*Article* **Effect of Temperature Sensor Numbers and Placement on Aeration Cooling of a Stored Grain Mass Using a 3D Finite Element Model**

**Benjamin Plumier 1,\* and Dirk Maier 2**

**Abstract:** Grain stored in silos in the United States of America is generally cooled with an aeration system to limit mold spoilage and insect infestation. Monitoring efficacy of aeration and real-time conditions of stored grain is generally done using temperature cables with fixed-spaced sensor locations that are hung from the roof of the silo. Numerous placement options exist in terms of the number of cables and their positions. However, little investigation has been done into the effects of cable placement on aeration system operation decisions and real-time monitoring of stored grain conditions. For a one-year period, the temperatures predicted by sensors in three recommended temperature cable configurations were evaluated for conditions in Ames, IA, USA. The average temperatures of each of the cable sensor configurations were lower than the average temperatures of the entire silo, with as much as an 11.4 ◦C difference. When sensor locations were used as inputs for aeration control, all cable sensor configurations predicted similar average temperatures. However, the temperature averages varied by as much as 3.6 ◦C depending on the temperature cable distribution chosen. Results demonstrated that temperature cables near the center or near the edges of the silos produce results that are not representative of the grain mass, resulting in less efficient aerations. Simulations were also conducted with randomized horizontal "wireless" sensor locations at fixed grain depths. The average temperatures were similar, but an increase in the number of sensors reduced variability between simulated storage years as the number of randomized sensors increased.

**Keywords:** aeration; finite element modeling; stored products; temperature sensors


*Article* **Control Techniques for Gob-Side Entry Driving in an Extra-Thick Coal Seam with the Influence of Upper Residual Coal Pillar: A Case Study**

**Shengrong Xie \*, Fangfang Guo and Yiyi Wu**

School of Energy and Mining Engineering, China University of Mining and Technology-Beijing, Beijing 100083, China<EMAIL_ADDRESS>(F.G<EMAIL_ADDRESS>(Y.W.)

**\*** Correspondence<EMAIL_ADDRESS>
**Abstract:** In multi-seam mining, the residual coal pillar (RCP) in the upper gob has an important influence on the layout of the roadway in the lower coal seam. At present, few papers have studied the characteristics of the surrounding rock of gob-side entry driving (GED) with different coal pillar widths under the influence of RCP. This research contributes to improving the recovery rate of the extra-thick coal seam under this condition. The main research contents were as follows: (1) The mechanical parameters of the rock and coal mass were obtained using laboratory experiments coupled with Roclab software. These parameters were substituted into the established main roof structure mechanics model to derive the breakage position of the main roof with the influence of RCP, and the rationality of the calculation results was verified by borehole-scoping. (2) Based on numerical simulation, the evolution laws of the lateral abutment stress in the lower working face at different relative distances to the RCP were studied. FLAC3D was used to study the whole space-time evolution law of deviatoric stress and plastic zone of GED during driving and retreating periods with various coal pillar widths under the influence of RCP. (3) The plasticization factor *P* was introduced to quantify the evolution of the plastic zone in different subdivisions of the roadway surrounding rock, so as to better evaluate the bearing performance of the surrounding rock, which enabled a more effective determination of the reasonable coal pillar width. The field application results showed that it was feasible to set up the gob-side entry with an 8 m coal pillar below the RCP. The targeted support techniques with an 8 m coal pillar could effectively control the surrounding rock deformation.

**Keywords:** residual coal pillar; gob-side entry driving; extra-thick coal seam; coal pillar size; surrounding rock control

#### **1. Introduction**

Gob-side entry driving (GED) with a small coal pillar (3~8 m) is widely used in China's mines due to the simple development processes, high resource recovery rate, and the ability of the coal pillar to isolate gangue, harmful gases, water, and fire in the adjacent gob [1,2]. China has many thick and extra-thick coal seams, and the GED has been successfully applied in many mines with extra-thick coal seams [3,4]. For the mining of extra-thick coal seams (>8 m), the characteristics mainly include [5–8]: (1) Strong-dynamic disturbance of abutment stress. Fully mechanized top coal caving mining is generally applied to the mining of extra-thick coal seams, with a large mining space and intense roof activity, resulting in severe abutment stress on the coal mass. (2) A thick top coal over the roadway. The roadway is usually developed along the floor line of the coal seam, the roof of the GED is composed of weak coal masses with a large thickness. The weak properties of the top coal mass seriously increase the difficulty of roadway control. (3) Disturbed by multiple mining-induced stresses. The GED is not only disturbed by the lateral abutment stress of the previous working face, but also advanced abutment stress of the present working face.

**Citation:** Xie, S.; Guo, F.; Wu, Y. Control Techniques for Gob-Side Entry Driving in an Extra-Thick Coal Seam with the Influence of Upper Residual Coal Pillar: A Case Study. *Energies* **2022**, *15*, 3620. https:// doi.org/10.3390/en15103620

Academic Editor: Sergey Zhironkin

Received: 18 April 2022 Accepted: 12 May 2022 Published: 15 May 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/).

Therefore, it is of great significance to study the selection principles of reasonable coal pillar width and surrounding rock control techniques of the GED in extra-thick coal seams.

Presently, many scholars around the world have conducted extensive research on the design of coal pillar width using various methods such as theoretical analysis, numerical simulation, and similar simulation. UDEC can effectively study the macro-mechanical behavior and crack propagation law during the rock failure process, so it is frequently applied to the analysis of the roadway failure mechanism. Bai et al. [9] studied how the failure mechanisms ofa7m coal pillar width during the formation process caused a large deformation. They investigated the propagation of shear and tensile cracks in the coal pillar of various widths and optimized the coal pillar width and support measures. Shi et al. [10] investigated the crack evolution mechanism of the gob-side entry for different conditions and proposed optimized-support parameters combined with roof-cutting measures. Gao [11–13] carried out a series of numerical simulations using the UDEC Trigon approach to focus on the roadway damage caused by squeezing failure and shear failure, and the effects of rock bolts in the roadway support were also evaluated. FLAC3D can simulate the mechanical behavior of geological materials and geotechnical engineering effectively and is widely used in underground coal mining activities [14–16]. He et al. [17–21] studied the stress distribution characteristics and plastic zone of roadway with different coal pillar widths in the process of coal mining and proposed a support technology for setting a reasonable width of coal pillars. The specific coal pillar width was designed and applied in engineering practices, which realized a good control effect of the roadway surrounding rock. Ma [22] conducted research on the stress distribution characteristics of the narrow coal pillar with different top coal heights of gob-side entry, and they believed that the thick top coal was not conducive to control of the roadway. Jiang [23] presented an approach for evaluating, designing, and optimizing EDG and yield pillar based on the results of numerical simulations and field practice. Han et al. [24–26] considered special geological situations to drive an entry along the gob, such as in an isolated island working face, in a deep soft-broken coal seam, or in a working face adopting roadway side sealing technology. The targeted control techniques of the surrounding rock were proposed and successfully applied in the field practices. These studies had made countless valuable explorations on the width of coal pillars and failure mechanisms and made certain innovations in the research methods. However, there were very few studies on the layout of the GED with the influence of a residual coal pillar (RCP) during the mining of multiple coal seams.

China has many multiple coal seam mining areas, such as Datong, Xinwen, Pingdingshan, etc. [27]. The mining of two adjacent coal seams tends to interact with each other due to the close distance [28,29]. The mining of the upper coal seam generally renders many section pillars, and lower coal seam roadways are usually designed with larger section coal pillars (usually over 20 m) to avoid safety hazards. That results in the waste of massive coal resources, especially for extra-thick coal seams. This paper focused on the layout selection of GED below the RCP. The lateral breakage position of the main roof under the influence of RCP was calculated by theoretical derivation and verified the rationality of the results using borehole-scoping. Under the influence of RCP, the whole space-time evolution laws of deviatoric stress and plastic zone of GED were studied during driving and retreating periods with different coal pillar widths using numerical simulation. The quantified index for the plastic zone of the surrounding rock of GED, plasticization factor P, was proposed. A coal pillar width of 8 m was finally determined. Based on the extent of the plastic zone of GED obtained by borehole-scoping, the targeted roadway support scheme was proposed. That provided a reference for promoting the coal recovery rate in similar geological conditions.

#### **2. Project Overview**

#### *2.1. Geological Conditions*

Nanyangpo coal mine is located in Shanyin County, Shanxi Province, China. The main mining seams are No. 4 and No. 6 coal seam with a spacing of 32 m. The average thickness of the No. 4 coal seam is about 3.0 m, which has been mined out, but 15 m section coal pillars are left between working faces. The current main mining seam is the No. 6 coal seam with an inclination of 3~5◦ and an average thickness of 9.6 m. The mining method is fully mechanized top coal caving mining of an extra-thick coal seam, the machine mining height is 3.5 m, and the caving height is 5~10 m. The geological column chart of the No. 6 coal seam is illustrated in Figure 1. The panel 26,102 is near the northern boundary of the mining field, and next to the panel 26,104 and 26,106 that have been mined out with a 20 m coal pillar left. To avoid wasting resources, Nanyangpo coal mine plans to develop the 26,102 tailgate along the gob with a small coal pillar. The layout of the panels is given in Figure 2.

**Figure 1.** Generalized stratigraphic column of the test site with the panel layout.

**Figure 2.** Detailed panel layout of the test site.

#### *2.2. Rock Mass Properties*

The physical and mechanical properties of rock mass are an important basis for the design of roadway support. The parameters obtained in the laboratory are also key data for further study on the theoretical model calculation and numerical simulation. The mechanical parameters of coal and rock mass in panel 26,102 are shown in Table 1.

**Table 1.** Properties of rock mass in panel 26,102.

Note: *σci* is the uniaxial compression strength, *σti* is the tensile strength, *ci* is cohesion, *Ei* is elastic modulus, *ϕi* is internal friction angle, and *νi* is Poisson's ratio.

It is difficult to get the mechanical parameters of coal and rock samples measured in the laboratory to reflect the actual physical and mechanical properties in engineering sites due to the absence of the original environment and structural characteristics of rock mass [30,31]. E. Hoek and E. T. Brown et al. [32,33] obtained the Hoek–Brown failure criterion by a large number of rock mechanics tests as well as field tests of rock masses after continuous revision and improvement. The parameters of coal and rock samples obtained in the laboratory were imported into Roclab software based on the Hoek–Brown strength criterion for calculation, and the revised parameters were obtained in Table 2, which were more in line with the engineering reality.

**Table 2.** Revised properties of rock mass by Hoek–Brown strength criterion.

Note: *mi* is the constant of the intact rock, *GSI* is the constant evaluating the fractured rock mass, and *D* is the disturbance factor.

#### **3. Breakage Position of the Main Roof**

*3.1. Influence of Upper Residual Coal Pillar*

The stress redistribution in the roof strata after mining of the upper coal seam resulted in the RCP bearing a larger load. Thereby, a certain range of stress elevation area was formed at the floor of the RCP. That meant the different positions of the lower working face will make the GED in different stress environments. Based on numerical simulation, the evolution laws of the lateral abutment stress in the lower working face at different relative

distances to the RCP were as shown in Figure 3. Taking the centerline of the upper RCP as the base point, the peak values of lateral abutment stress were 23.2 MPa and 25.5 MPa when the edge of the working face (EWF) was −37.5 m and 37.5 m from the centerline of the RCP separately, with an increase of 0.9% and 10.9%. The peak positions of lateral abutment stress were both about 6 m from the gob. When EWF was −17.5 m and 17.5 m from the centerline of the RCP separately, the peak values of lateral abutment stress were 25.7 MPa and 24.7 MPa with an increase of 11.7% and 7.4%, and the peak positions of lateral abutment stress were both about 6 m from the

gob. When EWF was −7.5 m and 7.5 m from the centerline of the RCP separately, the peak values of lateral abutment stress were 29.9 MPa and 29.3 MPa with an increase of 30.0% and 27.4%, and the peak positions of lateral abutment stress were about 7 m and 6 m from the gob. When EWF was −2.5 m and 2.5 m from the centerline of the RCP separately, the peak values of lateral abutment stress were 28.3 MPa and 28.0 MPa with an increase of 23.0% and 21.7%, and the peak positions of lateral abutment stress were about 11 m and 7 m from the gob. The peak values of lateral abutment stress of working face showed an "M-shaped" tendency, and the peak depth first increased and then declined.

**Figure 3.** The evolution laws of the lateral abutment stress in lower working face at different relative distances to the RCP. (**a**) Overview under different distances. (**b**) Peak stress ratio and distribution of peak locations. (Note: the gob side and virgin coal side in the figure were both relative to the lower working face).

The above distribution laws indicated that the lateral abutment stress was most significantly affected when the edge of the lower working face was located directly below the residual coal pillar, and the peak growth coefficient was 1.20~1.30. Additionally, when EWF of the lower coal seam was 2.5 m away from the centerline of RCP, the peak growth coefficient of the lateral abutment stress was 1.22. The influence of the stress disturbance of RCP was relatively diminished when a small coal pillar was set up to drive the entry along the gob in such conditions. Consequently, the superimposed disturbance impact of RCP and mining-induced stress on GED should be fully considered to prevent the roadway from instability.

#### *3.2. Mechanical Model of the Main Roof*

The main roof controls the upper weak strata of the coal seam. The fracture morphology, the hinge status, and the stability of key blocks after main roof breakage greatly impact the stress distribution of the surrounding rock [34,35]. There is a 15 m coal pillar left in No. 4 coal seam at 32 m above the edge of 26,104 gob in Nanyangpo coal mine. The disturbance of RCP on the stress of floor will inevitably affect the load distribution pattern of the main roof in No. 6 coal seam. Based on the distribution regulation of the lateral abutment stress obtained by numerical simulation in lower working face under the influence of RCP, the mechanical model of the elastic foundation beam is established, as shown in Figure 3. The load curve shows the abutment stress before the lateral breaking of the main roof. The main roof above the virgin coal area is simplified as the elastic foundation beam under the pressure of overburden rock, and the hanging part of the main roof in gob is assumed as the cantilever beam structure.

The rock-beam of the overhanging part in the gob is taken for forces analysis as in Figure 4b. According to ∑ *Fy* = 0:

$$\begin{array}{rcl} F &=& \int\_{1}^{l\_1} (\frac{(q\_1 - q\_2)\mathbf{x}}{l\_1} + q\_2) d\mathbf{x} \\ &+ \int\_{l\_2}^{l\_2} (q\_1 - \frac{(q\_1 - q\_c)\mathbf{x}}{l\_2 - l\_1}) d\mathbf{x} + \int\_{l\_2}^{l\_3} q\_c d\mathbf{x} \end{array} \tag{1}$$

Then, the shear force *Q*0 and the bending moment *M*0 at *x* = 0 are

$$\begin{array}{rcl} Q\_0 &= F = \int\_0^{l\_1} (\frac{(q\_1 - q\_2)\mathbf{x}}{l\_1} + q\_2) d\mathbf{x} \\ &+ \int\_{l\_1}^{l\_2} (q\_1 - \frac{(q\_1 - q\_2)\mathbf{x}}{l\_2 - l\_1}) d\mathbf{x} + \int\_{l\_2}^{l\_3} q\_c d\mathbf{x} \\\ M\_0 &= \int\_0^{l\_1} (\frac{(q\_1 - q\_2)\mathbf{x}}{l\_1} + q\_2) \mathbf{x} d\mathbf{x} \\ &+ \int\_{l\_1}^{l\_2} (q\_1 - \frac{(q\_1 - q\_2)\mathbf{x}}{l\_2 - l\_1}) d\mathbf{x} + \frac{q\_2(l\_3 - l\_2)(l\_3 + l\_2)}{2} \end{array} \tag{2}$$

where *l*1 is the horizontal distance from the center of RCP to the edge of the lower virgin coal; *l*2 is the influence range of RCP on the stress distribution of the main roof; and *l*3 is the hanging length of the main roof in the gob, which is approximately equal to the periodic weighting step of the working face.

*q*1 = *K*1*γH*, *q*2 = *K*2*γH*, where *K*1 and *K*2 are the stress increase coefficient, *H* is the buried depth of the main roof stratum. *q*c is the uniform load on the overhanging part of the main roof, the expression is as follows:

$$q\_{\varepsilon} = \frac{E\_1 h\_1^3 (\gamma\_1 h\_1 + \gamma\_2 h\_2 + \dots + \gamma\_n h\_n)}{E\_1 h\_1^3 + E\_2 h\_2^3 + \dots + E\_n h\_n^3} \tag{3}$$

The beam structure model with immediate floor, virgin coal, and immediate roof as the elastic foundation is shown in Figure 4c. Taking the edge of virgin coal as the origin, the relationship between subsidence *y* and stress *p* is as follows:

$$p = -ky\tag{4}$$

$$k = \frac{E\_i}{(1 - \nu\_i^{-2})h\_i}, \frac{1}{k} = \sum\_{i}^{n} \frac{1}{k\_i} \tag{5}$$

**Figure 4.** Mechanical model of elastic foundation beam. (**a**) Structural diagram before basic roof breaking. (**b**) Analysis of rock-beam of the overhanging part. (**c**) Analysis of rock-beam with the elastic foundation.

where *k* is the foundation stiffness coefficient, which is related to Poisson's ratio *νi* and

The elastic foundation beam differential equation is

$$\frac{d^4y}{dx^4} + \frac{ky}{EI} = \frac{q(x)}{EI}, \; \beta = \sqrt[4]{\frac{k}{4EI}} = \frac{1}{L} \tag{6}$$

where *L* is the characteristic length of the beam and *EI* is the bending rigidity of the main roof rock mass.

Then, the differential equation of elastic foundation beam is

$$\frac{d^4y}{dx^4} + 4\beta^4 y = \frac{q(x)}{EI} \tag{7}$$

According to the calculation of elastic foundation beams [36], the general solution of the deflection equation is

$$\begin{array}{ll} y & = \ y\_0 \phi\_1(\beta x) + \theta\_0 \frac{1}{\beta} \phi\_2(\beta x) - M\_0 \frac{1}{EI\beta^2} \phi\_3(\beta x) \\ & - Q\_0 \frac{1}{EI\beta^3} \phi\_4(\beta x) + \left< g\_{(x)} \right>\_t \end{array} \tag{8}$$

where the Kralof function, defined to simplify the calculation, is as follows:

$$\begin{array}{l} \phi\_1 = clh\beta x \cos\beta x\\ \phi\_2 = \frac{1}{2} (clh\beta x \sin\beta x + \sin\beta x \cos\beta x)\\ \phi\_3 = \frac{1}{2} sh\beta x \sin\beta x\\ \phi\_4 = \frac{1}{4} (clh\beta x \sin\beta x - \sin\beta x \cos\beta x)\\ ch\beta x = \frac{c^{\beta x} + c^{-\beta x}}{2}\\ sh\beta x = \frac{c^{\beta x} - c^{-\beta x}}{2} \end{array} \tag{9}$$

where *y*0, *θ*0, *M*0, *Q*0 are the deflection, angle of rotation, bending moment, and shear force at *x* = 0; *g*(*x*) *t* denotes the correction term that should be added when *x* > *t*.

For the solution of the elastic foundation beam under the loading condition as shown in Figure 4c, the deflection equation of the elastic foundation beam with a correction term section can be given according to the superposition principle:

thickness *hi* of the weak rock layers below the main roof.

$$\begin{array}{ll} \mathsf{g} &= \frac{q\_{2}}{k} + \left(\frac{q\_{3} - q\_{2}}{k l\_{1}} - \frac{q\_{2}}{k}\right) \frac{1}{\mathcal{J}} \phi\_{2}(\beta \mathbf{x}) - \frac{q\_{3} - q\_{2}}{k l\_{1}} \mathbf{x} \\ &+ \left\langle \begin{bmatrix} \frac{q\_{2} - q\_{3}}{k l\_{4}} + \frac{(q\_{3} - q\_{0})}{(l\_{5} - l\_{4}) k} \end{bmatrix} \mathbf{x} + \frac{2q\_{2} - q\_{3}}{k} \phi\_{1}[\beta(\mathbf{x} - l\_{4})] \right. \\ &+ \left[\frac{(q\_{0} - q\_{3})l\_{4}}{(l\_{5} - l\_{4}) k} - \frac{q\_{2}}{k}\right] - \left[\frac{q\_{2} - q\_{3}}{k l\_{4}} + \frac{(q\_{3} - q\_{0})}{(l\_{5} - l\_{4}) k} \right] \frac{1}{\mathcal{J}} \phi\_{2}[\beta(\mathbf{x} - l\_{4})] \Big\rangle\_{l\_{4}} \\ &\left\langle \frac{(q\_{0} - q\_{3})l\_{5}}{(l - l\_{4})k} - \frac{(q\_{0} - q\_{3})}{(l - l\_{5} - l\_{4}) k} \mathbf{x} + \frac{(q\_{0} - q\_{3})}{(l - l\_{5} - l\_{4}) k} \frac{1}{\mathcal{J}} \phi\_{2}[\beta(\mathbf{x} - l\_{5})] \right\rangle\_{l\_{5}} \end{array} \tag{10}$$

where the load on the main roof at *l*5 length in virgin coal approximately equals the original rock stress, that is *q*0 = *γH*, *q*3 = *K*3*γH*, which is approximately considered that the peak value of the abutment stress of the main roof above the virgin coal with a length of *l*4 from the edge of gob. *K*3 is the stress disturbance coefficient factor.

The equation for the deflection, angle of rotation, bending moment, and shear force of the main roof before breakage are given:

$$\begin{array}{rcl} y &=& y\_0 \phi\_1(\beta x) + \theta\_0 \frac{1}{\beta} \phi\_2(\beta x) - M\_0 \frac{1}{EI\beta^2} \phi\_3(\beta x) \\ &- Q\_0 \frac{1}{EI\beta^3} \phi\_4(\beta x) + \mathcal{g} \\ \theta &=& -y\_0 \beta \phi\_4 + \theta\_0 \phi\_1 - M\_0 \frac{1}{EI\beta^2} \phi\_2 - Q\_0 \frac{1}{EI\beta^2} \phi\_3 + \mathcal{g}' \\ M &=& y\_0 \cdot 4EI\beta^2 \phi\_3 + \theta\_0 \cdot 4EI\beta \phi\_4 + M\_0 \cdot \phi\_1 + Q\_0 \frac{1}{\beta} \phi\_2 + \mathcal{g}'' \\ Q &=& y\_0 \cdot 4EI\beta^3 \phi\_2 + \theta\_0 \cdot 4EI\beta^2 \phi\_3 - M\_0 \cdot 4\beta \phi\_4 + Q\_0 \phi\_1 + \mathcal{g}''' \end{array} \tag{11}$$

The boundary conditions are:

$$\begin{array}{rcl} \mathcal{M}\_{0} &=& \left[ q\_{1} \{ 4l\_{1}^{2} + 5l\_{2}^{2} + 5l\_{1}l\_{2} \} + q\_{2}l\_{1}^{2} \right.\\ &+ q\_{c} \{ l\_{1}^{2} - 5l\_{2}^{2} + 3l\_{3}^{2} + l\_{1}l\_{2} \} \Big] / 6\\\ Q\_{0} &=& \left[ q\_{1} (l\_{2} - l\_{1}) + 2q\_{2}l\_{1} + q\_{c} (2l\_{3} - l\_{1} - l\_{2}) \right] / 2\\ &y|\_{x=l} = 0, \theta|\_{x=l} = 0 \end{array} \tag{12}$$

According to the numerical simulation results and the geological conditions of the mine, the engineering parameters are given as follows: the depth of the 26,104 working face is about 250 m; the dip length of the working face is 240 m; and the periodic weighting step is 20~25 m.

The immediate roof, coal seam, and immediate floor are considered as the elastic foundation of the main roof beam. The foundation coefficient *k* is 0.0380 GPa; the elastic modulus, bending modulus, and bending stiffness of main roof are 21.78 GPa, 67.03 m4, and 1460 GN·m2, respectively; *l*1 = 2.3 m, *l*2 ≈ *l*3 ≈ 20~25 m; *l*4 ≈ 6 m, *l*5 approximately equals to 22 m, *l*5 ≥ 3 *L*, taken as 3 *L* = 60 m; *K*1, *K*2, *K*3 were approximated to 1.2, 0.7, 1.5; *q*0 = 5.85 MPa, *q*1 = 7.02 MPa, *q*2 = 4.10 MPa, *q*3 = 8.78 MPa.

Combining (9)–(12) and substituting the data, the breaking position of main roof is 4.4~5.8 m from the edge of the gob.

#### *3.3. Borehole-Scoping in the Main Roof*

Borehole-scoping can accurately visualize the lithology, thickness, delamination, cracks, and fractures of the overlying rock strata. To verify the theoretical derivation result of the lateral breakage position of the main roof after panel 26,104 retreated, the borehole-peeping stations were arranged in the test section of 26,102 tailgate to observe the cracks propagation in the roof strata. A total of 23 holes were drilled in 4 groups with a total depth of over 640 m. The site construction is shown in Figure 5.

The obvious crack was defined by a width over 5 mm and a length over 10 mm, and the remarkable crack was defined by a width over 10 mm and a length over 100 mm. The distribution of cracks propagation in the roof strata of 26,102 tailgate is shown in Figure 6, indicated that: (1) most of the cracks were distributed between 10.55 m and 21.43 m above the coal pillar, that is, the range of main roof thickness. (2) Remarkable longitudinal cracks with intersecting circumferential cracks occurred at depths of 19.49 m in BII and 13.93 m in BII, indicating that the main floor had been fractured. The borehole wall within 1 m of the

two cracks sites was relatively intact again, so it could be deduced that the fracture line of the main roof was located above the coal pillar. (3) Based on the angles and depths of the boreholes where the two remarkable cracks were observed, the position of fracture line in the main roof was 5~6 m away from the edge of the gob, which was consistent with the theoretical calculation result.

**Figure 5.** Site construction. (**a**) Layout of part boreholes in roof. (**b**) Crawler drill.

**Figure 6.** The distribution of roof strata cracks in 26,102 tailgate. Note: taking the representation of 45◦-T35 m as an example, 45◦ denotes the borehole angle, and T35 m represents the full length of the borehole; taking the description of (**a**) 68◦-19.62 m as an example, (**a**) denotes the serial number of the camera image of borehole, 68◦ denotes the borehole angle, and 19.62 m represents the detection depth.

#### **4. Discussion and Analysis of the Simulation Results**

#### *4.1. Global Model for the Pillar Width*

The model selects the 26,102 tailgate as the test roadway, which is adjacent to the gob of panel 26,104. The *X*-axis is 180 m in the length direction of working face, the *Y*-axis is 120 m in the advancing direction and the *Z*-axis is 110 m in the vertical direction. The roadway section is rectangular, and the dimension is 5.2 m × 3.5 m (width × height). The boundary displacement of the model is constrained in horizontal and bottom. The upper of the model is subjected to a stress of 4.05 MPa equivalent to the self-weight of the overburdened rock, and the lateral pressure coefficient is 1.2. The model is calculated using the Mohr-Coulomb model, while the double-yield model is used for the gob, and the mechanical parameters of each rock formation are taken from Table 2.

#### *4.2. Double-Yield Model of the Gob*

With the retreating of the working face, the broken blocks are backfilled in the gob after the main roof periodic fracture. The modulus coefficient of the gangue in the gob will increase significantly after compacted. The densely compacted gob can bear part of the abutment pressure, which effectively weakens the stress concentration in the coal pillar. Therefore, the double-yield model in FLAC3D can be used to simulate the compaction and hardening process of waste in the gob [37]. The overburden pressure parameters and mechanical parameters of gangue in the gob can be obtained by the Salamon formula, expressed as follows [38]:

$$
\sigma = E\_0 \varepsilon / \left( 1 - \frac{\varepsilon}{\varepsilon\_{\text{max}}} \right) \tag{13}
$$

where *E*0 is the initial tangential modulus, GPa; *σ* is the stress of gangue in the gob, MPa; *ε* is the bulk strain of the compressed gangue in the gob; *ε*max is the maximum bulk strain. The values are given:

$$\begin{array}{l} \varepsilon\_{\text{max}} = (b - 1) / b \\ E\_0 = \frac{1.039 v\_c^{1.042}}{b^{\prime \prime}} \\ b = (h + h\_c) / h\_c \end{array} \tag{14}$$

where *b* is the crushing expansion coefficient of gangue; *σ*c is the gangue compressive strength; *h* is the mining height of the coal seam; *h*c is the height of the roof caving zone in the gob.

According to the engineering situations of Nanyangpo coal mine, the average mining height of coal seam during retreating period is 9.6 m, and the height of the roof caving zone in the gob is about 34.6 m. The values can be substituted into Equations (13) and (14) to derive the double-yield model parameters of the gangue in the gob as shown in Table 3.

**Table 3.** Double-yield model parameters of gangue in the gob.

The specification of the established unit sub-model of the gob is 1 m × 1 m × 1 m, and a constant velocity of 10−5 m/s is applied to the upper surface of the model to determine the mechanical properties of the gob by trial-and-error method. when the parameters are set to the gangue with a density of 1000 kg/m3, bulk modulus of 11.12 GPa, shear modulus of 5.20 GPa, internal friction angle of 5◦, the stress–strain curve of the numerical simulation matches well with that of the theoretical calculation, as shown in Figure 7.

**Figure 7.** Numerical simulation inversion of goaf parameters.

#### *4.3. Discussion and Analysis of the Simulation Results*

#### 4.3.1. Results with Various Coal Pillar Widths

Deviatoric stress is the synthesis of horizontal, vertical, and tangential stresses, and represents the distribution of shear stress in the material subjected to loads, revealing that the essential force of rock failure is mainly caused by shear stress. This stress index is gradually adopted by an increasing number of papers [39,40].

The size of the coal pillar affects the state of deviatoric stress distribution, the failure range of the plastic zone and the deformation extent of the roadway surrounding rock in GED. In this paper, we define the plasticization factor *P* as a parameter to characterize damage of the surrounding rock in GED, which is given as

$$P\_1 = \frac{S\_1}{S\_p}, P\_2 = \frac{S\_2}{S\_r}, P\_3 = \frac{S\_3}{S\_\varepsilon}, P\_4 = \frac{S\_4}{S\_\varepsilon} \tag{15}$$

where *P*1, *P*2, *P*3, and *P*4 are the plasticization factors of coal pillar subdivision (I), top coal subdivision (II), virgin coal upper corner subdivision (III), and virgin coal subdivision (IV) in GED, respectively. *S*1, *S*2, *S*3, and *S*4 are the areas of plastic zone in I, II, III, and IV. *Sp*, *Sr*, and *Se* are the cross-section area of I, II, and gob-side entry, respectively.

As shown in Figures 8 and 9, with the increase of coal pillar width, the peak zone of deviatoric stress in the virgin coal area of GED tended to be gradually reduced, which contrasted with the coal pillar area. With the width of coal pillar being 4~6 m, the surrounding rock of gob-side entry was in a lower stress environment. While the plastic zone in II was coalesced with that of the overlying rock and the high stress of the surrounding rock was mostly concentrated in the virgin coal area, indicating that the bearing capacity of coal pillar was poor in such conditions. The range of plastic zone of GED decreased rapidly with a coal pillar of 8 m. The peak zone of deviatoric stress in coal pillar progressively enlarged, while virgin coal side continued to diminish up to similar peak values on both sides. It indicates that the bearing capacity of the coal pillar was enhanced and initiated to sustain the overburden stress in concert with the virgin coal. The coal pillar was in a high stress state with a width of 10~15 m. Additionally, the main bearing body of the overlying load gradually transformed from the virgin coal side to the coal pillar. At this time, the plastic zone of GED was small with a high stability of the surrounding rock, while a significant amount of the coal resources were wasted.

With the increase of coal pillar width, the plasticization factor in four subdivisions of the entry gradually declined. Correspondingly, the peak value of deviatoric stress in the coal pillar area grew swiftly to a stable state, while the value in virgin coal area continued to get smaller. When the width of coal pillar is 8 m, the *P* were not more than 80% in I and II, and 60% in III and IV. The peak deviatoric stress in coal pillar area was roughly coincident with that of the virgin coal area.

#### 4.3.2. Results with Disturbance of Panel Retreating

Mining-induced stress exerted an essential influence on the stability of the surrounding rock in GED. Exploring the distribution characteristics of the deviatoric stress and plastic zone of the surrounding rock advanced of the working face could provide the necessary basis for the surrounding rock control.

As shown in Figures 10 and 11:

(1) In the vicinity of the working face, the peak and range of deviatoric stress in coal pillar were much larger than that in virgin coal area, and the peak ratio was 1.41, indicating that the stress in the surrounding rock was mainly sustained by the coal pillar. *P* were more than 90% in I and II, and 100% in III and IV, representing that extensive damage occurred in the surrounding rock of GED, that meant reinforced-supported measures should be implemented to avoid destabilization of the coal pillar.

The above analysis suggested that the plasticization factor *P* in III and IV were more significantly troubled by mining-induced stress of the working face. *P* firstly dropped from both much more than 100% to 40% and 60% in III and IV, respectively. The deviatoric stress in coal pillar and virgin coal area first increased rapidly to the peak value, and then gradually tended to be stable. That indicated that the main bearing body of overburden stress was "the coal pillar → the virgin coal area → the collaboration of coal pillar and virgin coal area".

#### 4.3.3. Coal Pillar Width Determination

Based on the theoretical calculation and field measurement results, the fracture position of the main roof is located in the range of 5~6 m above the virgin coal from edge of the gob. That means the fracture line of the main roof is above the coal pillar when the width is 8 m, which contributes to maintaining the stability of the surrounding rock of the gob-side entry. The overlying strata load of GED is performed by virgin coal area in cooperation with coal pillar. The plasticization factors of coal pillar subdivision and top coal subdivision are both less than 80%, which are within the control of support. The surrounding rock in GED, advanced 60 m of the working face, is obviously affected after the panel retreated, and the influence on the stress and plastic zone of the surrounding rock tends to be stable. Therefore, it is feasible to set up an 8 m coal pillar with targeted support techniques to maintain the stability of the surrounding rock of GED.

#### **5. Surrounding Rock Control Techniques**

*5.1. Cracks Distribution of the Coal Body*

The results of borehole-scoping in 26,102 tailgate are shown in Figure 12:

(3) In the roof of GED. The coal body was severely damaged at a depth of 1.00 m from upper corner of coal pillar rib, 1.95 m from the roof of the roadway, and 1.76 m from the upper corner of virgin pillar rib. The coal body gradually turned intact with the distances exceeding 1.76 m, 4.72 m, and 3.97 m, respectively.

**Figure 8.** Characteristics of the surrounding rock with various coal pillar widths. (**a**) Deviatoric stress. (**b**) The plastic zone.

**Figure 9.** Distribution of the plasticization factor in subdivisions.

**Figure 10.** The distribution characteristics of surrounding rock advanced of the working face.

#### *5.2. Support Principles*

Based on the crack distribution of the surrounding rock, the compressive stress zone of bolts should cover the fracture zone, and the length of anchor cables should be greater than the depth of plastic zone, which played an effective role in improving the stress situations of the surrounding rock. The control techniques of the surrounding rock with the cooperative bearing of bolts and anchor cables were proposed, and the support principles are shown in Figure 13. The main contents were as follows: (1) Shallow bearing area of bolts. The 2.4 m prestressed bolts were used in the shallow part of the surrounding rock, supplemented by high strength plates and W-shaped steel straps, to provide combined compressive stress for the fractured zone, which contributed to preventing support components from failure caused by the surrounding rock falling of the fractured zone. (2) Deep bearing zone of anchor cables. The 8.0 m high prestressed anchor cables were used in the roof and were embedded in the stable strata by passing through the top coal of about 6 m thick in the roof. The 5.0 m high prestressed anchor cables were used in both ribs to pass through the fracture zone and anchored into the relatively intact coal body. The high pre-stressed anchor cable could increase the shear resistance of the coal body, making the surrounding rock of the roadway to reinstate the state of three-dimensional stress in a certain extent, and exert the bearing capacity of the deep surrounding rock of the roadway.

**Figure 11.** The stress distribution advanced of the working face in GED.

**Figure 12.** Borehole-scoping for the coal body in 26,102 tailgate.

**Figure 13.** Cooperative bearing structure of bolts and anchor cables.

#### *5.3. Technical Measures*

Based on the field geological conditions, the support scheme of 26,102 tailgate is shown in Figure 14. The specific support parameters are as follows:

**Figure 14.** Detailed support parameters. (**a**) Entry section. (**b**) Support pattern in roof. (**c**) Support pattern in coal pillar rib. (**d**) Support pattern in virgin coal rib.

The roof and two ribs were supported by high-strength steel bolts 20 mm in diameter and 2400 mm in length. The inter-row spacing of the bolts was 1000 mm × 900 mm. A row of anchor cables is arranged for every two rows of bolts. The anchor cables 17.8 mm in diameter and were used for roadway support with a length of 8000 mm in roof and 5000 mm in two ribs. The inter-row spacing for the cables was 1100 mm × 2700 mm. The bottom one of each row of anchor cables was replaced by a high-strength steel bolt 20 mm in diameter and 2400 mm in length for both ribs. The bolts and anchor cables were connected with steel ladder beams made of 16 mm round steel in both ribs and with W-shaped steel straps with length × width × thickness of 5000 mm × 300 mm × 3 mm in the roof. Bolts in the roof and rib corners were installed at a 15◦ incline. Bolt plates of W-shaped steel straps were selected for the bolts in both ribs. A metal mesh 4 mm in diameter was used in the roadway section to prevent broken coal body from falling.

#### **6. Application and Analysis**

Field measurements of ground response can effectively and comprehensively reflect the working status of the support system and verify the effect of the roadway support scheme, which contributes to the stability of the roadway support. Three stations were set up in the 200 m test section of 26,102 tailgate, and each station included one group deformation observation of GED and one group of bolts and anchor cables forces monitoring. The station layout is shown in Figure 15.

**Figure 15.** Layout of the field measurements.

Taking the monitoring results of the typical station II as an example, the surface deformation of the roadway surrounding rock is shown in Figure 16. With the combined support of high-strength prestressed bolts and anchor cables, there was no considerable deformation and damage in GED during roadway driving and retreating period. The surrounding rock deformation in the roof, virgin coal rib and coal pillar rib finally stabilized by 140 mm, 96 mm, and 105 mm separately after 25 days of roadway development. During the driving period, the maximum deformation of roadway in roof, virgin coal rib, and coal pillar rib advanced of the working face was 296 mm, 272 mm, and 251 mm, respectively.

**Figure 16.** Deformation observation of the surrounding rock. (**a**) During roadway development. (**b**) During working face retreating.

At each monitoring section, the forces of bolts and anchor cables were monitored in the roof and two ribs, respectively. The measured bolts and anchor cables of the coal pillar rib, roof and virgin coal rib were numbered B1, B2, B3 and C1, C2, C3, respectively. The initial prestressing forces of bolts and anchor cables were 72~80 kN and 175~190 kN, which were 40~45% and 34~37% of the breaking load separately. The monitoring results are shown in Figure 17.

**Figure 17.** Anchor cable (bolt) force at different distances from the working face.

When the monitoring section was 60 m away from the working face, the forces of B1, B2, and B3 bolts were about 51%, 49%, and 45% of their breaking load (179 kN), and the forces of C1, C2, and C3 anchor cables were about 37%, 40%, and 43% of their breaking load (520 kN). The growth of bolts and anchor cables forces was less than 10%, indicating that the roadway section was weakly affected by the mining-induced stress at a distance of over 60 m from the working face. The disturbance impact of abutment stress was rapidly intensified once the monitoring section was less than 60 m from the working face. The forces of B1, B2, and B3 bolts were about 78%, 74%, and 70% of their breaking load, and the forces of C1, C2, and C3 anchor cables were about 75%, 78%, and 69% of their breaking load with 20 m from the working face. The forces of bolts and anchor cables were far less than its upper breaking limit, indicating that they were in good working condition. The 20 m area of GED advanced of the working face was supported by single hydraulic props, which contributed to avoiding the instability of the surrounding rock caused by intense disturbance of mining-induced stress. Therefore, the monitoring of mining pressure in this area could not be highlighted. Field measurements of ground response showed that the combined control techniques of bolts and anchor cables with an 8 m coal pillar achieved effective control of the roadway surrounding rock under the influence of upper residual coal pillar.

#### **7. Conclusions**

Based on theoretical calculation, numerical simulation, and field measurements, the evolution laws of the lateral abutment stress in lower working face at different relative distances to the RCP were studied, as well as the whole space-time evolution law of deviatoric stress and plastic zone of GED during driving and retreating periods with various coal pillar widths under the influence of RCP. The targeted support techniques with an 8 m coal pillar were proposed. The conclusions were as follows:

**Author Contributions:** Conceptualization, S.X. and F.G.; methodology, F.G. and Y.W.; software, F.G. and Y.W.; data curation, F.G. and Y.W.; writing—original draft preparation, F.G.; writing—review and editing, S.X. and F.G.; supervision, S.X.; project administration, S.X., F.G. and Y.W. All authors have read and agreed to the published version of the manuscript.

**Funding:** This work is financially supported by the National Natural Science Foundation of China (Grant No. 52074296), the Fundamental Research Funds for the Central Universities (Grant No. 2022YJSNY18), the National Natural Science Foundation of China (Grant No. 52004286), the Fundamental Research Funds for the Central Universities (Grant No. 2022XJNY02), and the China Postdoctoral Science Foundation (Grant No. 2020T130701, 2019M650895), all of which are gratefully acknowledged.

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

#### **References**

**Yunbing Hou, Junqi Cui \* and Ruipeng Liu**

School of Energy and Mining Engineering, China University of Mining and Technology-Beijing, Beijing 100083, China

**\*** Correspondence<EMAIL_ADDRESS>
**Abstract:** Gas control in the heading face of a coal roadway is an important and difficult point in coal mining in China. On the basis of analyzing the disadvantages of high gas control cost and long drainage period in the existing mine heading face, a long-distance pre-drainage method of long-distance drilling is proposed to control the gas in the heading face so as to improve the tunneling speed. Applied to the engineering geological conditions of Changcun coal mine, the technology is studied in detail. First, a gas migration model considering permeability changing with time is established, and the model is put into the numerical simulation software to study the variation law of permeability and gas pressure under the conditions of single borehole and multi-borehole drainage. The results show that with the increase of drainage time, the permeability around the borehole increases gradually, the gas pressure decreases gradually, and the permeability at the borehole boundary increases the most, reaching 1.2 times the initial permeability. In the process of multi-borehole drainage, there will be mutual influence between boreholes, but with the increase of borehole spacing, the degree of this influence gradually decreases. Second, according to the results of numerical simulation, a reasonable gas drainage scheme is designed and applied in the field. The field application shows that the technology has a good gas drainage effect, the gas drainage concentration and flow are at a high level for a long time, the drilling cuttings quantity is always lower than the critical value, and the excavation length of roadway increases by more than 50 m per month. These results indicate that this technology is a promising method to realize the safe and rapid excavation of a mine coal roadway.

**Keywords:** directional long borehole; long-distance; heading face; permeability evolution; gas pressure evolution; gas drainage

**1. Introduction**

Gas disaster is one of the main disasters in coal mines in China. The occurrence of gas accidents not only affects the safety production of the mine, but also seriously affects the life safety of workers. Therefore, it is necessary to control the coal seam gas [1–3]. Coal seam gas drainage is an effective method to prevent the occurrence of gas accidents. At the same time, the use of the extracted gas not only increases the total amount of energy, but also is conducive to environmental protection [4,5].

According to the different location of drainage, coal seam gas drainage can be divided into working face drainage and heading face drainage. Among them, the heading face is difficulty of gas control, which often leads to abnormal gas emission accidents due to a poor drainage effect [6,7]. At present, gas drainage in the heading face mainly includes cross-measure borehole gas drainage technology in the floor rock roadway and in-seam borehole gas drainage technology in the heading face [8–10]. Cross-measure borehole gas drainage technology is to extract coal seam gas by drilling in the rock roadway under the coal seam (Figure 1a). This method has the advantages of high safety, good borehole structure stability and long drainage time [11–13]. However, due to the need to dig out a

**Citation:** Hou, Y.; Cui, J.; Liu, R. Study on the Long-Distance Gas Pre-Drainage Technology in the Heading Face by Directional Long Borehole. *Energies* **2022**, *15*, 6304. https://doi.org/10.3390/en15176304

Academic Editor: Sergey Zhironkin

Received: 16 August 2022 Accepted: 25 August 2022 Published: 29 August 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/).

rock roadway, there are some shortcomings, such as high cost, large engineering quantity, long construction period and so on [14]. In-seam borehole gas drainage technology is to extract the coal seam gas by drilling along the coal seam in the heading face (Figure 1b). Compared with the cross-measure borehole technology, this method does not need to excavate a rock roadway, which reduces the construction cost. However, the length of one-time pre-drainage is short, and can only reach 80–100 m. After passing the inspection, the excavation can be carried out, and then the next drainage cycle is carried out, that is "Drilling-Drainage-Inspection-Excavation-Drilling" [15,16], which leads to the slow excavation speed of roadway.

**Figure 1.** Schematic diagram of gas drainage technology in the heading face. (**a**) Cross-measure borehole gas drainage technology. (**b**) In-seam borehole gas drainage technology.

With the development of technology, directional long borehole gas drainage technology has been gradually applied in coal mines. Wang et al. [17] summarized the application of directional long boreholes in China's coal mines, including pre-mining and post-mining drainage, and found that it can effectively improve the gas drainage efficiency. Lu et al. [18] conducted the gas drainage testing of directional long boreholes in Daning Coal Mine, China, which effectively controlled the gas outburst accident risk of the mine. Compared with the ordinary borehole drainage method, the proportion of coal seam drilling is higher, and the drainage effect is better. Wang et al. [19], Li et al. [20] and Hao et al. [21] studied the gas drainage in goaf by directional long boreholes instead of a separate drainage roadway, analyzed and determined the reasonable horizon of directional long boreholes in the roof, and found that this method can effectively solve the problem of gas overrun in the upper corner of the working face, and can replace the roadway-based gas drainage, thereby saving much work.

Based on the advantages of the directional long borehole and the disadvantages of existing ordinary borehole gas drainage in the heading face, a new method of longdistance gas drainage in the heading face by using the directional long borehole is proposed. Through the construction of boreholds by drilling in the completed roadway (not floor rock roadway), the coal seam gas near the pre-excavation roadway is extracted. The construction diagram is shown in Figure 2. Compared with the cross-measure borehole gas drainage technology, this method does not need to excavate a rock roadway, and has the advantages of low cost. Compared with the in-seam borehole gas drainage technology in the heading face, the drainage distance is longer, which increases the length of what was once the excavation roadway, and can extract the coal seam gas for a long time. Taking the return air roadway of working face 2302 in Changcun coal mine as an example, this paper establishes

a gas migration model considering the change of permeability with drainage time, studies the gas migration law and borehole layout of directional long borehole drainage, carries out industrial testing on site, and analyzes the drainage effect.

**Figure 2.** Schematic diagram of gas pre-drainage in the heading face with directional long borehole. (**a**) Construction borehole at the gateway. (**b**) Construction borehole at preparation roadway.

#### **2. Project Summary and Basic Parameter Test**

The Changcun coal mine is located in the Changzhi City, Shanxi Province, China. It is a large modern mine of Shanxi Lu'an environmental protection energy development Co., Ltd. (Changzhi, China). The main mining area is 3# coal seam. The geological structure is mainly fold. The strata strike nearly north-south and incline westward with an inclination of 3–6◦. In the east, it is mainly a tilt structure with near east-west undulation. In the west, there are nearly north-south folds. The geological structure provides a good environment for coal seam gas storage. The gas content of the coal seam is generally 8–10 m3/t, and the gas content is large.

#### *2.1. Project Summary*

2302 working face is located in No. 23 mining area of the mine, mining 3# coal seam; the thickness of the coal seam is generally 4.84–7.32 m, and the average thickness is 6.09 m. The average depth of the coal seam is about 500 m. The roadway layout adopts the "two air intakes and one return" mode; the 2302 belt transportation roadway and auxiliary transportation roadway provide for entry of the air, and the 2302 return air roadway returns the air. A 2302 belt transportation roadway and auxiliary transportation roadway have been excavated, and 2302 return air roadway is the pre-excavation roadway. The mine location and roadway layout of 2302 working face are shown in Figure 3. According to the field measurement data, the gas content of 2302 working face is 8.5 m3/t, the gas pressure is 0.35 MPa, and the permeability is 8.09 × 10−8 m2. The attenuation coefficient of borehole gas flow is 0.1726–0.3025 d−1. According to the difficulty degree of coal seam drainage table (Table 1), the gas drainage in 2302 working face is classified as difficult.

The gas pressure is produced by the thermal movement of free gas. Changcun coal mine has high gas content and low gas pressure, which indicates that the coal seam has strong adsorption capacity, and the adsorption content of coal seam gas accounts for more. Gas desorption is a slow process. Research shows that under natural conditions, when the coal particle size is 1 cm, the time required for desorption of 90% of the gas is 15 years [22]. Therefore, for the coal seam with strong adsorption capacity, gas drainage should be carried out for a long time even under negative pressure.

**Figure 3.** Location of Changcun Coal Mine and roadway layout of 2302 working face.

**Table 1.** Difficulty degree of coal seam gas drainage.

#### *2.2. Basic Parameter Test*

The coal samples were taken on site, and the coal samples were processed into standard samples and pulverized with different particle sizes, respectively. The coal samples are used to test the mechanical parameters, industrial analysis and gas basic parameters of coal. The test results are shown in Tables 2 and 3.

**Table 2.** Mechanical parameters of coal samples.

**Table 3.** Industrial analysis and gas basic parameters.

#### **3. Gas Migration Equation in Coal Seam**

#### *3.1. Control Equation of Coal Seam Permeability Considering Gas Pressure Variation and Gas Adsorption and Desorption*

The research shows that coal seam permeability *k* can be expressed as an exponential function related to effective stress [23]:

$$k = k\_0 \exp\left(-3\mathcal{C}\_f \Delta \sigma\_\varepsilon\right) \tag{1}$$

where *k0* is the initial permeability, m2; *Cf* represents cleat volume compressibility, MPa−1; Δ*σe* is effective stress variation, MPa.

The cleat volume compressibility *Cf* can be expressed by the following formula [24]:

$$\begin{cases} \mathsf{C}\_{f} = \frac{1}{\mathsf{K}\_{p}}\\ \mathsf{K}\_{p} = \eta \cdot \mathsf{K} = \eta \cdot \frac{E}{3(1 - 2\nu)} \end{cases} \tag{2}$$

where *Kp* is pore bulk modulus, GPa; *η* is porosity of coal seam, %; *K* is elastic modulus of coal matrix, GPa; *E* is elastic modulus of coal, GPa; *υ* is Poisson's ratio of coal.

There are three changes in the process of drilling and gas drainage. Firstly, the coal seam is disturbed in the process of drilling, and the stress around the borehole redistributes, resulting in the change of effective stress. Secondly, the gas in a free state is extracted under negative pressure, and the decrease of gas pressure leads to the increase of effective stress, resulting in the compression of coal pores and the decrease of gas flow channels. Thirdly, the decrease of gas pressure promotes gas desorption and coal matrix shrinkage, resulting in the increase of coal fracture channels. Therefore, the variation of effective stress can be expressed as three parts:

$$
\Delta \sigma\_{\mathfrak{c}} = \Delta \sigma\_{\mathfrak{s}} + \Delta \sigma\_{\mathfrak{p}} + \Delta \sigma\_{\mathfrak{x}} \tag{3}
$$

where Δ*σs* is effective stress variation caused by stress redistribution, MPa; Δ*σp* is effective stress variation caused by gas pressure drop, MPa; Δ*σx* is effective stress variation caused by gas desorption, MPa.

Coal is a kind of soft elastic-plastic body. After stress redistribution, the tangential stress around the borehole will be higher than the strength of coal, so that the coal will be destroyed and a plastic zone and elastic zone will appear. Assuming that the coal is in the limit equilibrium state in the plastic zone, the tangential stress around the borehole can be described by the following formula [25,26]:

$$\Delta\sigma\_{t} = \begin{cases} c \cdot \cot\eta \cdot \left[ \left( \frac{1 + \sin\eta}{1 - \sin\eta} \right) \cdot \left( \frac{2x}{R\_{0}} + 1 \right)^{\frac{2 \sin\eta}{1 - \sin\eta}} - 1 \right], x \le H\\ \sigma\_{0} \cdot \left[ 1 + \frac{4H^{2}}{\left( 2x + R\_{0} \right)^{2}} \right] - \frac{4H^{2}}{\left( 2x + R\_{0} \right)^{2}} \cdot c \cdot \cot\eta \left[ \left( \frac{2H}{R\_{0}} \right)^{\frac{2 \sin\eta}{1 - \sin\eta}} - 1 \right], x > H \end{cases} \tag{4}$$

where *σ0* is initial stress, MPa; *c* is cohesion of coal, MPa; *ϕ* is internal friction angle, ◦; *x* is distance from coal body to borehole wall, m; *R0* is borehole diameter, m; *H* is distance from plastic zone boundary to borehole wall, m. *H* can be expressed as follows:

$$H = \frac{R\_0}{2} \left\{ \left[ \frac{\sigma\_0 \cdot (1 - \sin \eta)}{c \cdot \cot \eta} + 1 \right]^{\frac{1 - \sin \eta}{2 \cdot \sin \eta}} - 1 \right\} \tag{5}$$

According to the coal seam buried depth of 500 m, the initial stress *σ0* is 12.5 MPa, the diameter of long hole is 113 mm. The tangential stress as shown in Figure 4 can be calculated by substituting the cohesion *c* and internal friction angle *ϕ* measured in the laboratory and original stress and borehole diameter into Equations (3) and (4). If the position where the stress changes by 5% is taken as the influence boundary of the borehole, it can be seen from Figure 4 that the radius of the influence range of the directional long

borehole is 0.313 m, which is smaller than the thickness of the coal seam. Therefore, the influence of the stress change around the borehole on the permeability of the coal seam can be ignored, and the change of the effective stress can be expressed as follows:

$$
\Delta \sigma\_{\mathfrak{e}} = \Delta \sigma\_{\mathfrak{p}} + \Delta \sigma\_{\mathfrak{x}} \tag{6}
$$

**Figure 4.** Distribution of tangential stress around borehole.

The effective stress variation caused by gas pressure drop can be expressed by the following formula:

$$
\Delta \sigma\_p = P\_0 - P \tag{7}
$$

where *P0* is initial gas pressure of coal seam, MPa; *P* is gas pressure of coal seam, MPa.

The effective stress variation caused by gas adsorption and desorption can be expressed by the following formula:

$$
\Delta \sigma\_x = \mathbf{K} \cdot \Delta \varepsilon\_x \tag{8}
$$

where *εx* is the coal bulk strain caused by gas adsorption and desorption, which satisfies the Langmuir equation. It can be expressed by the following formula [24]:

$$
\varepsilon\_x = \frac{\varepsilon\_l \cdot P}{P + P\_l} \tag{9}
$$

where *εl* is Langmuir volume strain constant; *Pl* is Langmuir pressure constant.

The control equation of coal seam permeability considering gas pressure variation and gas adsorption and desorption can be obtained by simultaneous Formulas (1), (6)–(9):

$$k = k\_0 \exp\left\{-3\mathcal{C}\_f (P\_0 - P) \cdot \left[1 - \frac{E}{3 \cdot (1 - 2\nu)} \cdot \frac{\varepsilon\_I P\_l}{(P + P\_l)(P\_0 + P\_l)}\right]\right\} \tag{10}$$

#### *3.2. Control Equation of Gas Flow*

Coal seams are porous media, in which gas seepage conforms to the mass conservation equation [27]:

$$\frac{\partial X}{\partial t} + \nabla(\rho V) = 0\tag{11}$$

where *X* is the gas content in unit volume coal, kg/m3; *t* is the time variable, s; *ρ* is the gas density in the coal seam, kg/m3; *V* is the gas seepage velocity, m/s.

There are two forms of gas in a coal seam, namely free state and adsorption state. Therefore, the gas content in a coal seam includes free gas content and adsorption gas content, which can be expressed by a gas state equation and Langmuir equation, respectively:

$$\begin{cases} X\_1 = \eta \rho \\ X\_2 = \frac{abp}{1+bp} \cdot \frac{100-A-W}{100} \cdot \frac{1}{1+0.31W} \\ X\_3 = X\_1 + X\_2 \end{cases} \tag{12}$$

where *X1* is the free gas content, kg/m3; *X2* is the content of adsorbed gas, kg/m3; *a* is the maximum gas adsorption constant per unit mass of coal, m3/kg; *b* is the adsorption constant of coal, MPa−1; *A* is ash content of coal, %; *W* is moisture content of coal, %.

Assuming that the gas is an ideal gas, the gas density in the coal seam is as follows:

$$
\rho = \frac{M\_\% P}{RT} \tag{13}
$$

where *Mg* is molecular weight of gas, 16 g/mol; *R* is the ideal gas constant, 8.314 J/(mol·K); *T* is the absolute temperature, K.

It is assumed that the flow of gas in the coal seam conforms to Darcy's law [8,27]:

$$V = -\frac{k}{\mu}\nabla P\tag{14}$$

where *k* is the permeability of coal seam, m2; *μ* is the dynamic viscosity of gas, Pa·s.

The control equation of gas flow can be obtained by simultaneous Formulas (11)–(14):

$$
\left[\frac{M\_{\ $}\eta}{RT} + \frac{ab}{\left(1+bp\right)^{2}} \cdot \frac{100-A-W}{100} \cdot \frac{1}{1+0.31W}\right] \frac{\partial P}{\partial t} - k \cdot \frac{M\_{\$ }}{\mu RT} \nabla \left(P \nabla P\right) = 0 \tag{15}
$$

It can be seen from Formulas (10) and (15) that gas drainage and permeability are influenced by each other. With gas drainage, gas pressure decreases and gas desorption occurs, which affects the permeability. Accordingly, the change of permeability will affect the results of gas extraction. The relationship between gas drainage and permeability is shown in Figure 5.

**Figure 5.** Relationship between gas drainage and permeability.

#### **4. Variation of Permeability and Gas Pressure around Borehole**

#### *4.1. Single Borehole Drainage*

4.1.1. Geometric Model and Parameter Setting

Since the directional borehole is parallel to the pre-excavation roadway after construction in the design area, the geological conditions of the area through which the borehole passes have little change, so the numerical calculation model can be simplified into a local three-dimensional model. It is assumed that gas is the only flowing gas in the coal seam and the coal seam is isotropic. According to the coal seam thickness and gas parameters of Changcun coal mine, the COMSOL numerical simulation software is used to establish the coal seam gas migration model. In order to reduce the influence of boundary effect, the length of the model is 70 m, the height is 6 m, and the diameter of the borehole is 113 mm. The model diagram is shown in Figure 6. At the same time, a 15 m long monitoring line is arranged from the borehole center to the depth of coal seam, and four monitoring points

are arranged at 1 m, 3 m, 5 m and 7 m away from the borehole center to monitor the change of gas parameters. There are no flow boundary conditions around the model, and constant pressure boundary conditions around the borehole simulate the drainage pressure, and the drainage pressure is 20 kPa. The PDE module of numerical simulation software is used to substitute Formulas (10) and (15) into the calculation, and the calculation parameters are listed in Table 4.

**Figure 6.** Numerical model.

**Table 4.** Numerical calculation parameters.

4.1.2. Mesh Independence Test of Model

In order to study the influence of model mesh on simulation results, it is necessary to test the mesh independence. The mesh is divided into normal, fine and finer, as shown in Figure 7. According to the theoretical formula, the permeability and gas pressure distribution law of drainage time is 10 d are simulated, as shown in Figure 8.

**Figure 7.** Mesh division types: (**I**) normal; (**II**) fine; (**III**) finer.

**Figure 8.** Cloud chart of permeability and gas pressure distribution of different mesh types at 10 d of drainage: (**I**–**III**) are the permeability distribution of mesh type from normal to finer; (**IV**–**VI**) are the gas pressure distribution of mesh type from normal to finer.

It can be seen from the figures that the mesh division type is from normal to finer, the permeability and gas pressure distribution around the borehole are almost unchanged. In order to more clearly analyze the influence of mesh division type on simulation results, the data measured by the monitoring line are plotted into a curve, as shown in Figure 9.

**Figure 9.** Permeability and gas pressure distribution curves of different mesh types at 10 d of drainage: (**a**) Permeability distribution curve. (**b**) Gas pressure distribution curve.

It can be seen from the figures that the three different mesh types have obtained the same observation curve, so it can be determined that the mesh type has no effect on the simulation results of single borehole drainage. In order to save the time of numerical simulation, the mesh type is selected as normal mode in the simulation of single borehole drainage.

#### 4.1.3. Permeability Variation Law

According to the theoretical formula, the numerical simulation software is used to calculate the distribution characteristics of permeability around the borehole when the drainage time is 50 d, 100 d, 150 d, 200 d, 250 d and 300 d respectively, so as to observe the variation law of permeability around the boreholes. The curves from the monitoring data of monitoring line and monitoring points are shown in the figures below.

Figures 10 and 11 show the change of coal permeability around the borehole at different drainage times. It can be seen from the figure that the permeability around the borehole is greater than the initial permeability. The further away from the borehole boundary, the smaller the permeability, and the closer to the initial permeability. At the same time, it can also be seen that with the increase of drainage time, the range of permeability around the borehole gradually increases. The monitoring data of monitoring line and monitoring points form curves, as shown in Figure 12. It can be seen from Figure 12a that the permeability of the borehole boundary increases by 1.2 times. With the increase of drainage time, the coal permeability also increases, but the permeability at the borehole boundary remains unchanged. It can be seen from Figure 12b that no matter how far the distance from the borehole is, with the increase of drainage time, the permeability gradually increases from the initial permeability, but the increase in the amplitude of permeability decreases with the increase of time. The closer to the borehole, the faster the permeability increases with time in the early stages.

**Figure 10.** Three-dimensional cloud chart of coal permeability distribution around boreholes at different drainage times: (**I**) 50 d; (**II**) 100 d; (**III**) 150 d; (**IV**) 200 d; (**V**) 250 d; (**VI**) 300 d.

**Figure 11.** Cloud chart of coal permeability distribution around borehole at different drainage times: (**I)** 50 d; (**II**) 100 d; (**III**) 150 d; (**IV**) 200 d; (**V**) 250 d; (**VI**) 300 d.

#### 4.1.4. Variation Law of Gas Pressure

Based on the theoretical formula, the numerical simulation software is used to calculate the distribution characteristics of gas pressure around the borehole when the drainage time is 50 d, 100 d, 150 d, 200 d, 250 d and 300 d, respectively, so as to observe the variation law of gas pressure around the borehole and to plot the curves from the monitoring data of monitoring line and monitoring points, as shown in the figure below.

Figures 13 and 14 show the cloud diagram of gas pressure distribution around the borehole at different drainage times. It can be seen from the figure that the gas pressure increases gradually from the borehole boundary to the depth of the coal seam. With the increase of drainage time, the range of gas pressure decrease gradually increases. The monitoring data of monitoring line and monitoring points are plotted into curves, as shown in Figure 15. It can be seen from the figure that with the increase of drainage time, the gas pressure decreases more. The closer to the borehole, the greater the gas pressure drop rate in the early stage of drainage.

**Figure 13.** Three-dimensional cloud chart of gas pressure distribution around boreholes at different drainage times: (**I**) 50 d; (**II**) 100 d; (**III**) 150 d; (**IV**) 200 d; (**V**) 250 d; (**VI**) 300 d.

**Figure 14.** Cloud chart of gas pressure distribution around borehole at different drainage times: (**I**) 10 d; (**II**) 50 d; (**III**) 100 d; (**IV**) 150 d; (**V**) 200 d; (**VI**) 300 d.

**Figure 15.** Monitoring data of gas pressure of monitoring line and monitoring points. (**a**) Variation law of gas pressure with monitoring line at different drainage times. (**b**) Variation of gas pressure with time at different distances from borehole.

#### *4.2. Multi-Borehole Drainage*

#### 4.2.1. Geometric Model Setting

Based on the single borehole drainage model, a multi-borehole drainage model has been established, as shown in Figure 16. Except for increasing the number of boreholes, other calculation parameters of the model have not changed. Because it is estimated that the drainage time of boreholes is more than 200 d, the gas drainage effect of boreholes with different spacing is studied based on the drainage time of 200 d. The spacing of boreholes is 4 m, 6 m, 8 m, 10 m, 12 m and 14 m, respectively. In the process of borehole drainage, the whole coal seam in the area covered by the borehole should be drained to meet the standard. Therefore, a 60 m long monitoring line is arranged at the top of the coal seam, that is, in the middle of the upper boundary of the model, to monitor the simulation results. At the same time, a monitoring point is arranged at the midpoint of the two boreholes, to monitor the simulation results in the middle of the boreholes.

**Figure 16.** Numerical model.

#### 4.2.2. Mesh Independence Test of Model

In order to study the influence of model mesh on simulation results, it is necessary to test the mesh independence. The mesh is divided into normal, fine and finer, as shown in Figure 17. According to the theoretical formula, the permeability and gas pressure distribution law of the simulation borehole spacing of 6 m and the drainage of 20 d are shown in Figure 18.

**Figure 17.** Mesh division types: (**I**) normal; (**II**) fine; (**III**) finer.

**Figure 18.** Cloud chart of permeability and gas pressure distribution of different mesh types at 10 d of drainage: (**I**–**III**) are the permeability distribution of mesh type from normal to finer; (**IV**–**VI**) are the gas pressure distribution of mesh type from normal to finer.

It can be seen from the figures that the mesh type varies from normal to finer, and there is no obvious change in the permeability and gas pressure distribution around the borehole. In order to more clearly analyze the influence of mesh type on simulation results, the data measured by the monitoring line are plotted into a curve, as shown in Figure 19.

**Figure 19.** Permeability and gas pressure distribution curves of different mesh types at 10 d of drainage: (**a**) Permeability distribution curve. (**b**) Gas pressure distribution curve.

It can be seen from the figures that the simulation results of three different mesh types are different between boreholes under the multi-borehole drainage. The mesh type is from normal to finer, the permeability between boreholes decreases, and the gas pressure increases. The maximum error is 0.093% from normal to fine mesh types, and 0.049% from fine to finer mesh types. The error of simulation results between different mesh types is small. Since it is necessary to determine the borehole spacing when simulating the multi-borehole extraction, in order to ensure the extraction effect and consider the error, the finer mesh type is selected for simulation.

#### 4.2.3. Permeability Variation Law

According to the theoretical formula, the numerical simulation software is used to calculate the distribution characteristics of permeability around the borehole under different borehole spacing when the drainage time is 200 d, so as to observe the variation law of permeability around the borehole., and to draw the curves from the monitoring data of monitoring line and monitoring point, as shown in the figures below.

Figures 20 and 21 shows the distribution of permeability around differently spaced boreholes. It can be seen from the figure that with the increase of borehole spacing, the range of permeability gradually increases, but the permeability between boreholes decreases with the increase of borehole spacing. The data monitored by the monitoring line and the monitoring point are drawn into curves, as shown in Figure 22. It can be seen from Figure 22a that with the increase of borehole spacing, the permeability between boreholes gradually changes from greater than at the top of the boreholes to less than at the top of the boreholes, indicating that the drainage boreholes will interact with each other, resulting in the increase of permeability between boreholes; however, with the increase of borehole spacing, the degree of interaction between boreholes gradually decreases. It can be seen from Figure 22b that when the drainage time is 200 d, and the spacing between boreholes is 6 m, 10 m and 14 m, that is, the distance between measuring point and boreholes is 3 m, 5 m and 7 m, the permeability ratios are 1.112, 1.100 and 1.088, respectively, and the permeability gradually decreases. Compared with 1.067, 1.056 and 1.046 at the same time and location of single borehole drainage, the increase is 4.22%, 4.17% and 4.02%, respectively, which can be concluded as above.

**Figure 20.** Three-dimensional cloud chart of permeability distribution around boreholes with different spacing: (**I)** 4 m; (**II**) 6 m; (**III**) 8 m; (**IV**) 10 m; (**V**) 12 m; (**VI**) 14 m.

**Figure 21.** Cloud chart of permeability distribution around boreholes with different spacing: (**I**) 4 m; (**II**) 6 m; (**III**) 8 m; (**IV**) 10 m; (**V**) 12 m; (**VI**) 14 m.

**Figure 22.** Monitoring data of permeability of monitoring line and monitoring point. (**a**) Variation law of permeability with monitoring line at different borehole spacing. (**b**) The law of permeability variation with time with two boreholes.

#### 4.2.4. Gas Pressure Variation Law

According to the theoretical formula, the distribution characteristics of gas pressure around the boreholes are calculated respectively by using numerical simulation software under different borehole spacing when drainage time is 200 d. The gas pressure variation law around the boreholes is observed, and the data monitored by the monitoring line and the monitoring point are plotted into curves, as shown in the following figure.

Figures 23 and 24 show the distribution of gas pressure around different boreholes according to their spacing. It can be seen from the figure that with the increase of borehole spacing, the gas pressure reduction range gradually increases, but the gas pressure between boreholes increases with the increase of borehole spacing. The data monitored by the monitoring line and the monitoring point are drawn into curves, as shown in Figure 25. It can be seen from Figure 25a that with the increase of borehole spacing, the gas pressure between boreholes gradually changes from less than the top of boreholes to more than at the top of boreholes, indicating that the drainage boreholes will interact with each other, resulting in the decrease of gas pressure between boreholes; with the increase of borehole spacing, the degree of interaction between boreholes gradually decreases. It can be seen from Figure 25b that when the drainage time is 200 d, and the distance between boreholes is 6 m, 10 m and 14 m, that is, the distance between measuring point and boreholes is 3 m, 5 m and 7 m, the gas pressure is 0.133 MPa, 0.154 MPa and 0.175 MPa respectively, and the gas pressure increases gradually. Compared with 0.213 MPa, 0.233 MPa and 0.253 MPa at the same time and position of single borehole drainage, the values decreased by 36.49%, 35.32% and 31.10%, respectively, similar to the findings above.

**Figure 23.** Three-dimensional cloud chart of gas pressure distribution around boreholes with different spacing: (**I**) 4 m; (**II**) 6 m; (**III**) 8 m; (**IV**) 10 m; (**V**) 12 m; (**VI**) 14 m.

**Figure 24.** Cloud chart of gas pressure distribution around boreholes with different borehole spacing: (**I**) 4 m; (**II**) 6 m; (**III**) 8 m; (**IV**) 10 m; (**V**) 12 m; (**VI**) 14 m.

**Figure 25.** Monitoring data of gas pressure of monitoring line and monitoring point. (**a**) Variation law of gas pressure with monitoring line at different borehole spacing. (**b**) The law of gas pressure variation with time with two boreholes.

For a high gas coal seam with gas pressure lower than 0.74 MPa, the effective range of borehole drainage is often defined by relative pressure, that is, the boundary of the effective drainage range is 51% decrease in gas pressure [28]. Therefore, the effective drainage boundary of gas pressure is 0.172 MPa. It can be seen from Figure 25, when the spacing between boreholes is 14 m, the gas pressure between boreholes will be greater than 0.172 MPa, that is, the drainage between boreholes will not meet the standard. In order to more clearly analyze the effective drainage range between boreholes, the range of gas pressure lower than 0.172 MPa at different borehole spacing is drawn, as shown in Figure 26. It can be seen from the figure that with the increase of borehole spacing, the range of gas pressure lower than 0.172 MPa also gradually increases (the red part in the figure). However, when the borehole spacing is 14 m, there is a non-standard drainage zone between boreholes.

**Figure 26.** Range of gas pressure lower than 0.172 MPa at different borehole spacing: (**I)** 4 m; (**II**) 6 m; (**III**) 8 m; (**IV**) 10 m; (**V**) 12 m; (**VI**) 14 m.

#### **5. Field Application**

#### *5.1. Method Statement*

According to the above numerical simulation results, the final borehole spacing is determined to be 12 m. In order to control the coal seam within 15 m on both sides of the pre-excavation roadway, four boreholes need to be arranged. According to the length of roadway, four drilling fields with the size of 8 m × 5 m × 5 m are designed and constructed in the 2302 auxiliary transportation roadway. The distance between drilling fields is 400 m. There are eight boreholes designed and constructed in 1# drilling field and four boreholes in 2–4# drilling fields. The borehole layout plan is shown in Figure 27. The sealing method of "two plugging, one injection and one row" is adopted, and the sealing depth of the borehole is 20 m to ensure tight sealing without air leakage. And each borehole in the drilling fields of 1–4# is equipped with a concentration measuring port and orifice flowmeter, which is convenient for real-time monitoring of gas concentration and gas drainage flow.

**Figure 27.** Borehole layout plan.

The drilling equipment used in the construction is shown in Figure 28, mainly including operation console, guidance system console, water pump, directional drill, rotary unit, front gripper, crawler, etc. The operation console controls the hydraulic system and water system of the drilling rig. The guidance system console displays the collected data of the unit in the borehole on the screen. The water pump is used to provide water for the downhole motor. A directional drill is used to control drilling direction. The rotary unit is used to rotate the drill pipe. The front gripper is used to push and dismantle the drill pipe, and also to guide the drill pipe. The crawler can enable the drilling rig to walk freely in the underground roadway.

**Figure 28.** Construction equipment. (**a**) Overall drawing of drilling rig. (**b**) Operation console. (**c**) Front gripper. (**d**) Water pump. (**e**) Crawler. (**f**) Guidance system console. (**g**) Rotating unit. (**h**) Directional drill.

#### *5.2. Effect Analysis*

After the completion of drilling construction, record the gas drainage flow and gas concentration after different drainage times. After the completion of drainage, test the effects on the coal seam within the scope of the pre-excavation roadway. At the same time, record the excavation speed of the heading face during the excavation, and analyze the drainage effect.

Figure 29 shows the data record during borehole drainage and 2302 return air roadway excavation. Figure 29a,b show the gas concentration and gas drainage flow of a single borehole in four drilling fields in the first 50 d. It can be seen from the figure that the gas concentration and gas drainage flow change little in the first 35 d, with gas concentration higher than 70% and gas drainage flow higher than 0.38 m3/min. The gas concentration is still higher than 65% and the gas drainage flow is still higher than 0.33 m3/min after 50 d of drainage, with little decrease and good drainage effect. Figure 29c shows the measurement of the drilling cutting quantity in the coal seam before the roadway excavation (after each excavation distance of the roadway, it is necessary to predict the outburst in front of the heading face, and the drilling cutting quantity is one of the important indicators). It can be seen from the figure that the drilling cutting quantity is 3.1–4.7 kg/m, which is lower than the critical value of outburst of 6.0 kg/m, and there is no danger of outburst, i.e., there is no gas dynamic phenomenon in the process of roadway excavation. Figure 29d shows the excavation length per month. It can be seen from the figure that the excavation length of the long-distance gas pre-drainage method is more than 80 m per month; compared with the in-seam ordinary borehole gas drainage technology (Figure 1b), the excavation length can be more than 50 m per month, and the excavation speed is greatly increased.

#### **6. Discussion**

This paper introduces and studies the technology of long-distance advance predrainage of gas in the heading face, and has achieved good application results in the field. This technology can enable drilling of boreholes to control the gas in the area of the roadway to be excavated by use of the roadway that has been excavated. This technology is applicable to the following situations: (1) The mine is a mine with high gas or coal and gas outburst. (2) The mine needs long-term gas drainage to solve the gas problem. (3) There is a shortage of mining and replacement, which affects the efficient production of mine. This technology can save a large amount of cost compared with the technology of driving bottom drainage, and can save a large amount of time compared with the technology of digging while performing drainage, which has great technical advantages.

#### **7. Conclusions**

In view of the shortcomings of the existing commonly used gas drainage methods in the heading face, this paper proposes to adopt the long-distance gas pre-drainage technology in the heading face by directional long borehole. The main conclusions are as follows:

**Author Contributions:** Y.H.: Supervision, funding acquisition, writing the review, and editing. J.C.: Writing the original draft, methodology, and software. R.L.: Writing—review and editing. All authors have read and agreed to the published version of the manuscript.

**Funding:** This work was financially supported by the National Natural Science Foundation of China (Grant No. 52074296, 52004286), the China Postdoctoral Science Foundation (Grant No. 2020T130701, 2019M650895), the State Key R&D Plan (2017YFC0804303) of the Ministry of Science and Technology of China.

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

**Informed Consent Statement:** Not applicable.

**Data Availability Statement:** Available from corresponding author.

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

#### **References**


**7. References**

**16**

Fan Yang

*United States*

*Qualcomm MEMS Technologies, Inc.*

**Large Area a-Si/µc-Si Thin Film Solar Cells**

Providing a sustainable and environment friendly energy source, photovoltaic (PV) power is becoming ever-increasingly important, as it decreases the nation's reliance on fossil-fuel generated electricity. Though widely regarded as a clean and renewable energy source, large scale deployment of PV is still impeded by the fact that the cost of PV energy is generally higher compared to grid electricity. Current development of PV technology is focused on two aspects: 1) improving the efficiency of PV modules and systems and 2) lowering the cost of delivered electricity through decreasing the manufacturing and installation cost. The merit of commercial solar cells aiming at terrestrial application is justified by the cost of unit PV power generation, dollar per watt (\$/Wp), where Wp stands

Since the first practical PV cell grown on Si wafer at the Bell Laboratory in 1954, PV technology has been developed for more than five decades and evolved three "generations" based on different PV materials. The first generation of solar cells use crystalline materials, where the cost of the bulk materials has hit the point that further cost reduction is very difficult (Green 2007). In contrast, the second generation cells use thin film materials, where the required amount of materials is merely a few percent of that of bulk materials, significantly reducing the fabrication cost of this type of cells. The emerging, third generation of PV technology applies new materials and novel device concepts aiming at even higher efficiency and lower cost. At this moment, the commercial PV market is dominated by the first and second generation PV modules, and the third generation cells are still under lab research. As shown in Fig. 1, the efficiency of thin film PV system has improved from ~4 % in 1995 to >11 % in 2010, and will keep increasing to ~12% by 2020, a three-fold improvement compared to the system efficiency back in 1995. During the same timeframe, the cost of thin film PV system drops from ~4 \$/Wp to ~0.5 \$/Wp. Crystalline PV systems, though with higher efficiencies, have higher cost, i.e. 2.5 times the cost of thin film PV system. From the cost and material supply point of view, thin film solar cells will have a long-term development and gradually take more market share from the crystalline

Many thin film materials can be used for PV cells, e.g., Si, CdTe, CIGS or the emerging organic/polymeric materials. Comparing to other materials, thin film Si, including

1. The PV active Si is the most abundant solid state element on the earth's shell, allowing

amorphous Si (a-Si) and microcrystalline Si (µc-Si), have the following characters:

for practically unlimited production of Si cells.

**1. Introduction**

cells.

for the peak power generated by the cells.

*of the 21st European Photovoltaic Solar Energy Conference*, pp. 625-628, Dresden, Germany.


**5.3 Calculation methods**

The calculation method for total volumetric mass transfer coefficient (kLa) follows the ASCE Standard for Measurement of Oxygen Transfer in Clean Water [18]. Then the obtained kLat is calibrated to a standard reference temperature of 20 °C by using Equation (5) [16, 18],

$$k\_L a(20) = 1/\gamma \times k\_L a(T) \times 1.024^{(20-T)}\tag{5}$$

where kLa (20) and kLa (T) (hr-1) are kLa at 20 °C and the actual water temperature of T °C, respectively, γ is the activity coefficient of salt concentration and γ = 8.8×10-6×Cz+1, Cz (mg l-1) represents the concentration of sodium sulfite solution.

The performance of oxygen mass transfer is assessed in terms of oxygen mass transfer efficiency, which is calculated based upon Equation (6) [19],

$$E\_A \text{(20)} = \frac{DO\_S \text{(20)} \cdot k\_L a \text{(20)} \cdot V \times 10^{-3}}{G\_S \text{(20)} \cdot \rho \cdot O\_w} \times 100 \tag{6}$$

where EA (20) refers to oxygen transfer efficiency at 20 °C, (20) *Lk a* (l/hr) is kLa at 20 °C, (20) *DOS* (mg/L) is liquid-phase saturated DO concentration at 20 °C, (20) *GS* (m3/hr) is air flow rate at 20 °C and 1 atm, *V* (m3) is effective capacity of water tank, ρ is air density at 20°C and 1 atm (ρ = 1.204 kg/m3), and Ow (-) is oxygen content in air (Ow = 0.233 O2-kg/air-kg).

Improvement of Oxygen Transfer Efficiency in

0

**6.1 Introduction**

Fig. 32.

3

E

A(20) (%)

6

9

exterior of the resulting liquid film with the air.

**6. Analysis of the performance of LFFA**

**6.2 Experimental conditions and methods**

experiment. The performance of the LFFA is thus evaluated.

Diffused Aeration Systems Using Liquid-Film-Forming Apparatus 365

As indicated above, water body can form a liquid film in the atmosphere. The LFFA can remarkably improve oxygen supply efficiency by simultaneously contacting the interior and

Liquid-film-type aeration Conventional aeration

4 6 8 10 12 14

Fig. 31. The comparison data of (20) *EA* of liquid film and conventional aeration systems

In order to evaluate the oxygen transfer performance of the LFFA, in this section, the overall oxygen transfer process is divided into two-step oxygen transfer. That is, in the conventional aeration system, oxygen transport process is split into bubble transfer and surface transfer. In contrast, in the liquid film aeration system, it is divided into bubble transfer and liquid film transfer. Oxygen transfer efficiency of every step is separately derived from the

By means of a recycling liquid film apparatus (4 cm in each individual airlift pipe diameter, 1 cm in effective height, 12.56 cm2 in cross-sectional area of the airlift part), liquid film aeration tests in a 53 cm deep, 80 L water tank with a surface area of 1510 cm2 are carried out applying an air flow rate of 6 L/min. As a control, under the otherwise identical experimental conditions, the conventional aeration test with air and diffused test with nitrogen are both investigated in this study. Their experimental apparatuses are shown in

Air flow rate (L/min)

## **5.4 Results and discussion**

As shown in Table 7 and Fig. 30, (20) *Lk a* of liquid film aeration system displays an evolving trend same as that of conventional aeration system. That is, in the air flow rate range of 6-12 L/min, it gradually increases with increasing air flow rate. Beyond the upper limit, it will decrease slowly. However, as shown in Table 7 and Fig. 31. The (20) *EA* values of both liquid film aeration system and conventional aeration system decrease with increasing air flow rate.

Table 7. Comparison of liquid film aeration system with conventional aeration system

Fig. 30. The comparative data of (20) *Lk a* of liquid film and conventional aeration systems

As compared with (20) *Lk a* and (20) *EA* of both aeration systems, at the air flow rate ranging from 6 to 12 L/min, the aeration efficiency of liquid film aeration system increases by 6.3- 14.3%. Particularly at the air flow rate of 6 L/min, (20) *EA* is still up to 6.64% even utilizing a very shallow aeration depth of 0.5 m. The efficiency of liquid-film aeration is enhanced by 14.3% relative to the conventional aeration system under the identical conditions.

As indicated above, water body can form a liquid film in the atmosphere. The LFFA can remarkably improve oxygen supply efficiency by simultaneously contacting the interior and exterior of the resulting liquid film with the air.

Air flow rate (L/min)

Fig. 31. The comparison data of (20) *EA* of liquid film and conventional aeration systems