Numerical Simulation and Experimental Study on Liquid-Filling Forming of 2A12 Aluminum Alloy Fairing

Abstract

To address the challenges of excessive local thinning, poor surface quality, and low production efficiency in traditional multi-pass deep-drawn aluminum alloy fairings, this study investigates the effects of process parameters—including liquid chamber pressure, holding force, and differentiated lubrication schemes—on the liquid-filled forming performance and wall thickness distribution of a 460 × 280 × 1.5 mm thin-walled 2A12 aluminum alloy fairing. Employing an integrated liquid-filled forming technique combining a flexible punch with a rigid die, the research combines numerical simulation with experimental validation. The study demonstrates good consistency between experimental results and numerical simulations. The optimal forming process parameters are liquid chamber pressure of 10 MPa, holding force of 1100 kN, and a lubrication scheme (friction coefficients of 0.01 for the flange and forming zones and 0.06 for the transition radius zone). Under these parameters, part wrinkling and cracking are effectively suppressed, achieving optimal wall thickness uniformity in the formed parts, with a maximum thinning rate of only 6.6%. The proposed liquid-assisted forming process and differentiated lubrication scheme provide a new technical pathway for high-precision manufacturing of thin-walled complex curved components made of 2A12 aluminum alloy. Compared to traditional multi-stage drawing processes, both forming efficiency and quality are significantly improved.

1. Introduction

In recent years, the rapid advancement of the aerospace industry has imposed increasingly stringent demands on the lightweighting, high-strength, and aerodynamic precision of aircraft structural components [1]. The fairing, a critical aerodynamic component at the aircraft’s end, is typically designed as a deep conical shape to effectively reduce air resistance during flight. In manufacturing, the fairing is classified as a deep conical thin-walled rotary part, exhibiting significant deformation during production [2]. Typically, the depth-to-diameter ratio of fairing components is substantial, causing intense plastic flow in the sheet metal during forming. While 2A12 aluminum alloy exhibits some plasticity in the annealed state, its elongation at room temperature fails to meet the demands of extremely complex deformation [3].

Traditional stamping processes for fairings typically employ multi-stage deep drawing, requiring multiple sets of rigid dies and intermediate annealing between operations, resulting in low production efficiency. Localized excessive thinning often occurs during conventional stamping, while intense friction between dies and sheet metal compromises surface quality [4]. In aerospace applications, stringent requirements for surface quality and dimensional accuracy exist, where even minor defects can lead to structural failure. Therefore, exploring efficient and precise forming processes is particularly critical [5,6]. Addressing these challenges, Amini et al. [7] conducted numerical simulations and experimental studies on wrinkling behavior during mechanical deep drawing of conical components. Results indicate that for every 0.2 increase in the draw ratio, the minimum forming pressure required to prevent flange wrinkling must be increased by 50%.

To address uneven wall thickness distribution during the forming process of thin-walled complex curved components, Yuan et al. [8] investigated the reverse deep drawing process for conical curved parts. Results indicate that using preforms with removed flange radii for reverse deep drawing can increase the minimum wall thickness of formed parts by 8.58%. Yang et al. [9] applied liquid-filled forming technology to a large-sized 2A12 aluminum alloy rocket fairing. By optimizing the matching of liquid chamber pressure and blankholding force loading curves, they controlled the part thinning rate within 15%, achieved surface gaps smaller than 0.2 mm, and eliminated the need for multiple annealing passes and manual trimming required by traditional processes. Zhao et al. [10] innovatively proposed a stamping-expansion forming process for ultra-thin S-shaped bellows diaphragms made of GH4169 high-temperature alloy. Research indicated that under parameters of 30 t drawing force and 5 MPa expansion pressure, the maximum deviation of the diaphragm was only 0.02 mm, achieving a 30% improvement in precision compared to traditional stamping forming, with a wall thickness reduction rate below 5%. Xu et al. [11] developed a composite forming process combining liquid-filled drawing and hydraulic bulging for deep cylindrical components made of GH536 high-temperature alloy used in aero-engine applications. Results indicated that spherical-bottom preformed blanks exhibited the most uniform wall thickness distribution after bulging. These studies demonstrate that the matching relationship between blank geometry, liquid chamber pressure, and blankholding force loading is critical for achieving high-quality wall thickness distribution during forming.

To address forming limits and high-precision mold-following for thin-walled complex curved components, Piccininni et al. [12] proposed a process parameter optimization method combining response surface methodology and multi-objective genetic algorithms for controlling forming quality of aluminum alloy sheets under non-uniform deformation conditions. This approach significantly improved mold-following accuracy for complex geometries without fracture. Akay et al. [13] conducted numerical simulations and experimental studies on double-sided pressure-assisted warm liquid filling forming to address the poor room-temperature formability of high-strength aluminum alloys. They found that optimizing the loading curve effectively suppresses wrinkling on the side walls of deep cavity components. Liu et al. [14] studied AA6061-T6 aluminum alloy hemispherical components, establishing a nonlinear mapping relationship between process parameters and forming quality using a BP neural network. The study revealed that employing a “pre-expansion” process while precisely controlling the thickness and strength of the upper auxiliary plate significantly improves the uniformity of wall thickness distribution in the lower target plate. These findings demonstrate that through process innovation and multi-parameter co-optimization using different methods, the forming limits and precision of complex thin-walled components can be enhanced.

This study investigates the deep drawing of a 460 × 280 × 1.5 mm thin-walled 2A12 aluminum alloy sheet using a flexible punch finite element model, followed by a full-liquid-filled forming experiment on a fairing. The study investigates the influence of process parameters—including fluid chamber pressure and holding force—on the forming performance and wall thickness distribution of the fairing component. It proposes differentiated lubrication strategies for distinct forming zones and reveals the impact of sheet anisotropy on the fluid-assisted forming process, providing technical guidance for high-precision manufacturing of complex thin-walled aluminum alloy components with curved surfaces.

2. Fairing Liquid-Forming Process

2.1. Geometric Features of Formed Parts

The molded part, as shown in Figure 1, has a length of 201 mm, a width of 136 mm, a height of 48 mm, a transition radius of 10 mm, a wall thickness of 1.5 mm, and a curved upper surface. Since the right side of the part is open, sealing presents significant challenges. Therefore, the “one mold, two parts” process solution was adopted for optimization. The target part achieves effective cavity sealing through symmetrical forming, ensuring stable pressure during the liquid filling process.

The maximum thinning rate ε can be calculated using Formula (1), where t₀ denotes the initial wall thickness of the sheet material, and t_min denotes the minimum wall thickness after fairing forming.

$$\mathsf{\epsilon }={\displaystyle \frac{{\mathrm{t}}_0-{\mathrm{t}}_{\min }}{{\mathrm{t}}_0}}\times 100\%$$

(1)

2.2. Liquid-Filling Forming Principle

This study employs an integrated fluid-filled forming process combining a flexible punch with a rigid die. A high-pressure emulsion fluid serves as a flexible punch, replacing the traditional rigid punch to achieve flexible sheet forming. The fluid-filled forming process for the fairing is illustrated in Figure 2. Figure 2a shows the raw blank positioned above the rigid die with a specific profile prior to forming; Figure 2b shows the pressing and sealing stage, where the mold clamping firmly secures the blank edges. Combined with raw rubber tape for auxiliary sealing, this achieves complete cavity sealing to prevent high-pressure liquid leakage; Figure 2c depicts the free expansion stage after mold closure, where high-pressure emulsion is injected into the cavity. As cavity pressure increases, the liquid exerts a force perpendicular to the blank surface, causing the blank to deform into the concave mold cavity; Figure 2d depicts the film-adhering forming stage. Under continuous uniform force, the blank undergoes composite deformation involving both bending and expansion. This ultimately achieves high-precision conforming between the blank and the rigid concave mold cavity, forming a fairing component that meets dimensional requirements.

2.3. Forming Equipment and Molds

The fairing forming equipment employed in this experiment was the MHF-400t servo-proportional CNC sheet metal liquid-filled forming testing machine (manufacturer: Jinan Meiste Testing Machine Co., Ltd., Jinan, China). Tailored to the geometric characteristics of the fairing component, a liquid-filled forming die was designed as shown in Figure 3. The forming die primarily consists of an upper die, a lower die, and a liquid-filled joint. The mold material is Cr12MoV steel, heat-treated to achieve a hardness of 56 HRC, providing excellent wear resistance and enabling repeated use. The lubricants employed in the experiment were general-purpose industrial white lithium grease (manufacturer: Shanghai Hutou Chemical Co., Ltd., Shanghai, China) and 0.02 mm thick high-temperature-resistant PTFE film (manufacturer: Zhejiang Juhua Co., Ltd., Quzhou, China). Sealing was achieved through rigid sealing assisted by raw rubber tape.

3. Materials and Methods

3.1. Experimental Materials

The experiment utilized cold-rolled 2A12 aluminum alloy plates with a wall thickness of 1.5 mm (Chinese National Standard GB/T 3190-2020 [15], Al-Cu-Mg series wrought aluminum alloy). Its chemical composition is shown in

Table 1. Chemical composition of 2A12 aluminium alloy.

Element	Si	Cu	Mg	Zn	Mn	Ti	Ni	Fe	Cr	Others	Al
Value	0.5	3.8~4.9	1.2~1.8	0.3	0.3~0.9	0.15	0.1	0~0.5	0.1	0.5	allowance

, with Cu and Mg as the primary alloying elements. Characterized by high strength, this alloy is suitable for manufacturing aerospace structural components.

Cold-rolled 2A12 aluminum alloy plates develop pronounced rolling texture during the rolling process. The uniaxial compressive stress imposed during rolling causes internal grains to preferentially align along the rolling direction, forming a typical rolling texture. Since the annealing temperature does not reach the full recrystallization temperature, the texture characteristics remain intact, resulting in significant in-plane anisotropy in mechanical properties [16]. Following the national standard “Test Method for Tensile Properties of Metallic Materials at Room Temperature” (GB/T 228.1-2021) [17], specimens were cut along three directions relative to the plate rolling direction: 0° (parallel), 45° (oblique), and 90° (perpendicular), as illustrated in Figure 4 [18]. Tensile tests were conducted using a computer-controlled testing machine to calculate the true stress and true strain of the 2A12 aluminum alloy plate. Engineering strain ${\mathsf{\epsilon }}_0$ and engineering stress ${\mathsf{\sigma }}_0$ were then determined using Equations (2) and (3) [19].

$${\mathsf{\epsilon }}_0={\displaystyle \frac{\mathrm{L}}{{\mathrm{l}}_0}}$$

(2)

$${\mathsf{\sigma }}_0={\displaystyle \frac{\mathrm{F}}{{\mathrm{A}}_0}}$$

(3)

In the formula, L represents displacement, ${\mathrm{l}}_0$ denotes the gauge length of the specimen, F represents load, and ${\mathrm{A}}_0$ denotes the original cross-sectional area of the specimen.

The true stress $\mathsf{\sigma }$ and true strain $\mathsf{\epsilon }$ are calculated using Equations (4) and (5). The true stress–strain curve is plotted using Origin software (v2018), as shown in Figure 5.

$$\mathsf{\epsilon }=\ln (1+{\mathsf{\epsilon }}_0)$$

(4)

$$\mathsf{\sigma }={\mathsf{\sigma }}_0(1+{\mathsf{\epsilon }}_0)$$

(5)

To describe the hardening behavior after neck-backing for finite element simulation, the strengthening stage was fitted using Equation (6).

$$\mathsf{\sigma }=\mathrm{K}{\mathsf{\epsilon }}^{\mathrm{n}}$$

(6)

where K is the strength coefficient (MPa), n is the work hardening index.

Due to factors such as the fiber orientation structure formed during rolling, the plasticity of the plate varies with direction. The plastic strain ratio r value reflects the plate’s resistance to thinning [20]. The plastic strain ratio r is defined as the ratio of the true strain ${\mathsf{\epsilon }}_{\mathrm{b}}$ in the width direction to the true strain ${\mathsf{\epsilon }}_{\mathrm{t}}$ in the thickness direction of a uniaxially stretched specimen:

$$\mathrm{r}={\displaystyle \frac{{\mathsf{\epsilon }}_{\mathrm{b}}}{{\mathsf{\epsilon }}_{\mathrm{t}}}}={\displaystyle \frac{\ln \frac{\mathrm{b}}{{\mathrm{b}}_0}}{\ln \frac{\mathrm{t}}{{\mathrm{t}}_0}}}$$

(7)

In the equation: ${\mathsf{\epsilon }}_{\mathrm{b}}$ denotes the true strain of the specimen in the width direction, ${\mathsf{\epsilon }}_{\mathrm{t}}$ denotes the true strain of the specimen in the thickness direction, $\mathrm{b}$ denotes the width of the specimen after stretching, ${\mathrm{b}}_0$ denotes the initial width of the specimen, $\mathrm{t}$ denotes the thickness of the specimen after stretching, and ${\mathrm{t}}_0$ denotes the initial thickness of the specimen.

Average plastic strain ratio $\overline{\mathrm{r}}$ and in-plane anisotropy index $\Delta \mathrm{r}$:

$$\overline{\mathrm{r}}={\displaystyle \frac{{\mathrm{r}}_0+2{\mathrm{r}}_{45}+{\mathrm{r}}_{90}}{4}}$$

(8)

$$\Delta \mathrm{r}={\displaystyle \frac{{\mathrm{r}}_0-2{\mathrm{r}}_{45}+{\mathrm{r}}_{90}}{2}}$$

(9)

The anisotropy coefficients of the resulting plates are shown in

Table 2. Anisotropy of 2A12 aluminum alloy.

Direction	R0°	R45°	R90°	$\overline{\mathbf{r}}$	$\mathsf{\Delta }\mathbf{r}$
Width	9.28	9.37	9.4	0.544	0.276
Wall thickness	1.32	1.31	1.36
RT	0.584	0.480	0.631

.

Table 2 shows that the plastic strain ratio r values for the 2A12 aluminum alloy plate differ significantly across three distinct directions. The r value is highest in the 90° direction, with a measured average of 0.631, indicating that the plate is most susceptible to deformation in this orientation. The plastic strain ratio r is smallest in the 45° direction, with an average measured value of 0.480. This indicates that the plate is resistant to deformation in this direction. The calculated $\overline{\mathrm{r}}$ ≈ 0.54 is significantly less than 1. This demonstrates that the 2A12 aluminum alloy plate is highly susceptible to thinning in the thickness direction during stretching and exhibits poor resistance to thinning. The calculated “Δr ≈ 0.13 > 0” indicates significant in-plane anisotropy. The positive “Δr” value signifies that ear formation is likely to occur in the 0° and 90° directions during the liquid-filled deep drawing of the fairing, necessitating process optimization to suppress this phenomenon. The material properties of 2A12 aluminum alloy are summarized in

Table 3. Material properties of 2A12 aluminum alloy.

Parameter	Yield Strength/MPa	Tensile Strength/MPa	Young’s Modulus/GPa	Density g/cm3	Poisson’s Ratio
Value	63	168	71	2.78	0.33

.

The 2A12 aluminum alloy sheet used is in a cold-rolled state, subjected to 60% cold rolling deformation followed by low-temperature annealing at 380 °C. During cold rolling, the sheet experiences unidirectional rolling forces, causing internal grains to elongate along the rolling direction and undergo preferred orientation, forming a rolling texture. This results in distinct differences in grain orientation between different directions of the sheet. The grain orientation at 90° (perpendicular to the rolling direction) exhibits a soft texture orientation, characterized by low resistance to dislocation slip and easily activated slip systems, thereby demonstrating superior plastic deformation capability. Conversely, the 45° orientation represents a hard texture orientation of the rolling texture, featuring high resistance to dislocation slip. Additionally, partial deformation twins readily form in this direction during cold rolling, where twin boundaries further impede dislocation slip and propagation, significantly reducing plastic deformation capacity. This ultimately manifests as markedly different r-values in the 0°, 45°, and 90° directions, resulting in in-plane anisotropy [21].

3.2. Calculation of Forming Process Parameters

The core process parameters for liquid forming (chamber pressure, blankholding force) require fundamental values determined through mechanical calculations to provide a basis for subsequent numerical simulations and experimental optimization. By integrating sheet metal properties and part geometric characteristics, initial yield pressure, forming pressure, and minimum blankholding force are calculated separately.

3.2.1. Calculation of Fluid-Filling Expansion Pressure

During sheet metal bulging, the material primarily experiences tensile stress due to the normal force exerted by the fluid chamber, as shown in Figure 6 (${\mathsf{\sigma }}_{\mathrm{r}}$ denotes radial stress, ${\mathsf{\sigma }}_{\mathsf{\theta }}$ denotes circumferential stress). When the hydraulic pressure reaches Py, the stress reaches the material’s yield strength, and the sheet begins to undergo plastic deformation. As the fluid chamber pressure increases to Pc, the sheet undergoes a combined bending and bulging deformation under the combined action of hydraulic pressure and die support, achieving complete conforming to the cavity. The pressure at which liquid fills the punch and the sheet begins to undergo plastic deformation is the initial yield pressure ${\mathrm{P}}_{\mathrm{y}}$:

$${\mathrm{P}}_{\mathrm{y}}={\mathsf{\sigma }}_{\mathrm{S}}{\displaystyle \frac{2\mathrm{t}}{\mathrm{D}}}$$

(10)

In the equation, ${\mathsf{\sigma }}_{\mathrm{S}}$ represents the yield strength of the 2A12 aluminum alloy at 63 MPa, D denotes the equivalent diameter of the sheet expansion at 94.8 mm, and t indicates the sheet thickness at 1.5 mm. Calculations yield ${\mathrm{P}}_{\mathrm{y}}$ as 2 MPa.

To ensure complete conforming of the sheet to the concave mold cavity, the forming pressure ${\mathrm{P}}_{\mathrm{c}}$ must be calculated. Assuming the sheet thickness remains unchanged after liquid-filled forming, ${\mathrm{P}}_{\mathrm{c}}$ is:

$${\mathrm{P}}_{\mathrm{c}}={\mathsf{\sigma }}_{\mathrm{s}}{\displaystyle \frac{\mathrm{t}}{{\mathrm{r}}_{\mathrm{c}}}}$$

(11)

In the formula, ${\mathsf{\sigma }}_{\mathrm{s}}$ represents the yield strength of the 2A12 aluminum alloy at 63 MPa, t denotes the sheet thickness of 1.5 mm, and r_c is the transition radius of 10 mm. Calculations yield ${\mathrm{P}}_{\mathrm{c}}$ as 9.45 MPa.

3.2.2. Backing Force Calculation

During sheet metal liquid filling forming, the pressure in the liquid chamber exerts an upward ejecting force on the die. The magnitude of this ejecting force equals the product of the liquid chamber pressure and the contact projection area. The holding force must counteract this ejecting force while ensuring reliable sealing of the blank edges. The actual holding force must exceed the minimum holding force. To prevent high-pressure fluid from dislodging the mold, causing flash or pressure leakage, the minimum mold holding force ${\mathrm{f}}_{\mathrm{b}}$ must be calculated:

$${\mathrm{f}}_{\mathrm{b}}={\mathrm{P}}_{\mathrm{c}}\times \mathrm{A}$$

(12)

In the formula, ${\mathrm{P}}_{\mathrm{c}}$ represents the static pressure, A denotes the projected area of the fairing component in contact with the liquid (28,274.3 mm²), and ${\mathrm{f}}_{\mathrm{b}}$ represents the force of 267.2 kN.

3.3. The Effect of Sheet Anisotropy on Liquid-Filling Forming

The in-plane anisotropy of the 2A12 aluminum alloy sheet significantly influences the liquid-filled forming process of fairings. Along the material flow direction, the 90° orientation exhibits strong plastic deformation capacity and rapid flow velocity, readily causing localized material accumulation in this direction. At 0°, material flow is slower, making it prone to localized thinning, and exacerbating uneven wall thickness distribution. Regarding thinning resistance, the average plastic strain ratio $\overline{\mathrm{r}}$ ≈ 0.54 indicates susceptibility to thickness reduction in the sheet’s thickness direction. Improper process parameter control may cause excessive local thinning or even fracture. Regarding edge curling, the in-plane anisotropy index Δr > 0 causes the fairing edges to curl easily in the 0° and 90° directions, compromising dimensional accuracy and surface quality [22].

To optimize subsequent process parameters, account for sheet anisotropy by adjusting cavity pressure, blankholding force, and lubrication schemes. This balances material flow rates across all directions, suppresses thinning and edge curling, and enhances forming quality.

4. Finite Element Models and Reliability Verification

4.1. Finite Element Model Development

This study established a finite element model for the liquid-filled forming process of a fairing using the ABAQUS/Explicit(v2021) finite element analysis platform, as shown in Figure 7. The model primarily consists of three components: the upper die, lower die, and blank. The dies are defined as discrete rigid bodies, while the blank is modeled as a deformable body. Shell elements S4R were employed for meshing with a mesh size of 5 mm, with approximately 8000 total elements for the billet, ensuring computational accuracy and efficiency.

The material constitutive relationship employs a fitted power-law hardening model (Equation (6)), incorporating true stress–strain data and anisotropy parameters for 2A12 aluminum alloy. Contact between the die and billet was modeled as surface-to-surface contact. Boundary conditions were set as follows: the lower die was fixed, while the upper die closed downward to achieve blankholding. Liquid filling and forming were achieved through cavity pressure loading. Both cavity pressure and blankholding force were applied via uniform linear loading, consistent with the loading method of the experimental equipment.

4.2. Model Reliability Validation

To validate the reliability of the finite element model, a liquid-filled fairing forming experiment was conducted on aluminum alloy sheet metal under conditions matching the simulation: liquid chamber pressure of 10 MPa, holding force of 500 kN, and friction coefficient of 0.01. The simulated fairing is shown in Figure 8a, while the experimentally formed fairing is depicted in Figure 8b. To compare the wall thicknesses of the simulated and experimental fairings, 12 characteristic measurement points were selected circumferentially around the part, and the wall thickness at each point was measured. As shown in Figure 9, the simulated wall thickness slightly exceeds the experimental results. This discrepancy primarily stems from the simulation employing an ideal uniform lubrication model, whereas actual experiments involve localized friction variations between the die and blank. This leads to more intense plastic flow in the actual sheet metal, resulting in more pronounced localized thinning. Alternatively, leakage of high-pressure emulsion during the experiment may cause localized pressure fluctuations, causing greater localized plastic deformation of the blank against the die than simulated values. Nevertheless, the wall thickness distribution trends obtained from simulation and experiment are consistent. The relative error in wall thickness reduction rate is only 5.3%, falling within the permissible error range [23], demonstrating the reliability of the finite element model.

5. Analysis of Finite Element Simulation Results

Based on the validated finite element model, single-factor simulation experiments were conducted for chamber pressure, blankholding force, and lubrication schemes to investigate the influence patterns of various process parameters on fairing forming quality, providing a basis for determining optimal process parameters. Figure 10 shows the regional division of the 2A12 aluminum alloy fairing part, comprising the bulging zone, flange zone, and transition filet zone. The flange zone primarily serves sealing and material feeding functions. Excessive friction hinders material flow, causing rupture; insufficient friction may lead to flange wrinkling and seal failure. The transition fillet zone bears the highest contact normal pressure, where the sheet undergoes severe bending and reverse bending deformation. During the early forming stage, the sheet in the bulging zone remains suspended, subjected solely to liquid pressure; in the later stage, it gradually conforms to the die bottom.

5.1. Effect of Liquid Chamber Pressure on Part Forming Quality

Chamber pressure is a critical process parameter in liquid-filled forming, directly determining part forming quality. Insufficient pressure prevents complete film adhesion, resulting in failed fairing part formation; excessive pressure causes excessive thinning or even part rupture. To investigate the influence of liquid chamber pressure on the forming quality of 2A12 aluminum alloy fairings, six simulation experiments were conducted at pressures of 2 MPa, 4 MPa, 6 MPa, 8 MPa, 10 MPa, and 12 MPa, while maintaining a holding force of 500 kN and a friction coefficient of 0.01.

Finite element analysis results are shown in Figure 11 (color bar STH denotes Sheet Thickness, unit: mm). Simulation results indicate that when liquid chamber pressure ranges from 2 to 4 MPa, the overall wall thickness distribution of the part is relatively uniform, with the sheet metal in a free expansion state. However, due to insufficient forming pressure in the liquid chamber, the sheet metal fails to contact the cavity, resulting in poor forming quality. At liquid chamber pressures of 6–8 MPa, the maximum wall thickness reached 1.54 mm and the minimum was 1.16 mm. As pressure increased, the thinning zone gradually expanded. Although the sheet metal contacted the concave mold cavity, the transition filet area failed to achieve full film adhesion, resulting in poor part forming quality. At 10 MPa liquid chamber pressure, the maximum wall thickness was 1.57 mm and the minimum was 1.37 mm. The fairing part exhibited a maximum thinning rate of 9%. The film adhesion quality in the transition radius area was satisfactory. As liquid chamber pressure increased, the applied load enabled the sheet metal to overcome yield strength and undergo plastic elongation, resulting in good part forming quality. At a liquid chamber pressure of 12 MPa, excessive pressure causes the pressure on the sheet metal and die sidewall to increase exponentially with rising pressure. This leads to a dramatic increase in friction resistance on the die sidewall, hindering material flow into the die cavity from the flange area. Increased hydraulic pressure enhances the degree of plastic deformation in the blank, improving wall thickness uniformity. However, when pressure exceeds the critical threshold, the friction resistance on the die sidewall surpasses the flow stress of the flange zone material. This prevents material replenishment to the deformation zone, forcing it to rely solely on its own thickness reduction to meet the die-fitting requirements. This may result in localized excessive thinning.

5.2. Effect of Bending Force on Part Forming Quality

The holding force is a critical parameter for the forming quality of fairing components, serving to seal the cavity emulsion and control the material feeding rate into the die cavity from the flange area. With the fluid chamber pressure maintained at 10 MPa and a friction coefficient of 0.01, six sets of holding force simulation experiments were conducted using holding forces of 300 kN, 500 kN, 700 kN, 900 kN, 1100 kN, and 1300 kN.

Finite element analysis results are shown in Figure 12. The findings indicate that at holding forces of 300 kN and 500 kN, the compressive stress on the flange area exceeds the critical buckling stress, causing circumferential buckling and wrinkling. The flange region exhibits irregular creasing patterns, with material accumulation and wrinkling at the edges, compromising the sealing integrity of the fluid chamber pressure. At press-holding forces of 1100 kN and 1300 kN, the maximum wall thicknesses were 1.55 mm and 1.54 mm, respectively, while the minimum wall thicknesses were 1.28 mm and 1.18 mm. The maximum thinning rates were 14.7% and 21.3%, respectively. As the press-holding force increased, the flange area flattened, and the wrinkling phenomenon disappeared. However, a deep blue area of excessive thinning appeared in the central region of the fairing bottom, causing the part’s maximum thinning rate to surge sharply. Excessive blankholding force caused the friction resistance at the flange surface to exceed the material’s yield strength. Consequently, the overhanging material could not receive external feed supplementation and had to rely solely on its own thickness reduction to conform to the die cavity. At holding forces of 700 kN and 900 kN, the maximum wall thicknesses were 1.58 mm and 1.56 mm, respectively, while the minimum wall thicknesses were 1.38 mm and 1.33 mm. The maximum thinning rate was controlled below 12%. The flange area remained flat without wrinkling, and the color transition at the part’s base was uniform. This approach not only avoided wrinkling issues but also mitigated the thinning trend in the overhanging zone through material supplementation.

5.3. Effect of Lubrication Conditions on Part Forming Quality

Lubrication conditions significantly influence the forming performance of fairing components. Optimal lubrication can reduce thinning rates and improve wall thickness distribution irregularities in formed parts. Adjusting lubrication conditions is achieved by modifying the friction coefficients across different regions of the sheet metal. The friction coefficient between the sheet and the die determines material flow resistance and stress transfer efficiency. Fairings feature deep cavities and complex curved surfaces, making uniform lubrication methods challenging to simultaneously address the conflicting demands of promoting material replenishment in the flange zone while preventing material accumulation at the flange center. To avoid wrinkling and cracking defects in fairing components, a differentiated lubrication approach is proposed during forming. Differential lubrication involves applying distinct lubricants to different zones on the blank surface and die, inducing varying material flow velocities across regions to mitigate excessive thickening and thinning.

To investigate the impact of lubrication conditions on the forming performance and wall thickness distribution of the 2A12 aluminum alloy fairing part, the cavity pressure was maintained at 10 MPa and the holding force at 500 kN. Two lubrication schemes were designed, as shown in

Table 4. Lubrication scheme design.

Group	Scheme	Flange Region	Transition Filet Region	Bulging Region
Group 1	#1	0.1	0.1	0.1
	#2	0.06	0.06	0.06
	#3	0.01	0.01	0.01
Group 2	#4	0.1	0.06	0.06
	#5	0.01	0.06	0.01
	#6	0.06	0.01	0.01

. The first group employed conventional uniform lubrication to examine the overall friction coefficient’s influence on blank feeding. The second group employed differential lubrication to analyze the control effect of friction forces in different zones on material flow.

The finite element analysis results are shown in Figure 13. From the first set of simulation results, when the overall friction coefficient of the part is 0.1, the maximum wall thickness is 1.577 mm, the minimum wall thickness is 1.376 mm, and the maximum wall thickness reduction rate is 8.2%. Due to high friction resistance preventing sheet material from flowing and replenishing from the flange area, the deformation pattern approximates pure bulging. When the overall friction coefficient of the part is 0.06, the maximum wall thickness is 1.584 mm, with a minimum wall thickness of 1.389 mm and a maximum wall thinning rate of 7.4%. As the friction coefficient decreases, material from the flange region can better flow into the transition radius and bulging zones. This reduces the maximum thinning rate of the fairing part, decreases the area of the most severely thinned region, and results in a more uniform wall thickness distribution throughout the fairing part. When the overall drag friction coefficient of the fairing part is 0.01, the maximum wall thickness is 1.598 mm, with a minimum wall thickness of 1.401 mm and a maximum wall thickness reduction rate of 6.6%. Although the area of severe wall thinning is minimized with further friction coefficient reduction, material accumulation occurs in the transition filet zone. Excessive local wall thickness in this zone may cause wrinkling defects during actual production, leading to poor process stability.

The second set of simulation results shows that in Scheme #4, the maximum wall thickness is 1.577 mm, the minimum wall thickness is 1.377 mm, and the maximum wall thickness reduction rate is 8.2%. Compared to the overall friction coefficient of 0.06 in Scheme #2, the friction coefficient in the flange zone increased, resulting in reductions in both maximum and minimum wall thicknesses. This occurs because the increased friction coefficient in the flange zone hinders material flow toward the center of the flange from both sides, mitigating excessive wall thickness in the flange center. Simultaneously, it slows material flow toward the transition radius zone, exacerbating wall thinning. In Scheme #5, the maximum wall thickness is 1.597 mm, the minimum wall thickness is 1.400 mm, and the maximum wall thickness reduction rate is 6.6%. Compared to the overall friction coefficient of 0.01 in #3, increasing the friction coefficient in the transition radius area reduces material flow from the transition radius to the bulging zone and from the flange zone to the transition radius. This alleviates excessive local wall thickness in the transition radius area, facilitating the formation of a better-performing fairing component. In Scheme #6, the maximum wall thickness is 1.586 mm, the minimum wall thickness is 1.389 mm, and the maximum wall thickness reduction rate is 7.4%. Compared to Scheme #5, it reduces the maximum wall thickness but increases the wall thickness reduction rate. By increasing the friction coefficient in the flange zone and decreasing it in the transition radius zone, material flow from both sides of the flange zone to its center becomes restricted. Simultaneously, material flow from the flange zone to the transition radius zone is also hindered. While this further mitigates excessive local wall thickness in the transition radius zone, it increases the maximum wall thickness reduction rate. The differentiated lubrication scheme regulates material flow velocity across zones, balancing feed rate with deformation rate. This ensures smooth material replenishment in the flange region while suppressing material accumulation in the transition radius zone, achieving optimal control over forming quality.

6. Forming Experiment

A liquid-filled fairing forming experiment was conducted on a 460 × 280 × 1.5 mm thin-walled 2A12 aluminum alloy sheet using the simulation parameters from Figure 13e. The experimental results are shown in Figure 14: the fairing part achieved full film adhesion with good surface quality, a maximum thinning rate of 6.6%, and relatively uniform wall thickness distribution. The comparison between simulated and experimental wall thicknesses is shown in Figure 15, confirming that the formed fairing part met the acceptance criteria.

The fairing’s profile clearance was less than 0.15 mm, with a surface roughness Ra of 0.8 μm, outperforming the traditional multi-pass deep drawing process (profile clearance > 0.2 mm, surface roughness Ra > 3.2 μm). The “one die, two parts” process combined with a differentiated lubrication scheme effectively resolved the sealing challenges and uneven wall thickness distribution issues of non-enclosed fairings. The forming process was stable, with no pressure leakage or forming defects. Compared to traditional deep drawing, the liquid-filled forming process reduced forming passes from 3 to 1, eliminating intermediate annealing and manual trimming operations, boosting production efficiency by over three times. The maximum wall thinning rate of the formed part decreased from 13.3% to 6.6%, with significantly improved surface quality and wall thickness uniformity. The liquid-filled integral forming process demonstrates substantial technical advantages in manufacturing thin-walled, complex curved components from 2A12 aluminum alloy [24,25].

7. Conclusions

This paper investigates the influence patterns of various process parameters on the forming performance and wall thickness distribution of the fairing. It reveals the effect of anisotropy in 2A12 aluminum alloy plates on liquid-filled forming, proposes differentiated lubrication schemes for different forming zones of the fairing, and determines optimal forming process parameters. The main research conclusions are as follows:

(1)

Cold-rolled 2A12 aluminum alloy sheets exhibit significant in-plane anisotropy. This arises from rolling texture formed during processing, resulting in maximum plastic deformation capability at 90° (r = 0.631) and deformation hardening at 45° (r = 0.480), with an average plastic strain ratio $\overline{\mathrm{r}}\approx 0.54$. The sheet exhibits poor thinning resistance, and since $\Delta \mathrm{r}>0$, ear formation is prone to occur in the 0° and 90° directions during liquid-filled forming. This anisotropy causes variations in material flow direction during forming, leading to uneven wall thickness distribution, which requires resolution through process parameter optimization.

(2)

Liquid chamber pressure and holding force are critical parameters for fairing liquid-assisted forming, determining sheet conforming and wall thickness distribution: Insufficient liquid chamber pressure prevents full sheet conforming to the die. Excessive pressure causes localized over-thinning; the optimal forming pressure is 10 MPa. Inadequate holding force leads to circumferential instability and wrinkling in the flange zone, resulting in seal failure. Excessive holding force prevents material replenishment in the flange zone and causes excessive thinning in the deformation zone. The optimal forming holding force is 1100 kN. At this parameter, the plastic deformation of the sheet matches the material replenishment rate, effectively suppressing part wrinkling and fracture defects.

(3)

The proposed differentiated lubrication scheme (friction coefficient of 0.01 for the flange and bulging zones, 0.06 for the transition radius zone) enables controlled material flow rates across forming regions compared to traditional uniform lubrication: Increasing the friction coefficient in the transition radius zone suppresses material buildup, effectively resolving the conflict between material feeding and wrinkle prevention inherent in uniform lubrication. The formed part exhibits no wrinkling or buildup defects, with uniform wall thickness distribution.

(4)

The 2A12 aluminum alloy fairing formed under optimal process parameters (chamber pressure 10 MPa, holding force 1100 kN, differentiated lubrication scheme) exhibits perfect mold conformity and excellent surface quality. The maximum thinning rate is only 6.6%, with a mold gap < 0.15 mm and surface roughness Ra = 0.8 μm. Compared to traditional multi-stage deep drawing processes, forming efficiency increased by over three times, with significantly improved wall thickness uniformity and forming accuracy.