moreover, the relationship between crack formation, pore formation, and the process parameters was studied and the crack propagation discrepancy with an increase in the number of deposition layers. It was concluded that crack formation was related to process parameters and the number of deposition layers. The cracks initiated in the 3rd layer are accounted for by residual stress accumulation and the deviations in the composition of Ti-47Al-2Cr-2Nb deposition layers. Furthermore, substrate (Ti-6Al-4 V) preheating at 200°C alleviated cracking. Finally, a good metallurgical bond between the substrate and Ti-47Al-2Cr-2Nb deposition layers was found.
The addition of yttrium (Y) to TiAl alloys (specifically class TNM) and process parameter optimisation dramatically affects the formability, and ultimately the cracking behaviour and control of L-PBF produced components. Gao et al. [76] fabricated TNM alloys with varying Y contents (0, 1, 2, 3, 4 wt.%) and investigated the mechanism of improved formability, cracking sensitivity, cracking behaviour and control mechanism by Y additions. Improvements in the formability of Y added-TNM alloys were assigned to lower melt viscosities and good laser energy absorption. The addition of 2, 3 and 4 wt.% Y to the TNM alloys coupled with a laser energy density greater than 7.00 J/mm2 formed crack-free samples. The obtained microstructure and phase constituents were reported to contribute to microcrack formation and control significantly. Lower Y additions resulted in coarse columnar grains, oxygen segregation at the grain boundaries with dominating brittle B2 phase with poor ductility. In contrast, higher Y additions (2–4 wt.%) refined equiaxed grains, enhanced the oxygen-scavenging effect (through the presence of Y2O3 particles), and decreased brittle B2 phase content at higher Y additions significantly improve the ductility.
Finally, adding Nb to γ-TiAl alloys was also reported to account for improved mechanical properties based. Ismaeel et al. [77] produced Ti-Al–Mn–Nb alloys on a TC4 substrate and studied the effects of different Nb contents on the microstructure and properties of the alloys. The phases obtained consisted of γ-TiAl and α2-Ti3Al and a consecutive microstructural change with increased Nb additions from near full dendrite to near lamellar. Also, adding 7 at.% of Nb resulted in improved alloy’s hardness, strength and plastic deformation. Moreover, the elevated temperature oxidation resistance and tribological properties were significantly improved.
4. Hot corrosion
4.1 Definition
Hot corrosion can be defined as a chemical degradation on the metallic surface of materials operating at high temperatures, enhanced by the presence of molten ash and gases containing elements such as sulphur (S), chlorine and sodium [78]. Such environmental elements during fuel combustion promote damage to the protective oxide film by forming contaminants such as V2O5 and Na2SO4 [79]. This degradation form was initially identified in the early 1950s on combustion engines and boilers [80] and has been explored in numerous research works [50, 81, 82, 83, 84, 85, 86, 87].
4.2 Characteristics
Hot corrosion exists as Type I (known as High-Temperature Hot Corrosion) or Type II (Low-Temperature Hot Corrosion), with the former occurring above 800–950°C and the latter at 600–750°C [88, 89]. The occurrence of either attack form is dependent on several parameters such as the composition of the alloy, contaminant, and gas. Furthermore, other vital parameters are temperature and temperature cycles, erosion processes and gaseous velocity [90, 91]. The main difference between high-temperature hot corrosion (HTHT) and low-temperature hot corrosion (LTHC) is the morphologies thereof. HTHC is distinguished by the occurrence of a non-porous protective scale, internal sulphidisation and chromium (Cr) depletion.
4.2.1 High-temperature hot corrosion (HTHC)-type I
This form of attack, also referred to as molten salt-induced corrosion, comprises a liquid-phase salt mixture deposit observed at high temperatures at the start of deposition [92]. Traditionally, according to Nicholls and Simms [93], HTHC has been detected in a temperature array between the surface deposit melting point and vapour deposition dew point for the deposit. Above this suggested temperature band, instability of dew point deposit exists, resulting in evaporation. A series of chemical reactions occur, initially attacking the oxide film and progress to deplete Cr present in the substrate [94]. Oxidation of the base material is then accelerated by Cr depletion, promoting a porous oxide scale formation.
An example of this could be the formation of thermodynamically unstable liquid sodium sulfate (Na2SO4) deposits. The marine environment mainly sources such deposits in sea salt form, followed by atmospheric contaminants such as volcanic discharges and fuel. During combustion, the present Na2SO4 can combine with pollutants present in air or fuel (such as chlorides, V and Pb) to form a blend of low melting temperature salts, further broadening the temperature range attack [94]. In the presence of sodium chloride (NaCl), the following reaction after combustion can be observed:
2NaCl+SO2+O2=Na2SO4+Cl2E1
HTHC can be classified into four stages from initiation up to failure [95]:
Stage I: Initial coating deterioration—roughening of the surface edges coupled with localised oxide layer disintegration and minor base metal layer depletion is observed. If the surface is left untreated, the condition will worsen. Surface recoating and stripping may be adequate to remedy this degree of damage.
Stage II: Oxide layer rapture—characterised by an acceleration and advancement in surface roughness compared to Stage 1 and the protective oxide layer’s failure. Although the mechanical integrity remains maintained, there is no way to salvage the component to its original state.
Stage III: Detrimental sulphidisation—depicted by massive scale build-up on the component’s surface and indications of liquid Na2SO4 under the protective layer. The structural integrity of the part is significantly affected, attack by S contaminants proceeds.
Stage IV: Catastrophic attack—failure of the component occurs due to the observed significant blistered scale penetrating much into the base metal. Structural rigidity is lost.
This corrosion damage is characterised by a uniform attack, internal sulphide phases, depletion zone beneath a relatively smooth scale–metal interface [80, 96].
4.2.2 Low-temperature hot corrosion (LTHC)-type II
Type II corrosion has been reported [97, 98, 99, 100, 101, 102] as a liquid-phase deterioration by a blend of molten nickel (Ni) or cobalt (Co)-containing sulphates such as Na2SO4-CoSO4 or Na2SO4-NiSO4 accountable for corrosion initiation and propagation. The corrosion initiation is achieved through oxide layer fluxing, while propagation is accelerated by the mass movement of reactive elemental components through liquids present in the corrosion pits [80]. Studies [103, 104, 105, 106] have shown that conversion from CoO and NiO occurs when SO3 in the gas reacts with the sulphates, attributing to the extensive usage of mixed Na2SO4-NiSO4 in recent LTHC research studies.
LTHC can be found in coated or uncoated compressor and turbine parts. For instance, the sometimes turbine blade’s uncoated internal cooling systems operating at temperatures of about 650–750°C may be prone to this corrosion type [107]. The external rim of uncoated turbine blades reaches temperatures of 400–800°C [108]. LTHC is distinguished by the pit’s appearance and the absence of a sulphide zone at the corrosion front, consuming all the S [96].
4.3 Mechanisms
Two HTHC mechanisms have been proposed, namely sulphide-oxidation and salt fluxing mechanisms [94]. Acidic and basic fluxing reactions, presented initially by Goebel and Pettit [109, 110], may be obtained and rely on the compositions of the alloy, oxide and underlying coating [93]. According to this model, fluxing occurs due to the decomposition of oxides into corresponding cations and O2− (known as acidic fluxing) or oxides with O2− forming anions (referred to as basic fluxing).
In acidic fluxing, oxide ions are donated to the deposit melt through dissolving the oxide scale [93]:
MO=M2++O2−E2
Acidic environments in molten deposits can be developed through two main processes, namely, alloy-induced and gas-phase acidic fluxing. Basic fluxing is achieved through the production of oxide ions in a Na2SO4 deposit. Such is obtained by removing S and oxygen from the residue through reactions with the alloy or underlying coating. Subsequently, the oxide scales (e.g., MO) produced can react with the oxide ions through reactions [93]:
MO+O2−=MO22−E3
A conventional model for LTHC was proposed by Luthra [111]. As suggested by the model, LTHC follows two stages, namely, formation of liquid-form sodium-cobalt sulphate and attack propagation through SO3 migration through the liquid salt. In nickel-based alloys, the mechanism suggested by Shih and associates [112] for LTHC is sulphidisation.
4.4 Laboratory testing techniques
An alloy’s resistance to hot corrosion can mainly be determined using four standard tests: the electrochemical, crucible, accelerated oxidation, and burner-rig [94, 113]. The crucible tests remain the most highly ranked test for hot corrosion, simply consisting of either suspending, depositing, or completely immersing the testing sample in molten salts at elevated temperatures, as presented in Figure 3. As far as TiAl alloys are concerned, less work has been carried out to understand the hot corrosion behaviour of such alloys [114, 115, 116].
Figure 3.
Configurations used in hot corrosion crucible testing [114].
Gas turbine environments can be precisely simulated by employing burner-rig tests [117, 118], shown in Figure 4. The salt is in aerosol or fog form and fuel oil/air is introduced into the testing chamber to generate the test environment [119]. Simmons et al. [120] indicated that hot corrosion is an electrochemical process since hot corrosion consists of electrochemical reactions in which the molten salt acts as the conductive media or electrolyte.
Figure 4.
Burner rig hot corrosion test schematic representation where (a) is an illustrates the experimental setup for Na2SO4(g) exposure, (b) is an image of the specimen plate for Na2SO4(g) tests in a crucible with the salt container and (c) is an ex-situ salt hot corrosion schematic diagram setup for experimental studies [119].
4.5 Prevention methods
Some of the approaches used to prevent hot corrosion include maintaining both fuel purity and composition, properly selecting structural alloys, employing coatings, cleaning hot parts and air filtering [94].
4.5.1 Fuel purity and composition
Initiation and propagation of hot corrosion are greatly affected by impurities such as vanadium (V), S, and various alkali earth metals [121]. This can be controlled by adding magnesium (Mg), Cr, barium and calcium to the combustion fuel to decrease corrosion rate. The presence of zinc (Zn) in the form of anodes in the combustion fuel or as part of the protective coating can significantly reduce the occurrence of LTHC. According to Hancock and associates [122], Zn drastically reacts with chloride ions (i.e., when excess NaCl is available) and transfers the chloride to the gas-salt interface to transform to chloride gas via sulphidisation.
4.5.2 Proper alloy selection
The addition of Cr to superalloys has effectively reduced the occurrence of hot corrosion [123]. Historically [121, 124], Cr (15 wt.% for Ni-based and 25 wt.% for Co-based alloys) has been added to superalloys to reduce HTHC. Much related to TiAl alloys, Garip and Ozdemir [125] studied the effect of Cr, Mo and Mn on the cyclic hot corrosion behaviour, and subsequently reported the beneficial effects of Cr and Mn additions on the hot corrosion properties of the investigated samples. Cr’s effect on corrosion resistance is attributed to the ability of Cr to form Cr2O3, stabilising the chemistry melt, preventing reprecipitation of the protective oxide scale. Contrarily, increased Cr additions to superalloys can compromise the high-temperature strength and ductility [113] by forming TCP phases. The alloy and oxide film adhesion has been reported to be improved by the addition of zirconium, yttrium, scandium, cerium and lanthanum [113]. Silicon (Si), platinum (Pt), hafnium, Ti, Al, and Nb [126] were also found to increase resistance to hot corrosion.
4.5.3 Protective coatings
Such as diffusion, overlay and thermal barrier (TBCs) coatings can be used on relatively resistant alloys to combat hot corrosion. An alloy’s surface enrichment by Al, Si or Cr achieves diffusion coatings. Various aluminide diffusion coatings (i.e., PWA70, MDC3V, PWA62, TEW LDC2, Elbar Elcoat 360 and Chromalloy RT22) have been developed and can be alloyed with Pt to improve cyclic oxidation at high temperatures [127]. Overlay coatings, commonly referred to as M (base metal)–Cr–Al–Y coatings, are designed for LTHC and HTHC surface protection. Overlay coatings with low Cr-high Al coatings are used for HTHC protection, while high Cr-low Al coatings are used for LTHC [94]. TBCs protects the substrate from gaseous flow caused by heat and consist of an external ceramic usually zirconia) and an oxidation-resistant bond-coat overlay. Other coatings include intelligent coatings like RT22 (Pt-aluminide) and Sermetal 1515 (a triple-cycle Si-aluminide treatment), have been reported [127].
Inexpensive alternatives include oxide-based glass and glass–ceramic coatings [128, 129]. Oxide-based glass and glass–ceramic coatings exhibit a remarkable combination of properties such as excellent chemical inertness, high-temperature stability and superior mechanical properties, which effectively can mitigate deterioration caused by hot corrosion. The introduction of halogens on the surface of the alloy encourages the preferential formation of aluminium halides at elevated temperatures. The aluminium halides are then converted to thin, continuous, and protective alumina oxides. Fluorine provides the best oxidation protection [130]. Further examples of surface modifications coating and methods studied on γ-TiAl alloys include magnetron sputtering [131], laser cladding [132], sol–gel [133], pack cementation [134], chemical vapour deposition [135], slurry [136], ion implantation [137].
4.5.4 Cleaning hot parts and air filtering
Motoring washes can be flooded with plain water [121] to prevent hot corrosion using specified procedures in the maintenance manual for the specific engine model. Also, high-efficiency filters can be used to filter out air containing high sodium contents [138].
4.6 Hot corrosion studies for TiAl alloys
Although much work has been devoted to understanding the hot corrosion kinetics of Ni-based and Co-based superalloys, TiAl alloys emerged to have sparked much interest in recent years [1, 56, 57, 139]. Historically, reported works utilised alloys produced using conventional methods; however, more attention has recently shifted to AM routes [70, 73, 140, 141, 142, 143, 144, 145, 146]. Despite much devotion to improving structure–property relations of TiAls, little work has been reported on the hot corrosion of additively manufactured TiAl.
Garip and Ozdemir [147] produced an alloy to the nominal at.% composition of Ti-48Al-10Cr using electric current activated sintering and studied the hot corrosion kinetics of the alloys in Na2SO4 salt for 180 h at 700–900°C. A severe hot corrosion attack was observed at 900°C (refer to Figure 5), with a porous and loose layer consisting of Na2Ti3O7, TiO2, Al2O3 traces of TiS phase.
Figure 5.
Cross-sectional SEM images showing oxide scale microstructures with EDS analysis points represented in at.%, after hot corrosion exposure at (a) 800°C and (b) 900°C for 180 h [147].
In a study led by Xiong et al. [67], bare alloys TiAl, TiAlNb, and Ti3AlNb, were severely damaged after exposure at 750°C in (Na, K)2SO4 + NaCl melts as compared to those coated with enamel or TiAlCr. The corrosion mechanism was described to be much related to self-catalysis of sulphidisation and chlorination of metallic components. The initial mass loss observed is due to chloride volatility via metallic component chlorination. Of the alloys investigated, TiAlNb exhibited the best corrosion resistance due to adhesive Al2O3 enriched scale formation. Lastly, the degradation acceleration of sputtered TiAlCr coating was reported to be due to the chlorination of Cr and Al.
Additions of Nb and Si to traditional TiAl coatings were found to improve the hot corrosion resistance of a Ti-6Al-4 V alloy. In the stated work, Dai et al. [148] investigated the corrosion mechanisms on a mass loss basis following exposure at 800°C in a 75 wt.% Na2SO4 + NaCl salt mixture. Increasing single Nb additions deteriorated the hot corrosion resistance of the coating. Comparatively, increasing single Si additions continued to improve hot corrosion resistance. However, additions of both Nb and Si simultaneously showed better resistance to corrosion than single element additions. The corrosion protection of both Nb and Si (as seen in Figure 6) was related to SiO2 and Al2O3 formation in the initial stages of hot corrosion. Secondly, Si additions were reported to promote the formation of a Na2O-Al2O3-TiO2-SiO2 enamel, hindering contact between the corrosive media and the oxide scales.
Figure 6.
Representative hot corrosion model of TiAl-xNbySi coating where (a) illustrates TiO2 and Al2O3 formation and (b) shows an acidic dissolution of TiO2 to form sodium titanates including NaTiO2 and Na2TiO3 [148].
Tang et al. [149] studied the effect of enamel coatings on γ-TiAl against hot corrosion at 900°C. The enamel coating remained stable in the (Na,K)2SO4 melts, thus effectively protecting it against hot corrosion attack. Silicon-based coatings have also been shown to protect TiAl alloys. Rubacha et al. [150] evaluated the hot corrosion resistance of silicon-rich coated Ti-46Al-8Ta (at.%) alloy in NaCl, Na2SO4 and a mixture of the two salts. The formation of an amorphous SiO2 layer with TiO2 (rutile) and α cristobalite crystals enhanced the hot corrosion resistance of the TiAl alloy. Furthermore, Wu and colleagues [151] studied the hot corrosion resistance of a SiO2 coated TiAl alloy in 75 wt.% Na2SO4 + 25 wt.% NaCl salt mixture at 700°C. The enhanced hot corrosion resistance of the TiAl alloy was attributed to the formation of a compact and adherent amorphous SiO2 embedded with Na2Si4O9 and cristobalite. The incorporation of Si in aluminide coatings has also provided long-term oxidation protection of γ-TiAl alloys at temperatures of 950°C by forming a continuous and uniform α-alumina oxide scale [152].
5. Oxidation
5.1 Definition
When metallic materials are exposed to elevated temperatures in air, oxidation occurs, resulting in the formation of oxide scales. The crystal structure of the individual metals significantly affects the oxidation rate of high-temperature applicative parts [153, 154].
5.2 Oxidation behaviour in TiAl alloys
The following reactions may occur when TiAl alloys are subject to an oxidising environment:
12Tis+O2g=TiOsE4
TiOs+12O2g=TiO2sE5
2Als+32O2g=Al2O3sE6
The ultimate oxidation resistance of alloyed TiAls is achieved by forming protective Al2O3, Cr2O3 and SiO2 scales due to their outstanding thermal stabilities. In contrast, the unfavourable formation of porous TiO2 with a high crack tendency is often observed [153]. Cobbinah et al. [155] found that 4 and 8 at.% Ta additions to Ti-46.5Al alloy promoted the significant formation of a consistent, non-porous Al2O3 layer at the metal-oxide boundary. Additionally, the layer operated as a diffusion barrier and preceded to outstanding oxidation resistance of the TiAl alloys.
In a study by Pan et al. [156], a comprehensive understanding is provided of the role of alloying on the oxidation resistance of TiAl alloys. Protection was related much to the formation of Ti3Sn layer diminishing oxygen diffusion inwardly, promoted by Sn additions. Moreover, spallation resistance was enhanced by the Al2O3 oxide pegs providing a mechanical locking. The effect of cathodically electrodepositing a SiO2 film on the oxidation resistance of a TiAl alloy was studied [157]. After 900°C exposure in air, the resultant alumina- and silicon-enriched glass-like oxide scale (in Figure 7) was reported, preventing oxygen diffusion leading to remarkably decreased alloy oxidation rates.
Figure 7.
Representation of a γ-TiAl alloy coated with E-SiO2 film (a) and after thermal oxidation test (b) [157].
Surface modification of TiAl alloys via anodising has sparked interest in many high-temperature oxidation studies [158, 159, 160, 161]. For instance, the oxidation behaviour and protection mechanisms of a TiAl alloy were studied by anodising in a methanol/NaF solution and produced an aluminium (Al)-and fluorine-enriched anodic film [162]. After 100 h exposure at 850°C, no evidence of cracking and spallation was displayed on the surface. The enhanced high-temperature oxidation resistance is mainly attributed to the halogen effect, generation of Al2O3 and oxidised Al–F species inhibiting external oxygen diffusion. Much effort has been devoted to developing coatings for γ-TiAl alloys, summarised in an evaluation by Pflumm et al. [130]. Amongst many available coating methods, Si-modified aluminide coatings produced via pack cementation have gained popularity. One such study [81] demonstrated that a continuous α-Al2O3 scale remained adherent after exposure to a temperature of 950°C for 3000 h.
5.3 Oxidation kinetics of TiAl alloys
When a metal operating at elevated temperatures is exposed to air, an oxide scale forms. As oxide scale formation proceeds, the metal’s weight change can be plotted against time. Several laws such as linear, parabolic, logarithmic or cubic can be observed when studying oxidation kinetics [163]. In as far as TiAl alloys are concerned, either linear or parabolic oxidation kinetics prevail. While the former offers no protection against high-temperature oxidation, the latter promotes diffusion-controlled oxide scale formation, improving much on the oxidation resistance of the base material. Parabolic oxidation follows and obeys the following law:
ΔmA2=kptE7
where Δm = change in weight (in mg), A = surface area (in cm2), t = time (in sec) and kp = parabolic oxidation rate constant (in mg2.cm−4.sec1).
The optimum oxidation protection governed by the parabolic law often results in a thick and continuous TiO2 and Al2O3 scale. As such, Swadźba et al. [48] investigated the short-term oxidation behaviour of a TiAl 48–2-2 alloy produced by AM at a temperature range of 750–900°C in air. At 900°C, a non-porous scale consisting of TiO2, Al2O3 and nitrates, exhibiting parabolic oxidation (in Figure 8), was observed.
Figure 8.
Mass change against time plots for (a) oxidation rate constant of the AM produced TiAl 48–2-2 alloy and (b)–(c) the power-law constant – n extrapolation [48].
Garip [164] likewise studied the oxidation kinetics at 900°C in air for 200 h for TiAl alloys produced via pressureless and resistance sintering. Both alloys exhibited a nearly parabolic oxidation response, with oxidation rate constants of the pressureless sintered alloy of 0.6391 mgn cm−2n h−1, 1.8 times higher than that of the alloy compacted using resistance sintering. Multi-layered oxide scales consisting of TiO2 and Al2O3 were obtained.
5.4 Effect of alloy modifications on the oxidation resistance of TiAl
Oxidation protection offered by forming a continuous Al2O3 scale followed by a multilayer of TiO2 + Al2O3 is limited, unfortunately, to the maximum service temperature of ~830°C. Above this temperature, the protection potential presented by the oxide scales formed severely deteriorates, limiting the high-temperature application potential for structural components [165]. The current trend in research is to improve the oxidation resistance of TiAl through alloy modifications.
Nb is one element used in many research works [86, 87, 88, 89, 90] to improve the oxidation resistance of TiAl alloys. Al activities are promoted by Nb additions and accelerate protective Al2O3 oxide film formation, limiting oxygen diffusion into the alloy [166]. Also, the α2 phase present in TiAl alloys is significantly decreased by Nb additions, decreasing its oxygen solubility [54]. Although Nb was primarily used for improving oxidation resistance [167], other high-temperature properties such as strength and creep resistance have been enhanced by the presence of Nb.
The creep resistance and the oxidation resistance of TiAl and its alloys can be enhanced by adding Si. The oxidation improvement is said to be achieved through the refinement of TiO2 particles, inducing refined and compact TiO2 scales on the surface [165]. Moreover, Si promotes Al diffusion into the oxide scale, stabilises Ti, reduces Ti4+ ions and impedes external Ti4+ ions diffusion [168].
The effect of adding molybdenum (Mo) alone to TiAls to improve on high-temperature oxidation is minimal. The protection of Mo-containing TiAl alloys is through the formation of inner oxide layers of TiO2 and Ti2AlMo near the substrate surface [165]. Unfortunately, Mo additions cannot alter the external oxide film formed (i.e., comprises of loose and porous TiO2 scales) and its characteristics. It is recommended in practice that the improvement of high-temperature oxidation cannot be derived from adding Mo alone; instead, the combination of Mo with other alloys can have a beneficial effect on the alloys’ resistance to oxidation [169].
Cr additions promote the formation of Cr2O3 oxides, which act as mass ion transport barriers [170], enhancing oxidation resistance. In addition, the Al content existing in the alloys can be significantly suppressed by Cr additions, promoting the formation of Al2O3 scales. Oxygen diffusion at elevated temperatures can be accelerated by Cr3+ ion doping in titanium oxide, improving oxygen vacancy concentration. Contrarily, the doping effect may impair the TiAl alloy’s oxidation by making Ti4+ interstitially occupying TiO2 sites, improving the potential energy with a noticeable decrease in diffusion activation energy, encouraging the diffusion of Ti4+ in TiO2 [171].
Zirconium (Zr) additions can also enhance oxidation properties by altering the characteristics of the oxide formed during the primary stages of oxidation and promote oxide grain nucleation [172]. As a result, the refinement of the oxide particles occurs, which can hinder oxygen diffusion. Rare earth metals have been reported to enhance the oxidation resistance of TiAl alloys. As discussed in detail in a research paper by Dai et al. [165], the protection mechanism is contributed by grain refinement, substrate purification, oxide adherence improvement and promotion of Al selective oxidation.
6. Conclusions
The need for materials to give excellent mechanical properties under high temperatures and extreme conditions such as TiAl is in demand. The use of such alloys would mean a reduction in pollution and noise levels for aero-based engines due to improved thermal efficiencies. There are challenges in producing such alloys using the conventional arc and induction melt casting techniques due to the extremely high melting temperatures of the alloys. The AM route, particularly L-PBF, presents an opportunity to produce such alloys. What is critical in such trials is the operating parameters during processing. This has a direct influence on the performance and mechanical properties of the alloys so produced. Hot corrosion and oxidation of TiAl alloys are of great concern in gas turbine engines. Hot corrosion can be classified into HTHC and LTHC, with particular reference to mechanisms and characteristics. Protection control methods may result in fewer catastrophic failures. The hot corrosion process must be either totally prevented or detected early to avoid catastrophic failure. A sound understanding of oxidation mechanisms and kinetics of TiAl alloys makes it easier to tailor oxidation-resistant alloys by alloy modifications.
Acknowledgments
This research work is based on the research supported wholly/in part by the National Research Foundation of South Africa (Grant number 130004).
Conflict of interest
The authors declare no conflict of interest.
\n',keywords:"titanium aluminides, oxidation, hot corrosion, additive manufacturing, laser powder bed fusion",chapterPDFUrl:"https://cdn.intechopen.com/pdfs/78880.pdf",chapterXML:"https://mts.intechopen.com/source/xml/78880.xml",downloadPdfUrl:"/chapter/pdf-download/78880",previewPdfUrl:"/chapter/pdf-preview/78880",totalDownloads:130,totalViews:0,totalCrossrefCites:0,dateSubmitted:"February 15th 2021",dateReviewed:"September 7th 2021",datePrePublished:"October 8th 2021",datePublished:null,dateFinished:"October 8th 2021",readingETA:"0",abstract:"This research paper summarises the practical relevance of additive manufacturing with particular attention to the latest laser powder bed fusion (L-PBF) technology. L-PBF is a promising processing technique, integrating intelligent and advanced manufacturing systems for aerospace gas turbine components. Some of the added benefits of implementing such technologies compared to traditional processing methods include the freedom to customise high complexity components and rapid prototyping. Titanium aluminide (TiAl) alloys used in harsh environmental settings of turbomachinery, such as low-pressure turbine blades, have gained much interest. TiAl alloys are deemed by researchers as replacement candidates for the heavier Ni-based superalloys due to attractive properties like high strength, creep resistance, excellent resistance to corrosion and wear at elevated temperatures. Several conventional processing technologies such as ingot metallurgy, casting, and solid-state powder sintering can also be utilised to manufacture TiAl alloys employed in high-temperature applications. This chapter focuses on compositional variations, microstructure, and processing of TiAl alloys via L-PBF. Afterward, the hot corrosion aspects of TiAl alloys, including classification, characteristics, mechanisms and preventative measures, are discussed. Oxidation behaviour, kinetics and prevention control measures such as surface and alloy modifications of TiAl alloys at high temperature are assessed. Development trends for improving the hot corrosion and oxidation resistance of TiAl alloys possibly affecting future use of TiAl alloys are identified.",reviewType:"peer-reviewed",bibtexUrl:"/chapter/bibtex/78880",risUrl:"/chapter/ris/78880",signatures:"Ntebogeng Mogale, Wallace Matizamhuka and Prince Cobbinah",book:{id:"10669",type:"book",title:"Corrosion: Fundamentals and Protection Mechanisms",subtitle:null,fullTitle:"Corrosion: Fundamentals and Protection Mechanisms",slug:null,publishedDate:null,bookSignature:"Dr. Fahmina Zafar, Dr. Anujit Ghosal and Dr. Eram Sharmin",coverURL:"https://cdn.intechopen.com/books/images_new/10669.jpg",licenceType:"CC BY 3.0",editedByType:null,isbn:"978-1-83968-606-1",printIsbn:"978-1-83968-605-4",pdfIsbn:"978-1-83968-607-8",isAvailableForWebshopOrdering:!0,editors:[{id:"89672",title:"Dr.",name:"Fahmina",middleName:null,surname:"Zafar",slug:"fahmina-zafar",fullName:"Fahmina Zafar"}],productType:{id:"1",title:"Edited Volume",chapterContentType:"chapter",authoredCaption:"Edited by"}},authors:null,sections:[{id:"sec_1",title:"1. Introduction",level:"1"},{id:"sec_2",title:"2. Titanium aluminide (TiAl) alloys: phase, microstructures and mechanical properties",level:"1"},{id:"sec_2_2",title:"2.1 Phase and microstructural evolutions",level:"2"},{id:"sec_2_3",title:"2.1.1 Phase evolutions",level:"3"},{id:"sec_3_3",title:"2.1.2 Microstructure-mechanical property relations",level:"3"},{id:"sec_6",title:"3. Additive manufacturing (AM) of TiAl",level:"1"},{id:"sec_6_2",title:"3.1 Process overview",level:"2"},{id:"sec_7_2",title:"3.2 Research milestones",level:"2"},{id:"sec_9",title:"4. Hot corrosion",level:"1"},{id:"sec_9_2",title:"4.1 Definition",level:"2"},{id:"sec_10_2",title:"4.2 Characteristics",level:"2"},{id:"sec_10_3",title:"4.2.1 High-temperature hot corrosion (HTHC)-type I",level:"3"},{id:"sec_11_3",title:"4.2.2 Low-temperature hot corrosion (LTHC)-type II",level:"3"},{id:"sec_13_2",title:"4.3 Mechanisms",level:"2"},{id:"sec_14_2",title:"4.4 Laboratory testing techniques",level:"2"},{id:"sec_15_2",title:"4.5 Prevention methods",level:"2"},{id:"sec_15_3",title:"4.5.1 Fuel purity and composition",level:"3"},{id:"sec_16_3",title:"4.5.2 Proper alloy selection",level:"3"},{id:"sec_17_3",title:"4.5.3 Protective coatings",level:"3"},{id:"sec_18_3",title:"4.5.4 Cleaning hot parts and air filtering",level:"3"},{id:"sec_20_2",title:"4.6 Hot corrosion studies for TiAl alloys",level:"2"},{id:"sec_22",title:"5. Oxidation",level:"1"},{id:"sec_22_2",title:"5.1 Definition",level:"2"},{id:"sec_23_2",title:"5.2 Oxidation behaviour in TiAl alloys",level:"2"},{id:"sec_24_2",title:"5.3 Oxidation kinetics of TiAl alloys",level:"2"},{id:"sec_25_2",title:"5.4 Effect of alloy modifications on the oxidation resistance of TiAl",level:"2"},{id:"sec_27",title:"6. Conclusions",level:"1"},{id:"sec_28",title:"Acknowledgments",level:"1"},{id:"sec_31",title:"Conflict of interest",level:"1"}],chapterReferences:[{id:"B1",body:'Mphahlele MR, Olevsky EA, Olubambi PA. Chapter 12—Spark plasma sintering of near net shape titanium aluminide: A review [Internet]. Cao G, Estournès C, Garay J, Orrù R, editors. Spark Plasma Sintering. Elsevier; 2019. p. 281-299. Available from: http://www.sciencedirect.com/science/article/pii/B978012817744000012X'},{id:"B2",body:'Reith M, Franke M, Schloffer M, Körner C. 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Available from: http://www.sciencedirect.com/science/article/pii/S0925838816319351'},{id:"B166",body:'Raji SA, Popoola API, Pityana SL, Popoola OM. Characteristic effects of alloying elements on β solidifying titanium aluminides: A review. Vol. 6, Heliyon. Elsevier Ltd; 2020. p. e04463.'},{id:"B167",body:'Kothari K, Radhakrishnan R, Wereley NM, Sudarshan TS. Microstructure and mechanical properties of consolidated gamma titanium aluminides. Powder Metall. 2007;50(1):21-27.'},{id:"B168",body:'JIANG H ren, WANG Z lei, MA W shuai, FENG X ran, DONG Z qiang, ZHANG L, et al. Effects of Nb and Si on high temperature oxidation of TiAl. Trans Nonferrous Met Soc China (English Ed. 2008 Jun 1;18(3):512-7.'},{id:"B169",body:'Pflumm R, Donchev A, Mayer S, Clemens H, Schütze M. High-temperature oxidation behavior of multi-phase Mo-containing γ-TiAl-based alloys. Intermetallics. 2014 Oct 1;53:45-55.'},{id:"B170",body:'Wei DB, Zhang PZ, Yao ZJ, Liang WP, Miao Q , Xu Z. Oxidation of double-glow plasma chromising coating on TC4 titanium alloys. Corros Sci. 2013 Jan 1;66:43-50.'},{id:"B171",body:'Zhou C, Yang Y, Gong S, Xu H. Mechanism of Cr effect for improvement of oxidation resistance of Ti-Al-Cr alloys. Acta Aeronaut Asronautica Sin. 2001;22(1):73-77.'},{id:"B172",body:'Gaddam R, Sefer B, Pederson R, Antti ML. Oxidation and alpha-case formation in Ti-6Al-2Sn-4Zr-2Mo alloy. Mater Charact. 2015 Jan 1;99:166-174.'}],footnotes:[],contributors:[{corresp:"yes",contributorFullName:"Ntebogeng Mogale",address:"ntebogeng.mogale@yahoo.co.za",affiliation:'
Vaal University of Technology, Johannesburg, South Africa
Vaal University of Technology, Johannesburg, South Africa
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",metaTitle:"What Does It Cost?",metaDescription:"Open Access publishing helps remove barriers and allows everyone to access valuable information, but article and book processing charges also exclude talented authors and editors who can’t afford to pay. The goal of our Women in Science program is to charge zero APCs, so none of our authors or editors have to pay for publication.",metaKeywords:null,canonicalURL:null,contentRaw:'[{"type":"htmlEditorComponent","content":"
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All of our IntechOpen sponsors are in good company! The research in past IntechOpen books and chapters have been funded by:
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We are currently in the process of collecting sponsorship. If you have any ideas or would like to help sponsor this ambitious program, we’d love to hear from you. Contact us at info@intechopen.com.
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All of our IntechOpen sponsors are in good company! The research in past IntechOpen books and chapters have been funded by:
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European Commission
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Bill and Melinda Gates Foundation
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Wellcome Trust
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National Institute of Health (NIH)
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Research Councils United Kingdom (RCUK)
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Foundation for Science and Technology (FCT)
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Chinese Academy of Sciences
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Natural Science Foundation of China (NSFC)
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German Research Foundation (DFG)
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Max Planck Institute
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Austrian Science Fund (FWF)
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Initial biochemical studies have been exclusively analytic: dissecting, purifying, and examining individual components of a biological system; in the apt words of Efraim Racker (1913 –1991), “Don’t waste clean thinking on dirty enzymes.” Today, however, biochemistry is becoming more agglomerative and comprehensive, setting out to integrate and describe entirely particular biological systems. The ‘big data’ metabolomics can define the complement of small molecules, e.g., in a soil or biofilm sample; proteomics can distinguish all the comprising proteins, e.g., serum; metagenomics can identify all the genes in a complex environment, e.g., the bovine rumen. 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Dr. Blumenberg’s research is focused on the epidermis, expression of keratin genes, transcription profiling, keratinocyte differentiation, inflammatory diseases and cancers, and most recently the effects of the microbiome on the skin. 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Thus proteomics, an area of research that detects all protein forms expressed in an organism, including splice isoforms and post-translational modifications, is more suitable than genomics for a comprehensive understanding of the biochemical processes that govern life. The most common proteomics applications are currently in the clinical field for the identification, in a variety of biological matrices, of biomarkers for diagnosis and therapeutic intervention of disorders. From the comparison of proteomic profiles of control and disease or different physiological states, which may emerge, changes in protein expression can provide new insights into the roles played by some proteins in human pathologies. Understanding how proteins function and interact with each other is another goal of proteomics that makes this approach even more intriguing. Specialized technology and expertise are required to assess the proteome of any biological sample. 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