Introduction

A horizontally reversible plough (HRP, Xinjiang, P. R. China), commonly-used around China, is a tillage tool with high performances; see Fig. 11. Compared to conventional ploughs, the higher performances of the HRP present the following two aspects2. First, and more common, is excavating soil by moving forward with high speed (V, km h−1). Second, with a unique advantage, is cutting soil by continuous shifting (n, rad s−1). The shifting, as denoted by the symbol ‘n’ in Fig. 1, actually indicates a tillage behaviour that the plough-body can be propelled, by hydraulic cylinder (HC), to alternatively excavate soil in-between the two different tillage positions, e.g., A and B in Fig. 1. Note that the shifting and moving behaviours of the plough-body are simultaneously implemented in the practical tillage operations. We refer the readers to the work of1,3 for details of HRP.

Fig. 1
figure 1

Tillage operations of the HRP with high-speed. The yellow dashed curves on the plough-surface denote the hypothesized flowing trajectories of soil.

However, the field trials have found that the continuous and alternative shifting behaviour during the high-speed tillage of the HRP is prone to produce severe friction between soil and plough-surface, in which the abrasion with hard soil particles is predominant at the share-point Section2. The intense wear rate not only affects the tillage quality, but also has serious consequences for the tool life4. Figure 2 schematically illustrates the worn lands, measured by a scanning electronic microscope (SEM), at the share-point section after the HRP finished tilling 50 ha of the sandy loam soil.

Fig. 2
figure 2

(A) Ploughshare of HRP, (B) worn share-point section, (C) SEM micrographs of (B), (D) denotes the material properties of the ploughshare, and (E) indicates the soil properties. The yellow dotted curves indicate the hypothesized trajectories of soil. The tool speed is 6 km h−1 and the operational depth is 0.27 m.

In order to improve the HRP performances, a lot of researchers have developed experimental, analytical, and numerical methods to investigate the interaction between soil and plough-surface5. The primary approaches include the traditional approaches, e.g., empirical/semi-empirical solutions6, dimensional technique7, and the commonly-used approaches, e.g., finite element analysis (FEA)8, discrete element method (DEM)9 as well as computational fluid dynamics (CFD)10,11. More detailed comparisons of these approaches are available in12.

Based on the aforementioned comparisons, we have ever applied the CFD approach for the flowing interaction between soil and ploughshare of the HRP in sandy loam soil at three different tillage speeds of 5, 6, 7 km h−1 and three different operational depths of 0.27, 0.315, 0.360 m13. The calculated is in very good agreement with the measured. However, further measurements have found that severe abrasion/attrition wear with white materials appeared at the share-point section; see Fig. 2C. Particularly, the energy dispersive spectroscopy (EDS) based measurements have shown that the white materials contained some sort of oxide, e.g., Fe2O3. Although this mechanical fatigue has been demonstrated to be caused by the flowing interaction between soil and tool, we are not sure whether the emerging white materials are involved in the thermal effects, caused by the high-speed shifting tillage operations. To the authors’ knowledge, no research publication is available on the investigation of the thermal effect on the performances of high-speed moldboard plough, nor is the underlying physical mechanism clear4. This research is aimed to fill this gap, particularly the thermal effect on the share-point wear in sandy loam soil in the high-speed shifting tillage operations of the HRP.

This paper extended our previous work to evaluate the thermal effect by using a combined multi-body dynamics analysis (MDA) and fluid–solid-thermal simulation. The calculated tillage stresses, tillage temperatures at the share-point section were compared with the published worn lands at the same tillage conditions (such as tillage speed and depth). The tillage scenario of the HRP addressed in this study is schematically illustrated in Fig. 1.

Materials and methods

Geometric model

The three-dimensional (3D) model of the HRP was constructed by a combined approach of feature-based modelling and virtual assembly in commercial software SolidWorks (Dassault Systemes SolidWorks Inc., USA). To achieve this accurately, we used the scanning technique to attain the geometrical data of all the components with complex geometries, e.g., the ploughshare. The overall dimensions of the HRP are 2.8 m length, 2.35 m width and 1.2 m height. The primary dimensions of the ploughshare are schematically illustrated in Fig. 3. More detailed information on the models of the HRP and the ploughshare is available in2, respectively.

Fig. 3
figure 3

Three-dimensional ploughshare model. (Min et al., 2019).

Tillage forces

In order to let the readers understand well the unique tillage behaviours of the HRP and also in order to calculate the tillage temperatures in the subsequent section of this study, it is essential to show the tillage forces on the plough-surface. However, unfortunately, the tillage-force equations for the traditional ploughs cannot be used for HRP due to the tillage force dependent of tool geometries, tillage behaviours, soil mechanical properties, and operational parameters (such as tillage depth and speed)8,14. For this reason, we based the geometries and the tillage performances of the HRP on the deduction of the tillage-force equations for the HRP tillage.

The classical tillage-force equation P for the traditional plough is15

$$P = P_{1} + P_{2} + P_{3}$$
(1)

where P1 is the frictional force, \(P_{1} = fG\),\(f\)—the friction coefficient, \(G\)—the weight of plough-body; P2—the soil deformation force, \(P_{2} = \frac{1}{4c}(\frac{a}{R})^{2} (\frac{1}{2} + \frac{aRtg\varphi }{L})(1 + \frac{2Rtg\varphi }{L}) + k_{a} a_{1} \frac{a}{L} + \frac{{bk_{a} }}{tg\varphi }(\frac{1}{B} + \frac{1}{{a_{1} }}) + \gamma a_{2}\), ka— the soil type coefficient, R—the curvature radius of plough-surface; P3—the draught force, \(P_{3} = \varepsilon abv^{{2}}\), \(a\)—the operational breadth, \(b\)—the operational depth, \(v\)—the towing velocity of a tractor.

Note that although Eq. (1) has been widely used for the traditional ploughs, there are two limitations if it is applied for the HRP tillage. The corresponding limitations are: (1) the plough-surface of HRP is different from those of the traditional ploughs, thereby the Eq. \(P_{2}\) inappropriate for HRP; (2) the soil type coefficient ka is varied frequently with the tillage time so that it cannot be accurately measured and, thus, may be inconveniently used in the practical application14.

For this reason, we modified the Eq. P2 and Eq. P3, respectively, based on the aforementioned analyses, as well as, the unique tillage behaviours and the geometries of the HRP. Due to manuscript length constraints, the relevant deductions are briefly presented below.

$$P_{2} = [K_{1} (\frac{1}{R} + K_{2} )^{2} + K_{3} ]ab$$
(2)

where K1—the soil compression coefficient, K2—the soil adhesion coefficient, and K3—the soil-steel surface friction coefficient.

$$P_{3} = \frac{1}{2}\rho ab\frac{{V^{2} }}{\sin \theta \cos \beta }$$
(3)

where \(\rho\)—the soil density, \(\theta\)—the horizontal element line angle, \(\beta\)—the projection of the included angle, between clod flight trajectory and horizontal line, in a plane perpendicular to the upper sideline of the plough-body; see Fig. 4.

Fig. 4
figure 4

Flowing trajectory of the soil clods in the soil-tool interaction.

Substituting Eq. (2) and Eq. (3) into Eq. (1), respectively, the tillage-force equation for the HRP tillage is

$$P = fG + [K_{1} (\frac{1}{R} + K_{2} )^{2} + K_{3} ]ab + \frac{1}{2}\rho ab\frac{{V^{2} }}{\sin \theta \cos \beta }$$
(4)

Currently, Eq. (4) has been demonstrated to be valid by the field trials in Xinjiang of China and has been successfully used in the previous studies3. Detailed information on this equation is available in14.

MDA and fluid–solid-thermal simulation

In this section, in order to accurately obtain the soil-tool interaction, particularly the tillage temperature and tillage stress, we applied the high-fidelity simulation framework of the combined MDA and fluid–solid-thermal model for all simulations under the actual tillage conditions of the HRP. The main advantages of this method are that the performances of mechanical components, as well as, their loads can be investigated on a system level16. The obtained loads can, in turn, be applied for the subsequent stress & strain analysis. Thus, such the approach would allow for a much more accurate simulations, in contract to the fluid–solid-thermal model that is merely used16.

The detailed steps of this approach in this study are: firstly, the MDA is performed to investigate the combined tillage operation of continuous shifting and moving at the three different working conditions; then, with the MDA solution the interaction load data between soil and tool are outputted at the specified discrete time-steps and transferred to the fluid–solid-thermal model of the ploughshare; finally, the coupled stress & strain, as well as the tillage temperature, at the share-point, is numerically predicted (Fig. 5).

Fig. 5
figure 5

Schematic diagram for the coupled MDA with fluid–structure-thermal simulations.

Soil and soil-tool interaction

Soil model

In this study the soil is sandy loam soil in Xinjiang of China (86°5′ ~ 87°8′ N, 43°7′ ~ 45°20′E). According to Table 1 and Fig. 2C, the soil was characterized as a Bingham material. This model has been demonstrated valid and successfully been used in our previous work13,17.

Table 1 Mechanical characteristics of soil (Zhu et al., 2016).

The common equation of the Bingham plastic model is:

$$\tau = \tau_{Y} + \mu \dot{\gamma }\;\;\;{\text{when}}\;\;\left| \tau \right| > \tau_{Y}$$
(5)
$$\dot{\gamma } = 0\;\;{\text{when}}\;\;\left| \tau \right| \le \tau_{Y}$$
(6)

where \(\tau\) is shear stress in Pa, \(\dot{\gamma }\)─shear rate in s−1, \(\tau_{Y}\)─yield strength in Pa, and \(\mu\)─plastic viscosity in Pa s.

For the Bingham fluid equations, there are the following two aspects to note: (1) the Bingham model presented above is regularized in numerical implementations to avoid the divergence of viscosity at zero rate of strain; this is because the original Bingham model cannot be used in numerical modeling without regularization due to the infinity of the viscosity when shear rate is zero. Currently, several approaches, including Tanner’s bi-viscosity model and Papanastasiou’s exponential model, are used for the regularization of the Bingham model; we refer the readers to the work of O’Donovan18 and Papanastasiou19 for details of the regularization. (2) The reasons why the Bingham model was used in this study are that the inertia effect may not be so relevant because of not so fast soil propagation and deposition during the HRP tillage, and this material model is its simplicity because only two parameters (yield strength and plastic viscosity), easily measured in the lab or in the field, are required19; For the fast soil propagation and deposition, the Bingham model may have poor capabilities in accurately model the surface pressure19.

In addition, concerning the volume fraction, i.e. soil, water and air, a multiphase fluid dynamics strategy should be used. For simplicity, Karmakar & Kushwaha tried using a single-phase laminar flow, demonstrating that this dynamics strategy is effective enough to analyze the soil-tool flow interaction because of its high molecular weight (Karmakar & Kushwaha, 2005).

Soil—tool dynamic interaction

The soil in the soil-tool interaction was modelled as a box of length × width × depth20, in which the length direction was prescribed as the direction of the towing speed V, whilst the width direction was prescribed as the direction of the shifting speed n (Fig. 1). Considering the actual tillage-scenario of HRP, we set the soil size (length × width × depth) as (14–20) m, 2.0 m, and 0.4 m, respectively, although the real soil-length may be so large due to a long-time tillage. Note that these soil sizes presented here are determined by the towing and shifting speeds of the plough-body as well as the operational depth. According to the practical tillage behaviors of the plough-body (Fig. 1), for all simulations with the MDA approach, any translation was fixed along the direction orthogonal to either upper or bottom surface, whilst the front, rear, left, and right surfaces were movable (Fig. 6).

Fig. 6
figure 6

Measurement on the heat partition coefficient on the soil/ploughshare interface. (A) Experimental setup, (B) location of plough-share, (C) heat conducting device, and (D) ploughshare.

Thermal distribution between soil and tool

From the viewpoint of heat transfer, it is necessary to determine the heat partition coefficient between soil and tool prior to calculating the thermal distributions at the share-point section. However, many of the current findings on heat partition are primarily focused on the metals. The heat partition between soil and tool (such as soil and plough-surface in agriculture) has not been considered in any detail.

Inspired by Mechanical Manufacturing theory, we proposed to couple inverse heat transfer in mechanical cutting technique with functional conversion to obtain the heat partition between soil and plough-surface. Inverse heat transfer is the method to determine the heat generation rate on cutting edge by minimizing the discrepancy between the measured and simulated temperatures at the same thermocouple locations21. The relevant temperature discrepancy can be determined by the function Obj in Eq. (7). Note that the cutting edge presented here is actually the share-point of the ploughshare.

$$Obj = \sum\limits_{i = 1}^{{n_{i} }} {\sum\limits_{j = 1}^{{n_{j} }} ( } \left. {T_{j}^{{t_{i} }} } \right|_{\exp } - \left. {T_{j}^{{t_{i} }} } \right|_{sim} )^{2}$$
(7)

where \(\left. {T_{j}^{{t_{i} }} } \right|_{\exp }\) is the measured temperature at thermocouple j at the time ti, \(\left. {T_{j}^{{t_{i} }} } \right|_{sim}\)—the calculated temperature at the same thermocouple location and time.

Figure 6 schematically illustrates the experimental setup for the heat partition on the soil-tool interface. Figure 7 shows the relevant flowchart of the test. The experimental set-up consists of a welding power source (ZX7-400 K, Shanghai, China), K-type thermocouples (Hefei, China), data acquisition modules (JY-DAM-AITC, Nanjing, China), sandy loam soil, a heat conducting device and a computer.

Fig. 7
figure 7

Flow chart of the inverse heat transfer solution of heat partition on soil/ploughshare interface.

For all measurements, ten thermocouples were placed within the zone ‘oab’, and particularly most of them were as close enough to the share-point section so that more accurate temperature could be obtained. The area of the zone ‘oab’ is approximately 0.0379 m2, same as the computational area in our previous studies13 , in order for further comparison. The heat conducting device, whose area is also 0.0379 m2, is tightly mounted onto the share-point section in order for better heat conduction. The material of heat conducting device is the same as that of the ploughshare, i.e., cast alloy steel. The sandy loam soil was created in accordance with its properties by using the approach in10. The working parameters in the test were three different tillage speeds (5, 6, 7 km h−1) and three different operational depths (0.27, 0.315, 0.360 m). The entire test last 4 h, which is also the time when 50 ha was finished tilling by HRP. The final measurements are list in Table 2. Considering the measurement errors, we chose 0.72 as the heat partition coefficient for all simulations in this study22. Due to manuscript length constraints, we refer the readers to the work of Feng Luo22 for the details of the heat partition measurements.

Table 2 Heat partition coefficients between ploughshare and sandy loam soil.

Multi-body dynamics analysis

The development of a multi-body dynamic model of HRP and its subsequent simulations have been discussed in detail previously3. The three-dimensional model of HRP is imported into MSC ADAMS motion software (Santa, CA, USA) in preparation for a multi-body dynamics analysis. During the HRP tillage, the moving parts are basic beam (BB), hydraulic cylinder (HC), plough-body, and reversing rod (RR); the immovable parts are main beam (MB) (Fig. 1). The dynamic analysis by the Newton Raphson iteration was performed to implement the combined tillage-operations of continuous shifting and moving forward. The loads on the plough-surface from MDA were subsequently exerted onto ploughshare for stress & strain analysis. In the entire tillage operations, the tillage behaviours of ploughshare were varied with the HC speeds, which ranged from 15 to 25 mm s−1 in 2.5 mm s−1 increments.

Fluid–solid-thermal coupling analysis

The three-dimensional model of ploughshare, as shown in Fig. 3 for the MDA, is imported into ANSYS 11 mechanical (ANSYS, Inc., Canonsburg, PA, USA) in preparation for the fluid-thermal-structural coupled analysis. All the analyses are involved in such the governing equations as fluid, solid, heat transfer and fluid-thermal-structural coupled boundary conditions. During this coupled analysis, the imposed tillage forces and temperature distributions at the share-point are obtained from the heat-flowing coupling analysis. The tillage-temperature field is obtained by simultaneous solving the heat convection equation in the fluid domain and the heat conduction equation in the solid domain.

The commercially available ANSYS ICEM was used for the mesh generation of the fluid–solid-thermal model. According to13,19, the tetrahedral elements, each solid of which has ten nodes, were used to mesh the entire ploughshare. Considering the significant effects of the share-point section on the soil flow, we refined the mesh there. Through the grid independence verification, the entire computational domain, including the fluid domain and the solid domain was divided into 6536 elements and 2876 nodes under the premise of the accuracy and the computational cost. The average skewness of the grid is 0.436. Note that the mesh used for stress analysis is the same as the mesh used for the heat-flow coupling analysis, and that the general grid interface connection is applied for the interface in-between flow domain and solid domain. Figure 8 schematically indicates the computational area and mesh division of the thermal-fluid–structure model as HRP is operated from right to left. Here the ploughshare is regarded as a stationary tool within a visco-plastic flowing domain13; the domain represents the soil; the yellow arrow zone, indicated by (A), denotes the real flow domain of the soil over the ploughshare; and the flow area is approximate 0.0379 m2, same as that in Sect. Thermal distribution between soil and tool.

Fig. 8
figure 8

Computational area and mesh division of the thermal-fluid–structure model for the soil-ploughshare interaction. (A) Schematic of solid element, (B) flow domain of the soil, and (C) meshed share-point.

Boundary conditions

For the cases examined in this study, as shown in Fig. 8, we prescribed the detailed boundary conditions for the MDA and the fluid-thermal-structural simulations below, respectively.

For the MDA, the loads exerted at the share-point under the real HRP tillage scenario are directly transmitted from the MDA solutions. According to3, the solutions were divided into 51 load steps for each tillage process, as shown in Fig. 1 i e., the plough-body tilling from the position A to B or from the position B to A at the three different working conditions. Since the shifting speed is controlled by the HC, it is necessary to prescribe the HC speed for the HRP tillage in advance. According to the real tillage operations of the HRP and13, in this study the HC speed was kept constant at 20 mm s−1, the moving speeds of the plough-body were 5, 6, and 7 km h−1, and the operational depths were 0.270, 0.315, and 0.360 m. Note that the effects of operational depths on the tillage loads could be determined through the relationship between moving speed and operational depth2. In addition, in order to obtain more accurate results, we considered the condition ‘none-reflect’ in the soil-tool interaction for all simulations. This approach has been demonstrated valid in calculating the exerted loads on the tool in infinite soil conditions13.

For the fluid-thermal-structural simulations (Fig. 8), the inlet speed component of soil was perpendicular to the share-point; the outflow pressure was exerted at the outlet; No slip appeared at either bottom or side of the ploughshare; and the free-surface grid movement was employed within the entire ploughshare surface13. Note that only conduction and convection were included in all simulations in accordance with the practical tillage operations of HRP.

All the simulations were focused on the coupled stresses and the tillage temperatures at the share-point section. Note that the entire calculation last for 40 h, which is the time when the 50 ha of sandy loam soil was finished tilling by HRP. As depicted previously, the SEM based measurements at the share-point section were achieved after the 50 ha of sandy loam soil was tilled by the HRP. It took Dell Precision T7920 Workstation almost one month to accomplish all the calculations in this study. Any calculation that was stopped and remained stable was considered to be converged as the residual of each equation was less than 1.0 × 10−5.

Results

MDA of the actual tilling scenario

Due to the symmetric geometry of ploughshare (Fig. 3) and the actual tillage operations of HRP, the tillage behaviours of the plough-body show periodicity2,14. Thus, the tillage forces at the share-point are believed to vary periodically with the tillage-time. For this reason, the tillage force variations at the share-point for 10 s are chosen to present at the three different operational speeds and depths in this section. Note that the 10 s indicates the time when the plough-body shifts from A to B or from B to A (Fig. 1).

Figures 9 and 10 schematically plot the tillage force variations on the plough-surface by the MDA approach. Since the tillage force at the share-point section accounts for about 83% of the one on the ploughshare, whilst the tillage force on the ploughshare accounts for about 50% of the one on the plough-surface2,14, the maximum tillage force at the share-point accounts for about 41.5% of the one on the plough-surface, as list in Table 3.

Fig. 9
figure 9

Tillage forces on the plough-surface at the different towing speeds.

Fig. 10
figure 10

Tillage forces on the plough-surface at the different operational depths.

Table 3 Maximum tillage forces on the share-point section.

As observed, for the constant operational depth and the three different towing speeds, the tillage forces have the same variation trends, always reaching the maximum values at 3 or 4 s and then gradually decreasing to about 4000 or 5000 N with the plough-body shifting. Similar variations can also be found in the tilling scenario of the constant moving speed and the three different operational depths. These results may be explained as due to the varying interaction between share-point and soil, or, the varying contact area between them. In detail, during the HRP tillage, soil is drawn into the ploughshare domain through the share-point and subsequently flows out its center section into the plough-breast by way of the plough-shank; see Fig. 8. Since both share-point and plough-shank are both soil-cutting sections of the plough-surface, so much energy there necessitates cutting, breaking, reducing and inverting soils. Consequently, the maximum tillage forces at any operational speed and depth occur at the first half of plough-body shifting (see the black arrows in Figs. 9, 10 and 11). In addition, there is another interesting finding that the tillage force at the initial interaction between soil and plough-body is rapidly increased and sharply decreased at any towing speed or operational depth (see the red arrows in Figs. 10 and 11). This phenomenon may be attributed to the mechanical failure of soil after the initial interaction between soil and share-point. At this time, the greatest stress concentration occurs at the share-point section because of rather small contact area there15.

Fig. 11
figure 11

Coupled stress and tillage-temperature distributions.

Fluid–solid-thermal simulation of the ploughshare

Based on the tillage forces obtained in Figs. 9 and 10 and Table 3, this section presents the fluid–solid-thermal simulations of the ploughshare at the three different towing speeds (such as 5, 6, and 7 km h−1) and the three operational depths (such as 0.270, 0.315, and 0.360 m). Since rather good qualitative agreement was observed for either coupled stress or tillage-temperature distribution over the ploughshare for all simulations, Fig. 11 only shows their distributions at the different working conditions. The red, green, and blue colors in the figure indicate the sections where the maximum, median and minimum coupled stresses or tillage temperatures over the ploughshare exist, respectively. As for the discrepancies in value, Tables 4 and 5 present the maximum coupled stresses and the highest tillage temperatures at the different working conditions.

Table 4 Maximum coupled stresses at the share-point section.
Table 5 Maximum tillage temperatures on the share-point section.

As observed, all the highest tillage temperatures and the maximum coupled stresses are located at the share-point section of the ploughshare. It follows that the maximum coupled stresses are accompanied with the highest tillage-temperatures. Meanwhile, either coupled stress or tillage temperature is gradually decreased with the distance, along the black dashed line L, away from the share-point. This line L is perpendicular to the share-point section. These simulations are in agreement with the practical distributions due to the fact that the share-point is the essential cutting edge of ploughshare13.

For better understanding the relationship between coupled stress and tillage temperature, Figs. 12, 13, 14 and 15 plot the variations on either coupled stress or tillage temperature at the share-point section with the distance at the different tillage conditions. The meaning of distance is the same as that in Fig. 11. As observed, at any operational depth or towing speed, both the maximum coupled stresses and the highest tillage temperatures appear at the share-point section. Either coupled stresses or tillage temperatures are gradually decreased with the distance away from the share-point increase. The maximum values in coupled stress and tillage temperature are increased with either operational depth or towing speed increase, respectively. For example, at the constant operational depth of 0.360 m, the maximum coupled stress is increased from 278.19 to 574.93 MPa, correspondingly the highest tillage temperature is increased from 100.10 to 204.50 °C, as the towing speed of the HRP is increased from 5 to 7 km/h.

Fig. 12
figure 12

Coupled stresses variations at the different operational depths.

Fig. 13
figure 13

Coupled stresses variations at the different towing speeds.

Fig. 14
figure 14

Tillage temperature variations at the different towing speeds.

Fig. 15
figure 15

Tillage temperature variations at the different operational depths.

Discussions

Based on the results in Section of MDA of the actual tilling scenario, it could be seen that the MDA approach allows many loading scenarios to be investigated, and that the generated tillage-force data from the MDA simulations could further be provided for the fluid–solid-thermal predictions. Thus, it is demonstrated once again that the previous conclusion in13 is true that the MDA approach could investigate the tillage performances on a system level, which, in turn, helps to accurately obtain the dynamic behaviors of the tillage tools.

According to the results in Section of the fluid–solid-thermal simulations, the obtained coupled stresses are compared with the calculated CFD stresses in13 (Fig. 16). Meanwhile, we also present the previously-obtained SEM views of the share-point at the three different working conditions (Figs. 17 and 18) in order to let the readers understand the comparisons well. The rectangles and arrows presented there are used to clearly show the abrasion/attrition wear variations and white materials at the share-point section, respectively.

Fig. 16
figure 16

Comparison of the coupled and the maximum CFD stresses on the share-point. (A) CFD stress distributions, and (B) the comparative values.

Fig. 17
figure 17

SEM views of the share-point at the different operational depths (Min et al., 2019).

Fig. 18
figure 18

SEM views of the share-point at the different tillage speeds (Min et al., 2019).

As shown in Figs. 11, 16, 17 and 18, the maximum coupled stresses and the highest tillage-temperatures in this study both appear at the share-point section of the ploughshare. The section is just where both greatest CFD based stresses and most severe wear rates appear. Meanwhile, the coupled stress, tillage temperature, CFD-based stress and wear rate are all increased with either towing speed or operational depth increase. Their increase trends are quite same with each other, but the coupled stresses are much larger than the CFD based ones at the same towing speed and operational depth. For example, at the constant operational depth of 0.36 m and the different towing speeds of 5, 6, and 7 km/h, the maximum CFD stresses are 23.2, 28.0, and 32.5 MPa, whilst the maximum coupled stresses are 278.2, 444.5 and 574.9 MPa. The maximum coupled stresses are approximately 11, 15, and 17 times as large as the largest CFD stresses, respectively.

These comparisons agree well with the idea, proposed by23,24,25, that the maximum force on the tool exerted by soil contributes to the most severe abrasive wear on its surface. Thus, coupled with our previous measurements (such as the severe abrasion/attrition wear plus some sort of oxide, e.g., Fe2O3), we conclude that the visible abrasion/attrition wear of the share-point with the white materials implies the combined effects of the thermal and mechanical consequences. That is, the thermal effects, caused by the high-speed tillage operation of the HRP in sandy loam soil, could aggravate the share-point wear. This may be explained as due to the following facts:

First of all, it is generally accepted that the wear mechanisms on tillage tools including the mouldboard plough are influenced by the material of the tool, opposing material (soil), environment (moisture, temperature), and dynamic factors (stress on sliding surface, sliding time, sliding speed, and sliding type)4. In this regard, the tool material, the soil type and the dynamic factor are the three main aspects affecting the wear mechanisms on tillage tools.

Then, from the viewpoint of the soil type, the sandy loam soil in the HRP tillage operations is well known to contain more particles and less water compared to the other soil categories in China2,14. These particles including rocks, gravel, and plant residues have different sizes and shapes. Mostly important, the number of the particles with the larger size in the sandy loam soil is greater than in the other soil categories. According to26,27, an increase in abrasive particle size causes the higher indentation depth on the plough-surface and, hence, the severe wear rate. This phenomenon indicates the more violent interaction appearing between soil and tool for the larger particle size and the same working conditions.

From the viewpoint of the dynamic factors, the unique dynamic behaviours of HRP-the continuous, alternative, and commuting high-speed tillage is accompanied with the transitional + rotational kinematic movements of soil and, consequently, the more irregular soil flow on the plough-surface2,14. The irregular soil flow is prone to produce the more intensive soil disturbance and as a result the larger load acting along the share-point28. Consequently, the greater soil stress concentration appears at the share-point section due to the more intensive stagnation point of fluid flow and the more abrupt load drop at the tool edge.

Surely, the aforementioned violent interaction and great tillage forces are accompanied with the elevated tillage temperatures at the share-point section. Although the tillage temperature obtained in this study is not too high (Table 5), the temperature always exists on the share-point section during the long-time tillage of the HRP, e.g., 4 h in this study. Since the tillage temperature at the share-point section frequently varies with the plough-body shifting, the high temperature gradients appear at the share-point section, which is prone to cause thermal runaway and thermal stress. The temperature stress is detrimental to the micro-shape or size shape of share-point. Long-term thermo-mechanical stresses impact the surface integrity at the share-point section, modify the surface residual stress distributions, and, hence, drive the fatigue rupture29.

Based on the above analyses, the underlying physical mechanism of the at the share-point wear in the high-speed shifting tillage operations of the HRP in sandy loam soil is that the violent sliding & rubbing, the greatly varied tillage temperature and the long-term thermo-mechanical stress gradient produce the severe abrasion/attrition wear with the white materials.

Conclusions

This paper extends our previous work by using a coupled fluid–solid-thermal simulation with MDA to verify the thermal effect, generated by the high-speed tillage operation of the HRP in sandy loam soil, on the share-point wear. The verification has been done primarily with respect to the relationship among the coupled stress, the tillage temperature and the wear land at the same working conditions.

The comparisons among the coupled stress, the tillage temperature and the wear land confirm that the thermal effects caused by the high-speed shifting operations of the HRP do aggravate the share-point wear. The severe abrasion/attrition wear with white materials at the share-point section may be caused by the combined effects of the violent sliding & rubbing, the greatly varied tillage temperature and the long-term thermo-mechanical stress gradient.

Overall important insights have been gained into the thermal effects in high-speed shifting tillage operations on the soil-engaging component, e.g., the share-point, adding to the knowledge-base usefully applicable to the design of the high-speed mouldboard. However, the numerical predictions should be further updated to investigate the quantitative relationship between the previously-measured worn lands and the commonly-calculated coupled stresses at the share-point section in sandy loam soil during the HRP tillage.