**Degradation of a Micro-Hybrid Dental Composite Reinforced with Polyaramide Fiber under the Influence of Cyclic Loads**

**Leszek Szalewski 1,**†**, Aneta Kami ´nska 2,**†**, Eliza Wallner 3, Justyna Batkowska 4, Tomasz Warda 5, Dorota Wójcik 6,\* and Janusz Borowicz <sup>6</sup>**


Received: 8 September 2020; Accepted: 12 October 2020; Published: 19 October 2020

**Featured Application: The article can help dentists in deciding whether to choose fiber-reinforced composite adhesive bridges as a treatment option. The results indicate that the choice of another therapeutic method should be considered in the case of increased occlusal forces (parafunction, bruxism, occlusal obstructions). The obtained results may become a starting point for research on new materials**/**fibers to increase the mechanical properties of adhesive composite bridges to obtain a material resistant to high occlusal forces.**

**Abstract:** Dental composites reinforced with glass fibers have a low tensile modulus and relatively low fatigue resistance. The aim of the study was to analyze the fatigue properties of a dental composite reinforced with polyaramide fibers under the influence of a cyclic, vertical load. For this purpose, we designed a thermoformable template, corresponding to the construction of adhesive bridges in the side section of the jaw. Fifty-four composite samples were made for the study. They were divided into three groups—control (K) and two experimental groups (R1 and R2). The experimental samples were subjected to cyclic fatigue using 75 N load. The number of cycles was 4690 and 20,100. The study used a three-point bending test. Statistical analysis showed a change in elasticity in groups related to the number of load cycles. The study showed that the samples from the control group required the greatest force to break in relation to those subjected to the work cycles. The maximum force in control (K) group was 738.1 N, R1—487.8 N, and R2—451.4 N. The determined algorithm showed a change in deflection associated with the increase of force value. The study did not show any relationship between the type of sample fracture and the number of load cycles.

**Keywords:** composite resins; compressive strength; fixed partial denture

#### **1. Introduction**

Increasing aesthetic requirements of patients has led to the development of composite dental resins. Usually dental composites are made up from the matrix and fillers that are connected with each other by so-called silanes. Modifications of these two components in the last 20 years have increased the use of dental composites. They are a material commonly used to reconstruct lost tooth tissues using aesthetic restorations [1,2].

Composite materials are based on methacrylate compounds. Their matrix is an organic photopolymerizating resin, which consists mostly of bisphenol A-glycidyl methacrylate (Bis-GMA), triethylene glycol dimethacrylate (TEGMA), and urethane dimethacrylate (UDMA). The inorganic phase is macro-, micro-, or nanofillers, based mostly on silicon compounds [3,4]. Additionally, the composite contains photo initiators and proadhesive agents—silanes. The micro-hybrid composites contain a mixture of at least two types of "glass" or quartz molecules, irregular in shape and similar in diameter (from 0.2 to 3 μm) and from 5 to 15% of small particles (0.04 μm). In these materials, the filler constitutes 60–70% by volume, i.e., about 77–84% by weight of the composite part [5].

In order to increase the field of application, as well as to broaden the indications for the use of composite materials, their combinations with fibers such as carbon, polyaramide, polyethylene, and glass are used [6,7]. Currently, such solutions are used in periodontology [8], endodontics [9], prosthetics [10], and orthodontics [11].

Both wholly aromatic polyamides or the shorter aromatic polyamids form aramids and stand for synthetic polyamides comprising >85% amide groups (–CO–NH–) bound directly to two aromatic rings. Such polymers are characterized as high-performance materials due to their very good mechanical strength and exceptional high thermal resistance. They are spun into fibers and used in advanced fabrics, such as sport and work protective clothing, bullet-proof armor, advanced composites in armament and aerospace industries, composites such as asbestos substitutes, and high-temperature insulation paper.

The aramid structure is based on rigid aromatic amide linkage. It is responsible for the exceptional properties of these materials. Highly directional and efficient interchain hydrogen bonds are established, giving basis to materials with a high tendency to crystallize and with extremely high cohesive energy density. At the same time, they are also responsible for the insolubility of the wholly aromatic polyamides, a disadvantage that prevents the expansion of the application field of these materials, whereas the improvement of the solubility is a topic of the present research interest [12]. Aromatic polyamides are obtained in reactions that form amide bonds between aromatic rings of high thermal stability and high strength. The outstanding rigid molecular chain structure, good orientation, and organization of the crystalline structure provides high strength and low elongation of the orientation of the molecular chains. This provides high tensile strength, impact, and differentiated thermal stability for various temperature ranges for an extended time [13]. The covalent bonds in the polymer are responsible for the high strength. However, man-made polymers generally do not exhibit the corresponding potential high modulus. High modulus and strength may result from structural perfection, crystallinity, and crystalline and amorphous orientation. It is well known that the highest elastic moduli reported from linear polymers are generally much smaller than theoretical values [14].

The bending strength of a prosthetic restoration is an extremely important feature whose assessment is necessary in specific cases of prosthesis use, for example, when it is necessity to increase its resistance to fracture during a long period of use or when a prosthesis with an elongated structure is made [15]. Chewing is the main factor that causes mechanical degradation of the composite resin. While chewing, the mandible moves repeatedly in the vertical and horizontal planes [16,17]. The mechanics of the chewing process adversely affect composite resins and lead to stress within the material structure [18]. During chewing, prosthetic restorations are subjected to loads with forces whose direction is consistent with the long axis of the tooth. The size of chewing force in the mouth is from 3 to 36—49 N on average—but it can also reach 1000 N [19,20]. During this process, the opposing teeth remain in mutual point contact, and the highest loads occur in the mouth only sporadically [21].

The number of cyclic contacts during chewing and swallowing per day depend on the number of meals, their consistency, the person's age and sex, the number of natural teeth, the type of prosthetic restorations, and the functional disorders of the masticatory apparatus that may significantly increase the number of contacts [22,23]. On average, single contacts last 0.3 s, and their number is variable, ranging from 658 to 2300 contacts a day [21,24,25].

The available literature has not yet explained all aspects of the degradation of composites nor, consequently, the problem related to the maximum time of the restoration use in real conditions. In general, tests have been carried out on standardized samples, reproducing cyclic loads and finite elements that were usually subjected to static loads [26–29].

The aim of the research was to determine and compare the effect of cyclic vertical loads on the strength of composite resins reinforced with high molecular weight polyaramide fiber.

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

#### *2.1. Preparation of the Test*

The study used a gypsum model of the jaw, from which a fragment covering the area of three teeth was dissected, between the second premolar and the second molar on the left side. The obtained model (model A) was duplicated in a silicone mass, and the obtained form was filled with a chemo-hardening, burnt-out acrylic resin, which was changed into a chromium–nickel alloy in the casting process. On the basis of the obtained model, we produced standard templates (Figure 1) that were used to make samples. Plates that were 1.5 mm thick were used to make the templates, which were placed in a pressure thermoforming device. The tiles were heated for 55 s, after which they were stretched on the prepared models. Next, in the gypsum model A, the first molar tooth was removed, and in the remaining two teeth, limiting this lack, a preparation was made within the crowns to obtain a model that was reproduced and turned into a chromium–nickel alloy (model B). Model B was used as a handle in which the samples were fixed during the test. In order to obtain the model on which the samples were made (C-model), before replacing with the chromium–nickel alloy, we blocked a 2-mm space for the future restored tooth (bridge span) with the use of a chemically curing resin. The obtained distance prevented the sample from leaning against the model during its possible deflection during the test (Figure 1). A chromium–nickel model (D) (Figure 1) of the opposed tooth (the first left mandible molar) was also made for model A, which allowed us to keep the proper contact of the opposing teeth. This model was used in the study as a contact element with the test samples.

**Figure 1.** The schema of sample preparation.

The study used a technical composite in A2 color and a polyaramide fiber with a width of 3 mm, which were used to made adhesive bridges.

#### *2.2. Specimen Preparation*

On model C, we pressed thin thermoformable films to insulate the material from the model walls. Then, the first layer of 1 mm thick composite was applied to the prepared place, on which a 17 mm polyaramide fiber was placed, corresponding to the range between the prepared areas in crowns of the teeth defining the gap (sample length). The template was then applied and polymerized using a 470 nm wavelength tube (Clear Blue LED 1200 mW/cm2) for 40 s for every 5 mm section. Subsequent layers of 0.5 mm thick composite were applied and condensed using standard dental applicators and polymerized in the same way until the template was filled. The excess material was removed. After polymerization and obtaining anatomical shapes, we released the samples from the models, and any excess of the composite material was removed using a composite milling cutter. The value of the width and height of connectors in the samples was 4 × 3 mm, respectively.

The research material consisted of anatomical samples (n = 54), which were divided into three groups: the control group (K, 21 samples) and two experimental groups (R1, R2), divided according to the number of cycles: 18 and 15 samples, respectively.

#### *2.3. Fatigue Strength Test*

The Zwick/Roell Z 2.5 (ZwickRoell GmBh & Co. KG, Ulm, Germany) testing machine was used in the test. During strength tests, we placed all samples in the holder, which was model C, with a distance of 11.5 mm between two end supports. When testing the control samples, we set the initial force to 5 N, and the traverse speed was 5 mm/min. The test lasted until the sample was destroyed, which was defined as a 20% drop in strength in relation to the achieved maximum force for a given sample.

Samples from R1 and R2 were subjected to cyclic loads of 75 N at a 90◦ angle using the tooth's opposed model. During this process, the traverse speed was 10 mm/min, and the sample holding time was 0.3 s. After this time, the tested object was relieved to zero force. Then, the cycle was repeated. The number of cycles in the R1 group was 4690, which is the average value corresponding to the number of weekly opposing teeth contact, and in the R2 group, we used 20,100 cycles, the number of which corresponds to monthly dental contacts [21]. After fatigue loads, the samples were subjected to a bending strength test in the same way as in group K. The obtained parameters were maximum force [N], deflection at maximum force [mm], deflection at yielding point [mm], and post-yield displacement [mm].

#### *2.4. Statistical Analysis*

The data were analyzed using the SPSS 20.0 PL statistical package (IBM, New York, NY, USA) [30], the Kolmogorov–Smirnov test (normality of distribution), and a one-way analysis of variance were used with Tukey's multiple comparison test. Spearman's nonparametric correlation coefficients and selected linear regression coefficients were also estimated. The distribution of the crack types recorded in individual samples was verified by a non-parametric χ<sup>2</sup> test.

#### *2.5. Microscope Observations*

After durability tests, microscopic observations of the work surfaces of fractures of samples were made to check for possible damage. Observations were made using a VEGA//LMU scanning electron microscope with 100 and 2000 magnification.

#### **3. Results**

The study showed that the samples from the control group required the greatest force to break in relation to those subjected to the work cycles. The study did not show statistically significant differences between the destructive force in the case of groups R1 and R2, despite the differences in the number of load cycles; however, the numerical differences were found in favor of group treated by smaller number of cycles. Moreover, tests from experimental groups (R1 and R2) were characterized by significantly lower elasticity (deflection) under the maximum force action. Samples from experimental groups were definitely more inflexible compared to the control samples. The reduction of deflection after cyclic loads may result from changes in the internal structure of the samples. Detailed results regarding strength parameters for particular groups are presented in Table 1.

**Table 1.** Strength parameters of samples included in the tests (mean ± standard error of the mean (SEM)).


a, b—differ significantly at ≤0.05; SEM—standard error of the mean.

During the strength tests, we found three types of cracks: longitudinal, transversal, and defragmentation (Figure 2). It was not demonstrated that the method of sample fracture depended on the number of performed fatigue cycles (χ2, *p* = 0.132). Cracks in external polyaramide fibers were observed only in one sample from group R2 (Figure 3). Photographs from the electron microscope show the fiber structure, which is characterized by an empty space inside. Effective forces are focused on the outer surface of the fiber. The presence of fine cracks in the area of the main crack was also noticed. The vertical application of force causes the emergence of cutting forces, leading to side cracks in various directions. It has been observed that with the increase in the number of fatigue cycles, the number of these cracks is much higher, which is related to the increase in the hardness of the composite material with the simultaneous increase in brittleness (Figures 4–6). Such a situation may result from the fact that the contact points occur on the unevenness of the teeth chewing surfaces and may be a consequence of the lack of perfect stiffness of the test sample system.

Table 2 presents correlation coefficients between the analyzed strength parameters. There was a statistically significant dependence (*p* ≤ 0.01) between the number of cycles and individual measured characteristics, whose values decreased considerably with the increase of the number of destructive cycles. It has been noted that with the increase of the maximum force, the flexibility of the sample characterized by material deflection increased significantly. Similarly, highly significant, positive relationships were found for all three plastic deformations of the analyzed samples.

**Figure 2.** The types of ruptures found during the experiment ((**A**) longitudinal, (**B**) transverse, (**C**) defragmentation).

**Figure 3.** View of sample R2 (magnification 1.99 k×).

**Figure 4.** View of control group sample (magnification 2.00 k×).

**Figure 5.** View of sample R1 (magnification 97×).

**Figure 6.** View of sample R2 (magnification 95×).


**Table 2.** The correlation coefficients between measured samples' characteristics.

\*\* correlation is significant at *p* ≤ 0.01 (one-sided).

Figure 7 illustrates the dependence of deflection at maximum force on the value of this force depending on the group treated by various numbers of working cycles. It is visible that each group demonstrated different reactions on breaking force. The most resistant was the control group not subjected to cyclic loads. The durability of other groups also varied according to number of cycles, with better resistance characterized by R1 group in comparison to the R2 group.

**Figure 7.** The relationship between deflection at maximum force and the maximum force value.

Figure 8 shows that the dependence of the deflection on the load value in groups K, R1, and R2 were distinguished on the basis of the fatigue cycle number. It is possible that materials in certain samples could behave the other way; however, we did not find any differences using standard error of the mean (SEM). All determined regression coefficients were statistically significant (*p* ≤ 0.01). Their properties resulted in the possibility of estimating the expected value of a variable. Therefore, on this basis, expected values of deflection were estimated under the influence of a specific load depending on the number of cycles (Table 3). Increasing the value of the acting force may lead to an increase in sample deflection in each of the studied groups.

**Figure 8.** Relationship of deflection from the load values in particular groups: K, R1, R2.


**Table 3.** Expected values of deflection depending on the specific load (mm).

#### **4. Discussion**

Fiber-reinforced composite materials are used as alternative materials for metal restorations. Modifications of composite resins, through various types of fibers improving the mechanical properties of restorations, allow for their use not only in prosthetic crowns, but also in solid partial dentures (FPD)—e.g., in adhesive bridges [31]. These additions differ from the total permanent dentures by the replacement of the crowns, on which the prosthetic bridge structure is based, with crown inserts, i.e., onlay and inlay. This construction of prosthetic bridges is an alternative for patients who do not agree to a complete permanent replacement due to the need of significant tooth tissue reduction, or when minimal tooth reduction is possible [32,33]. In the construction of bridges, the connectors between the individual elements of the bridge are most vulnerable to destruction. Solid partial dentures of fiber-reinforced composite resins are made by placing the fiber in the structure and surrounding it with a composite resin. During their fatigue cyclic tests, Lobhauer et al. [34] observed the slow spread of cracks in brittle materials, such as composite resins. The literature lacks studies describing polyaramid fiber due to the fact that this is a new material in dentistry. Research results derived from the study on composites reinforced with polyaramid fibers provided by Selvaraj et al. acknowledge that the addition of polyaramid into composite impact/increases composite strength [35].

Previous studies on dental composites used to make prosthetic bridges were mainly based on standardized samples. Meanwhile, such samples subjected to cyclic loads show lower values of bending strength (30–50%) than those obtained in static studies and are considered more sensitive in assessing the effectiveness of clinical materials [29]. When testing the strength of dental composites, Papadogiannis et al. [28] stated that fatigue strength is associated not only with the type of filler, but also with the resin matrix. Kuroda et al. [27] used standardized samples with dimensions of 3.0 × 4.0 × 40 mm in their tests, which underwent cyclic fatigue loads with a force of 100 N. They noted that the strength of fiber-reinforced dental composites increases the bending strength. They noted better strength of composites reinforced with glass fiber under the influence of increasing force compared to the unreinforced control sample. However, these studies did not take into account the physiological interactions between the opposing teeth during chewing. In a study by Nobuhisa et al. [26], who analyzed the FEM model of a three-point conventional bridge reinforced with fiber glass, the authors applied a load of 629 N in the rebuilding the lack of the first molar in the mandible. The significant improvement in connector rigidity under vertical load conditions causing twisting and bending movement was stated. As a consequence, the stiffness of connections between individual elements of the sample structure improved, which significantly reduced the deflection of the span because the stresses generated by the vertical load were transferred to the reinforcing fiber. However, no relationship was found between the number of fatigue cycles and bending strength. Similar dependences on the better strength of samples reinforced with polyaramide fiber were recorded in the present research, whereby the actual strength of the manufactured bridges was tested with a strictly established number of loading cycles. Although the experimental groups required significantly lower force in order to break, the destructive force was still in the reported range of habitual and maximal biting forces in molar teeth, which are 300 N and 500 N, respectively [36,37].

The authors are aware that the tests were carried out in laboratory conditions and do not perfectly reflect the conditions in the oral cavity (the authors did not use any bonding system between sample

and the chromium–nickel model); nevertheless, the demonstrated differentiation indicates the validity of further research in this direction.

#### **5. Conclusions**

Insertion of numerous and fine polyamide fiber fragments into the composite resin may contribute to the elimination of cracks in the material itself, limiting their propagation in restorations. Such a reinforcement could contribute to extending the life of these additions. The improvement of mechanical properties of dental composites, including flexural strength, is of particular importance in the rehabilitation of patients with functional disorders of the masticatory apparatus, where chewing forces are very high.

Studies based on the analysis of anatomically reversible elements require the introduction of laboratory validation to allow comparing test results. The conditions for dental materials in the oral cavity are difficult. In addition, the large anatomical variability of elements reconstructed with the use of dental materials excludes the possibility of developing a conclusion on the basis of standard solutions.

The deflection of adhesive bridges predicted in the study indicates the need for further research in this field with the use of finite element modelling. The polyaramide fiber used in the research can replace the previously used glass fiber to strengthen the construction of prosthetic bridges.

**Author Contributions:** L.S.: conceptualization, data curation, formal analysis, methodology, writing—original draft; A.K.: conceptualization, data curation, formal analysis, methodology, writing—original draft; E.W.: data curation, methodology, resources, writing—original draft; J.B. (Justyna Batkowska): data curation, formal analysis, visualization, writing—original draft; T.W.: conceptualization, data curation, methodology, visualization, writing—review and editing; D.W.: data curation, formal analysis, methodology, visualization, writing—original draft; J.B. (Janusz Borowicz): conceptualization, formal analysis, supervision, writing—review and editing. All authors have read and agreed to the published version of the manuscript.

**Funding:** This research received no external funding.

**Conflicts of Interest:** The authors declare no conflict of interest.

#### **References**


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## *Article* **Fracture Properties of Concrete in Dry Environments with Di**ff**erent Curing Temperatures**

**Zhengxiang Mi 1,2, Qingbin Li 2, Yu Hu 2,\*, Chunfeng Liu <sup>3</sup> and Yu Qiao <sup>3</sup>**


Received: 12 June 2020; Accepted: 7 July 2020; Published: 9 July 2020

**Abstract:** This paper investigated the fracture properties of concrete in dry environments with different curing temperatures (5, 20, 40, and 60 ◦C). For each curing condition, the key fracture parameters of concrete were tested using wedge splitting specimens at five different ages (3, 7, 14, 28, and 60 d). The results show that in dry environments, the effective fracture toughness and fracture energy of concrete exposed to elevated temperatures increased at a relatively high growth rate at an early age. Nevertheless, the growth speed of effective fracture toughness and fracture energy decreased more quickly at elevated temperatures in the later stages. As a result, the concrete cured at higher temperature exhibited lower ultimate values of fracture parameters, and vice-versa. Namely, a temperature crossover effect was found in the effective fracture toughness and fracture energy of concrete under dry environments. Considering the early growth rate and ultimate values of fracture parameters, the optimum temperature suitable for concrete fracture properties development under dry condition was around 40 ◦C.

**Keywords:** concrete; fracture properties; dry environments; different curing temperatures; temperature crossover effect

#### **1. Introduction**

Concrete is used in a wide variety of structures, which are usually exposed to changing temperature and moisture content. The development of concrete mechanical and physical characteristics is significantly affected by the curing conditions. Harsh environmental conditions not only make the on-site casting and quality control process difficult but also accelerate setting, promote uneven distribution, reduce the later-age strength, increase the likelihood of cracking both before and after hardening and deteriorate the durability [1–5]. In order to construct the concrete structures in extreme weather conditions (such as extremely hot and dry conditions, cool and damp conditions), we must pay high attention to the concrete deterioration caused by harsh climate.

As we all know, the temperature is a very important factor that affects the concrete properties, and extensive studies have been conducted thus far; but the majority of those studies were related to the influence of temperature on concrete mechanical parameters, such as compressive strength [6], flexural-tensile strength [7], splitting-tensile strength [8], and Young's elastic modulus [9]. In general, temperature has a double effect on the development of concrete mechanical parameters. A higher curing temperature can significantly promote the development of concrete elastic modulus and strength in the early stages, but it also decreases the ultimate values of elastic modulus and strength. This behavior is called the temperature inversion phenomenon or temperature crossover effect [10]. Furthermore, apart from the Young's elastic modulus and strength, the elevated curing temperature

will also deteriorate the microstructure and durability of concrete. For example, Sharp et al. [11] found that the microstructures of concrete exposed to 60 ◦C exhibited larger apparent porosity than those exposed to 10 ◦C after one year, which was a significant reason that the concrete cured at 60 ◦C exhibited the lower ultimate strengths. Jiang et al. [12] demonstrated that with the increasing of curing temperature, the autogenous shrinkage of concrete increases, and the cracking resistance of concrete becomes worse. Especially, under high-temperatures curing conditions, the resistance to chloride irons permeability was weakened [13]; further, after prolonged exposure, the concrete exposed to higher temperature showed anomalous expansion and accompanied by micro-cracks, leading to a serious deterioration of the concrete structure durability [14].

The humidity is another important environmental factor affecting the concrete properties, and the concrete is usually placed at reduced humidity in engineering practice, resulting in that the free water inside the concrete evaporates into the surrounding at lower humidity. An insufficient moisture supply owing to this evaporation and diffusion not merely worsens the hydration reaction and microstructure of concrete but also drastically deteriorates the mechanical and fracture properties [15]. For example, Flatt et al. [16] theoretically proved that C3S stop hydrating when the internal relative humidity is less than 80%. Sagrario et al. [17] noticed that as the curing humidity decreases, the length of the chemical reaction chain forming the C-S-H gel increases, leading to a lower degree of hydration. Nawa et al. [18] reported that the strength at 60% relative humidity was 41% smaller than that in water after 7 d. Clearly, mechanical and physical properties of concrete measured under saturation conditions are larger than that of concrete under the realistic circumstance, and the influence of ambient humidity should be taken as seriously as that of temperature.

The fracture characteristics of concrete can dramatically affect the structural response of concrete members in the harsh environment, and the influence of ambient temperature and humidity on the fundamental fracture parameters of concrete should be studied. This is useful for a comprehensive understanding of the fracture properties of concrete exposed to realistic conditions, which is the key premise for assessing crack stability and predicting concrete structure failure based on fracture mechanics. However, so far, less information is available regarding the influence of environmental conditions on fracture properties of concrete, and some of the conclusions are contradictory. For instance, Buyukozturk et al. [19] clarified that due to the action of the pore water pressure inside the specimen, the fracture energy of concrete increases with reducing moisture. Mi et al. [20] observed that the value of fracture energy of concrete under dry conditions was smaller than that under saturated curing conditions. Besides, Yu et al. [21] obtained fracture energy gain data for concrete exposed to 14, 23, and 35 ◦C, and they clarified that concrete fracture energy improved dramatically with increasing temperature. Unfortunately, they did not investigate the fracture energy of concrete after 20 d. Analogously, Li et al. [22] observed that the fracture energy of full-graded concrete cast in hot weather was larger than that of concrete cast in cold weather from 28 d to 180 d. From the above results, it can be seen that fracture energy of concrete increased with increasing curing temperature. Conversely, Huang [23] clarified that the fracture energy of concrete stored at elevated temperature in summer was much smaller than that at 20 ◦C after 56 d, and the former was only half of the latter. This indicated that the elevated temperature deteriorated the fracture properties of concrete in later ages, which was in line with the results observed by Mi et al. [24]. Further, Mechtcherine [25] demonstrated that concrete fracture energy was less dependent on the temperature from 2 ◦C to 50 ◦C. It is clear that there is no consensus concerning the effect of ambient conditions on concrete fracture properties. What's more, there are very litter or no studies explaining the influence of temperature and humidity coupling on the fracture properties of concrete.

Therefore, the current paper presents the investigation of the fracture characteristics of concrete in dry environments with different curing temperatures. The fracture properties of concrete including fracture toughness and fracture energy were measured under a constant relative humidity of 50% with four different temperatures. The information about the fracture properties of concrete under hot-dry (or cold-dry) environments was not available in previous publications. The experimental

data is helpful for a deeper and comprehensive understanding of the fracture properties of concrete under extreme climate environment, which can greatly enhance the numerical simulation of cracking behavior of concrete members in harsh environments and promote the development of cracking design specifications for concrete structures.

#### **2. Experimental Program**

#### *2.1. Materials and Mix Proportions*

The cement employed was Portland cement with the grade of P.I 42.5 satisfying the specifications of ASTM C150. The bulk density of cement equals 1450 kg·m−3. The fly ash accounted for 25% of the cementitious materials, which is mainly for saving cement and reducing the heat of hydration. The chemical composition of cementitious materials was measured using XRD, and the results were given in Table 1. The coarse aggregate was crushed basalt, physical properties and gradation of which met the specifications of ASTM C33/C33M. The fine aggregate used was natural river sand, and its fineness modulus was 2.52. Further, the sand accounted for 39% of the total aggregate by weight, and this proportion remained unchanged. After comprehensively considering the fluidity and strength, the effective water to binder ratio of 0.4 and a 468 kg·m−<sup>3</sup> cementitious content was selected. The concrete mix proportions were 1:0.40:1.50:2.34 (cementitious:water:sand:gravel). The slump of the concrete is 85 mm. Moreover, an external vibrator was utilized for concrete compaction.

**Table 1.** Chemical composition of cementitious materials (weight/%).


It is noted that before mixing, all the mixture constituents for concrete were reserved in the chambers with humidity and temperature previously set at 50% and 20 ◦C for not less than one day to ensure all the specimens having the same initial condition.

#### *2.2. Curing Condition*

In this study, the temperatures used were 5 ◦C, 20 ◦C, 40 ◦C, and 60 ◦C, corresponding to cold, normal, hot, and extremely hot weather, respectively. As for the curing humidity, it was set at 50% RH (relative humidity) for each temperature. The temperature and humidity during curing were controlled automatically using specialized equipment, and the control accuracy of temperature and relative humidity was ±2 ◦C and ±2% RH, respectively. Five test ages were designed for each curing condition, and three replicates fracture samples were poured for each age. The experimental data were evaluated and analyzed using averages of three identical samples. However, it should be emphasized that for a given variable, if the value of a specimen deviated from the average by more than 10%, then an additional specimen was conducted to improve the credibility of the results.

#### *2.3. Specimen Preparation*

The notched wedge splitting specimens are employed for investigating concrete fracture properties. The advantages of using it are that its self-weight has an insignificant effect on test results and the ratio of ligament length to sample size is considerably large [26]. However, to avert the unexpected failure of the sample and improve the stability of the sample, the double supports were used instead of the traditional sample with one-line support [27]. The test principle is shown in Figure 1a, where the vertical load is converted into the horizontal splitting force through a wedge-loading fixture. In the present test, the wedge angle is 15◦.

**Figure 1.** The fracture test specimen: (**a**) principle of wedge splitting test method; (**b**) the dimension of the specimen.

The dimension of the wedge splitting specimen was 120 × 300 × 300 mm (thickness × width × height), as shown in Figure 1b. The height of the specimen was greater than the recommended height in the specification in order to measure the fracture energy and fracture toughness independent of specimen size. The key purpose of this research is to investigate fracture properties of concrete under dry environments instead of the size effect, so a single specimen was adopted. In the concrete casting process, a steel plate was inserted into the tensile surface of the sample to form a notch. The thickness and initial length of the notch were 2 mm and 120 mm, respectively.

The specimen was poured in steel mold designed and manufactured according to experimental requirements, and each specimen was vibrated for about 1 min on a vibrator to eliminate air bubbles and to increase the density. After pouring, the sample was kept at natural conditions, and the pouring surface was covered with a layer of waterproof plastic film to prevent desiccation. Approximately 4.5 h after setting, the steel plate was cautiously pulled out. Subsequently, the sample was moved into the corresponding environmental box and kept there for 24 h. Afterward, the steel mold was removed and the concrete transferred into the same box again, where it was maintained until the test age. This process guarantees that the concrete has sufficient strength to avert damaging when placed, moved, and de-mold. Besides, this procedure also ensures that when concrete fracture properties begin to progress, all specimens achieve the same characteristics for better comparison. On the contrary, if the specimen is immediately transferred to the corresponding environment box after casting, the concrete already has different hydration degrees at the onset of its fracture properties development, resulting in difficulties in comparison [9]. Finally, this process is closer to practice, where concrete structures are usually cured for some time before being exposed to the working environment.

In addition to fracture specimens, additional cubes with a side length of 100 mm were cast using the same batch of concrete to determine the elastic modulus and strength for each research temperature. The cubes' size and test procedure are consistent with the specification of DL/T 5150-2001 [28]. Although the size of the specimens differs with that recommended by ASTM C 496, test specification conforms to the ASTM C 496 [29]. For each investigated curing condition, eight identical cubes were poured, four of which were used to measure the strength, and the remaining four were employed to evaluate the elastic modulus.

#### *2.4. Testing*

The wedge splitting specimens were tested according to ASTM E 1820 on an Instron servo-hydraulic testing machine. The fracture experiment set-up is shown in Figure 2. During loading, the main deformation of crack mouth opening displacement (CMOD) was monitored in real-time by a clip-gauge mounted at the mouth of the notch. The capacity and accuracy of the clip-gauge were 5 mm and ±0.2 μm, respectively. The tests were controlled by the CMOD. To start the test, the specimen was preloaded with 0.2 kN. Subsequently, the Instron testing machine was driven by an initial displacement criterion of 0.04 mm·min−1. In the post-peak range, the loading rate was gradually increased up to 0.08 mm·min<sup>−</sup>1. The Instron testing machine stops the loading when the load drops to 95% of the peak load. The main reason for choosing this loading scheme was to make the specimen failed within 30 to 40 min. During the test, the information on load and CMOD was collected twice per second. Due to the use of a testing machine with large stiffness, the crack propagated steadily during the test.

**Figure 2.** The fracture experiment set-up.

The elastic modulus and compressive strength at 28 d were measured using a universal testing machine with a range of 100 tons and a linear variable displacement transducer. To start the test, the sample was preloaded with 5 kN. Thereafter, the testing machine was driven by the displacement criterion of 0.2 mm·min−1. The elastic modulus and compressive strength were determined by averaging the test results of the four identical cubes. Table 2 presents the measured values of elastic modulus and compressive strength under different curing conditions.

**Table 2.** Mechanical properties of concrete under different temperature at 28 d.


#### **3. Results**

#### *3.1. Load-CMOD Curves*

Figure 3 presents the complete load-CMOD curves for concrete in a dry environment with different temperatures, which was used to obtain the peak load and calculate the fracture parameters such as fracture energy. Owing to the heterogeneity of concrete, the load-CMOD curves measured by the same set of specimens are different. For each investigated temperature and age, only the averaged curve of the same group of specimens was presented to avoid confusion. The overall load-CMOD curve is smooth, demonstrating that the unloading process was slow after the peak load and the test was carried out in a stable test state.

**Figure 3.** Load-crack mouth opening displacement (CMOD) curves under dry environments with different temperatures: (**a**) at 5 ◦C; (**b**) at 20 ◦C; (**c**) at 40 ◦C; (**d**) at 60 ◦C.

As presented in Figure 3, at the beginning of loading, there was a linear relationship between load and CMOD, indicating that concrete was in the stage of linear elasticity. The specimen did not undergo significant nonlinear deformation before the crack initiation or reaching the cracking load. For each curing condition, the initial slope increased as the concrete become older, which was attributed to the increase of the concrete strength and stiffness with time elapsing. Further, with the increase of curing time, the peak load grew markedly under each investigated curing temperature. For instance, for the specimens exposed to 5 ◦C, the peak load was 2.683 kN after 3 d, 2.906 kN after 7 d, 3.156 kN after 14 d, 3.427 kN after 28 d, and 3.911 kN after 60 d. The peak load increased by 14.1% from 28 d to 60 d, indicating that there was still a lot of un-hydrated cement inside the concrete under this curing condition. A similar phenomenon was also observed when the concrete kept at high temperature, but the rate of growth was lower than that of the concrete exposed to low temperature. Typically, for the concrete exposed to 60 ◦C, the peak load improved from 3.565 kN after 3 d of hydration to 4.491 kN after 60 d of hydration, where the peak load hardly increased beyond 28 d hydration.

Figure 3 also reveals that the peak load improved with the growing curing temperature during an early age, but this situation did not hold in later ages, with the greatest ultimate values of peak load corresponding to the specimens kept at 40 ◦C. In detail, at 3 d, the peak load improved from 2.683 kN at 5 ◦C to 2.723 kN at 20 ◦C, 3.366 kN at 40 ◦C, and finally to 3.565 kN at 60 ◦C with a final ascent of 32.9%. However, at 60 d, the peak load of concrete stored at 5 ◦C, 20 ◦C, 40 ◦C and 60 ◦C was 3.911 kN, 4.092 kN, 4.781 kN, and 4.491 kN, respectively. The peak load of concrete exposed to 60 ◦C was only 12.9% and 8.9% greater than that of concrete kept at 5 ◦C and 20 ◦C, respectively, and even was 6.5% smaller than that of concrete exposed to 40 ◦C.

Besides, the post-peak curvature of the curve, which represented the brittleness of the concrete, was related to the curing condition and age. This can be attributed to the improvement of the interfacial transition zone (ITZ), in which the stress concentration was decreased and more new hydration products were accumulated as the curing age was prolonged. Consequently, the bridging effect of the aggregate decreased and the fracture mode of concrete gradually transformed from the interfacial transition zone to aggregate being broken. Particularly, with further strengthening of ITZ at 60 d of curing age, the fracture path at high temperature was directly through the aggregate, which was similar to that of 28 d. As a result, the characteristic of load-CMOD curves in the unloading zone at 28 and 60 d followed the same pattern.

#### *3.2. Fracture Energy*

Fracture energy (*G*F) represents the energy required to generate new cracks per unit area during crack propagation, which is applicable to energy release rate criterion. The simplest method to calculate fracture energy is the fracture work method. Thus, the fracture energy can be evaluated using Equation (1). Note that for a specimen that is not large, the effect of its weight is negligible. Thus, the own weight of the specimen was ignored when evaluating the fracture energy. Besides, the fracture experiment can never reach the zero load level [30]. Therefore, the far-tail constant value was usually used to determine the true fracture energy [31]. In the present paper, the fracture experiments did not stop before the unloading zone of the load-CMOD curve appeared a full tail. The part that was not measured was extended inversely by the fit-best expression in the evaluation of *G*F.

$$G\_{\rm F} = \frac{W}{B\left(h - a\_0\right)}\tag{1}$$

where *W* is the fracture work which equals the area below the measured splitting force against CMOD curve; *a*<sup>0</sup> is the length of pre-crack; *B* and *h* represent the thickness and height of wedge splitting specimen, respectively.

Figure 4 shows the fracture energy for concrete exposed to dry environments with different temperatures as a function of the age. In this figure, each point denotes the mean value of fracture energy evaluated by the three companion wedge splitting specimens, and the error bars are not presented to avoid displaying a pell-mell plot. In any case, the difference between one replicate and the mean of the same group was not more than 10% concerning the mean.

**Figure 4.** Fracture energy for concrete exposed to dry environments with different temperatures.

It can be seen from Figure 4 that the concrete fracture energy increased rapidly during the first 3 d, irrespective of the curing conditions investigated. Over time, fracture energy kept increasing at a reduced rate and gradually stabilized when the age was large enough. For instance, in the case of 5 ◦C, the fracture energy was 130.4 N·m−<sup>1</sup> after 3 d of hydration, 157.2 N·m−<sup>1</sup> after 7 d of hydration, 174.4 N·m−<sup>1</sup> after 14 d of hydration, 189.4 N·m−<sup>1</sup> after 28 d of hydration, and 204.7 N·m−<sup>1</sup> after 60 d of hydration. The value of fracture energy at 60 d was 8.1% greater than that at 28 d. It means that the fracture energy continuously improved with the increase of age and the growth rate was still relatively great. In other words, in the case of 5 ◦C, 60 d of curing time were not sufficient to make the concrete achieve the stable state and there was a lot of un-hydrated cement inside the concrete. A similar phenomenon was observed under elevated temperature, but the fracture energy tended to reach a plateau value in a shorter time. Typically, for 60 ◦C, fracture energy was 173.7 N·m−<sup>1</sup> after 3 d of hydration, 194.4 N·m−<sup>1</sup> after 7 d of hydration, 207.7 N·m−<sup>1</sup> after 14 d of hydration, 220.8 N·m−<sup>1</sup> after 28 d of hydration, and 221.7 N·m−<sup>1</sup> after 60 d of hydration. Obviously, in this case, the growth rate of fracture energy decreased significantly after 14 d and even stopped increasing after 28 d. At high curing temperature, the shorter stable time is mainly due to a higher temperature accelerating the hydration reaction rate of cement.

Figure 4 also reveals that in the early stage, the fracture energy of concrete increased significantly with the increasing temperature; but after 28 d of hydration, the curing temperature had the opposite effect on the fracture energy, that is, the larger later-age value of fracture energy was measured by concrete exposed to the lower temperature, and vice-versa. For instance, at 3 d, the fracture energy of concrete exposed to 60 ◦C was larger by 10.9%, 20.9%, and 32.9% compared to that of specimens exposed to 40 ◦C, 20 ◦C and 5 ◦C, respectively. However, for the specimens kept at elevated temperatures, the growth rate of the fracture energy decreased faster with the extension of curing time. As a result, the fracture energy difference between each temperature gradually decreased with time elapsing. In detail, after 28 d, the fracture energy at 60 ◦C reached 220.8 N·m<sup>−</sup>1, which was 4.7%, 10.2%, and 16.6% higher than that of concrete at 40 ◦C, 20 ◦C and 5 ◦C, respectively. Further, after about 41 d of hydration, the fracture energy curves of concrete kept at 60 ◦C and 40 ◦C crossed each other, indicating that from 41-day onwards, the fracture energy of concrete exposed to 40 ◦C was larger than that of concrete exposed to 60 ◦C. Moreover, it can be easily inferred from the growth rate of fracture energy at 20 ◦C and 5 ◦C that, with the further increase of age, the curves at these two temperatures must intersect that at a higher temperature, and the values of fracture energy at these two temperatures certainly exceed that of concrete kept at the higher temperature. The maximum fracture energy limit was measured by concrete cured at 5 ◦C. Accordingly, the fracture energy of concrete cured at different temperatures has a temperature crossover effect. Considering the early growth rate and later-age value of fracture energy, around 40 ◦C was the optimum curing temperature propitious to the fracture energy development of concrete in dry environments. This optimal temperature was related to the faster hydration rate and higher hydration degree of the cementitious material, as well as to the microstructure of dense concrete with finer pores distribution [32].

#### *3.3. E*ff*ective Fracture Toughness*

Fracture toughness characterizes the cracking resistance of concrete, which is usually used for the stress intensity factor criterion. A smaller value of fracture toughness indicates that the concrete is apt to fracture abruptly before the occurrence of obvious unrecoverable deformation. The effective fracture toughness (*K*IC) is calculated with the following equation [33]:

$$K\_{\rm IC} = \frac{F\_{\rm symax}}{B\sqrt{h}} \cdot \frac{3.675[1 - 0.12(a - 0.45)]}{(1 - a)^{3/2}} \; , \; a = \frac{a\_{\rm c}}{h} \; , \tag{2}$$

where *F*smax is the horizontally component of peak load; the meaning of *h* and *B* is the same as in Equation (1); *a*c is the critical effective crack length, which is evaluated by the following equation [33]:

*a*<sup>c</sup> = (*h* + *h*0) ⎛ ⎜⎜⎜⎜⎝ 1 − 13.18 9.16 + *CMOD*c·*E*·*B*/*F*smax ⎞ ⎟⎟⎟⎟⎠ <sup>−</sup> *<sup>h</sup>*<sup>0</sup> , (3)

where *h*<sup>0</sup> is the thickness of clip gauge holder; *CMOD*<sup>c</sup> is the crack opening displacement corresponding to *F*smax; *E* is the tensile Young's modulus.

Figure 5 depicts the test results of effective fracture toughness against age under four curing conditions investigated. Although each point in Figure 5 denotes the mean values of the three replicate samples, the error bars are not shown again for the sake of brevity. Interestingly, the scattering of data observed in effective fracture toughness was smaller when compared with the case of fracture energy.

**Figure 5.** Effective fracture toughness for concrete exposed to dry environments with different temperatures.

During the unstable propagation, redundant cracks appeared and the propagation path of the main crack became more random; however, the energy consumed by these extra cracks was difficult to calculate, so these energies were ignored when evaluating fracture energy. As a result, the fracture energy produced larger randomness.

As can be presented in Figure 5, for each curing condition investigated, the effective fracture toughness was closely related to age and its value increased with age, which was the same as fracture energy with age. In the beginning, effective fracture toughness improved significantly with the growing age, because at this stage the reaction rate of cement hydration was quite fast. The growth rate increased by degrees and then reduced after maximizing. Subsequently, effective fracture toughness increased slightly with age at a reducing rate and finally reached a constant as the age was further extended. Typically, for 5 ◦C, the effective fracture toughness continuously increased from 0.889 MPam0.5 after 3 d of hydration to 1.043 MPam0.5 after 7 d of hydration, 1.113 MPam0.5 after 14 d of hydration, 1.204 MPam0.5 after 28 d of hydration, and finally 1.306 MPam0.5 after 60 d of hydration, with a significant increase of 0.417 MPam0.5 or 46.9%. Similarly, for 60 ◦C, the effective fracture toughness was 1.164 MPam0.5 at the age of 3 d, 1.304 MPam0.5 at the age of 7 d, 1.395 MPam0.5 at the age of 14 d, 1.490 MPam0.5 at the age of 28 d, and 1.509 MPam0.5 at the age of 60 d. The fracture toughness at the age of 60 d was only 1.3% greater than that at the age of 28 d. This also demonstrated that it takes a shorter time for the effective fracture toughness to reach the constant value under higher temperatures.

Figure 5 also shows that, as observed in the fracture energy, there was also a temperature crossover effect in the effective fracture toughness of concrete in dry environments. As the curing temperature increased, concrete effective fracture toughness improved dramatically during the early ages, but this situation did not hold after 28 d, in which greater ultimate value of effective fracture toughness was acquired from the concrete kept at a lower temperature, and vice-versa. Specifically, at the age of 3 d, the fracture toughness of concrete kept at 60 ◦C was 1.164 MPam0.5, which was 5.7% greater than at 40 ◦C, 21.1% greater than at 20 ◦C, and 30.9% greater than at 5 ◦C. Evidently, in the early days, higher temperatures promoted the rapid development of concrete effective fracture toughness. However, at the elevated temperature, the increasing rate of fracture toughness of concrete decreased faster with

the extension of curing time. As a result, a larger later-age value of the effective fracture toughness was measured by concrete kept at a lower temperature. For example, for 60 ◦C, effective fracture toughness increased very litter beyond 14 d and almost reached a plateau value. In contrast, the specimens kept at a lower temperature (5 ◦C and 20 ◦C) had a higher increasing rate of effective fracture toughness. Representatively, at the age of 60 d, the fracture toughness of the concrete kept at 60 ◦C was only greater 15.5% and 4.4% when comparing concrete kept at 5 ◦C and 20 ◦C, respectively, and was smaller 2.3% when comparing concrete kept at 40 ◦C. In other words, the difference in the fracture toughness between different temperatures was decreasing as the curing age was prolonged. The fracture energy curves of concrete kept at 40 ◦C and 60 ◦C crossed each other at about 43 d, demonstrating that the effective fracture toughness of concrete cured at 40 ◦C was larger than that of concrete at 60 ◦C beginning from the 43rd day. Considering the increasing rate of effective fracture toughness at 5 ◦C and 20 ◦C, we can conclude that with the further extension of curing time, the curves at 5 ◦C and 20 ◦C will doubtlessly cross the curves at the higher temperature. This clearly showed that the optimum temperature propitious to the fracture energy development of concrete in dry environments was about 40 ◦C, which was the same as that for fracture energy.

#### **4. Discussion**

As mentioned earlier, during the first 3 d, the *G*<sup>F</sup> and *K*IC of concrete improved quickly with increasing age owing to the fast hydration rate of cement in this stage. The initial growth rate of concrete fracture parameters increased progressively and then reduced after maximizing, accompanying a steeper curve. As the curing temperature increased, the duration for this stage became shorter. In the later stage, the rate of hydration went down, resulting in that the concrete fracture parameters developed overtime at a smaller growth rate. In general, there should be the third stage, in which the fracture parameters of concrete tended to their stable values, as the age was further extended. However, at low temperatures, the third stage was not seen or was not notable because this study only investigated the age of up to 60 d, meaning that there was a lot of un-hydrated cement inside the concrete under these curing temperatures. This variation of concrete fracture parameters regarding the age was consistent with other independent investigations [34]. For example, Beygi et al. [35] found that with the increase of test age from 3 to 90 d, *G*<sup>F</sup> and *K*IC increase from 0.961 MPam0.5 to 1.448 MPam0.5 and from 99.7 N·m−<sup>1</sup> to 126.5 N·m<sup>−</sup>1, respectively. Lee et al. [36] demonstrated that the fracture energy shows a rapid increase during earlier and then starts to converge to a certain extent of 173.1 N·m−<sup>1</sup> at 28 d. The main reason for the increase of concrete *G*<sup>F</sup> and *K*IC with age was the strengthening of the ITZ between the paste and aggregate. ITZ is the most sensitive area inside the concrete, where the largest number of micro-cracks are formed [37]. In other words, the failure of concrete usually depends on the strength of its weakest area. At an early age, there was a lot of un-hydrated cement in the ITZ, and it resulted in high porosity in this zone [38]. The bridging effect of the aggregate was very strong causing the cracks to propagate along with the interface, and a large number of aggregates were extracted from the matrix, as presented in Figure 6a. As the reaction of cement hydration progressed, an increasing number of hydration products accumulated in the ITZ and the ITZ pores were stuffed with hydration products, indicating that the size and content of pores in ITZ and cement paste decreased. The ITZ and cement paste also became stronger. As a result, the bridging effect of the aggregate became weaker and a growing number of aggregates were broken. The fracture mode under load also changed from around aggregate or bond zones through the aggregate directly, as presented in Figure 6b. Further, the aggregate strength was greater than that of the ITZ and cement paste, demonstrating that more energy was needed in order to get over the enhanced ITZ and cement paste and the aggregates. Consequently, the greater values of fracture energy and effective fracture toughness were obtained at later ages. However, as the strength of the ITZ and cement paste further improved, the fracture pattern did not alter dramatically, where an only increasing number of aggregates were extracted. Therefore, effective fracture toughness and fracture energy did not increase significantly, but gradually tended to their stable values.

(**a**) (**b**) **Figure 6.** The fracture surface of concrete at different ages: (**a**) at 3 d; (**b**) at 60 d.

More importantly, as is evident from the data in Figures 5 and 6, the effective fracture toughness and fracture energy of concrete cured at elevated temperatures developed at a faster speed at the early stage. For instance, after 3 d of hydration, the *G*<sup>F</sup> and *K*IC of concrete cured at 60 ◦C was 20.9% and 21.1% larger respectively than that of concrete cued at 20 ◦C. This promotion effect of temperature on the early fracture parameters of concrete is related to the accelerated hydration reaction rate of cement, and the finer pore distribution and denser microstructure [39]. Over time, however, the rate of growth of the fracture parameters of concrete kept at high temperatures also decreased at a higher rate, which resulted in these samples obtained the smaller ultimate values of *G*<sup>F</sup> and *K*IC than those of samples kept at reduced temperatures. For example, after 60 d of hydration, the difference in the *G*<sup>F</sup> and *K*IC between the samples exposed to 60 ◦C and 20 ◦C was reduced to 4.2% and 3.3%, respectively. Even the values of *G*<sup>F</sup> and *K*IC at 60 ◦C were less than that at 40 ◦C beginning from the 43rd day and 41st day, respectively. Further, the growth rate of fracture parameters at 5 ◦C and 20 ◦C also indicated that with the further extension of curing time, the curves at 5 ◦C and 20 ◦C will doubtlessly cross the curves at the higher temperature. Namely, although a higher temperature accelerated the rate of hydration reaction of cement and promoted the development of *G*<sup>F</sup> and *K*IC in the early stage, it harmed the development of the ultimate value of fracture parameters of concrete. The *G*<sup>F</sup> and *K*IC of concrete showed a temperature crossover effect, and the optimum temperature propitious to their development was approximately 40 ◦C. Such a trend was in line with previous observations found by other scholars with a slightly different method. Oswaldo et al. [40] claimed that curing at high temperatures was conducive to rapid strength gain, but at later ages, the greatest strength was measured by concrete cured at 20 ◦C. Boubekeur [41] found that the concrete compressive strength cured at 50 ◦C was 7% smaller than that of concrete stored at normal temperature after 28 d. Lee et al. [42] demonstrated a 15% reduction in the concrete compressive strength cured at 60 ◦C after 14 d compared to that cured at 40 ◦C. Nasir et al. [43] clarified that after 28 d the tensile strength improved with the pouring temperature up to 32 ◦C, but decreased these properties as a further increase in temperature.

The adverse effect of high-temperature on the later-age fracture properties of concrete makes us wonder: What happened at the microscopic level at a high-temperature? We concluded that this adverse effect of high-temperature on the effective fracture toughness and fracture energy of concrete was related to some different mechanisms. Firstly, this phenomenon is caused by the non-homogeneous dispersion of hydration products inside the pores of the hardened cement paste [44]. The hydration reaction of cement was accelerated at elevated temperature, so there was neither sufficient time for the hydration products to disperse surrounding nor sufficient time for them to precipitate and distribute evenly. Consequently, more and more hydrated products were concentrated surrounding the hydrating or un-hydrated particles of cement, resulting in a more non-homogeneous distribution of hydrated products and causing the concrete to become a porous structure [45]. Secondly, the change in the C-S-H apparent density at high-temperature was another important reason for the decrease of later-age values of concrete fracture parameters [46]. Gallucci et al. [47] clarified that there was a remarkable enhancement in the C-S-H apparent density in the situation of elevated temperature, which came from the different assembly of the C-S-H. Owing to this increase, the filling space of C-S-H was reduced, leading to that the microstructure of the concrete was much coarser and porous, which had a detrimental effect on the concrete fracture parameters. Other mechanisms that resulted in the temperature crossover effect were cumulative of compact and low infiltration hydration products surrounding the cement particles and the delayed diffusion of the hydrate in cement secondary hydration reaction [40]. Namely, the elevated temperature might have resulted in the rapid evaporation of uncombined water in the concrete, which restrained the further spread of cement and limited the progress of hydration reaction, thereby restricting the forming of more C-S-H and the development of fracture parameters. Finally, the deviation of thermal expansion coefficients among the concrete components was also related to the negative effect [48]. In the situation of high-temperature, the concrete would expand in volume. The deformation of each constituent of concrete is not equal because of the difference in thermal expansion coefficients, which will cause a great stress build-up in the air void concentration area or the ITZ. When this stress exceeds the matrix tensile strength, micro-cracks begin to emerge in weak areas, especially those at the interface of the cement paste and the aggregate [49]. What is worse, due to the capillary action, these micro-cracks will become an ideal moisture path and increase the mobility of moisture inside the concrete, which can generate a pressure of expansion of the water vapor due to swelling of water vapor more quickly than getting away from the sample, leading to further micro-cracking. As a result, the fracture mode of concrete at elevated temperature has altered due to higher porosity and micro-cracking, which means that a decreasing number of aggregates were broken, as presented in Figure 7. In short, all the above factors, in particular the high porosity, together deteriorated the properties of concrete at later-age and induced a reduction in the ultimate values of the breaking fracture and the fracture toughness of the concrete at high temperature. On the other hand, concrete stored under low temperature (5 ◦C and 20 ◦C) continuously underwent the favorable effect, with a very slow hydration rate, which left the dissolved ions enough time to diffuse and resulted in a more uniform hydration products distribution and lower coarse porosity [50].

**Figure 7.** The fracture surface of concrete at different temperatures: (**a**) at 5 ◦C; (**b**) at 60 ◦C.

#### **5. Conclusions**

In the current research, the fracture properties of concrete in dry environments with different temperatures were investigated using the wedge splitting specimens. Four different temperatures (5 ◦C, 20 ◦C, 40 ◦C, and 60 ◦C) were considered, which represented cold, normal, hot, and extremely hot weather, respectively. The complete load-CMOD curves of specimens were obtained at 3, 7, 14, 28, and 60 d. The key fracture parameters of concrete were also evaluated. The following conclusions can be drawn from the experimental data:

1. In a dry environment, the gain in fracture energy of the concrete strongly depended on the curing temperature. An elevated temperature was beneficial to the rapid development of concrete fracture energy at an early age. However, this situation did not hold at later ages, in which a higher value of the fracture energy at a later age was obtained by concrete stored at lower temperatures, and vice versa. Generally, at 60 d, the fracture energy of concrete cured at 60 ◦C was only larger 7.7% and 3.2% when comparing concrete stored at 5 ◦C and 20 ◦C, respectively, and even was smaller 2.8% when comparing concrete stored at 40 ◦C. Considering the early growth rate and later-age value of fracture energy, the optimum curing temperature for fracture energy of concrete under dry condition was around 40 ◦C.


It should be noted that the aforementioned conclusions are only valid for the concrete studied in this paper, and the general conclusions for other types of concrete need further study.

**Author Contributions:** Conceptualization, Z.M.; formal analysis, Z.M.; funding acquisition, Q.L. and Y.H.; investigation, Z.M.; methodology, Z.M.; project administration, C.L. and Y.Q.; resources, C.L. and Y.Q.; supervision, Y.H.; writing—original draft, Z.M.; writing—review and editing, Q.L. and Y.H. All authors have read and agreed to the published version of the manuscript.

**Funding:** This research was funded by National Natural Science Foundation of China (51979145, 51839007), the Fundamental Research Funds for the Central Universities, CHD (300102219305), the Open Research Fund Program of State Key Laboratory of Hydroscience and Engineering (sklhse-2020-D-01), and the Research Projects of China Three Gorges Corporation (Contract numbers: BHT/0806 and WDD/0427).

**Conflicts of Interest:** The authors declare no conflicts of interest.

#### **References**


© 2020 by the authors. Licensee MDPI, Basel, Switzerland. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC BY) license (http://creativecommons.org/licenses/by/4.0/).

#### *Article*

## **Chemical and Mechanical Roughening Treatments of a Supra-Nano Composite Resin Surface: SEM and Topographic Analysis**

**Francesco Puleio 1,\*, Giuseppina Rizzo 1, Fabiana Nicita 1, Fabrizio Lo Giudice 1, Cristina Tamà 1, Gaetano Marenzi 2, Antonio Centofanti 1, Marcello Ra**ff**aele 3, Dario Santonocito <sup>3</sup> and Giacomo Risitano <sup>3</sup>**


Received: 21 May 2020; Accepted: 26 June 2020; Published: 28 June 2020

**Abstract:** Background: Repairing a restoration is a more advantageous and less invasive alternative to its total makeover. The aim of this study was to analyze the effects of chemical and mechanical surface treatments aimed at increasing the roughness of a supra-nano composite resin. Methods: 27 cylindrical blocks of microhybrid composite were made. The samples were randomly divided into nine groups (n = 3). The samples' surface was treated differently per each group: acid etching (35% H3PO4, 30 s and 60 s), diamond bur milling, sandblasting and the combination of mechanical treatment and acid etching. The samples' surface was observed by a scanning electron microscope (SEM) and a confocal microscope for observational study, and surface roughness (Ra) was recorded for quantitative analysis. Results: The images of the samples sandblasted with Al2O3 showed the greatest irregularity and the highest number of microcavities. The surfaces roughened by diamond bur showed evident parallel streaks and sporadic superficial microcavities. No significant roughness differences were recorded between other groups. The difference in roughness between the control group, diamond bur milled group and sandblasted group was statistically significant. (p < 0.01). Comparison between the diamond bur milled group and the sandblasted group was also significant (p < 0.01). Conclusion: According to our results, sandblasting is the best treatment to increase the surface roughness of a supra-nano composite.

**Keywords:** surface roughness; microhybrid composite; sandblasting; surface treatment; composite repair; minimal invasive dentistry

#### **1. Introduction**

Over the past few years, the quality of direct and indirect restorations has improved in terms of the adhesion strength, longevity and composition of the resin matrix and filler [1].

However, like all dental materials, the composites undergo deterioration processes as a consequence of mechanical (cyclic fatigue and wear) or thermal stresses and chemical degradation (enzymatic, hydrolytic and acidic) [2,3].

Fractures, marginal bacterial infiltration, dentin treatments, teeth or restoration color changes, indirect restorations or endodontic post detachment can also compromise the result and require a makeover [4–9].

In these cases, complete replacement can be a long and expensive procedure with the possibility of further healthy dental tissue loss and an increased risk of pulp exposure. Therefore, repairing the restoration by preserving parts or rebonding an indirect restoration can be an advantageous alternative as these techniques are less invasive and allow the prolongation of the efficiency of conservative therapies over time [5,10–16].

The prognosis of a repair and maintenance treatment depends on the adhesion strength achieved between the old restoration and the new composite material layer [10].

In clinical practice, the adhesion techniques and mechanisms at the dental tissues–composite or composite–composite interfaces are different [17].

The adhesion strength between two composite surfaces depends on their chemical composition and characteristics such as the roughness, conditioning procedures and ability to become wet of the polymerized surface [10].

Furthermore, contrary to what happens in the composite layering technique, during reparation, the material integration process is hindered by the relative lack of unpolymerized monomers [18–20].

After intraoral photopolymerization, the conversion rate from monomer to polymer is between 45% and 70% because not all monomers participate in the polymerization and the remaining monomers are available for new adhesion [21–23].

Considering that the number of unsaturated double bonds in monomers decrease over time, the adhesion strength reduces by between 25% and 80%, and consequently the effectiveness of the repair process is time-related [18–20].

The adhesion of new composite to an old restoration is achieved with a new bond established by residual monomers and micromechanical retention that exploits the surface irregularities of the old restoration.

In order to improve this type of adhesion, various surface treatments have been described in the literature such as surface roughening with diamond burs of different granulometry, sandblasting with aluminum oxide or silica sand, acid etching and the application of hydrogen peroxide or silane.

A systematic review of the data still appears insufficient to indicate the best method for repairing Bis-GMA-based resins [11,14,18,21,22]. The aim of this study is to analyze the effects of chemical and mechanical surface treatments and their combination on the roughness of a supra-nano composite resin surface.

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

27 cylindrical blocks (height 4 mm and diameter 6 mm) of supra-nano composite col. A1 (Estelite Sigma Quick, Tokuyama Dental, Japan) were made using a silicone mold matrix.

Considering a power level of 85% with a type-I error = 0.05, for parameter roughness (Ra), three samples for each independent group (surface-treated and independent control) were necessary.

The technical characteristics of the composite resin provided by the manufacturer were: a resin matrix (bisphenol A-glycidyl methacrylate (Bis-GMA) and triethylene glycol dimethacrylate (TEGDMA)) and particles of reinforcement (71% of the total volume) formed by spherical particles of silica and zirconia, with sizes ranging from 0.1 to 0.3 μm (average size of 0.2 μm).

The cylindrical blocks were obtained through two vertical composite increments of 2 mm inside the silicone matrix. Using a LED curing lamp (Valo, Ultradent, South Jordan, UT, USA) the composite layers were light-cured for 20" at a distance of 1 mm at 3.200 mW/cm2. In order to prevent the inhibition of polymerization due to the presence of oxygen, and to create a homogeneous surface, the last layer of composite was covered by a glass plate before light-curing.

To make the samples' surface uniform, all the composite blocks were polished under 4× magnification (EyeMag Pro S, Zeiss, Oberkochen, Germania), using Soft-Lex (3M ESPE, St Paul, Minnesota) coarse-grained, medium, fine and superfine discs for 10 s each. After each step the samples were washed with distilled water and air dried.

The samples were randomly divided into 9 groups (n = 3). The surface of the samples from each group was treated with different roughening protocols:

A) Control group, no surface treatment.

B) Etching for 30 s.

C) Etching for 60 s.

D) Roughening with diamond bur.

E) Roughening with diamond bur and etching for 30 s.

F) Roughening with diamond bur and etching for 60 s.

G) Sandblasting.

H) Sandblasting and etching for 30 s.

I) Sandblasting and etching for 60 s.

Etching was performed with 35% orthophosphoric acid (Ultra-Etch, Ultradent) for 30 s in groups B, E and H or 60 s in groups C, F and I. After etching, each sample was washed with distilled water and air dried.

Milling was carried out with a diamond bur with granulometry 151 μm (6837 KR Komet) mounted on a handpiece for 3 s (groups D, E and F)

Sandblasting (20 s) was performed with an intraoral sandblaster using Al2O3 50 μm (Dentoprep, Rønvig Dental).

The observational analysis of the treated surface of the samples was carried out with an SEM (Phenom G2, Phenom, Eindhoven, the Netherlands), at 2.100x magnification, with 5 kV voltage and secondary electrons (SE). This SEM does not require any treatment of the sample surfaces [24,25].

The samples were then observed and analyzed with a confocal microscope (Leica DCM 3D, Leica Microsystems) in order to measure roughness. In each sample an area of 2.5 mm2 was selected using a systematic random sampling protocol for stereological and morphometrical analysis and the roughness (Ra) was calculated using the following formula:

$$Ra = \frac{1}{l} \int\_0^l |Z(\mathbf{x})| d\mathbf{x}$$

For each experimental group the mean and quantitative parameter standard deviation (SD) were calculated.

Data were analyzed using the Student's t-Test and differences of p < 0.05 were considered statistically significant.

The statistical analysis was performed using the SPSS 17.0 for Windows package and the Prism software package (GraphPad, La Jolla, CA, USA).

#### **3. Results**

#### *3.1. SEM Analysis*

The SEM observation showed that the different types of roughening protocol exhibited no substantial differences among samples of the same group.

Comparisons between groups showed different surface morphologies, displayed in Figure 1.

**Figure 1.** SEM images (2100×). Voltage: 5kV. Type of electron: secondary electron (SE). Groups A, B and C show a surface without irregularities; groups D, E and F show evident parallel streaks resulting from the bur action (arrows indicates the parallel streaks); groups G, H and I, the sandblasted samples, show the highest number of microcavities (arrows indicate microcavities).

The surface of the control group samples (group A) was homogeneous, without irregularities and without microcavities. Etching with 35% orthophosphoric acid for 30 s or 60 s (groups B and C) did not cause any observational modification of the composite surface. The surface roughened by diamond bur (group D) showed evident parallel streaks resulting from the bur action. The images of the sandblasted samples (group G) show greater irregularity and a high number of microcavities. Etching using 35% orthophosphoric acid for 30 s and 60 s (groups E, F, H and I) as an additional treatment to milling and sandblasting caused no change on the sample surface.

In all experimental groups, the optical confocal microscope analysis shows three-dimensional topographic images similar to the SEM observation (Figure 2).

**Figure 2.** Confocal microscope. Three-dimensional surface topography. Red expresses the peaks; blue shows the depressions. Groups A, B and C show a surface with low irregularities; groups D, E and F show a moderate presence of surface irregularities; groups G, H and I show the highest presence of surface irregularities.

#### *3.2. Profilometric Analysis*

The mean values of roughness (Ra) and the standard deviation (SD) for each experimental group are shown in Figure 3.

**Figure 3.** Roughness (Ra) mean values of experimental group.

There were no statistically significant differences in Ra between groups A (94.2 ± 44.7 nm), B (78.6 ± 23.3 nm) and C (60.7 ± 35.3 nm).

The same result was obtained by comparing the values for groups D (1647.8 ± 471.9 nm), E (1812.3 ± 300.7 nm) and F (1603.8 ± 280.4 nm), and those for groups G (2955.6 ± 572.9 nm), H (2777.8 ± 447.6 nm) and I (2855.6 ± 494 nm).

The comparison of Ra between group A (control) and groups D and G indicated statistically significant differences (p < 0.01). Additionally, the difference in Ra between groups D and G was statistically significant (p < 0.01).

#### **4. Discussion**

The introduction into clinical practice of build-up and indirect restorations made of composite resins requires the knowledge of adhesion mechanisms, especially when the surfaces are made of already polymerized composite [26]. This analysis can also be linked to the necessity of following a conservative approach in order to increase the longevity of direct restorations and the possibility of further repairing processes [27]. The treatments used most often to enhance the adhesive's action are chemical and mechanical surface roughening [14,18,19,21,26]. In this study, two mechanical and one chemical technique of surface conditioning and their combination were used. The mechanical techniques consisted of surface roughening by the action of a diamond bur or sandblasting with Al2O3. The chemical conditioning technique instead consisted of an etching procedure on the samples' surface with 35% orthophosphoric acid for 30 s or 60 s [28].

The images obtained with SEM and three-dimensional topography with an optical confocal microscope in all experimental groups agree, showing that among samples of the same groups the surface treatments cause substantially comparable morphological alteration. This highlights how the response of the same type of material to various techniques does not change and is therefore predictable.

The control group (treated exclusively by polishing with abrasive discs) presented a smooth, homogeneous surface without any irregularities or microcavities.

The surfaces roughened by a bur were irregular, with a limited number of microcavities distant from each other and with parallel streaks resulting from the action of the bur. The sandblasted samples showed greater irregularity and a high number of microcavities.

The comparison between the observations for these latter groups, according to previous studies, confirms the hypothesis that sandblasting with Al2O3 is the most suitable treatment for increasing the micromechanical retention of composite surfaces [29].

In our research, the treatment of etching alone, regardless of its duration, does not cause a significant improvement of composite roughness. Furthermore, there are not any substantial modifications on the surface even when the etching action is carried out on a previously mechanically conditioned surface. This outcome is independent of the duration of acid application. The observational data, therefore, seem to exclude any positive impact of the etching, according to Loomans et al., who stated that 35% orthophosphoric acid is not able to cause significant alterations to the resin filler. This component can be modified only by more aggressive acids, such as hydrofluoric acid [28].

The morphological findings (SEM and confocal) are confirmed by the profilometric analysis quantifying the average roughness (Ra) of the experimental groups.

The profilometric analysis (Figure 2) shows variable roughness among the experimental groups. The color scale from red to blue expresses the difference in height at different points of the same sample. To understand these data it is important to note that it is almost impossible to obtain a perfectly flat sample by hand, and this is particularly evident in the A, B and C groups, where the analysis shows the lowest Ra values but red and blue areas are visible. This color expression afforded the researchers a visual way of understanding the surfaces' pattern and without comparing their height. Groups A, B and C obtained the lowest Ra values; moreover, there were no significant differences in the average roughness between surfaces treated with etching only and the control group. Even when the effect of the etching on milled or sandblasted surfaces was evaluated, the variation in roughness was not

statistically significant. The concordance between the observations shows how chemical treatment does not change the roughness of a smooth or previously mechanically roughened surface.

The average roughness increased significantly when comparing the control group to the samples (D, E and F) whose surface was roughened by the action of a diamond bur. This statistical evaluation was also valid for the sandblasted groups (G, H and I).

The comparison between the groups treated with mechanical conditioning shows that the sandblasted samples had significantly higher average roughness. These data agree with the observational findings and confirm that sandblasting is the best treatment possible for roughening a composite surface [28].

SEM observation allows us to directly observe the surface morphology to understand the modification related to different tools and the combination of materials and techniques [30]. The principal advantage of this technique is the possibility to directly observe samples without any major surface modification, especially samples with irregular and reflective surfaces, where producing bi-dimensional images does not give any information regarding surface roughness. In particular, the SEM used in this study gives a precise and reliable surface image not requiring any surface metallization prior to observation. The confocal laser microscope with profilometric analysis gives a 3D objective evaluation of the surfaces, providing a visual color scale that shows differences in height among different points of the sample and so provides quantitative data on the sample microsurface [31]. The combination between SEM analysis and the confocal laser provides the possibility to obtain and order numerical categories such as the smoothness, roughness and waviness of irregular and reflective surfaces along with a precise and detailed bi-dimensional image [32].

The main limitation of this study is related to the experimental samples that evaluated just one supra-nano composite. There is a need to compare the efficacy of this surface treatment on other composites, considering that different composite materials with different physical-chemical characteristics might respond to surface treatments in a specific manner.

In our experimental sample the specimens were not subjected to thermocycling procedures due to the in vitro experimental setting of the research. The thermocycling process gives the researchers the possibility to perform tests and evaluate the samples in a setting more similar to the clinical world. In the literature it has been described how aged composite shows lower adhesion values when compared to non-aged composite, probably due to hydrolytic degradation in the resin matrix that occurs in the oral environment, and due to the reduction of free radicals available that can react chemically with a fresh composite [20,33,34].

Moreover, further study is necessary to evaluate the bond strength of the composite surface and its effect.

#### **5. Conclusions**

Considering the results of our research, it is possible to conclude that among the treatments used to increase the roughness of a supra-nano composite surface, the most effective treatments are mechanical ones. Within this category of conditioning, sandblasting creates the best environment for the microretention of the adhesive system.

The use of orthophosphoric acid does not result in efficient surface roughness: the SEM images show a different appearance after 60 s acid etching; however, there is no statistically significant difference. According to our results, it can be concluded that sandblasting should be carried out to increase the surface roughness of an already polymerized supra-nano composite.

This treatment makes it possible to integrate the chemical forces generated between the monomers contained both in the adhesive and in the old composite surface with mechanical microretention.

**Author Contributions:** Conceptualization, F.P.; methodology, G.R.; software, M.R., D.S.; formal analysis, C.T., F.N.; investigation, G.M., A.C.; data curation, F.L.G.; writing—review and editing, F.P.; supervision G.R. All authors have read and agreed to the published version of the manuscript.

**Funding:** This research received no external funding.

**Conflicts of Interest:** The authors declare no conflict of interest.

#### **References**


© 2020 by the authors. Licensee MDPI, Basel, Switzerland. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC BY) license (http://creativecommons.org/licenses/by/4.0/).

*Article*

## **Assessment of the Di**ff**erent Types of Failure on Anterior Cantilever Resin-Bonded Fixed Dental Prostheses Fabricated with Three Di**ff**erent Materials: An In Vitro Study**

**Adolfo Di Fiore 1,\*, Edoardo Stellini 1, Gianpaolo Savio 2, Stefano Rosso 3, Lorenzo Grai**ff **1, Stefano Granata 1, Carlo Monaco <sup>4</sup> and Roberto Meneghello <sup>3</sup>**


Received: 19 May 2020; Accepted: 15 June 2020; Published: 17 June 2020

**Abstract:** background: resin-bonded fixed dental prosthesis (RBFDP) represents a highly aesthetic and conservative treatment option to replace a single tooth in a younger patient. The purpose of this in vitro study was to compare the fracture strength and the different types of failure on anterior cantilever RBFDPs fabricated using zirconia (ZR), lithium disilicate (LD), and PMMA-based material with ceramic fillers (PM) by the same standard tessellation language (STL) file. Methods: sixty extracted bovine mandibular incisives were embedded resin block; scanned to design one master model of RBFDP with a cantilevered single-retainer. Twenty cantilevered single-retainer RBFDPs were fabricated using ZR; LD; and PM. Static loading was performed using a universal testing machine. Results: the mean fracture strength for the RBFDPs was: 292.5 Newton (Standard Deviation (SD) 36.6) for ZR; 210 N (SD 37.6) for LD; and 133 N (SD 16.3) for PM. All the failures of RBFDPs in ZR were a fracture of the abutment tooth; instead; the 80% of failures of RBFDPs in LD and PM were a fracture of the connector. Conclusion: within the limitations of this in vitro study, we can conclude that the zirconia RBFDPs presented load resistance higher than the maximum anterior bite force reported in literature (270 N) and failure type analysis showed some trends among the groups

**Keywords:** zirconia; digital dentistry; lithium disilicate; resin bonded bridge; fracture; adhesive restorations; CAD/CAM; PMMA

#### **1. Introduction**

Traumatic loss [1], or congenital absence of one anterior maxillary incisor [2] in adolescents, requires immediate treatment with temporary or definitive solutions for aesthetic and functional reasons. Resin-bonded fixed dental prosthesis (RBFDP) represents a highly aesthetic and conservative treatment option to replace a single tooth in a younger patient, before implant treatment becomes available [3] or after orthodontic treatment [4]. In literature, survival rates of RBFDPs were 87.7% in medium-term observation [5]. The main factors of failure include debonding [6], secondary caries on

the abutment tooth [5], and fracture of the retainers [7]. These causes are determined by two prosthetic characteristics of RBFDPs: retainer designs [3,8] and properties of the materials [9]. The framework designed with two-retainer has been the most used by clinicians and dental technicians because it was considered with higher fracture resistance than the frameworks with one retainer (cantilever). However, several studies demonstrated that two-retainer RBFDPs have a higher fracture rate and lower survival rate than cantilever RBFDPs [10–15]. Traditionally, the framework of RBFDPs is made of metal alloy, but different metal-free materials with more aesthetics and bond strengths are available. Several authors showed excellent longevity of zirconia [16,17] and lithium ceramic [18] cantilever RBFDPs over 10–20 years with few mechanical complications. However, the mechanical complications were different, according to the material. Debonding rate of 8% and one loss of restoration was revealed by Kern et al. [17] for zirconia cantilever RBFDPs; no debonding was recorded by Sailer et al. [18] for lithium disilicate cantilever RBFDPs. However, the assessment of the mechanical performance of prosthetic materials through clinical trials is difficult because of several conditions and characteristics of the patients.

In the last few years, digital technologies have been introduced in dentistry to improve patient comfort, decrease operative time, and reduce clinical treatment [19–22]. The use of scanners, computer-aided design (CAD) software, and computer-aided manufacturing (CAM) machines have opened many possibilities to replicate and fabricate dental prosthesis in different materials.

Therefore, uniform fabrications of cantilever RBFDPs with different materials by the same standard tessellation language (STL) file allow knowing the real performance of these materials, and the possible adverse comportments on the abutment tooth in extreme conditions.

The purpose of this in vitro study was to compare, using the universal testing machine, the fracture strength and the different types of failure on anterior cantilever RBFDPs fabricated using zirconia, lithium disilicate, and PMMA-based material with ceramic fillers by the same STL file.

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

Sixty extracted bovine mandibular incisives were stored in physiological saline solution at a temperature between 5 ◦C and 10 ◦C [23]. They were embedded in autopolymerizable methacrylate resin block of dimensions 35 × 50 × 14 mm (ProBase Cold, Ivoclar Vivadent, Bologna, Italy), with the cement–enamel junction above 1 mm and orthogonal the resin base. After horizontal preparation of 1–1.5 mm using a diamond chamfer bur on lingual surface near the cementum–enamel junction of each tooth, air-polishing was performed (S2, EMS) with sodium bicarbonate-based powders (EMS) on 60 teeth for 30 s at distance of about 3 centimeters on total surface of each tooth (Figure 1).

**Figure 1.** Air-polishing used to clean incisive embedded in resin block.

All teeth were imported in virtual environment by laboratory scanner (Smart Big Open Technology, Rezzato (BS), Italy). Using a Cad software (Exocad DentaCad, Darmstad, Germany), one master model of RBFDP with a cantilevered single-retainer was designed (Figure 2).

**Figure 2.** Design of master model of cantilever BFDP.

The RBFDP presented the same parameter setting, shape, and size of the cantilever tooth, independently from the anatomical morphology of the bovine mandibular incisives. The retainer wings were fabricated with a uniform thickness of 0.7 mm. The connectors were designed with a height of 3 mm and a width of 1.5 mm [24,25]. Twenty RBFDPs with a cantilevered single-retainer design for the groups were fabricated with 3 different materials: zirconia (Katana ML, Kuraray, Milano, Italy), lithium disilicate (IPS e.max Press LT A2, Ivoclar Vivadent, Bologna, Italy), and PMMA-based material with ceramic fillers (HIPC, Bredent GmbH, Senden Germany). Each RBFDP has been created by the same CAD design with different procedures based on the type of material. The twenty RBFDPs in HIPC were milled by a computer numerical control machine (Roland DWX-50, Roland DG Corporation, Osaka, Japan). The excessive material was removed with tungsten bur (Komet Dental; H250NEX). All of the samples were polished using a polisher machine (Sirio Dental, Meldola (FC), Italy). Twenty zirconia RBFDPs were milled by a CNC machine (Roland DWX-50, Roland DG Corporation, Osaka, Japan), polished using diamond burs, and sintered (Nabertherm LHT 01/17 D, Nabertherm, Lilienthal, Germany) for 12 h at 1450 ◦C following the manufacturing instructions. The 20 RBFDPs in lithium disilicate were fabricated with wax lost technique. A wax block (CeraWax, Co.N.Ce.P.T, Busseto (PR), Italy) was used to mill the RBFDPs. A sprue was waxed onto the cantilever tooth of all RBFDPs. They were then embedded in a universal investment material (IPS PressVest Premium investment, Ivoclar Vivadent, Bologna, Italy), with 2 RBFDPs (Figure 3).

**Figure 3.** Cantilever RBFDPs milled in wax.

The wax was removed in a heated furnace (Sirio SR 750-In Fire, Sirio Dental S.r.l., Meldola (FC), Italy) and the RBFDPs were pressed with lithium disilicate glass-ceramic (IPS e-max Press LT A2, Ivoclar Vivadent, Bologna, Italy) in the machine-calibrated furnace (Luxor Press-SR 862, Sirio Dental S.r.l., Meldola (FC), Italy) following manufacturer recommendations (Figure 4).

**Figure 4.** RBFDPs in lithium disilicate fabricated with wax lost technique.

The sprue was removed using a metallic disk and all RBFDPs in lithium disilicate were polish with diamond burs. We obtained 60 comparable RBFDPs with a cantilever design without preparation of the abutment in three different materials. The thickness of the wings of all RBFDPs were measured using a caliper. All of the abutment teeth were treated with the same procedure: 15 s of etching (Scotchbond Universal Etching, 3M, Milano, Italy), 15 s of rinsing with water, 15 s of drying, and the bonding (Scotchbond Universal Adhesive, 3M, Milano, Italy) were applied for 20 s and polymerized for 60 s using a lamp (Valo, Ultradent, South Jordan, UT, USA). All of the RBFDPs were cemented with dual-cured resin cement (Relyx Ultimate, 3M, Milano, Italy) and polymerized for 60 s using led lamp (400–500 nm), but with different procedures. Air abrasion was performed with 50 μm Al2O3 particles at a 2.5 bar pressure for 15 s at a distance of 10 mm on the wings of RBFDPs in zirconia. After drying, the wings were cleaned with 96% isopropanol for 3 minutes. The bonding (Scotchbond Universal Adhesive, 3M, Milano italy) was applied on the wings for 20 s and polymerized for 60 s. Air abrasion was performed with 110 μm Al2O3 particles at a 2.5 bar pressure for 15 s, at a distance of 10 mm on the wings of RBFDPs in PMMA-based material with ceramic fillers. After drying, the primer (Visiolink, Bredent, Seden, Germany) was applied on the wings and polymerized for 60 s. Instead, the wings of RBFDPs in lithium disilicate were etching with hydrofluoric acid (Ceramic Etching gel, Ivoclar Vivadent, Bologna Italy) for 60 s, 60 s of rinsing, and 30 s of drying. The bonding (Scotchbond Universal Adhesive, 3M, Milano, Italy) was applied for 20 s and polymerized for 60 s. All procedures were performed according to the manufacturer's instructions. The samples were numbered through the engraving on the resin of a code indicating the material. All samples were mounted on a 30◦ angled support of a universal testing machine (Acumen 3, MTS Systems Corporation, Eden Praire, MN, USA) and a static loading at a crosshead speed of 1.5 mm/min, until failure to the incisal edge of the pontic tooth was performed to simulate real situation (Figure 5). The load was transferred through a 6-mm diameter steatite ball in the middle of the incisal edge of the pontic tooth until failure [25].

Four types of failure were recorded: debonding of RBFDPs, fracture of the connector, debonding of the RBFDPs with fracture of the abutment tooth, and fracture of the abutment tooth with the RBFDP still bonded. All of the methods were synthetized in the following flowchart (Figure 6).

**Figure 5.** A sample mounted on a 30◦ angled support of the universal testing machine.

**Figure 6.** Flowchart of the methods.

The Kruskal–Wallis test with a post hoc analysis using Dunn's test was used to compare the three groups. The level of statistical significance was set as α =0.05 and statistical power of 80%. A statistical software (SPSS v16.0; SPSS Inc., Chicago, IL, USA) was used for the analysis.

#### **3. Results**

The mean fracture strength of the RBFDPs in zirconia was 292.5 N (SD 36.6) (Figure 7a), 210 N (SD 37.6) for lithium disilicate (Figure 7b), and 133 N (SD 16.3) for PMMA-based material with ceramic fillers (Figure 7c).

**Figure 7.** Graphic load-extension of the sample: (**a**) number 3 in zirconia, (**b**) number 5 in lithium disilicate, and (**c**) number 2 in PMMA-based material with ceramic fillers.

In the group of zirconia RBFDPs, all of the failures were fractures of the abutment tooth (Figure 8a); in the group lithium disilicate 80% of failures were fractures of the connector and 20% debonding of the RBFDPs, with fractures of the abutment tooth (Figure 8b). In the group of PMMA-based material with ceramic fillers, 80% of failures were fractures of the connector (Figure 8c) and 20% debonding. Kruskal–Wallis generated a *P*-value < 0.001, identifying statistically significant differences between groups. Dunn's post hoc tests indicated, however, that the difference was statistically significant between the group RBFDPs in zirconia and PMMA-based material with ceramic fillers (P = 0.003), instead no differences were found between RBFDPs in zirconia and lithium disilicate (P = 0.43), and RBFDPs in lithium disilicate and PMMA-based material with ceramic fillers (P = 0.09).

**Figure 8.** Different types of failure: (**a**) fracture of the abutment tooth in the group RBFDP in zirconia, (**b**) fracture of the connector in the group RBFDP in lithium disilicate, and (**c**) fracture of the connector in the group RBFDP in PMMA-based material with ceramic fillers.

#### **4. Discussion**

RBFDP is an aesthetic and conservative treatment option to replace a single tooth in a younger patient; therefore, it is a technique sensitive procedure because it requires proper planning of the clinical case and choice of materials. This in vitro study investigated the different types of failure after static fracture strength tests on RBFDPs fabricated using zirconia, lithium disilicate, and PMMA-based material with ceramic fillers. Zirconia, lithium disilicate, and PMMA-based material with ceramic fillers presented the mean values static fracture strength of 292.5 N (SD 36.6), 210 N (SD 37.6), and 133 N (SD 16.3), respectively. Moreover, different types of failure were collected according to the materials. The 80% of the RBFDPs made in lithium disilicate and PMMA-based material with ceramic fillers presented fracture of the connector; 20% of debonding and all of the RBFDPs in zirconia presented fracture of the abutment tooth.

The results can be explained by the different elastic modulus of materials. The lithium disilicate and PMMA-based material with ceramic fillers presented lower elastic modulus and, consequently, a greater deformation until the fracture of the materials. Another possible explanation of the fracture of the connector was recognizable to the superior bond strength, with respect to the materials. The authors highlight that the lithium disilicate is not indicated to fabricate RBFDPs by manufacturer; however, the results showed a mean fracture strength of 210 N. Therefore, clinical studies are necessary to evaluate the performance of this material used to fabricate RBFDPs. Instead, all the RBFDPs in zirconia presented a fracture of the abutment tooth. The reasons may be attributable to the high elastic modulus of the material or high bond strength of resin luting cement, with respect to the abutment tooth.

Opposite results were observed in clinical trials [16–19] compared to this in vitro study. The reasons could be several; however, the load strength of the RBFDP is different in each patient. Moreover, the pontic elements may not have static and dynamic contacts. In this in vitro study, the uniform fabrication of all the RBFDPs with different materials by the same STL file, and the strength applied through a universal machine until failure, allowed to know the mechanic performance of these prostheses with respect to an in vivo study.

In literature, the same authors in two in vitro studies evaluated the influence of framework design [14] and mode of loading on the fracture strength [26] on cantilever RBFDPs fabricated in pre-sintered aluminum-oxide blocks (In-Ceram alumina blanks). The microscopic examination revealed that 58% of the specimens fractured at the connector only, exactly at the framework-to-veneer interface, 17% fractured at the connector, including veneered parts of the pontic, and 25% fractured at the retainer only. The possible difference of the performance should be explained that, in this research, the authors used bovine teeth. Physiologic occlusal forces for adults in the anterior region were determined to be in a range of 10 to 35 N [27]. Maximal incisive biting forces may vary up to 270 N, primarily depending on facial morphology and age [28]. Compared to the maximal incisive biting forces, only zirconia RBFPDs reached values higher than the physiologic values. Even though the mean fracture strength of lithium disilicate RBFDPs was lower compared to the maximal incisive biting force, it is possible to use the material to fabricate RBFDPs. However, the clinicians should evaluate, with attention, the static and dynamic contacts of the clinical case before use of lithium disilicate RBFDP. Whereas, PMMA-based material with ceramic filler RBFDPs represents a cheap and aesthetic solution for a temporary prosthesis for patients waiting for implant treatments. Therefore, the choice of a correct material, according to therapeutic needs, is crucial for the survival of RBFDPs.

Long-term clinical evidence is needed to evaluate the performance of these materials, especially for use of polymer with ceramic fillers. The drawbacks of this in vitro study were the lack of clinical conditions, as artificial aging, dynamic loading, and of physiologic tooth mobility; however, the uniform fabrications of all the RBFDPs with different materials by the same STL file and the same laboratory conditions allowed to know the real performance of this material, with respect to an in vivo study where the RBFDPs have different shapes and dimensions.

#### **5. Conclusions**

Within the limitations of this in vitro study, we can conclude that the zirconia RBFDPs presented load resistance higher than the maximum anterior bite force; furthermore, the lithium disilicate RBFDPs showed a mean fracture strength similar the anterior bite force.

**Author Contributions:** Conceptualization, A.D.F. and E.S.; methodology, R.M.; validation, S.G. and G.S.; formal analysis, S.G. and S.R.; investigation, G.S. and L.G.; data curation, L.G.; writing—original draft preparation, A.D.F. and E.S.; writing—review and editing, C.M. All authors have read and agreed to the published version of the manuscript.

**Funding:** This research received no external funding.

**Acknowledgments:** The authors thank Giorgio Calgaro for materials used for experiments and technical support.

**Conflicts of Interest:** The authors declare no conflict of interest.

#### **References**


© 2020 by the authors. Licensee MDPI, Basel, Switzerland. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC BY) license (http://creativecommons.org/licenses/by/4.0/).

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