OPC concrete is more susceptible to acid attack due to the presence of hydration products, namely Ca(OH)2. The Ca(OH)2 on the surface of OPC concrete is consumed by the reaction with acid and turned into gypsum, which is soft and porous and causes the surface of OPC concrete to deteriorate. Subsequently, gypsum would undergo distractive reaction with tricalcium aluminates within the cement matrix, resulting in the formation of calcium sulphoaluminate (ettringite), which has large volume and causes expansive deterioration mechanism. The reaction of OPC with sulfuric acid is presented in Figure 16. OPC concrete is more susceptible to acid attack due to the presence of hydration products, namely Ca(OH)2. The Ca(OH)2 on the surface of OPC concrete is consumed by the reaction with acid and turned into gypsum, which is soft and porous and causes the surface of OPC concrete to deteriorate. Subsequently, gypsum would undergo distractive reaction with tricalcium aluminates within the cement matrix, resulting in the formation of calcium sulphoaluminate (ettringite), which has large volume and causes expansive deterioration mechanism. The reaction of OPC with sulfuric acid is presented in Figure OPC concrete is more susceptible to acid attack due to the presence of hydration products, namely Ca(OH)2. The Ca(OH)2 on the surface of OPC concrete is consumed by the reaction with acid and turned into gypsum, which is soft and porous and causes the surface of OPC concrete to deteriorate. Subsequently, gypsum would undergo distractive reaction with tricalcium aluminates within the cement matrix, resulting in the formation of calcium sulphoaluminate (ettringite), which has large volume and causes expansive deterioration mechanism. The reaction of OPC with sulfuric acid is presented in Figure 16. 16. Fly ash-based geopolymers do not contain hydration products; they produce N-A-S-H gel, which is acid resistant and has less surface deterioration. Fly ash also has low calcium content, resulting in the formation of less gypsum and ettringite that cause expansive deterioration mechanism. The change in colour of geopolymer from grey to slightly brown is due to the reaction of iron (II) oxide with sulphuric acid to produce iron (II) sulphate, as supported by previous researchers [46].

the reaction with acid and turned into gypsum, which is soft and porous and causes the

**Figure 16.** Chemical reaction of OPC with sulfuric acid. **Figure 16.** Chemical reaction of OPC with sulfuric acid.

**Figure 16.** Chemical reaction of OPC with sulfuric acid.

**4. Conclusions and Future Work**

**4. Conclusions and Future Work**

**Figure 16.** Chemical reaction of OPC with sulfuric acid. **4. Conclusions and Future Work** In this study, the best formulation for synthesis of fly ash based geopolymer with highest compressive strength in terms of concentration of NaOH, SS/SH ratio, and S/L Fly ash-based geopolymers do not contain hydration products; they produce N-A-S-H gel, which is acid resistant and has less surface deterioration. Fly ash also has low calcium content, resulting in the formation of less gypsum and ettringite that cause expansive deterioration mechanism. The change in colour of geopolymer from grey to slightly brown is due to the reaction of iron (II) oxide with sulphuric acid to produce iron (II) sulphate, as supported by previous researchers [46].

In this study, the best formulation for synthesis of fly ash based geopolymer with

### **4. Conclusions and Future Work**

In this study, the best formulation for synthesis of fly ash based geopolymer with highest compressive strength in terms of concentration of NaOH, SS/SH ratio, and S/L ratio was determined. In addition, the durability of both fly ash based geopolymer and OPC concrete in an acid environment were tested and compared. Based on the results obtained from the analysis and experimental data, the following conclusions can be drawn:

Because the mechanical properties and durability of fly ash based geopolymer concrete are higher compared to OPC based concrete, it is highly recommended to use this material for rigid pavement application, as it suffers from less deterioration under acid attack compared to OPC concrete, and this could lower the maintenance cost required. This demonstrates that fly-based geopolymer is more suitable for use in rigid pavement applications due to its high mechanical performance and durability, which could result in lower rigid pavement maintenance costs. It is beneficial to expand future research on the study of using fly ash based geopolymer as rigid pavements in terms of quality, effect of environment temperature, and long-term durability when exposed to harmful environmental conditions.

**Author Contributions:** Conceptualization, M.F.M.T. and M.M.A.B.A.; data curation, M.F.M.T., M.M.A.B.A. and S.Z.A.R.; formal analysis, M.F.M.T., M.M.A.B.A. and S.Z.A.R.; investigation, M.F.M.T., M.M.A.B.A., S.Z.A.R., M.R.M.H., C.M.R.G. and A.A.K.; methodology, M.F.M.T., M.M.A.B.A., S.Z.A.R., M.R.M.H., C.M.R.G. and A.A.K.; project administration M.F.M.T. and M.M.A.B.A.; validation, M.R.M.H., A.V.S. and P.V.; writing—review and editing, M.F.M.T., M.M.A.B.A., S.Z.A.R., M.R.M.H., A.V.S. and P.V. All authors have read and agreed to the published version of the manuscript.

**Funding:** This study was supported by the Center of Excellence Geopolymer and Green Technology (CEGeoGTECH) UniMAP and Faculty of Technology Mechanical Engineering, UniMAP. This work was supported by the TUIASI's internal grants program (GI\_PUBLICATIONS/2021), financed by the Romanian Government. The authors wish to thank the Ministry of Education, Malaysia, for their financial support of this study through Fundamental Research Grant Scheme (FRGS), FRGS/1/2014/SG06/UNIMAP/02/20.

**Institutional Review Board Statement:** Not applicable.

**Informed Consent Statement:** Not applicable.

**Data Availability Statement:** Not applicable.

**Acknowledgments:** The authors wish to deliver their gratitude to the funding support from TU-IASI's internal grants program (GI\_PUBLICATIONS/2021) sponsored by the Romanian Government, Ministry of Education, Malaysia through Fundamental Research Grant Scheme (FRGS), FRGS/1/2014/SG06/UNIMAP/02/20 for funding and Centre of Excellence Geopolymer and Green Technology for their countless inputs and the best technical support.

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

### **References**

## *Article* **Preparation of Fly Ash-Ladle Furnace Slag Blended Geopolymer Foam via Pre-Foaming Method with Polyoxyethylene Alkyether Sulphate Incorporation**

**Ng Hui-Teng 1,2 , Heah Cheng-Yong 1,3,\* , Liew Yun-Ming 1,2 , Mohd Mustafa Al Bakri Abdullah 1,2 , Catleya Rojviriya 4 , Hasniyati Md Razi 5 , Sebastian Garus 6 , Marcin Nabiałek 7 , Wojciech Sochacki 6 , Ilham Mukriz Zainal Abidin 5 , Ng Yong-Sing 1,2 , Agata Sliwa ´ 8 and Andrei Victor Sandu 9,10,\***

**Abstract:** This paper uses polyoxyethylene alkyether sulphate (PAS) to form foam via pre-foaming method, which is then incorporated into geopolymer based on fly ash and ladle furnace slag. In the literature, only PAS-geopolymer foams made with single precursor were studied. Therefore, the performance of fly ash-slag blended geopolymer with and without PAS foam was investigated at 29–1000 ◦C. Unfoamed geopolymer (G-0) was prepared by a combination of sodium alkali, fly ash and slag. The PAS foam-to-paste ratio was set at 1.0 and 2.0 to prepare geopolymer foam (G-1 and G-2). Foamed geopolymer showed decreased compressive strength (25.1–32.0 MPa for G-1 and 21.5– 36.2 MPa for G-2) compared to G-0 (36.9–43.1 MPa) at 29–1000 ◦C. Nevertheless, when compared to unheated samples, heated G-0 lost compressive strength by 8.7% up to 1000 ◦C, while the foamed geopolymer gained compressive strength by 68.5% up to 1000 ◦C. The thermal stability of foamed geopolymer was greatly improved due to the increased porosity, lower thermal conductivity, and incompact microstructure, which helped to reduce pressure during moisture evaporation and resulted in lessened deterioration.

**Keywords:** polyoxyethylene alkyether sulphate foam; blended geopolymer foam; high temperature; pre-foaming method

### **1. Introduction**

Today, the excellent lightweight and mechanical properties, good thermal and chemical stability, and low environmental impact have led to the successful use of porous geopolymer

**Citation:** Hui-Teng, N.; Cheng-Yong, H.; Yun-Ming, L.; Abdullah, M.M.A.B.; Rojviriya, C.; Razi, H.M.; Garus, S.; Nabiałek, M.; Sochacki, W.; Abidin, I.M.Z.; et al. Preparation of Fly Ash-Ladle Furnace Slag Blended Geopolymer Foam via Pre-Foaming Method with Polyoxyethylene Alkyether Sulphate Incorporation. *Materials* **2022**, *15*, 4085. https:// doi.org/10.3390/ma15124085

Academic Editor: Andrea Sorrentino

Received: 11 May 2022 Accepted: 6 June 2022 Published: 8 June 2022

**Publisher's Note:** MDPI stays neutral with regard to jurisdictional claims in published maps and institutional affiliations.

**Copyright:** © 2022 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 (https:// creativecommons.org/licenses/by/ 4.0/).

or geopolymer foam as a catalyst, thermal insulator, adsorbent and filter [1]. Geopolymer foam can be made by mixing aluminosilicate material (natural resource or industrial waste), foaming agent and alkali solution. There are numerous methods that can be used to prepare geopolymer foam, such as replica, sacrificial filler, pre-emulsification, solid impregnation, direct foaming and pre-foaming [1–3]. The pre-foaming method involves a two-step process, which is more complex compared to the commonly used direct foaming method, as it only involves a one-step process. In the direct foaming method, the foaming agent is directly added to the geopolymer mixture. In the pre-foaming method, the foaming agent is mixed with water to generate foam through a foam generator, and the generated foam is then added to the geopolymer mixture.

Several studies have been focused on the production of geopolymer foam by adopting the pre-foaming method, with the incorporation of polyoxyethylene alkyether sulphate (PAS) foaming agent. PAS is a brown liquid extracted from coconut oil, and it can be easily and quickly dissolved in water to generate foam [4]. Based on previous studies, Ibrahim et al. [5] produced fly ash geopolymer foam with PAS foam-to-paste ratios ranging from 0.5 to 2.0, revealing that the compressive strength of the foamed geopolymer (17.8–4.1 MPa) decreased with increasing foam-to-paste ratio. Mastura et al. [6] and Ibrahim et al. [4] produced fly ash geopolymer foam with a PAS foam-to-paste ratio of 1.0, reporting that the compressive strength of the foamed geopolymer was in the range of 9.0 MPa to 19.3 MPa. On the other hand, Tiong et al. [7] also produced concrete foam using PAS foam.

The past literature examined the influence of PAS foam addition on geopolymer at room temperature, as well as the influence of PAS foam addition on geopolymer at high temperatures. Mastura et al. [2] studied the effect of high temperature exposure on the strength performance of fly ash geopolymer foam with a PAS foam-to-paste ratio of 1.0. The results showed that the compressive strength of the foamed geopolymer reduced with the increase of temperature from 200 ◦C (33.3 MPa) to 800 ◦C (18.6 MPa).

Moreover, in previous studies [8,9], geopolymer foams were mainly based on a single precursor. Very little research has been investigated on geopolymer foams based on blended aluminosilicates. For instance, Zhang et al. [10] prepared fly ash-ground granulated blast furnace slag geopolymer foam via a pre-foaming method using diluted aqueous. They concluded that the geopolymer foam based on blended precursors outperformed those produced using single precursor in terms of thermo-mechanical properties. Thus far, there are no studies on the preparation of PAS-geopolymer foams using blended precursors to investigate their performance at room temperature and high temperatures.

Consequently, the PAS-geopolymer foams produced in this study were made from a mixed precursor of fly ash and ladle furnace slag. The physical, mechanical and thermal performance of fly ash-slag blended geopolymer foams were evaluated at room temperature (29 ◦C) and high temperatures (200–1000 ◦C). The microstructural, functional group and phase analyses of fly ash-slag blended geopolymer foams were also assessed. The PAS foamto-paste ratio was fixed at 1.0 and 2.0 to prepare geopolymer foam. Unfoamed geopolymer was also prepared to compare with foamed geopolymer. Investigation of the influence of PAS foam addition on the thermal properties of fly ash-slag blended geopolymer foam is a worthwhile effort.

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

### *2.1. Raw Materials*

Fly ash and ladle furnace slag were mixed as a blended precursor to prepare geopolymer foam. Table 1 demonstrates the chemical composition of fly ash and slag using X-ray fluorescence. Three bulks of fly ash and slag were analysed. SiO2 and Al2O3 were the main compounds of fly ash, with a total amount of 84.30 wt.%, which was within the suitable composition range (83.70–87.21 wt.%) to produce geopolymers [2,11]. The low CaO content demonstrated that the fly ash was Class F fly ash based on ASTM C618. On the other hand, CaO, SiO2 and Al2O3 were the main compounds of slag, with a total amount of 87.19 wt.%., which was similar to other studies (73.00–83.15 wt.%) for geopolymer preparation [12,13].

**Table 1.** Chemical composition of fly ash and slag.

Polyoxyethylene alkyether sulphate (PAS) was utilised as a foaming agent to prepare geopolymer foam. The PAS consists mainly of 11.0 wt.% of PdO and 78.6 wt.% of SO3. By using the pre-foaming method, the PAS was prepared in foam before being mixed with the geopolymer mixture, as in accordance with ASTM C796. The foam was produced by PAS and water through a foam generator with a constant pressure of 0.5 MPa, as shown in Figure 1. The mechanism of PAS foam can be divided into four stages, as illustrated in a schematic diagram in Figure 2. The PAS molecule had both hydrophilic (water soluble) and hydrophobic (water insoluble) substances [14]. When PAS was mixed with water, the head of the PAS molecule favoured water while its tail resisted water. As a result, micelles formed, causing bubbles to form in the water medium. Moreover, a bubble can form in the air medium as seen in Stage IV. *Materials* **2022**, *15*, x FOR PEER REVIEW 4 of 18

**Figure 1.** Preparation of PAS foam using a foam generator. **Figure 1.** Preparation of PAS foam using a foam generator.

lution was placed at room temperature for 24 h before use.

mation in (**III**) water and (**IV**) air.

**Figure 2.** (**I**) Structural formulas of PAS and water, (**II**) PAS and water mixture, PAS foam for-

Sodium hydroxide pellets (HmbG®, Sigma-Aldrich, Taufkirchen, Germany) and distilled water were mixed together to prepare a sodium hydroxide solution. The main constituents of the sodium silicate solution (South Pacific Chemicals Industries Sdn. Bhd., Selangor, Malaysia) were 30.1% SiO2, 9.4% Na2O, and 60.5% H2O. Sodium hydroxide and sodium silicate solutions were mixed to prepare an alkaline activator, and the mixed so-

**Figure 1.** Preparation of PAS foam using a foam generator.

**Figure 2.** (**I**) Structural formulas of PAS and water, (**II**) PAS and water mixture, PAS foam formation in (**III**) water and (**IV**) air. **Figure 2.** (**I**) Structural formulas of PAS and water, (**II**) PAS and water mixture, PAS foam formation in (**III**) water and (**IV**) air.

Sodium hydroxide pellets (HmbG®, Sigma-Aldrich, Taufkirchen, Germany) and distilled water were mixed together to prepare a sodium hydroxide solution. The main constituents of the sodium silicate solution (South Pacific Chemicals Industries Sdn. Bhd., Selangor, Malaysia) were 30.1% SiO2, 9.4% Na2O, and 60.5% H2O. Sodium hydroxide and Sodium hydroxide pellets (HmbG®, Sigma-Aldrich, Taufkirchen, Germany) and distilled water were mixed together to prepare a sodium hydroxide solution. The main constituents of the sodium silicate solution (South Pacific Chemicals Industries Sdn. Bhd., Selangor, Malaysia) were 30.1% SiO2, 9.4% Na2O, and 60.5% H2O. Sodium hydroxide and sodium silicate solutions were mixed to prepare an alkaline activator, and the mixed solution was placed at room temperature for 24 h before use.

#### sodium silicate solutions were mixed to prepare an alkaline activator, and the mixed solution was placed at room temperature for 24 h before use. *2.2. Preparation of Unfoamed and Foamed Geopolymers*

Fly ash, slag and alkaline solution were mixed together to prepare the unfoamed geopolymer, and the geopolymer was designated as G-0. The mixture was stirred with a blender (IKA-RW-20 digital, Taufkirchen,Germany) at a constant mixing speed of 1000 rpm for 15 min before being cast into a cubic (5 × 5 × 5 cm) plastic mould. The moulded mixture was placed at room temperature for 1 day, and it was further placed in an oven (60 ◦C) for another 1 day. The curing regime was selected in accordance to Bakharev [15], who stated that pre-curing 1 day before oven-curing for another day would lead to excellent compressive strength. The demoulded sample was stored at room temperature for 28 days before being subjected to high temperatures.

On the other hand, G-0 and PAS foam were mixed to prepare the geopolymer foam. The PAS foam-to-paste ratio was set at 1:1 and 2:1, and the PAS-to-water ratio was fixed at 1:10. The selection of these ratios was based on the optimal ratios obtained by Ibrahim

et al. [5]. The foam, PAS and water were measured in volume (ml). Based on the PAS foamto-paste ratio of 1.0 and 2.0, the resulting geopolymer foam was designated as G-1 and G-2, respectively. The geopolymer foam also used the similar curing method as the unfoamed geopolymer. The constant formulations (sodium hydroxide concentration, solid:liquid, sodium silicate:sodium hydroxide and fly ash:slag ratios) were chosen as tabulated in Table 2, owing to the optimal ratios achieved in previous work [16].

**Table 2.** Details of sample preparation.

#### *2.3. High Temperature Exposure*

After 28 ambient-curing days, the unfoamed and foamed geopolymers were sintered to high temperatures (Table 2) in a Muffle furnace, with a heating rate of 10 ◦C/min and a soaking time of 2 h. The unheated geopolymer sample was also prepared as a control sample to compare with the heat-treated geopolymer sample.

#### *2.4. Test and Characterisation Method*

The water absorption (dry and immersion weights) and apparent porosity (dry, immersion and suspended weights) of geopolymer were measured based on ASTM C642. The bulk density (mass and volume) of geopolymer was measured in accordance with BS EN 12390-7. The volume and mass changes of geopolymer were measured before and after exposure to high temperatures.

The pore size distribution of geopolymer was determined by Synchrotron radiation Xray tomographic microscopy, equipped with Drishti software and Octopus Reconstruction software. The specimen was prepared in a small piece (0.2 cm × 0.2 cm × 0.5 cm) for analysis.

The thermal conductivity of geopolymer was measured by a KD2 Pro Thermal Properties Analyser (Decagon Devices Inc., Pullman, WA, USA). The transient line heat source method was applied based on IEEE 442-1981 and ASTM D5334. The pre-calibration of the sensor was performed according to DB1175 before each measurement.

The compressive strength of geopolymer before and after exposure to high temperatures was tested using a mechanical tester (Instron Machine Series 5569). Three samples were tested at each exposure temperature.

The morphology of fly ash, slag and geopolymer can be visualised by a scanning electron microscope (SEM, JEOL JSM-6460 LA). Fly ash and slag were prepared in power form while geopolymer was prepared in a small piece (1 × 1 × 1 cm) for analysis. The specimens were coated with platinum using a coater (JEOL JFC 1600) before analysis.

The functional group of fly ash, slag and geopolymer was identified using Fourier transform infrared spectroscopy (FTIR, Perkin Elmer Spectrum RX1 Spectrometer), with a resolution of 4 cm−1 and a scan range of 650 cm−1 to 4000 cm−1 . The sample was prepared in powder form for analysis.

The crystalline phase of fly ash, slag and geopolymer was carried out using an X-ray diffractometer (XRD, D2-Phaser Bruker), with 10–80◦ 2θ, 2◦/min scan rate and 0.02◦ scan step. The phase patterns of specimens were analysed by X'Pert HighScore Plus software, and the specimens were prepared in powder form for analysis.

#### **3. Results and Discussion**

#### *3.1. Effect of PAS Foam-to-Paste Ratio*

The bulk density of geopolymer can be determined by the changes in water absorption and apparent porosity, as shown in Figure 3. Geopolymer with higher porosity was expected to exhibit higher water absorption but lower bulk density. G-0 had a bulk density of 2.09 g/cm3 , water absorption of 7.4% and apparent porosity of 17.3%. The inclusion of PAS foam increased the water absorption and porosity to 17.1–20.3% and 27.9–29.1%, respectively, while reducing the bulk density to 1.69–1.77 g/cm3 in geopolymer foam. When the PAS foam was added to the geopolymer, the pores in the sample would increase directly due to the increased bubbles caused by the PAS foam addition (Figure 1), subsequently reducing the bulk density. A similar explanation was observed by Ibrahim et al. [5], who reported that the bulk density of fly ash geopolymer foam decreased from 1.65 g/cm3 to 1.20 g/cm3 with increasing PAS foam content. The bulk density of the fly ash-slag blended geopolymer foam achieved in this work was greater than in their work, which was due to the high-density slag addition [16]. *Materials* **2022**, *15*, x FOR PEER REVIEW 7 of 18

**Figure 3.** Bulk density, water absorption and apparent porosity of G-0, G-1 and G-2. **Figure 3.** Bulk density, water absorption and apparent porosity of G-0, G-1 and G-2.

The pore size distribution and total porosity of G-0, G-1 and G-2 are plotted in Figure 4. G-0 (Figure 4a), G-1 (Figure 4b) and G-2 (Figure 4c) were mostly made up of open pores, while minorly made up of closed pores. However, foamed geopolymer (14.81% for G-1 and 15.94% for G-2) exhibited a greater total porosity than G-0 (10.45%), supporting the bulk density and apparent porosity results (Figure 3). This was because of the large pore diameter of 500–1000 µm (~78%) present in the foamed geopolymers but absent in G-0. The pore size distribution and total porosity of G-0, G-1 and G-2 are plotted in Figure 4. G-0 (Figure 4a), G-1 (Figure 4b) and G-2 (Figure 4c) were mostly made up of open pores, while minorly made up of closed pores. However, foamed geopolymer (14.81% for G-1 and 15.94% for G-2) exhibited a greater total porosity than G-0 (10.45%), supporting the bulk density and apparent porosity results (Figure 3). This was because of the large pore diameter of 500–1000 µm (~78%) present in the foamed geopolymers but absent in G-0.

**Figure 4.** Total porosity and pore size distribution of (**a**) G-0, (**b**) G-1 and (**c**) G-2. **Figure 4.** Total porosity and pore size distribution of (**a**) G-0, (**b**) G-1 and (**c**) G-2.

The addition of PAS foam decreased the thermal conductivity of G-0 from 1.05 W/(m∙K) to 0.70–0.81 W/(m∙K), as demonstrated in Figure 5. The results were complied with bulk density and apparent porosity values, as shown in Figure 3. When a geopolymer has a higher porosity and a lower density, its thermal conductivity is always lower [17]. The trend was further supported by pore size distribution results (Figure 4) as the foamed geopolymers had larger pore sizes up to 1000 µm but G-0 had lower pore sizes The addition of PAS foam decreased the thermal conductivity of G-0 from 1.05 W/(m·K) to 0.70–0.81 W/(m·K), as demonstrated in Figure 5. The results were complied with bulk density and apparent porosity values, as shown in Figure 3. When a geopolymer has a higher porosity and a lower density, its thermal conductivity is always lower [17]. The trend was further supported by pore size distribution results (Figure 4) as the foamed geopolymers had larger pore sizes up to 1000 µm but G-0 had lower pore sizes up to 500 µm.

up to 500 µm.

*Materials* **2022**, *15*, x FOR PEER REVIEW 9 of 18

**Figure 5.** Thermal conductivity of G-0, G-1 and G-2.

The relationship between compressive strength, bulk density, apparent porosity and thermal conductivity of G-0, G-1 and G-2 is displayed in Figure 6. The addition of PAS foam decreased the compressive strength of G-0 from 40.5 MPa to 26.3 MPa (G-1) and 21.5 MPa (G-2), as shown in Figure 6a. As stated earlier, pores were induced in the geopolymer after adding PAS foam, thereby weakening the sample structure and resulting in poor strength. The explanation was further supported by Figure 6b, as the compressive The relationship between compressive strength, bulk density, apparent porosity and thermal conductivity of G-0, G-1 and G-2 is displayed in Figure 6. The addition of PAS foam decreased the compressive strength of G-0 from 40.5 MPa to 26.3 MPa (G-1) and 21.5 MPa (G-2), as shown in Figure 6a. As stated earlier, pores were induced in the geopolymer after adding PAS foam, thereby weakening the sample structure and resulting in poor strength. The explanation was further supported by Figure 6b, as the compressive strength had a strong correlation with apparent porosity, bulk density and thermal conductivity, with correlation coefficients (R2 ) of 0.93881, 0.98106 and 0.97550, respectively. According to the functional classification of International Union of Laboratories in Construction Materials, Systems and Structures (RILEM), a sample with a density of 1.44– 1.84 g/cm3, compressive strength of >17.0 MPa and thermal conductivity of 0.4–0.7 W/(m∙K) can be used as a structural lightweight concrete [17]. Based on the compressive strength, density and thermal conductivity values achieved in this work, only G-2 (compressive strength of 21.5 MPa, density of 1.69 g/cm3 and thermal conductivity of 0.7 W/(m∙K)) met the RILEM requirements for the application.

ductivity, with correlation coefficients (R2) of 0.93881, 0.98106 and 0.97550, respectively.

**Figure 6.** (**a**) Compressive strength and (**b**) correlation curve between compressive strength, bulk density, apparent porosity and thermal conductivity of G-0, G-1 and G-2. **Figure 6.** (**a**) Compressive strength and (**b**) correlation curve between compressive strength, bulk density, apparent porosity and thermal conductivity of G-0, G-1 and G-2.

The compressive strength of the fly ash-slag blended geopolymer foam achieved in this work (21.5–26.3 MPa) was higher than the work done by Ibrahim et al. [5], who adopted the same mixing ratios of PAS foam-to-paste and PAS-to-water as in this work to produce fly ash geopolymer foam (10.2–15.8 MPa). The inclusion of slag promoted a combined matrix of calcium (CASH) and sodium (NASH) aluminosilicate hydrate in the According to the functional classification of International Union of Laboratories in Construction Materials, Systems and Structures (RILEM), a sample with a density of 1.44–1.84 g/cm3 , compressive strength of >17.0 MPa and thermal conductivity of 0.4–0.7 W/(m·K) can be used as a structural lightweight concrete [17]. Based on the compressive strength, density and thermal conductivity values achieved in this work, only G-2 (compressive strength of 21.5 MPa, density of 1.69 g/cm3 and thermal conductivity of 0.7 W/(m·K)) met the RILEM requirements for the application.

The compressive strength of the fly ash-slag blended geopolymer foam achieved in this work (21.5–26.3 MPa) was higher than the work done by Ibrahim et al. [5], who adopted

density, apparent porosity and thermal conductivity of G-0, G-1 and G-2.

**Figure 6.** (**a**) Compressive strength and (**b**) correlation curve between compressive strength, bulk

this work (21.5–26.3 MPa) was higher than the work done by Ibrahim et al. [5], who adopted the same mixing ratios of PAS foam-to-paste and PAS-to-water as in this work to produce fly ash geopolymer foam (10.2–15.8 MPa). The inclusion of slag promoted a combined matrix of calcium (CASH) and sodium (NASH) aluminosilicate hydrate in the

the same mixing ratios of PAS foam-to-paste and PAS-to-water as in this work to produce fly ash geopolymer foam (10.2–15.8 MPa). The inclusion of slag promoted a combined matrix of calcium (CASH) and sodium (NASH) aluminosilicate hydrate in the blended geopolymer foam, while fly ash geopolymer foam only had a NASH matrix [10], thereby obtaining higher compressive strength in the blended geopolymer foam compared to the pure geopolymer foam. blended geopolymer foam, while fly ash geopolymer foam only had a NASH matrix [10], thereby obtaining higher compressive strength in the blended geopolymer foam compared to the pure geopolymer foam. *3.2. High Temperature Performance*

#### *3.2. High Temperature Performance* Figure 7 demonstrates the mass, volume and bulk density of G-0, G-1 and G-2 before

Figure 7 demonstrates the mass, volume and bulk density of G-0, G-1 and G-2 before and after exposure to high temperatures. Increasing temperature caused the bulk density of all geopolymers to gradually reduce up to 600 ◦C, and the bulk density was relatively stable between 600 ◦C and 1000 ◦C (Figure 7a). The changes in bulk density were varied by the mass and volume. The mass (Figure 7b) and volume (Figure 7c) of all geopolymers were reduced with increasing temperature up to 600 ◦C, and the mass and volume remained almost constant above 600 ◦C. The reduction in mass and volume at temperatures below 600 ◦C was affected by the loss of physically and chemically bonded water in the sample structure [11]. A geopolymer sample with a higher mass or lower volume was expected to have a higher bulk density. and after exposure to high temperatures. Increasing temperature caused the bulk density of all geopolymers to gradually reduce up to 600 °C, and the bulk density was relatively stable between 600 °C and 1000 °C (Figure 7a). The changes in bulk density were varied by the mass and volume. The mass (Figure 7b) and volume (Figure 7c) of all geopolymers were reduced with increasing temperature up to 600 °C, and the mass and volume remained almost constant above 600 °C. The reduction in mass and volume at temperatures below 600 °C was affected by the loss of physically and chemically bonded water in the sample structure [11]. A geopolymer sample with a higher mass or lower volume was expected to have a higher bulk density.

**Figure 7.** (**a**) Bulk density, (**b**) mass and (**c**) volume of G-0, G-1 and G-2 before and after exposure to high temperatures. **Figure 7.** (**a**) Bulk density, (**b**) mass and (**c**) volume of G-0, G-1 and G-2 before and after exposure to high temperatures.

At high temperatures, G-0 (1.97–2.03 g/cm3) had a higher bulk density than the foamed geopolymers (1.57–1.70 g/cm3). This was due to G-0 obtaining a higher remaining

120.75 cm3). The lower residual mass and volume of foamed geopolymer were due to the increased dehydration caused by the high-moisture PAS foam addition. In terms of bulk density loss, the density loss of the fly ash-slag blended geopolymer foam in this work

At high temperatures, G-0 (1.97–2.03 g/cm3 ) had a higher bulk density than the foamed geopolymers (1.57–1.70 g/cm3 ). This was due to G-0 obtaining a higher remaining mass (230.25–247.60 g) and volume (116.41–121.96 cm3 ) after exposure to high temperatures compared to the foamed geopolymer (mass of 173.43–200.78 g and volume of 110.46–120.75 cm3 ). The lower residual mass and volume of foamed geopolymer were due to the increased dehydration caused by the high-moisture PAS foam addition. In terms of bulk density loss, the density loss of the fly ash-slag blended geopolymer foam in this work (4.94–7.43%) was comparatively lower than that of the fly ash geopolymer foam (17.65–28.66%) produced by Mastura et al. [2], verifying that the blended geopolymer foam showed better structural stability than the pure geopolymer foam at high temperatures. *Materials* **2022**, *15*, x FOR PEER REVIEW 11 of 18 (4.94–7.43%) was comparatively lower than that of the fly ash geopolymer foam (17.65– 28.66%) produced by Mastura et al. [2], verifying that the blended geopolymer foam showed better structural stability than the pure geopolymer foam at high temperatures.

> G-1 and G-2 also had greater water absorption and apparent porosity than G-0 at high temperatures, as plotted in Figure 8. The porosity and water absorption of G-0, G-1 and G-2 showed parallel and similar trends, increasing from 29 ◦C to 1000 ◦C. The increased porosity from 29 ◦C to 600 ◦C was associated with the abovementioned dehydration of the sample, leading to the formation of pores in G-0, G-1 and G-2. The slight increase in porosity between 600 ◦C and 1000 ◦C might be related to the recrystallisation effect, as described in the phase analysis below. The induced new crystals caused the geopolymer matrix to slightly open and form pores in the unfoamed and foamed geopolymers. G-1 and G-2 also had greater water absorption and apparent porosity than G-0 at high temperatures, as plotted in Figure 8. The porosity and water absorption of G-0, G-1 and G-2 showed parallel and similar trends, increasing from 29 °C to 1000 °C. The increased porosity from 29 °C to 600 °C was associated with the abovementioned dehydration of the sample, leading to the formation of pores in G-0, G-1 and G-2. The slight increase in porosity between 600 °C and 1000 °Cmight be related to the recrystallisation effect, as described in the phase analysis below. The induced new crystals caused the geopolymer matrix to slightly open and form pores in the unfoamed and foamed geopolymers.

**Figure 8.** (**a**) Water absorption and (**b**) apparent porosity of G-0, G-1 and G-2 before and after being subjected to high temperatures. **Figure 8.** (**a**) Water absorption and (**b**) apparent porosity of G-0, G-1 and G-2 before and after being subjected to high temperatures.

Figure 9 depicts the compressive strength changes of G-0, G-1 and G-2 before and after being subjected to high temperatures. The compressive strength of all geopolymers was identical at below 600 °C, increasing from 29 °C to 200 °C and decreasing from 200 °C to 600 °C. The temperature of 200 °C caused better dissolution of slag, fly ash and alkali activator, thereby increasing the compressive strength [18]. The presence of pores (Figure 8) and bulk density loss (Figure 7) caused by the dehydration process resulted in a decrease in compressive strength at 200–600 °C. Beyond 600 °C, the trend in compressive strength of G-0 was different from that of G-1 and G-2. The strength of G-0 increased slightly at 800 °C due to the dense microstructure obtained, while the strength of G-1 and G-2 was continuously reduced at 800 °C due to the loose microstructure obtained, as shown in the following microstructural analysis. At 1000 °C, the strength of all geopolymers increased slightly was affected by phase change, as shown in the phase analysis below, which was attested by Murri et al. [13] for unfoamed geopolymer. Figure 9 depicts the compressive strength changes of G-0, G-1 and G-2 before and after being subjected to high temperatures. The compressive strength of all geopolymers was identical at below 600 ◦C, increasing from 29 ◦C to 200 ◦C and decreasing from 200 ◦C to 600 ◦C. The temperature of 200 ◦C caused better dissolution of slag, fly ash and alkali activator, thereby increasing the compressive strength [18]. The presence of pores (Figure 8) and bulk density loss (Figure 7) caused by the dehydration process resulted in a decrease in compressive strength at 200–600 ◦C. Beyond 600 ◦C, the trend in compressive strength of G-0 was different from that of G-1 and G-2. The strength of G-0 increased slightly at 800 ◦C due to the dense microstructure obtained, while the strength of G-1 and G-2 was continuously reduced at 800 ◦C due to the loose microstructure obtained, as shown in the following microstructural analysis. At 1000 ◦C, the strength of all geopolymers increased slightly was affected by phase change, as shown in the phase analysis below, which was attested by Murri et al. [13] for unfoamed geopolymer.

At high temperatures, G-0 (36.9–43.1 MPa) had a greater compressive strength than the foamed geopolymers (25.1–36.2 MPa). The explanation was consistent with the higher residual bulk density, mass and volume (Figure 7) and lower porosity (Figure 8) obtained in G-0 compared to the foamed geopolymers. The addition of PAS foam degraded the compressive strength of the control sample, resulting in the unfoamed geopolymer pos-

at the same temperature exposure. An exception occurred in G-1 at 800 °C due to strength loss (4.6%). As stated earlier, foamed geopolymer had a higher porosity than the un-

sessing better heat resistance than the foamed geopolymer.

temperatures.

**Figure 9.** (**a**) Compressive strength and (**b**) strength changes of G-0, G-1 and G-2 with respect to unexposed samples. **Figure 9.** (**a**) Compressive strength and (**b**) strength changes of G-0, G-1 and G-2 with respect to unexposed samples.

foamed geopolymer. High porosity allowed moisture to be removed during heating without deteriorating the structure of the foamed geopolymer, while pressure was generated in the unfoamed geopolymer during the moisture removal process and resulted in a loss of strength. Therefore, the compressive strength of the foamed geopolymer at high temperatures was greatly improved due to the increased pores, especially in the case of G-2. The increase in porosity improved the thermal resistance of geopolymer foam [19]. In addition, the lower thermal conductivity value (Figure 5) of foamed geopolymer, particularly G-2, facilitated the heat transfer and reduced the damaging effect of heat at elevated

Mastura et al. [2] heated fly ash geopolymer foam at temperatures ranging from 200 °C to 800 °C, obtaining a compressive strength in the range of 18.6 MPa to 33.3 MPa and a strength loss in the range of 29.3% to 60.4%. The fly ash-slag blended geopolymer foam produced in this work achieved better compressive strength (25.1–36.2 MPa) than their work even at high temperatures up to 1000 °C. The strength in this work even increased At high temperatures, G-0 (36.9–43.1 MPa) had a greater compressive strength than the foamed geopolymers (25.1–36.2 MPa). The explanation was consistent with the higher residual bulk density, mass and volume (Figure 7) and lower porosity (Figure 8) obtained in G-0 compared to the foamed geopolymers. The addition of PAS foam degraded the compressive strength of the control sample, resulting in the unfoamed geopolymer possessing better heat resistance than the foamed geopolymer.

(4.2–68.6%) at high temperatures instead of losing strength, which was also different from their work. Thus, the geopolymer foam made with blended precursors was more stable than the geopolymer foam made with a single precursor at high temperatures. The influence of various foaming techniques and foaming agents on the thermo-mechanical performance of geopolymer foam has been widely investigated. Geopolymer foam prepared by varying foaming methods and foaming agents resulted in different thermo-mechanical properties. For instance, Bai et al. [8] added hydrogen peroxide (H2O2) into geopolymer by adopting the direct foaming method, obtaining a compressive strength in the range of 4.5 MPa to 20.4 MPa at high temperatures. Le-ping et al. [9] used the direct foaming method to incorporate alumina powder (AP) into geopolymer, obtaining compressive strengths ranging from 6.3 MPa to 13.8 MPa at high temperatures. This work and Mastura et al. [2] work used the pre-foaming method to add PAS foam into geopolymer, with compressive strength of 18.6–36.2 MPa at high temperatures, which However, referring to Figure 9b, G-0 lost compressive strength (2.2–8.7%) between 400 ◦C and 1000 ◦C, while foamed geopolymer gained compressive strength (4.2–68.6%) at the same temperature exposure. An exception occurred in G-1 at 800 ◦C due to strength loss (4.6%). As stated earlier, foamed geopolymer had a higher porosity than the unfoamed geopolymer. High porosity allowed moisture to be removed during heating without deteriorating the structure of the foamed geopolymer, while pressure was generated in the unfoamed geopolymer during the moisture removal process and resulted in a loss of strength. Therefore, the compressive strength of the foamed geopolymer at high temperatures was greatly improved due to the increased pores, especially in the case of G-2. The increase in porosity improved the thermal resistance of geopolymer foam [19]. In addition, the lower thermal conductivity value (Figure 5) of foamed geopolymer, particularly G-2, facilitated the heat transfer and reduced the damaging effect of heat at elevated temperatures.

showed better strength and thermal performance than the geopolymer foam incorporated with H2O2 or AP. This was because PAS was an amphiprotic substance that was strongly hydrophilic and easily dissolved in water during the pre-foaming process, yielding and resulting in the formation of homogeneous air bubbles and pores [20]. However, when H2O2 or AP interacted with the geopolymer matrix, pores of varying sizes and shapes were formed during the hydrogen gas generation [3]. Thus, PAS-geopolymer foam prepared by the pre-foaming method achieved a more stable pore size and shape when com-Mastura et al. [2] heated fly ash geopolymer foam at temperatures ranging from 200 ◦C to 800 ◦C, obtaining a compressive strength in the range of 18.6 MPa to 33.3 MPa and a strength loss in the range of 29.3% to 60.4%. The fly ash-slag blended geopolymer foam produced in this work achieved better compressive strength (25.1–36.2 MPa) than theirwork even at high temperatures up to 1000 ◦C. The strength in this work even increased(4.2–68.6%) at high temperatures instead of losing strength, which was also different from their work. Thus, the geopolymer foam made with blended precursors was more stable than the geopolymer foam made with a single precursor at high temperatures.

The influence of various foaming techniques and foaming agents on the thermomechanical performance of geopolymer foam has been widely investigated. Geopolymer foam prepared by varying foaming methods and foaming agents resulted in different thermo-mechanical properties. For instance, Bai et al. [8] added hydrogen peroxide (H2O2) into geopolymer by adopting the direct foaming method, obtaining a compressive strength in the range of 4.5 MPa to 20.4 MPa at high temperatures. Le-ping et al. [9] used the direct foaming method to incorporate alumina powder (AP) into geopolymer, obtaining compressive strengths ranging from 6.3 MPa to 13.8 MPa at high temperatures. This work and Mastura et al. [2] work used the pre-foaming method to add PAS foam into geopolymer, with compressive strength of 18.6–36.2 MPa at high temperatures, which

showed better strength and thermal performance than the geopolymer foam incorporated with H2O2 or AP. This was because PAS was an amphiprotic substance that was strongly hydrophilic and easily dissolved in water during the pre-foaming process, yielding and resulting in the formation of homogeneous air bubbles and pores [20]. However, when H2O2 or AP interacted with the geopolymer matrix, pores of varying sizes and shapes were formed during the hydrogen gas generation [3]. Thus, PAS-geopolymer foam prepared by the pre-foaming method achieved a more stable pore size and shape when compared to the H2O2- or AP-geopolymer foam prepared by the direct foaming method, resulting in PAS-geopolymer foam outperforming H2O2- or AP-geopolymer foam at high temperatures. *Materials* **2022**, *15*, x FOR PEER REVIEW 13 of 18 pared to the H2O2- or AP-geopolymer foam prepared by the direct foaming method, resulting in PAS-geopolymer foam outperforming H2O2- or AP-geopolymer foam at high temperatures.

#### *3.3. Microstructural Analysis 3.3. Microstructural Analysis*

The SEM images of fly ash and slag were taken at a magnification of × 300, as shown in Figure 10. The fly ash particles (Figure 10a) were spherical, while the slag particles (Figure 10b) were presented in block shape. On the other hand, the SEM images of unheated G-0 (Figure 10c) and G-2 (Figure 10d) were taken at a magnification of × 500. G-0 and G-2 were chosen due to the greatest difference in compressive strength between G-0 and G-2 (Figure 6). Generally, the geopolymer matrix in G-0 and G-2 showed some pores, cracks, remaining fly ash and slag particles. A more compact microstructure was observed in G-0 as compared to G-2. The larger pores and cracks of G-2 were caused by the pores induced by PAS foam addition, as supported by bulk density and porosity results (Figures 3 and 4), thereby achieving a low compressive strength. The SEM images of fly ash and slag were taken at a magnification of × 300, as shown in Figure 10. The fly ash particles (Figure 10a) were spherical, while the slag particles (Figure 10b) were presented in block shape. On the other hand, the SEM images of unheated G-0 (Figure 10c) and G-2 (Figure 10d) were taken at a magnification of × 500. G-0 and G-2 were chosen due to the greatest difference in compressive strength between G-0 and G-2 (Figure 6). Generally, the geopolymer matrix in G-0 and G-2 showed some pores, cracks, remaining fly ash and slag particles. A more compact microstructure was observed in G-0 as compared to G-2. The larger pores and cracks of G-2 were caused by the pores induced by PAS foam addition, as supported by bulk density and porosity results (Figures 3 and4), thereby achieving a low compressive strength.

**Figure 10.** SEM images of (**a**) fly ash, (**b**) slag, unheated (**c**) G-0 and (**d**) G-2 (FA denoted fly ash). **Figure 10.** SEM images of (**a**) fly ash, (**b**) slag, unheated (**c**) G-0 and (**d**) G-2 (FA denoted fly ash).

Figure 11 depicts the SEM images of G-0 and G-2 after being heated to 800 °C and 1000 °C. The temperatures of 800 °C and 1000 °C were selected based on the different compressive strength trends of G-0 and G-2 at 800 °C (Figure 9). At 800 °C, the geopolymer matrix in G-0 was observed to be connected and smooth with a small number of pores and residual particles (Figure 11a), leading to a slight improvement in the compressive strength. However, the compressive strength of G-2 was further degraded at 800 °C due to the loose matrix obtained with a large number of pores (Figure 11b). At 1000 °C, the disappearance of residual particles and the appearance of the solidified melt in the geo-Figure 11 depicts the SEM images of G-0 and G-2 after being heated to 800 ◦C and 1000 ◦C. The temperatures of 800 ◦C and 1000 ◦C were selected based on the different compressive strength trends of G-0 and G-2 at 800 ◦C (Figure 9). At 800 ◦C, the geopolymer matrix in G-0 was observed to be connected and smooth with a small number of pores and residual particles (Figure 11a), leading to a slight improvement in the compressive strength. However, the compressive strength of G-2 was further degraded at 800 ◦C due to the loose matrix obtained with a large number of pores (Figure 11b). At 1000 ◦C, the disappearance of residual particles and the appearance of the solidified melt in the geopolymer matrix [21]

polymer matrix [21] most likely contributed to the increase in compressive strength of G-0 (Figure 11c) and G-2 (Figure 11d). The presence of rod and/or plate anorthite crystal

most likely contributed to the increase in compressive strength of G-0 (Figure 11c) and G-2 (Figure 11d). The presence of rod and/or plate anorthite crystal (Figure 11c') was also expected to enhance the compressive strength of G-0 at 1000 ◦C. Anorthite was typically discovered in a high-calcium geopolymer system at 1000 ◦C [22]. *Materials* **2022**, *15*, x FOR PEER REVIEW 14 of 18

**Figure 11.** SEM images of G-0 and G-2 after exposed to (**a**,**b**) 800 °C and (**c**,**c'**,**d**) 1000 °C (FA denoted fly ash). **Figure 11.** SEM images of G-0 and G-2 after exposed to (**a**,**b**) 800 ◦C and (**c**,**c'**,**d**) 1000 ◦C (FA denoted fly ash).

### *3.4. Phase Analysis*

activation of fly ash and slag.

*3.4. Phase Analysis* Figure 12 illustrates the phase patterns of fly ash, slag, unexposed, and temperatureexposed G-0 and G-2. Fly ash had an amorphous broad hump in the range of 15° to 35° 2θ, with some peaks of mullite, quartz and hematite. Slag was mainly crystalline with peaks of calcio-olivine, merwinite, magnetite and calcium aluminium oxide. Unheated G-0 and G-2 (Figure 12b,c) had an amorphous broad hump between 16–38° 2θ, also had some peaks of mullite, quartz, hematite, calcio-olivine and calcium aluminium oxide, which came from raw fly ash and slag. In addition, some new peaks of calcite and calcium silicate hydrate (CSH) appeared in the unheated samples, which were caused by alkali Figure 12 illustrates the phase patterns of fly ash, slag, unexposed, and temperatureexposed G-0 and G-2. Fly ash had an amorphous broad hump in the range of 15◦ to 35◦ 2θ, with some peaks of mullite, quartz and hematite. Slag was mainly crystalline with peaks of calcio-olivine, merwinite, magnetite and calcium aluminium oxide. Unheated G-0 and G-2 (Figure 12b,c) had an amorphous broad hump between 16–38◦ 2θ, also had some peaks of mullite, quartz, hematite, calcio-olivine and calcium aluminium oxide, which came from raw fly ash and slag. In addition, some new peaks of calcite and calcium silicate hydrate (CSH) appeared in the unheated samples, which were caused by alkali activation of fly ash and slag.

The rise in temperature caused the (i) crystalline peaks to change, increase, reduce, disappear and/or decompose and (ii) amorphous phase (geopolymer matrix) to reduce and/or decompose (Figure 12b,c). Calcite and CSH vanished and decomposed at 800 °C, as supported by Shaikh [23] for unfoamed geopolymer. The combined decomposition of the geopolymer matrix, calcite and CSH caused nepheline, albite and anorthite to form at 800 °C and 1000 °C. The presence of anorthite was also confirmed by the SEM image in

**Figure 12.** Phase patterns of (**a**) fly ash and slag, unexposed and temperature-exposed (**b**) G-0 and (**c**) G-2. **Figure 12.** Phase patterns of (**a**) fly ash and slag, unexposed and temperature-exposed (**b**) G-0 and (**c**) G-2.

*3.5. Functional Group Identification* Figure 13 plots the IR spectra of fly ash, slag, unexposed and temperature-exposed G-0 and G-2. The main band of raw fly ash was located at 1031 cm−1, corresponding to the asymmetric stretching vibration of Si-O-X (X = Si or Al) [24]. The band at 775 cm−1 was expressed as the symmetric stretching vibration of Si-O-Si [25]. Moreover, the main band of raw slag was located at 858 cm−1, indicating the stretching vibration of Ca-O and Si-O [26]. The bands at 2168 cm−1, 2025 cm−1 and 1418 cm−1 were verified as the stretching vi-The rise in temperature caused the (i) crystalline peaks to change, increase, reduce, disappear and/or decompose and (ii) amorphous phase (geopolymer matrix) to reduce and/or decompose (Figure 12b,c). Calcite and CSH vanished and decomposed at 800 ◦C, as supported by Shaikh [23] for unfoamed geopolymer. The combined decomposition of the geopolymer matrix, calcite and CSH caused nepheline, albite and anorthite to form at 800 ◦C and 1000 ◦C. The presence of anorthite was also confirmed by the SEM image in Figure 11c'. The addition of PAS foam did not cause any changes to the phase patterns of geopolymer at room temperature and high temperatures.

#### brations of O-C-O [27]. The bands at 2805 cm−1 and 3325 cm−1 were denoted as the stretching vibrations of OH and H-O-H [28]. *3.5. Functional Group Identification*

matrix was formed [32].

The presence of O-C-O (~2160, ~2020 and ~1420 cm−1) and OH and H-O-H (~3250 cm−1) stretching vibrations in the unheated G-0 and G-2 originated from raw fly ash and slag. New bands of calcite [29] and H-O-H bending vibration [30] appeared at ~2350 cm−1 and ~1650 cm−1, respectively, in unheated G-0 and G-2. Calcite disappeared and decomposed at 800 °C, as proven by the phase analysis in Figure 12. The wavenumber of fly ash at 1031 cm−1 and 775 cm−1 shifted to ~980 cm−1 and ~680 cm−1, respectively, in unheated G-0 and G-2, inferring the NASH matrix was formed [28,31]. The wavenumber of slag at 858 cm−1 shifted to a higher value in the unheated G-0 and G-2 (~860 cm−1), implying the CASH Figure 13 plots the IR spectra of fly ash, slag, unexposed and temperature-exposed G-0 and G-2. The main band of raw fly ash was located at 1031 cm−1 , corresponding to the asymmetric stretching vibration of Si-O-X (X = Si or Al) [24]. The band at 775 cm−1 was expressed as the symmetric stretching vibration of Si-O-Si [25]. Moreover, the main band of raw slag was located at 858 cm−1 , indicating the stretching vibration of Ca-O and Si-O [26]. The bands at 2168 cm−1 , 2025 cm−1 and 1418 cm−1 were verified as the stretching vibrations of O-C-O [27]. The bands at 2805 cm−1 and 3325 cm−1 were denoted as the stretching vibrations of OH and H-O-H [28].

The rise in temperature changed the intensity of the absorption bands of G-0 and G-2. The main band at ~980 cm−1 widened at high temperatures, especially at 1000 °C, indicating that the high temperatures led to the formation of structural disorder, as proven by the literature for unfoamed geopolymer [33,34]. The intensity of ~1650 cm−1 reduced at high temperatures, inferring that the geopolymer had dehydrated [19]. The intensity at

~860 cm−1 vanished at 1000 °C, implying that the cross-linking in the CASH matrix was lessened [35]. The addition of PAS foam did not create any new bands in the geopolymer

at room temperature and high temperatures.

**Figure 13.** IR spectra of fly ash, slag, unexposed and temperature-exposed G-0 and G-2. **Figure 13.** IR spectra of fly ash, slag, unexposed and temperature-exposed G-0 and G-2.

**4. Conclusions** This paper compares the thermo-mechanical performance of fly ash-ladle furnace slag blended geopolymer with and without PAS foam addition at 29–1000 °C. The PAS foam-to-paste ratio of 1.0 and 2.0 was used to prepare geopolymer foam. The incorporation of PAS foam degraded the compressive strength of unheated G-0 (40.5 MPa) by 35.0– 46.9%, and also decreased the heat-treated G-0 (36.9–43.1 MPa) by 7.6–33.9%. Nonetheless, when compared to unheated samples, the compressive strength of the heat-treated G-0 decreased by 2.2–8.7% at 400–1000 °C, whereas the compressive strength of the heat-The presence of O-C-O (~2160, ~2020 and ~1420 cm−1 ) and OH and H-O-H (~3250 cm−1 ) stretching vibrations in the unheated G-0 and G-2 originated from raw fly ash and slag. New bands of calcite [29] and H-O-H bending vibration [30] appeared at ~2350 cm−1 and ~1650 cm−1 , respectively, in unheated G-0 and G-2. Calcite disappeared and decomposed at 800 ◦C, as proven by the phase analysis in Figure 12. The wavenumber of fly ash at 1031 cm−1 and 775 cm−1 shifted to ~980 cm−1 and ~680 cm−1 , respectively, in unheated G-0 and G-2, inferring the NASH matrix was formed [28,31]. The wavenumber of slag at 858 cm−1 shifted to a higher value in the unheated G-0 and G-2 (~860 cm−1 ), implying the CASH matrix was formed [32].

treated G-1 and G-2 increased by 4.2–68.5% at the same temperature exposure. The higher porosity, lower thermal conductivity and lower connectivity of the geopolymer matrix aided dehydration and reduced deterioration of the geopolymer structure, resulting in a significant improvement in the strength loss of foamed geopolymer. **Author Contributions:** Conceptualisation, N.H.-T. and H.C.-Y.; methodology, N.H.-T.; software, N.H.-T. and C.R.; validation, N.H.-T.; formal analysis, N.H.-T., S.G. and M.N.; investigation, N.H.- T.; resources, M.M.A.B.A.; writing—original draft preparation, N.H.-T.; writing—review and editing, The rise in temperature changed the intensity of the absorption bands of G-0 and G-2. The main band at ~980 cm−1 widened at high temperatures, especially at 1000 ◦C, indicating that the high temperatures led to the formation of structural disorder, as proven by the literature for unfoamed geopolymer [33,34]. The intensity of ~1650 cm−1 reduced at high temperatures, inferring that the geopolymer had dehydrated [19]. The intensity at ~860 cm−1 vanished at 1000 ◦C, implying that the cross-linking in the CASH matrix was lessened [35]. The addition of PAS foam did not create any new bands in the geopolymer at room temperature and high temperatures.

L.Y.-M.; data curation and visualisation, H.M.R., I.M.Z.A., N.Y.-S., W.S., A.Ś. and A.V.S.; supervision,

#### H.C.-Y. and L.Y.-M. All authors have read and agreed to the published version of the manuscript. **4. Conclusions**

**Funding:** This publication was supported by Gheorghe Asachi Technical University of Iasi TUIASI Internal Grants Program (GI\_Publications/2021), financed by the Romanian Government. **Institutional Review Board Statement:** Not applicable. **Informed Consent Statement:** Not applicable. This paper compares the thermo-mechanical performance of fly ash-ladle furnace slag blended geopolymer with and without PAS foam addition at 29–1000 ◦C. The PAS foam-to-paste ratio of 1.0 and 2.0 was used to prepare geopolymer foam. The incorporation of PAS foam degraded the compressive strength of unheated G-0 (40.5 MPa) by 35.0–46.9%, and also decreased the heat-treated G-0 (36.9–43.1 MPa) by 7.6–33.9%. Nonetheless, when compared to unheated samples, the compressive strength of the heat-treated G-0 decreased by 2.2–8.7% at 400–1000 ◦C, whereas the compressive strength of the heat-treated G-1 and G-2 increased by 4.2–68.5% at the same temperature exposure. The higher porosity, lower thermal conductivity and lower connectivity of the geopolymer matrix aided dehydration

and reduced deterioration of the geopolymer structure, resulting in a significant improvement in the strength loss of foamed geopolymer.

**Author Contributions:** Conceptualisation, N.H.-T. and H.C.-Y.; methodology, N.H.-T.; software, N.H.-T. and C.R.; validation, N.H.-T.; formal analysis, N.H.-T., S.G. and M.N.; investigation, N.H.-T.; resources, M.M.A.B.A.; writing—original draft preparation, N.H.-T.; writing—review and editing, L.Y.-M.; data curation and visualisation, H.M.R., I.M.Z.A., N.Y.-S., W.S., A.S. and A.V.S.; supervision, ´ H.C.-Y. and L.Y.-M. All authors have read and agreed to the published version of the manuscript.

**Funding:** This publication was supported by Gheorghe Asachi Technical University of Iasi TUIASI Internal Grants Program (GI\_Publications/2021), financed by the Romanian Government.

**Institutional Review Board Statement:** Not applicable.

**Informed Consent Statement:** Not applicable.

**Data Availability Statement:** All the data is available within the manuscript.

**Acknowledgments:** The authors wish to gratefully acknowledge Sultan Azlan Shah Power Station, TNB Janamanjung Sdn Bhd, Seri Manjung, Perak and Southern Steel Berhad, Perai, Penang for supplying the fly ash and ladle furnace slag, respectively, for this research.

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

#### **References**

## *Article* **Freeze–Thaw Effect on Road Concrete Containing Blast Furnace Slag: NMR Relaxometry Investigations**

**Liliana Maria Nicula 1 , Ofelia Corbu 1,2,3,\* , Ioan Ardelean 4 , Andrei Victor Sandu 3,5,6,\* , Mihai Iliescu 1 and Dorina Simedru 7**

**Abstract:** The present work investigates the effect of freeze–thaw cycles on the porosity of three mixtures of road concrete containing blast furnace slag in comparison with two mixtures made with conventional materials. The main technique used in our investigations is nuclear magnetic resonance (NMR) relaxometry. This permitted the extraction of information with respect to the freeze–thaw effect on pore-size distribution, which influences both the mechanical strength and the molecular transport through the material. Moreover, by using this technique, the structure of the air voids was analyzed for the entire pore system in the cement paste and the aggregate particles. The samples under study were first dried in a vacuum oven and then saturated with water or cyclohexane where the distribution of the transverse relaxation times of the protons was recorded. The NMR relaxation measurements were performed on samples extracted from specimens maintained at 300 freeze–thaw cycles and on control samples extracted from specimens kept in water during the freeze–thaw period. Scanning Electron Microscopy (SEM) was used to analyze the microstructure of concrete samples in order to obtain information about the pore sizes and the distance between them. The results from the NMR relaxation measurements were consistent with those obtained by using standard techniques for determining the porosity and the freeze–thaw resistances. The investigations made it possible to establish the optimal composition of blast furnace slag that can be incorporated into road concrete compositions. This non-invasive technique can also complete standard techniques for assessing the porosity and the progress of internal cracks during the freeze–thaw test.

**Keywords:** road concrete; porosity; (NMR) relaxometry; freeze–thaw; blast furnace slag

## **1. Introduction**

The resistance to freezing is a factor that needs to be considered when designing road concrete structures because the freeze–thaw phenomenon produces deterioration of the internal structure and on the surface layer of the concrete. If the concrete is exposed to a dry environment the structure is not usually affected by the freeze–thaw phenomenon. However, due to the fact that the road concrete is in contact with water, the water is absorbed by the concrete in the structure of the pores and it generates a high internal

**Citation:** Nicula, L.M.; Corbu, O.; Ardelean, I.; Sandu, A.V.; Iliescu, M.; Simedru, D. Freeze–Thaw Effect on Road Concrete Containing Blast Furnace Slag: NMR Relaxometry Investigations. *Materials* **2021**, *14*, 3288. https://doi.org/10.3390/ ma14123288

Academic Editor: Jeong Gook Jang

Received: 10 May 2021 Accepted: 10 June 2021 Published: 14 June 2021

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tension, which leads to internal structure damage and produces irreversible changes of its properties [1–3].

The most important factor influencing freeze–thaw resistance is the distribution and dimensions of the pores [1,4]. The free spaces filled with water in the fresh cement paste or the spaces resulting from the reaction of cement particles during the process of hydration of the cement paste represents the fraction of total volume unoccupied by solid matter and is known as porosity. By mixing components, the cement particles are spread in the total volume of materials and, upon contact with water, reaction takes place between the mineral components and the solution forming around the cement particles. The dispersion of cement particles as evenly as possible is improved by the superplasticizers used [5]. These particles are transformed in colloids by dispersing the solid substance in the liquid due to the forces of attraction at the molecular and atomic level, but the water continues to penetrate inside these colloids. The pore size does not change much during the mixing and dormancy stage in which the activity of the chemical compounds slows down, but there is a change in their size that commences during the hardening stage [6]. The hardened cement paste in the cooling and densifying stage consists of hydrates of different compounds in the form of a gel, calcium hydroxide crystals Ca(OH)2, and other minor components, such as un-hydrated cement, and free spaces left in the hydration process [7]. Moreover, additional pores may result in the concrete application stage due to improper compaction. The free spaces that have not been filled with hydration products form capillary pores and the free spaces formed in gel as interconnected spaces define the gel pores. These gel pores can be classified into intra-C-S-H and inter-C-S-H pores [8].

The total volume of voids in concrete can occupy between 1% and 10% of the total volume and a porous concrete at the upper limit indicates low strength [7]. The capillary pores form an interconnected system spread irregularly throughout the mass of the paste [7,9], with sizes ranging from 50 nm and up to 5 µm [10]. The dimensions of the gel pores are much smaller than the capillary pores (approximately 1 ÷ 50 nm) and, therefore, are classified as micropores. The gel pores absorb water even at very low values of ambient humidity [7,9] due to the hygroscopic property of the cement paste and the presence of submicroscopic pores. The spaces with larger dimensions of 5 ÷ 10 µm are considered as entrained air [10].

The PN-EN 480-11 standard [11] describes the air-voids system in the hardened concrete by the following parameters: total air content A (%), specific surface α (mm2 ), spacing factor L (mm) and the micropores content A300 (%). The spacing factor (L) of the air pores can be determined using the simplified Powers' model, which assumes that all air bubbles have the same diameter and are distributed in the corners of a cube. Another approach for determining the spacing factor (L) of the air pores is based on the Pilleo concept, which considers both the air voids system in the cement paste and in the aggregate particles; the principles of the method were described by Wawrze´nczyk J. and Kozak W. in their paper [12]. The sustainability requirements limit the spacing factor to 0.20 mm in Norway and Denmark and to 0.25 mm in Canada [13]. Yuan J. [14] experimented with the variation of concrete pores in the action of freeze–thaw cycles using X-ray tomography (CT). The results showed that there was no large variation in the amounts and volume of pores located in the range (0.5 mm3–20 mm3 ), but there were major changes for pores in another range. Tracz T. and Zdeb T. [15] studied the effects of hydration and carbonation on porosity and permeability at 90 days and up to 2 years of preservation on pastes made of ordinary Portland cement with different ratios of *w*/*c*. The results highlighted changes in capillary porosity caused by subsequent hydration advances and the carbonation phenomenon, while the depth of permeability did not change significantly during the monitored period.

Due to the pozzolanic and hydraulic properties, supplementary cementitious materials SCM may be used as a substitute for cement in the designed composition. They influence the properties of fresh concrete by their reaction with water or calcium hydroxide CH and forms additional hydrating compounds. Supplementary cementitious type materials may have pozzolanic properties by chemically reacting with calcium hydroxide to form additional hydrated calcium silicate C-S-H, which is a hydrating compound that has the greatest contribution to the development of concrete strength [6].

In this context the granulated and ground blast furnace slag GGBS in combination with ultrafine silica SF are used in the concrete industry for different applications because they have strength, durability, economic and environmental benefits.

Rao S.K. et al. [16] investigated the effect of GGBS on mechanical properties and the abrasion resistance of RCC for a level of substitution of cement with GGBS from 10% to 60%. Roller compacted concrete (RCC) is a stiff mixture of traditional concrete components that is proportioned with higher aggregate content and lower cementitious material content than compared to conventional concrete [6]. The test results showed that after the age of 28 days the increase in GGBS content led to an increase in the abrasion resistance of RCC at all replacement levels. Aghaeipour A. et al. [17] investigated the effects of GGBS as a substitute for cement in different percentages (20%, 40% and 60%) regarding water absorption, permeability and freeze–thaw resistance in RCC, with applications in road infrastructures in mind. The results not only showed a reduction in water absorption but also in mechanical resistances of concretes with GGBS compared to those without GGBS content. The depth of water penetration under pressure was lower for substitution levels up to 40%, after which it was higher than the value obtained from the control mixture. The lowest losses of resistance after 300 freeze–thaw cycles were registered in mixture with 60% substitution of GGBS, which is lower than the control mixture. Limbachiya V. et al. [18] made concrete paving blocks using GGBS and silica fume SF to replace the ordinary Portland cement OPC. High content of SiO2 in SF and of CaO in GGBS increased the durability properties of concrete paving blocks, which allowed the successful replacement of cement up to 40%. However, it was observed that the levigation properties of the blocks with GGBS and SF decreased, but within acceptable limits regarding the soil protection at levigation. In addition to improving the mechanical properties, the reduction in cracks developed in concrete during the different stages of hardening and densification is just as important for the durability of road pavements. Crack growth restriction has been studied in the works [19,20] by using hybrid fibers, such as steel-polyvinyl-polypropylenecalcium carbonate CaCO3, in cementitious compositions. The research conducted by Lam M.N.-T. et al. [21,22] investigated the feasibility of using slag aggregate resulting from electric arc furnace (FEA) to partially replace natural coarse aggregates and fly ash to partially replace cement.

The investigation methods regarding the effect of the supplementary cementitious materials are varied and depends on the envisaged properties of the manufactured materials. Moreover, research aims to establish the links between microscopic and macroscopic properties [23]. Nedunuri S.S.S.A. et al. [24] used water saturation and mercury porosimetry to study the evolution of pore structure as a function of the hardening age of Portland cement and of those that are partially replaced with SCM, such as SF, GGBS and fly ash. Majhi R.K. et al. [25] determined the rate of water absorption by capillary pores using the sorptivity test on concrete with GGBS and aggregates recycled from concrete in its composition. The microstructural characteristics extracted from scanning electron microscopy and X-ray diffraction, confirmed the results obtained by such classical methods. Renato J. et al. [26] studied the porosity of concrete using two non-destructive methods, which are X-ray microtomography and digital scanning. The porosity results obtained through the non-destructive methods were compatible with those obtained through the standard mercury intrusion porosimetry MIP test. However, the study by Diamond S. [27] drew attention to the fact that the MIP method is inadequate for measuring the pore size due to the method's inability to penetrate the C-S-H gel pores; however, it is rather suitable for assessing the apparent porosity. The analysis of the pore structure was performed in the study [28] using the fractal theory, which allows a quantitative evaluation based on the fractal size of the pores. By the fractal model proposed by Jin S. et al. [29], a regression relationship was obtained between the fractal size and the durability factor. The characteristic parameters of the pore structure of the recycled concrete with residual fibers were

investigated in the study conducted by Zhou J. et al. [30] by using the mercury intrusion porosimetry test and fractal theory. The results showed that the pore structure was mainly influenced by the water–cement ratio. The morphology of the pores developed in the first 7 days of hardening of the concrete was analyzed in the paper [31] relative to two different compositions. The first is made of simple concrete (PC) and the second is a mixture of highperformance reinforced concrete (SFRC). The internal voids of the mixtures in combination with the geometric parameters were identified by computed tomography (CT) scanning and the data were analyzed using digital image processing software (DIP) [32].

Kowalczyk R.M. et al. [33] highlighted, via nuclear magnetic resonance NMR relaxometry technique, the mechanism of water–isopropanol exchange in cement pastes. The results showed that the isopropanol revealed the presence of capillary pores better than revealed in the experiments on samples saturated with water. On the other hand, isopropanol draws water out from the gel pores but does not replace it in the same amount. One of the non-destructive methods commonly used to assess internal cracks resulting from the freeze–thaw phenomenon is the method of measuring the ultrasonic pulse propagation time (UPTT) in compliance with the Technical Report CEN/TR 15177:2006 [34]. The increase in propagation time of the wave through concrete shows a greater number of voids and internal cracks.

The great advantage of the NMR techniques is that they are completely non-invasive and, in addition, they permit the study of materials without prior preparation of samples. In the present work, NMR relaxometry [5,8,35–41] is used, again, to determine the residual water within the pores of the material and the effects of the freeze–thaw cycles on the porous system. Using this technique, information can be obtained about the porous structure of concrete containing a mixture of cement paste and aggregates, which is similar to the approach of Wawrze´nczyk J. and Kozak W. [12] and based on the Philleo concept. The NRM investigations on the effects of the freeze–thaw cycles complement the results obtained by measuring the physical characteristics (density, water absorption and content of permeable pores) of hardened road concrete determined by standard methods. The porosity results obtained from this experiment were compared to the results obtained in a previous experiment [42] on the compressive strengths and the loss of compressive strength after 300 freeze–thaw cycles. The microstructure of the concrete (in this case the pore sizes and the distance between them) can be investigated and analyzed using Scanning Electron Microscopy (SEM). The porosity induced by freeze–thaw cycles was analyzed for three mixtures, prepared with blast furnace slag and compared with two conventional mixtures used in road pavements. The investigation methods aimed to obtain more accurate conclusions for the selection of the optimal mixture of the blast furnace slag that would ensure durability for road concrete.

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

### *2.1. Materials*

The Portland cement used for preparing the samples is of the CEM I 42,5R type and it was supplied by Lafarge Holcim S.A. The performances stated by the manufacturer are in compliance with the technical specification from SR EN 197-1:2011 [43]. The specified cement surface is 4385 cm2/g and the specific weight is 3.00 g/cm3 . The granulated blast furnace slag 0/12.5 mm obtained by sudden cooling in water was supplied by the company Liberty (Galat,i, Romania). After grounding to under 69 µm, the granulated and ground blast furnace slag GGBS reached the specific surface of 5770 cm2/g and specific weight of 2.77 g/cm3 . The admissibility conditions of the GGBS blast furnace slag were in compliance with the technical specifications from SR EN 15167-1:2007 [44]. The high content of 95% vitreous mass and the 1.15 ratio between calcium oxide and silicon oxide demonstrates good activity. The value of the activity index at 28 days was 74.94%. The values obtained for the alkalis content of 1% and the 0% calcination loss show that the GGBS used does not generate expansion by its use in concrete mixtures. Table 1 demonstrates the oxide composition for cement and (GGBS).

**Table 1.** Oxide content for cement CEM I 42.5R and (GGBS) (%).

The aggregate proportions used in the preparation approach were 32% sand (sand dimension of 0/4 mm) and 68% coarse aggregate with a maximum grain dimension of 25 mm. The selection of aggregates was in compliance with the national standard NE 014:2002 [45] for road coating with role of wear and the physical properties were in compliance with SR EN 12620:2003 and SR EN 12620+A1:2008 [46]. The granulometric curve of the total aggregate mixture followed the framing within the granulometric area that is permitted by the standard NE 014-2002 [45]. The coarse aggregates were made of crushed river gravel (4/8) and crushed quarry rock (dacite from the category of igneous rocks) that were (8/16) and (16/25) mm, respectively. The use of crushed aggregates enables better adhesion between the aggregate and the cement paste due to the irregular surface, which is beneficial for the resistance to bending. The crushed strength of 14% (expressed by the Los Angeles coefficient) and the wear resistance of the aggregates of 6% (expressed by the micro-Deval coefficient) demonstrate superior performance compared to the standard SR EN 12,620 [46] result of tests performed by the manufacturer. The artificial aggregate from ungranulated blast furnace slag (ACBFS) is a crystalline byproduct resulting from the solidification in air of molten blast furnace with applications in road pavements [47]. The characteristics of natural sand and aggregates (ACBFS) were determined by the manufacturer and are presented in Table 2.

**Table 2.** Characteristics of natural sand (NA) and of aggregates from blast furnace slag (ACBFS), 0/4 mm.

The additives used here are commercially available from Badische Anilin und Soda Fabrik (BASF, Ludwigshafen, Germany). As a superplasticizer, we used MasterGlenium SKY 527 and air training additive Master Air 9060 (BASF, Germany) and both possess characteristics in compliance with SR EN 934-2+A1:2012 [48]. The water in the concrete mixture was taken from the water supply system of the city Cluj-Napoca in compliance with SR EN 1008:2003 [49].

#### *2.2. Sample Preparation*

The parameters of designing the road concrete complied with the minimum requirements for the severe exposure class (XF4) in accordance with SR EN 206-1:2002 and SR EN 206+A1:2017 [50]. The classification in conditions of exposure according to the XF4 class requires a minimum cement dosage of 350 kg/m3 , maximum *w*/*c* ratio of 0.45 and a minimum strength class of C35/45 with entrained air [50,51]. A number of five mixtures were prepared with different quantities of materials presented in Table 3. The first two control mixtures were prepared using Portland cement type CEM I 42,5R utilizing natural aggregates and the following three mixtures were prepared using granulated and crushed blast furnace slag, under 63 µm, and aggregates from crushed blast furnace slag at a dimension of 0/4 mm in different proportions. The below observations were used:

**S 360**, control mixture of 360 kg/m3 Portland cement dosage and natural aggregates;

**S 414**, control mixture of 414 kg/m3 Portland cement dosage and natural aggregates;

**S 54/20**, 360 kg/m3 (cement) + 54 kg/m3 (GGBS) and 20% (ACBFS)\_0/4 mm + 80% (NA)\_0/4;

**S 54/40**, 360 kg/m3 (cement) + 54 kg/m3 (GGBS) and 40% (ACBFS)\_0/4 mm + 60% (NA)\_0/4;

**S 54/60**, 360 kg/m3 (cement) + 54 kg/m3 (GGBS) and 60% (ACBFS)\_0/4 mm + 40% (NA)\_0/4.

**Table 3.** Material quantities in mixtures.

Compared to the control mixture S 360, the blast furnace slag (GGBS) was brought as a contribution to the cement mass in a percentage of 15% and, compared to the control mixture S 414, the GGBS substituted the cement in a percentage of 13%. The ground granulated blast furnace slag was used in road concrete mixtures as a binder with supplementary constituents with cement characteristics similar to those described in the study [52]. Natural sand NA was substituted in percentages of 20%, 40% and 60% with crushed blast furnace slag aggregates ACBFS with dimensions of 0/4 mm. In this experiment the water content and additive content was adjusted to maintain the consistency of the concrete at the most appropriate values of 20 mm.

#### *2.3. Methods*

#### 2.3.1. Sampling, Preservation and Preparation of Samples for Testing

For each mixture, 8 cubic specimens with a side of 150 mm and 3 specimens with a side of 71 ± 1.5 mm, as shown in Table 4, were prepared and kept in the air for 24 h after which they were stripped and immersed in water at a temperature of 20 ± 2 ◦C.

At the age of 7 days, all specimens were removed from water and stored in the climatic chamber at a temperature of 20 ± 2 ◦C and at a humidity of 65% ± 5% up to the age of 50 days. Next, the cubic specimen with the side of 71 mm was kept in water until the age of 100 days and the abrasion resistance was then determined; the values of are found in paper [53]. From the age of 100 days to the age of 150 days they were kept in the air in the climate chamber. The density, water absorption and content of permeable pores were determined from the average of the results for the 3 cubic specimens for each type of mixture with the side of 71 mm and at the age of 150 days in compliance with the standard ASTM C642-2006 [54]. The nuclear magnetic resonance measurements were performed on concrete cores extracted from two intact 150 mm cubes for each mixture; one tested for 300 freeze–thaw cycles and one control specimen maintained in water until the age of 150 days. For the NMR measurements, the cubes with apparent density close to the average of the 3 cubes tested for the freeze–thaw in the study were selected [42] by the method of the loss of compressive strength according to SR 3518:2009 [55]. The thermostatic chamber was set to maintain a temperature of (−17 ± 2) ◦C for 4 h for the freeze cycle, up to (20 ± 2) ◦C for the thaw cycle for 4 h and the humidity of RH 95%. Four days before the commencement of the test, the specimens were placed in a water tank at a temperature of (20 ± 5) ◦C for saturation and the control specimens were kept in water. Those intended for the freeze–thaw cycles were introduced in the thermostatic chamber, such as in Figure 1.

**Table 4.** Number of samples prepared for testing in each composition.

**Figure 1.** Control specimens kept in water and specimens underwent the freeze–thaw test (150 mm × 150 mm × 150 mm).

After 300 repeated freeze–thaw cycles, the specimens were tested for compression and the compressive strength losses were calculated, for the specimens tested with freeze–thaw cycles compared to the control specimens, with Equation (1) as follows:

$$\eta\_{\text{fl}} = \frac{f\_{cmwater} - f\_{cmfreez}}{f\_{cmwater}} \times 100 \text{ (\%)}\tag{1}$$

where:

*ηn*—the compressive strength loss after "*n*" freeze–thaw cycles; *f cm water*—the average compressive strength of the samples maintained in water during "*n*" freeze–thaw cycles;

*f cm freeze*—the average compressive resistance of the samples maintained in the thermostat chamber during "*n*" freeze–thaw cycles.

Compressive strength at 150 days was determined by applying a uniform and continuous force increasing in increments of 0.5 MPa/s on a loading machine type Advantest 9

of 300 tf. in accordance with SR EN 12390-3:2002 [56] and is calculated as by Equation (2) as follows:

$$f\_{\mathcal{L}} = \frac{F}{A\_{\mathcal{L}}} \text{ (MPa)}\tag{2}$$

where:

*fc*—compressive strength in N/mm2 ,

*F*—the maximum load in N;

*Ac*—cross-sectional area of the section of rupture in mm2 .

#### 2.3.2. NMR Relaxometry Approach for Determining the Pore Size Distribution

The NMR relaxometry techniques are completely non-invasive and allow the investigation of cement samples without special prior preparation. These have been used successfully in the recent years to extract information about the porous structure of cementbased materials and to extract information with respect to the evolution of water inside the pores during the hydration process [5,8,35–40]. In the NMR relaxometry of cement-based materials, both the longitudinal relaxation time and the transverse relaxation time can be monitored [5,8,35–40]. Due to the fact that longitudinal relaxation measurements are slower and are more difficult to apply to the systems that evolve during the measurement process, it is preferable for many applications to measure the transverse relaxation time. However, in the case of transverse relaxation measurements, it is important to take into account the effects that diffusion in internal gradients could have on the accuracy of the measurement and so it is preferable that they be performed in low magnetic fields and with pulse sequences that could minimize internal gradients effects [5,8,40].

A technique often used in NMR applications on cement-based materials is known as the Carr-Purcell-Meiboom-Gill (CPMG) technique [57]. This technique allows the rapid and robust measurement of the transverse relaxation rate (1/T2) of the nuclear spins and if the time interval between radio frequency pulses is short, the influence of internal gradients can be neglected. Under these conditions, there is a relationship [5,8,37–39] between the transverse relaxation rate and the surface-volume (*S*/*V*) ratio of the pores that is described as follows.

$$\frac{1}{T\_2} = \varepsilon \frac{S}{V} \tag{3}$$

In the above equation, ε represents the surface relaxivity, which is a constant that depends on the interaction of molecules with the surface, the intensity of the magnetic field in which the NMR experiment is performed and the content of magnetic impurities of the pore surface. Note that in Equation (3) we neglected the contribution of bulk relaxation to the phenomenon of nuclear relaxation because the bulk relaxation time is much longer than the one induced by the surface. Based on Equation (2) we can establish a direct proportionality between the relaxation time and the pore size. This proportionality allows us to discover the distribution of pore dimensions if we know the distribution of relaxation times and the relaxivity of the surface. Even if the relaxivity of the surface is not known, information can be obtained about the relative distribution of pore sizes. This distribution can be extracted from the CPMG echo series if a numeric Laplacian transformation is applied to this series [58,59].

In the present work, in order to highlight the effect of the freeze–thaw cycles on the relative distribution of pore sizes, the specimens were investigated by NMR relaxometry at the age of 150 days and after the completion of the 300 freeze–thaw cycles. The test period in the study was extended from the requirements of national standard NE014-2002 [45] from 100 to 300 freeze–thaw cycles and is similar to the number of freeze–thaw cycles provided by the standard ASTM C666/C666M-03 [60]. A laboratory exposure of concrete samples to an extended number of repeated freeze–thaw cycles can simulate, as closely as possible, the damage caused by the freeze–thaw phenomenon in concrete exposed to real weather, as experienced in the literature [61], where the samples were tested in a laboratory for up to 1000 freeze–thaw cycles. From the core of intact cubes kept in the

thermostatic chamber and the control cubes kept in water during this period, cylindrical samples with a length of 20 mm and a diameter of 9.5 mm were extracted, such as in Figure 2. To eliminate interpretation errors, the mass of the samples extracted from the cubes was the same for each mixture of 1.06 g. It was possible to take samples for testing at these small sizes because the proportion of maximum size aggregates has a relatively small volume of the total volume of aggregates. The samples were dried in an oven at 105 ± 5 ◦C for 24 h to eliminate water from the pores. They were then inserted into 10 mm NMR tubes which were sealed to prevent molecular exchange with water from atmospheric air. After measuring the dry samples, they were dried again and immersed in water for 48 h at 22 ± 2 ◦C. After this interval, the samples were lightly buffered with filter paper, then inserted into the NMR tubes and sealed. The third set of measurements was performed on samples saturated with cyclohexane and follows the same procedure as in the case of samples saturated with water. The saturation with cyclohexane was performed to better highlight the inter-C-S-H pores and capillaries, as is demonstrated in the other study [8]. Note that the intra-C-S-H pores are highlighted by the presence of water molecules that cannot be removed without destroying the material [8].

**Figure 2.** Samples ready for the NMR measurements.

The NMR measurements were performed with a low field instrument MinispecMQ20 (Bruker, Karlsruhe Germany) using the CPMG technique. Before each measurement, the samples were permitted to reach the thermal equilibrium at a temperature of 35 ◦C. A number of 2000 spin echoes were recorded in each experiment and the time between two echoes was maintained at 0.1 ms to reduce the effects of internal gradients on the measurements. Relaxation time distributions were extracted from the CPMG echo series using a Laplace numerical inversion [59].

#### 2.3.3. Scanning Electron Microscopy (SEM)

Small pieces of samples were taken from the interior part of the crushed samples and analyzed at room temperature using a VEGA3 SBU electronic microscope with an Energy Dispersive Spectrometer Quantax EDS from Bruker in order to obtain information about the pore's sizes and the distance between the pores. Fragments from the 150 mm side cubes remaining after the NMR measurements were used from the control sample set, which was kept in water during the freeze–thaw test.

2.3.4. Density after Immersion and Boiling, Water Absorption and Proportion of Permeable Pores

The density after immersion and boiling, water absorption and permeable pore content were obtained by a method in accordance with ASTM C 642-2006 [54] that was used in another study [62]. The water-saturated samples with the dry saturated surface were weighted in air (b) and then dried in an oven at 100–110 ◦C until a constant mass (*a*) was reached. The samples were then placed in a suitable container, covered with tap water and boiled for 5 h. They were then allowed to cool to room temperature (22 ± 2 ◦C). The moisture from the surface was removed with a towel and the dry mass of the surface-dry sample suspended in air after immersion and boiling was determined (*c*), with the mass

measured in air. The apparent mass (*d*), after immersion and boiling, was measured in water with the hydrostatic balance, such as in Figure 3. The absorption after immersion (*m*1) and absorption after immersion and boiling (*m*2) were calculated with the following equations.

$$\left[m\_1 \right. = \left[\left(b - a\right)/a\right] \times 100\tag{4}$$

$$m\_2 = \left[\left(c - a\right)/a\right] \times 100\tag{5}$$

**Figure 3.** Hydrostatic weighing of cubes with 71 mm side.

Dry mass density, *ρ*1, and density after saturation and boiling, *ρ*2, were calculated with the following equations:

$$
\rho\_1 = \left[ a/(c-d) \right] \times \rho\_w \tag{6}
$$

and

$$
\rho\_2 = \left[ a/(a-d) \right] \times \rho\_w \tag{7}
$$

For water density, *ρw*, the value of 0.998 g/cm3 was used. The permeable pore proportion was calculated with the equation:

$$P\_0 = \left[ (\rho\_2 - \rho\_1) / \rho\_1 \right] \times 100,\text{ specified in } \% \tag{8}$$

which is in accordance with ASTM C 642-2006 [54] standard.

#### **3. Results and Discussion**

#### *3.1. Relative Distribution of Pore Sizes*

The distribution of transverse relaxation times for the gel pores obtained from the Laplace numerical inversion, for each composition described in Table 3, is shown in Figures 4 and 5. Three peaks can be distinguished in all cases and the position of the peaks can be associated with the three types of pores. The values recorded on the horizontal axis for the transverse relaxation time T2 are proportional to the pore size (see Equation (3)) and the values recorded on the vertical axis show the probability of having a certain pore dimension. The measuring unit on the y-axis is arbitrary (a.u. = arbitrary units). For a direct comparison of the curves, the same scale was used at each measurement performed.

Figure 4 depicts the samples dried according to the procedure described above, both control samples (Figure 4a) and those subjected to the freeze–thaw process (Figure 4b). A small amount of water (the first peak) was observed in the gel-like pores (intra C-S-H) with dimensions up to 2 nm [8], which could not be removed during drying process. In the mixture S 360, the distribution of transverse relaxation times corresponding to the gel pores (intra C-S-H) for the control samples, preserved in water (Figure 4a), shows a maximum placed in the range 0.02–0.3 ms and indicated a maximum probability density of 0.0045 for a value of T2 of 0.08 ms. At T2, the values slightly shifted to the right and

the maximum probability density decreased to 0.003 for the mixtures S 54/20 and S 54/40 followed by the mixtures S 414 and S 54/60, with an intensity inferior to 0.002. However, for the samples subjected to 300 freeze–thaw cycles (Figure 5b), the transverse relaxation time distribution interval shifted to the right (for S 414, S 54/20, S 54/40) at the higher pore size values between 0.02–0.6 ms and the most accentuated values were recorded in the mixtures S 360 and S 54/60 between 0.06–0.6. For all mixtures stored in the refrigerator (f-t), a maximum probability density close to the value recorded for the control mixtures, 0.003, was registered. The results of the experiment show that, after the freeze–thaw test, the pore size was higher compared to the control mixtures, which indicated the appearance of microcracks even in the intra C-S-H pores.

**Figure 4.** Distribution of the transverse relaxation time T2 after drying: (**a**) control samples kept in water (w); (**b**) samples subjected to 300 freeze–thaw cycles (f-t).

**Figure 5.** Distribution of the transverse relaxation time T2 after water saturation: (**a**) control samples kept in water (w); (**b**) samples subjected to 300 freeze–thaw cycles (f-t).

By introducing the samples into the water, the clear increase in the area of the peaks corresponding to the gel pores (intra C-S-H and inter C-S-H) can be observed in Figure 5.

The pore distribution continues with several lower intensity peaks, which correspond to capillary pores, cracks or are artefacts of the numerical inverse Laplace. In the control samples (Figure 5a) the distribution of the transverse relaxation times was within the range 0.03–1.10 ms and for the mixtures tested for freeze–thaw (Figure 5b) the peak shift to the right was recorded at higher values, in the range 0.1–1.20 ms, as can be observed in Figure 6, which shows the starting and ending values for T2 in each mixture corresponding to the gel pores as well as the position of the maximum (a.u.) (broken line).

The relative distribution of pore sizes in mixture S 54/20 and S 54/40 was kept close to the distribution in mixture S 360 and lower than the distribution in the control mixture S414 in both control samples and the samples tested for freeze–thaw. However, for the 54/60 mixture, the distribution range shifted to higher values compared to the two reference mixtures in both the control specimens and in the specimens that underwent the freeze–thaw text.

**Figure 6.** The distribution range of the gel pores for: (**a**) the control samples; (**b**) for samples subjected to freeze–thaw cycles. Both measurements are performed on water-saturated samples.

Figure 7a shows the maximum probability density recorded in both control samples kept in water (w) and in samples tested in water-saturated freeze–thaw (f-t). For the watersaturated control samples (Figure 7a), the maximum probability density with the highest value was registered in the mixture S 414 (of 0.023) and the lowest value (of 0.012) was obtained in the mixture S 360. In all blast furnace slag mixtures, the maximum probability density value corresponding to the gel pores is higher than the reference mixture S 360, but lower than in the control mixture S 414. For all water-saturated samples tested at 300 freeze– thaw cycles (Figure 7a), it was observed that the maximum value of the probability density corresponding to the peak of the gel pores increases with the increase in the percentage of substitution with blast furnace slag aggregate and it is above the level in the control mixtures S 360 and S 414. The largest difference in the maximum value of the probability density compared to the control sample was recorded in the mixture S 54/60 maintained at freeze–thaw.

For the better monitoring of the capillary pores, it is preferable that they are saturated with cyclohexane because a higher degree of filling can be obtained in the case of cyclohexane than in the case of water [8]. This is due to the significantly lower contact angles in the case of cyclohexane than water relative to the mineral surfaces inside the cement paste [8,63]. Figure 8 shows the distribution of relaxation times and Figure 7b shows the maximum probability density corresponding to the peak of the capillary pores in the samples saturated with cyclohexane.

It is observed in Figure 7b that the evolution of the maximum probability density of the capillary pores has a tendency similar to that of the gel pores from the water-filled samples (Figure 7a); in all compositions with blast furnace slag, the intensity of the capillary pores increases with the increasing level of substitution of slag aggregate, respectively. In the control samples, the shortest range of distribution of capillary pores was recorded in the mixture S 360 situated between 1.0–75 ms and the largest range was recorded in the mixture S 54/60 between 2.2–150 ms, according to Figure 9a.

**Figure 8.** Transverse relaxation time distribution T2 after cyclohexane saturation: (**a**) control samples kept in water (w); (**b**) samples subject to 300 freeze–thaw cycles (f-t).

**Figure 9.** The capillary pores distribution range: (**a**) for the control samples (w); (**b**) for the samples tested in freeze–thaw (f-t). Both measurements were made on cyclohexane-saturated samples.

In samples subjected to 300 freeze–thaw cycles, the transverse relaxation time distribution T2 is slightly shifted to the right in mixture S 360 between 1.0–90 ms and more accentuated in mixture S 54/60 between 7.0–200 ms, as can be seen in Figure 9b. In the S 54/60 mixture, the peak corresponding to the capillary pores does not tend towards the minimum intensity in the position of T2 (200 ms), which indicated the presence of the larger capillary pores or internal cracks in concrete.

A longer relaxation time (Figure 9b) and a higher intensity (Figure 9b) in the S 54/60 mixture show an increase in the size and the proportion of capillary pores in the samples tested at freeze–thaw (f-t) compared to the control specimens kept in water (w).

This change in the dimension of capillary pores from the samples subjected to freeze– thaw (f-t) is explained by the fact that the phenomenon of frost water enters the structure of concrete capillary pores, freezes at about −0.5 ◦C and the resulting volume of the ice formed is approximately 9% higher compared to the water initially absorbed [1–3]. Thus, repeated freeze–thaw cycles lead to an increase in the size of the capillary pores and the appearance of cracks.

#### *3.2. Pores Size*

SEM images obtained for the investigated samples are presented in Figure 10. In order to obtain a better view of the pores, all samples were measured using the same parameters but at different magnifications. The pores radius and the distance between pores were measured automatically by SEM software after selecting the shape most appropriate (circle or ellipse) to the shape of the pores. It can be observed that the radius of the pores ranges

from ~33.84 µm for S 360 to ~2.95 µm for S 54/60. The distance between pores ranges from ~915.06 µm for S360 to ~63.64 µm for S 54/40. For S 54/60, no pores were found in the close vicinity of the selected pore to measure the distance. Pore radii with dimensions over ~2.95 µm indicate their framing at capillary pores in accordance with Reference [10]. The results obtained for the S 360 mixture suggest a lower pore density because the largest pore distance was identified. These results can be correlated with those obtained from the technique (NMR), which identified the lowest density of capillary pores on the control sample S 360 (Figure 7a). For the mixture S 54/20 and S 54/40 the distance between the pores was shorter compared to the mixture S 360, which suggests a higher pore density. It is observed that the measurements (NMR) show the same increase in the probability density in the compositions S 54/20 and S 54/40 compared to the mixture S 360 (Figure 7a).

**Figure 10.** SEM images of S360 (**a**), S414 (**b**), S54/20 (**c**), S54/40 (**d**) and S54/60 (**e**).

#### *3.3. Density, Water Absorption and Permeable Pores Content of Hardened Concrete*

Following the determinations performed in compliance with [54] on cubes with side 71 mm, the proportion of permeable pores (P0), dry density (ρ1) and density after saturation and boiling (ρ2) are illustrated in Figure 11a, respectively. The reported results represent the average of three samples tested in each mixture and the standard deviation (SD) and the coefficient of variation (CoV) are shown in Table 5.

**Figure 11.** (**a**) Dry density (ρ1 ), density after immersion and boiling (ρ2 ) and content of permeable pores (P0 ) at the age of 150 days; (**b**) the relationship between the loss of compressive strength after 300 freeze–thaw cycles (η300) and the permeable pore content (P0 ) of the control samples.

**Table 5.** Standard deviation (SD) and coefficient of variation (CoV) calculated for density and permeable pore content.

Evaluation of compressive strengths and loss of compressive strengths (η300) after 300 freeze–thaw cycles determined on cubes with a 150 mm side, according to SR 3518:2009 [55], are found in the other studies [42] and included in Table 6. The compressive strengths for the control samples and for the samples tested at 300 freeze–thaw cycles recorded a standard deviation (SD) in the range (0.15–5.53) MPa and the coefficient of variation (CoV) between (0.2–9.1) %. It is observed that the scattering of results is below the accepted limit of 15%, while having a reasonable quality in the range of 5% to 10% as suggested in the study [64].

Analyzing the results from Table 6, it is observed that the compressive strengths determined on the control specimens have a similar evolution with the results of the densities presented in Figure 11a. Furthermore, the value of compressive strength losses after 300 freeze–thaw cycles tends to evolve similarly to that of the permeable pore content and develops a second order polynomial relation having a very good correlation coefficient (R value), shown in the diagram in Figure 11b.

**Table 6.** The average compressive strengths, resistance losses after 300 freeze–thaw cycles [42], standard deviation and coefficient of variation.

In Figure 12, it is observed that the compressive strengths determined on the control specimens and on specimens maintained at 300 freeze–thaw cycles maintain their evolution trend with the content of permeable pores (Figure 12a) with the relative size of the capillary pores evaluated by NMR technique (Figure 12b,c). With the reduction in the content of permeable pores (P0) and the transverse relaxation time (T2), for which the probability density (au) recorded the maximum value (broken line in Figure 9a,b), the result is the increase in mechanical strengths. The polynomial relation of order two, derived by regression, developed between the compressive strength and the content of permeable pores and the relative size of the capillary pores determined on the control samples, possesses a value of the correlation coefficient very close to 0.8779 and 0.8567 (Figure 12a,b), which confirms the concordance results obtained by standard methods and NMR technique. Moreover, the confirmation of the results obtained by the NMR technique can be appreciated from the correlation coefficient 0.9931 (having the standard error of only 6.9 × 10−3 ) obtained from the polynomial relationship developed between the compressive strength and the transverse-room relaxation time (T2) for the maximum (a.u.) capillary pores and are measured on samples maintained at 300 freeze–thaw cycles (Figure 12c).

**Figure 12.** Relationship between compressive strength (*fcm*) and porosity: (**a**) relative to the permeable pore content (P0 ) in the control samples; (**b**) relative to the relative pore size for the position in which the transverse relaxation time (T2 ) records the maximum density in the control samples; (**c**) relative to (T2 ) for the position in which it records the maximum density at the samples kept at freeze–thaw (f-t).

The permeable pore content obtained for the mixture S 54/20 is lower compared to the two control mixtures S 360 and S 414. The mixture S 54/40 is above the level of the control mixture S 360 and below the level of S 414. However, at mixture S 54/60, in which the substitution with artificial aggregates was 60% the porosity of the concrete increased the most, it was above the level of the two control mixtures. For blast furnace slag mixtures, the

mechanical strengths decreased as the level of substitution with slag aggregates increased and the lowest values were registered for the S 54/60 mixture. The cause that led to the decrease in the mechanical strengths was the increase in capillary porosity, which in turn was influenced by the amount of water used and by the increase in water/binder ratio; this is a fact shown in the specialty literature [6]. The increase in the specific surface of the ground furnace blast slag and the increase in the porosity of the slag aggregate led to the increase in the water requirement in the preparation of the mixtures S 54/40 and S 54/60 in order to obtain a similar consistency to the reference mixtures; the results are in compliance with the specialty literature [65]. The crushed slag aggregates have an angular shape, which reduces the workability of the concrete and have a higher porosity than natural sand [47]. Similarly, the grinding fineness of the blast furnace slag was higher than that of cement; the specific surface in furnace blast slag was 5770 cm2/g and in cement it was 4385 cm2/g. However, for the mixture S 54/20, the proportions of blast furnace slag used led to the increase in the mechanical resistances and to the decrease in the porosity, obtaining better performances than the control mixture S 414 at the same water/binder ratio. For this mixture, the use of slag as a binder (GGBS) and aggregates (ACBFS) in low dosages was advantageous.

#### **4. Conclusions**

The connectivity of gel (intra C-S-H and inter C-S-H) and capillary pore network was investigated for the first time on road concrete samples tested for repeated freeze– thaw cycles through low field NMR relaxometry. The experiments rely on monitoring the distribution of water and cyclohexane molecules saturating the pores. The relative distribution of gel and capillary pores, determined through NMR, was conducted on samples taken from control specimens kept in water during the freeze–thaw test period and on samples extracted from specimens tested at 300 freeze–thaw cycles. The content of permeable pores was obtained on cubes with a side of 71 mm, made of the same composition and tested at the same age. For the three blast furnace slag compositions, the same amount of (GGBS), 54 kg/m3 , but different percentages (20%, 40%, 60%) of crushed aggregates (ACFBS) were used. Different dosage of Portland cement, 360 and 414 kg/m3 , were used for the two compositions made with conventional materials. The conclusions of the experiments are summarized in the following:

**Author Contributions:** Conceptualization, writing original draft and investigation, L.M.N.; Writing original draft, project administration and scientific supervision, O.C.; Methodology, investigation, data curation and validation, I.A.; Methodology, investigation, data curation and validation (SEM), D.S.; Data curation, validation and writing—reviewing and editing, A.V.S.; Resources and investigation, M.I. All authors have read and agreed to the published version of the manuscript.

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

**Institutional Review Board Statement:** Not applicable.

**Informed Consent Statement:** Not applicable.

**Data Availability Statement:** The data presented in this study are available on request from the corresponding author.

**Acknowledgments:** I.A. acknowledges the financial support of the Romanian CNCS—UEFISCDI, project number PN-III-P4-ID-PCE-2020-0533.

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

### **References**

## *Article* **Selected Properties of Densified Hornbeam and Paulownia Wood Plasticised in Ammonia Solution**

**Przemysław Mania \* , Karol Hartlieb, Grzegorz Mruk and Edward Roszyk**

Department of Wood Science and Thermal Techniques, Faculty of Forestry and Wood Technology, Pozna ´n University of Life Sciences, Wojska Polskiego 38/42, 60-627 Pozna ´n, Poland<EMAIL_ADDRESS>(K.H<EMAIL_ADDRESS>(G.M<EMAIL_ADDRESS>(E.R.) **\*** Correspondence<EMAIL_ADDRESS>Tel.: +48-618-487-448

**Abstract:** The aim of the study was to densify samples of Paulownia Clone wood in vitro 112 and hornbeam (*Carpinus betulus* L.) by compression in the radial direction. Before the specimens were densified, they were subjected to plastic treatment in an ammonia solution. After densification, the compressive strength in the radial direction and the determination of the Brinell hardness in all three anatomical directions of the wood were determined. The wood swelling in humid air (98% RH) and liquid water was also determined. Paulownia wood density increased by about 280% and hornbeam wood density by 40%. The Brinell hardness parallel to the fibres increased by 49 and 390%, perpendicular by 80 and 388% for hornbeam and Paulownia, respectively. A significant increase in the compressive strength of wood in the radial direction was also observed. Densified hornbeam wood exposed to water showed a high swelling value of 153, while Paulownia wood exhibited 107%.

**Keywords:** densified wood; ammonia; mechanical properties; Brinell hardness; Paulownia; hornbeam

#### **1. Introduction**

It is known that the density of wood is related to its mechanical properties, i.e., the strength of the wood increases proportionally with the increase in its density. Densification, i.e., the thickening of the wood structure, improves the mechanical properties of lowdensity species and replaces hard species. Wood species with high hardness can also be densified, and their properties are even more improved [1]. The first concept of wood compaction appeared in 1900, but it consisted only of wood compression and did not consider plasticising treatment. Research on increasing wood density has been carried out for many years. It is said that this effect is achieved by compressing the wood in the transverse direction, which improves its mechanical properties [2–6]. In the 20th Century, patents for densification through wood compression began to appear [1,7]. These ideas involved mechanically compressing the wood tissue after subjecting it to steam under high pressure. However, this method did not capture the "memory effect" of wood, i.e., the post-deformation effect [8,9], which led to the "desire" for the wood to return to its original dimensions. According to Jakes et al. [8], hemicelluloses are responsible for dimension stabilisation of lignocellulosic materials. Lignin is the substance responsible for the shape memory effect. Delignification, leading to a reduction in the content of the components that make up the matrix encrusting the cellulose skeleton (hemicellulose and lignin), causes a significant decrease in post-deformation recovery because the hemicellulose in the wood is responsible for shape stabilisation. Lignin plays a substantial role in the quasielastic recovery mechanism [6,10–13]. Wood subjected to densification treatment has better physical and mechanical properties, but the increase in density does not always go hand in hand with improving these properties. This is permanent damage to the cell wall during such significant deformation of the wood [3,14].

The thermo-hygro-mechanical (THM) wood thickening process depends on the type of wood, both in terms of the degree of densification and the properties of the modified

**Citation:** Mania, P.; Hartlieb, K.; Mruk, G.; Roszyk, E. Selected Properties of Densified Hornbeam and Paulownia Wood Plasticised in Ammonia Solution. *Materials* **2022**, *15*, 4984. https://doi.org/10.3390/ ma15144984

Academic Editors: Andrei Victor Sandu and Joel Pawlak

Received: 11 May 2022 Accepted: 15 July 2022 Published: 18 July 2022

**Publisher's Note:** MDPI stays neutral with regard to jurisdictional claims in published maps and institutional affiliations.

**Copyright:** © 2022 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 (https:// creativecommons.org/licenses/by/ 4.0/).

wood. This is due to the structural diversity and composition of the chemical cell walls of individual types of wood. Species with a lower density can be densified more easily than species with a higher value [15,16]. This is related to porosity and susceptibility to deformation. For this reason, when the timber is compressed in the radial direction, the earlywood zones first deform, and then the latewood zones. In this case, deformations of the early and latewood zones are inversely proportional to the value of their modulus of elasticity. Usually, the compaction of the wood tissue contributes to the improvement of the mechanical properties of the wood, even more than the increase in the density of the wood. Sometimes, however, the increase in the value of the mechanical parameters is smaller than the increase in wood density caused by its compaction.

One method that ensures plasticisation and fixation in one step is treatment with ammonia. Ammonia penetrates deeper into wood than water. This causes significant plasticisation of the wood, and the permanent displacement of particles becomes possible as the cellulose changes its morphology as new bonds are formed. The effect of ammonia is to swell the wood tissue, dissolve the low-molecular substances it contains, and split higher polymerised hemicelluloses, which is manifested by increasing the solubility of wood in water. Part of the ammonia is permanently bound to the lignin. Along with the destruction of hemicelluloses, some changes occur in the cellulose itself, and some of the bonds between lignin and carbohydrates are broken. Ammonia penetrates not only amorphous areas, but also, it can be assumed, partially in ordered areas. Due to the washing out of soluble substances and the deep penetration of ammonia and water particles into the cell membrane structure, an additional capillary system appears, and the inner surface of the wood is increased. Microfibrils gain greater freedom of movement from each other, manifested externally in the advanced plasticisation of the wood tissue [17–21].

*Paulownia Clone* in vitro, commonly known as Oxytree, is a new species that has appeared in Poland for several years. Apart from the tree's physical and mechanical properties and the environmental impact of the tree, the species has not been thoroughly researched. Furthermore, not many studies related to the compression of this wood structure have been carried out. The aim of the study was to densify wood after its plasticizing in ammonia solution and to determine its selected properties. It was decided to determine the hardness of the wood in all anatomical directions and the compressive strength in the direction of densification. The dimensional stability of the material in water and humid air was also determined. Therefore, it was decided to compare the densification process after the wood was plasticised with ammonia on two completely different types of wood: the heaviest type of wood naturally occurring in Poland and one of the lighter ones, which is popular in Poland. For this purpose, hornbeam and Paulownia wood were used.

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

The test samples were obtained from 10 Paulownia wood planks, 24 mm thick, purchased from a local seller, originating from Shantung Province (China). The boards were machined and cut into samples with lengths: of 22 mm in the tangential and longitudinal directions and 106 mm in the radial direction. The panels were selected so that the growth rings were as parallel as possible to the longer edge of the specimen. The moisture content (MC) of the samples obtained was about 12%. Hornbeam wood was the second experimental material, as hornbeam is the heaviest species that grows naturally in Poland. The material for the tests was taken from 60 mm-thick hornbeam logs. The log was cut into one-meter sections and mechanically processed to obtain samples of the exact dimensions as above. The drawing showing the model cut out with the dimensions marked is shown below (Figure 1). The densification process was applied to 70 samples of hornbeam and Paulownia wood. The exact number of samples was used as the reference material. Thirty samples were then used for the hardness determination, 30 for the compressive strength of the wood, and 10 for thickness swelling.

of the wood, and 10 for thickness swelling.

**Figure 1.** Sample shape and its dimensions. **Figure 1.** Sample shape and its dimensions.

The density of the samples was determined before densification and after the modifications made to measure the degree of compaction as a percentage. Density was determined using the ISO 13061-2 standard [22]. Rectangular samples were measured in three directions with a calliper with an accuracy of 0.01 mm and then weighed on a laboratory The density of the samples was determined before densification and after the modifications made to measure the degree of compaction as a percentage. Density was determined using the ISO 13061-2 standard [22]. Rectangular samples were measured in three directions with a calliper with an accuracy of 0.01 mm and then weighed on a laboratory balance with a measurement accuracy of 0.001 g. Density was calculated as the ratio of its mass and volume.

samples were then used for the hardness determination, 30 for the compressive strength

balance with a measurement accuracy of 0.001 g. Density was calculated as the ratio of its mass and volume. The samples were plasticised in a 25% ammonia solution. It consisted of heating the pieces in a boiling aqueous 25% ammonia solution for seven hours. The samples were placed in 6 L round-bottomed flasks. The flasks containing 2.5 L of the ammonia solution The samples were plasticised in a 25% ammonia solution. It consisted of heating the pieces in a boiling aqueous 25% ammonia solution for seven hours. The samples were placed in 6 L round-bottomed flasks. The flasks containing 2.5 L of the ammonia solution and the pieces were connected to a reflux condenser and placed in a heating mantle to bring the solution to a boil. The boiling point of the solution was 110 ◦C. After 7 h, the samples were taken out of the solution and inserted into the clamps in a hot-plasticized state.

and the pieces were connected to a reflux condenser and placed in a heating mantle to bring the solution to a boil. The boiling point of the solution was 110 °C. After 7 h, the samples were taken out of the solution and inserted into the clamps in a hot-plasticized state. A particular device was designed to densify the wood. The device was designed in such a way as to limit the possibility of the sample buckling on either side (Figure 2). Therefore, the samples after densification had a rectangular shape. Compression took place in the radial direction of the wood on a Zwick ZO50TH (Zwick/Roell, Ulm, Ger-A particular device was designed to densify the wood. The device was designed in such a way as to limit the possibility of the sample buckling on either side (Figure 2). Therefore, the samples after densification had a rectangular shape. Compression took place in the radial direction of the wood on a Zwick ZO50TH (Zwick/Roell, Ulm, Germany) strength testing machine. The maximum deformation that the samples could withstand without cracking and failure was 85 mm for paulownia and 55 mm for hornbeam wood samples. This compression of the samples was achieved by applying a stress of 2.5 N/mm2 . This is the value beyond which the stress rapidly increased, as proven by preliminary tests. After pressing the specimens, the clamps were clamped by tightening the screws to prevent elastic recovery. The samples clamped in this way were left in the laboratory to dry. *Materials* **2022**, *15*, x FOR PEER REVIEW 4 of 9

**Figure 2.** Drawing of the device for densification of wood samples. **Figure 2.** Drawing of the device for densification of wood samples.

After six weeks of conditioning, the densified wood samples were subjected to compression and a Brinell hardness test. The hardness of the tested materials was determined in three anatomical directions. A steel ball with a diameter of 10 mm was pressed into the

samples and for paulownia wood, 500 N and 150 N, respectively. The maximum and minimum diameters of the residual indentation were measured with a Brinell magnifier to

> 2 (−√2−2)

where *F* is the force acting on the ball (N), *D* is the diameter of the pressed ball (mm), and

Compressive strength was determined in the direction of densification, so in the radial direction. The stress at the proportionality limit (so-called compressive strength perpendicular to the grain or relative strength (RcR in the radial direction)) was determined. Compression tests were conducted using a numerically controlled test machine, Zwick

The maximum wood swelling degree was determined on five densified hornbeam and Paulownia wood samples. The maximum swelling was determined only in the radial direction. The oven-dried samples were measured in the radial direction with a calliper with an accuracy of 0.01 mm; then, the samples were placed in a desiccator above the water surface for 14 days (RH = 98%). After this time, the samples were again measured in this direction with an electronic calliper. Similar measurements were made on five consecutive samples, with the difference that the samples were placed in a beaker with distilled water for 14 days. Thanks to this, it is possible to evaluate the behaviour of such wood in humid air and water. The maximum swelling degree was calculated as the ratio

The experimental data were analysed using the STATISTICA 13.3 software (TIBCO Software Inc., Palo Alto, CA, USA) with the analysis of variance (ANOVA). Significant differences between mean values of the parameters describing paulownia and hornbeam wood properties were determined using Tukey's HSD test. The comparison tests were performed at a 0.05 significance level. Identical superscripts, e.g., a, b, c, denote no signif-

The results of the determined density of hornbeam and Paulownia wood before and

(MPa) (1)

the accuracy of 0.1 mm. The hardness was calculated from the following formula:

=

d is the average residual indentation (mm).

Z050TH (Zwick/Roell, Ulm/Germany).

**3. Results and Discussion**

after densification are presented in Table 1.

of the increase in the size of the sample to its initial size.

icant difference between mean values of the investigated properties.

After six weeks of conditioning, the densified wood samples were subjected to compression and a Brinell hardness test. The hardness of the tested materials was determined in three anatomical directions. A steel ball with a diameter of 10 mm was pressed into the surface of the samples with a force of 1000 N for densified and 500 N for control hornbeam samples and for paulownia wood, 500 N and 150 N, respectively. The maximum and minimum diameters of the residual indentation were measured with a Brinell magnifier to the accuracy of 0.1 mm. The hardness was calculated from the following formula:

$$HB = \frac{2F}{D\pi\left(D - \sqrt{D^2 - d^2}\right)} \text{ (MPa)}\tag{1}$$

where *F* is the force acting on the ball (N), *D* is the diameter of the pressed ball (mm), and d is the average residual indentation (mm).

Compressive strength was determined in the direction of densification, so in the radial direction. The stress at the proportionality limit (so-called compressive strength perpendicular to the grain or relative strength (RcR in the radial direction)) was determined. Compression tests were conducted using a numerically controlled test machine, Zwick Z050TH (Zwick/Roell, Ulm/Germany).